Microsoft Word - 03-3290_s_ETASR_V10_N2_pp5361-5366
Engineering, Technology & Applied Science Research Vol. 10, No. 2, 2020, 5361-5366 5361
www.etasr.com Mangi et al.: Parametric Study of Pile Response to Side-by-Side Twin Tunneling in Stiff Clay
Parametric Study of Pile Response to Side-by-Side
Twin Tunneling in Stiff Clay
Naeem Mangi
Department of Civil Engineering
Quaid-e-Awam University of
Engineering, Science & Technology
Nawabshah, Pakistan
naeem08ce30@gmail.com
Daddan Khan Bangwar
Department of Civil Engineering
Quaid-e-Awam University of
Engineering, Science & Technology
Nawabshah, Pakistan
daddan@quest.edu.pk
Hemu Karira
Department of Civil Engineering
Mehran University of Engineering &
Technology
Khairpur Mir’s, Sindh, Pakistan
engr.hemu07civil@gmail.com
Samiullah Kalhoro
Department of Civil Engineering
Quaid-e-Awam University of Engineering, Science & Technology
Nawabshah, Pakistan
samiullahkalhoro63@gmail.com
Ghulam Rasool Siddiqui
Department of Civil Engineering
Mehran University of Engineering & Technology
Khairpur Mir’s, Sindh, Pakistan
ghulamrasoolsiddiqui@gmail.com
Abstract—A three dimensional coupled-consolidation numerical
parametric study was carried out in order to gain new insight of
single pile response to side-by-side twin tunneling in saturated
stiff clay. An advanced hypo plasticity (clay) constitutive model
with small-strain stiffness was adopted. The effects of relative to
the pile tunnel depths were investigated by simulating the twin
tunnels near the pile at various depths of tunnels, namely near
the pile shaft, adjacent to the pile toe, and below the pile toe. It
was found that the second tunneling in each case resulted in a
larger settlement than the one due to the first tunneling with a
maximum percentage difference of 175% in the case of twin
tunneling near the mid-depth of the shaft. This occurred due to
the degradation of clay stiffness around the pile during the first
tunneling. Conversely, the first tunneling-induced bending
moment was reduced substantially during the second tunneling.
The most critical location of twin tunnels relative to the pile was
found to be below the pile toe.
Keywords-twin tunneling; pile foundation; parameteric study
I. INTRODUCTION
Tunnel excavation-induced stress relief and ground
movements inevitably cause additional deformation and stress
on adjacent pile foundations. It is a major concern for designers
and engineers to evaluate the adverse effects on the existing
piles. To understand the pile–soil–tunnel interaction
mechanism, many field monitoring studies and centrifuge
model tests [1-3] and analytical solutions and numerical
modeling studies [4-7] have been conducted. They all
concluded that tunneling adjacent to existing pile foundations
caused pile settlement, additional axial load on piles, and
induced bending moments along the piles, which is
unfavourable for piled foundations. Their magnitudes are likely
depended on the relative locations of tunnels and piles.
However, most previous studies have focused on the effects of
a single tunnel on single piles and pile groups. In fact, twin
tunneling is particularly favoured when developing
underground transportation systems [8]. To obtain a
satisfactory numerical model of the single pile response to side-
by-side twin-tunneling, the analysis needs to take account of
the soil’s small strain non-linearity. In view of the
aforementioned issues, this study aims to systematically
investigate the settlement and load transfer mechanism of an
existing single pile due to side-by-side twin tunnels in saturated
stiff clay. To achieve these objectives, a three-dimensional
coupled-consolidation numerical parametric study was carried
out by varying the twin tunnel depths relative to the pile.
II. THREE DIMENSIONAL COUPLED CONSOLIDATION
ANALYSIS
To investigate single pile response to side-by-side twin
tunneling in stiff saturated clay, a three-dimensional coupled
consolidation numerical parametric study was conducted. To
facilitate validation of the numerical model, the tunnel
diameter, pile diameter, embedded length, and clear distance
between the pile and the tunnel were identical (in prototype
scale) to those in the centrifuge test [9]. Figure 1 shows the
elevation view of the configuration of a typical numerical
simulation in which twin tunnels were excavated adjacent to
the pile toe. The diameter of each tunnel (D) was 6m. The
embedded length (Lp) and diameter (dp) of the pile were 18m
and 0.8m respectively. The modeled pile represents a
cylindrical reinforced concrete (grade 40, reinforcement
ratio=1) with a bending moment capacity of 800kNm. The
center-to-center distance between each tunnel and the pile was
5.5m (0.92D). It is worth noting that, in reality, high-rise
buildings are unlikely to be built on a single pile. This
hypothesized study is a more virtual case [3]. This
simplification was made to understand the settlement and load
transfer mechanism more clearly. The length of each model
tunnel (along its longitudinal direction) is 72m, which is
Corresponding author: Naeem Mangi
Engineering, Technology & Applied Science Research Vol. 10, No. 2, 2020, 5361-5366 5362
www.etasr.com Mangi et al.: Parametric Study of Pile Response to Side-by-Side Twin Tunneling in Stiff Clay
equivalent to 12D. Each tunnel excavation was simulated in 28
steps. At each step, the tunnel advanced a distance of 2.5m
(0.42D) [2]. The time increment of one day for each step was
adopted in the finite element analysis. A monitoring section
was selected at the transverse centreline of the pile (i.e. y/D=0)
as a reference for the tunnel advancements. The parametric
study consisted of nine different numerical simulations (in
total) in which the existing single pile was located between the
twin tunnels, which were excavated one after the other on
either side of the pile at various depths of tunnels (zt) relative to
the pile length (Lp), namely near the pile shaft (zt/Lp=0.67 and
0.83), adjacent to the pile toe (zt/Lp=1.00) and below the pile
toe (zt/Lp=1.17, 1.33, 1.50, 1.67, 1.83 and 2.00). In addition to
these simulations, a pile load test (L) was conducted
numerically in “greenfield” conditions (i.e. with no tunnels
present) to obtain the ultimate capacity of the pile in stiff clay.
Based on this, the working load was then calculated with a
factor of safety of 3.0. The obtained working load was applied
to the pile in the parametric analysis simulating twin tunneling.
Table I summarises the conducted numerical simulations.
Fig. 1. Configuration of a typical numerical run representing case of
zt/Lp=1.00
TABLE I. NUMERICAL SIMULATIONS SUMMARY
Description of numerical run zt/Lp C/D
Twin tunneling near the pile shaft
0.67 1.5, 1.5
0.83 2.0, 2.0
Twin tunneling next to the pile toe 1.00 2.5, 2.5
Twin tunneling below the pile toe
1.17 3.0, 3.0
1.33 3.5, 3.5
1.50 4.0, 4.0
1.67 4.5, 4.5
1.83 5.0, 5.0
2.00 5.5, 5.5
zt=tunnel depth, Lp=pile length, C/D=cover to diameter of tunnel ratio
III. FINITE ELEMENT MESH AND BOUNDARY CONDTIONS
Figure 2 shows an isometric view of a typical finite element
mesh (for the case of zt/Lp=1.0).The size of the mesh for each
numerical simulation is 72m×72m×60m. These dimensions
were sufficiently large to minimize the boundary effects in the
numerical simulation because a further increase in the
dimensions of the finite element mesh did not lead to any
change in the computed results. Regarding the element size in
the mesh, it is found that further halving the adopted mesh size
leads to a change in the computed results of no more than only
0.2%, suggesting that the mesh is sufficiently fine. Eight-noded
hexahedral brick elements were used to model the soil and the
pile. Four-noded shell elements were adopted to model each
tunnel lining. Roller and pin supports were applied to the
vertical sides and the base of the mesh, respectively. Therefore,
movements normal to the vertical boundaries and in all
directions of the base were restrained. The water table was
assumed to be at the ground surface. Initially, the pore water
pressure distribution was assumed to be hydrostatic. Free
drainage was allowed at the top boundary of the mesh. The
tunnel lining was assumed to be continuous and impervious.
Fig. 2. Finite element mesh and boundary conditions of a typical
numerical analysis (zt/Lp=1.00)
Interaction between the piles and surrounding soil was
modeled by the surface-to-surface contact provided in Abaqus
software package [10]. The surface-to-surface contact
considers the shape of both the master and slave surface in the
contacting region and allows pile-soil friction. The penalty
approach was used for tangential contact and the normal
behavior was modeled as hard contact with no normal relative
displacement between the pile and the surrounding soil. The
interface was modelled by the Coulomb’s friction law, in which
the interface friction coefficient (µ) and limiting displacement
(γlim) are required as input parameters. A limiting shear
displacement of 5mm was assumed to achieve full mobilization
of the interface friction equal to µ×p', where p' is the normal
effective stress between two contact surfaces, and a typical
value of µ for a bored pile of 0.35 was used [11]. The tunneling
stress release process was modelled by the “element death”
technique which is widely used in finite element analysis. In
this technique, elements and nodes can be deactivated and
activated. In this study, the volume loss is predefined before
tunneling by specifying the area of the annulus gap between the
tunnel lining and the excavated soil. Pattern of the non-uniform
displacement boundaries was determined according to the
displacement controlled model (DCM) in [5]. The soil inside
the segment is excavated by deactivating soil elements inside it
Ground water table
Working load
(1.93 MN)
Existing single pile
(Lp=18 m)
Stiff clay
Note:- Lp is embedded length of pile
zt is depth of tunnel axis from the ground surface
zt=18 m
11 m
First
Tunnel
Second
Tunnel
2.5 m 2.5 m
Engineering, Technology & Applied Science Research Vol. 10, No. 2, 2020, 5361-5366 5363
www.etasr.com Mangi et al.: Parametric Study of Pile Response to Side-by-Side Twin Tunneling in Stiff Clay
and by specifying zero horizontal displacement at the tunnel
face of the tunnel segment to be excavated. In the meantime,
the shell elements representing the tunnel lining are activated.
IV. CONSTITUTIVE MODEL AND MODEL PARAMETERS USED
IN FINITE ELEMENT ANALYSIS
A basic hypoplastic model was developed to capture the
nonlinear behavior (upon monotonic loading at medium-to
large-strain levels) of granular materials [12]. The basic model
consists of five parameters (Table II). To account for the strain
dependency and path dependency of the soil stiffness (at small
strains), authors in [12] further improved the basic hypoplastic
model by incorporating the concept of intergranular strain,
which requires five additional parameters. The hypoplastic clay
model with small strain stiffness has been implemented in the
finite element software package Abaqus through a user-defined
subroutine. The coefficient of lateral earth pressure at rest, Ko
was estimated by the relevant equation in [14]. The concrete
pile and tunnel lining were assumed to be linear elastic with
Young's modulus of 35GPa and Poisson's ratio of 0.25. The
thickness of the lining was taken as 0.25m. The unit weight of
concrete was taken as 24kN/m
3
. The parameters for the piles
and the tunnel lining are summarized in Table III.
TABLE II. ADOPTED MODEL PARAMETERS OF KAOLIN CLAY
Description Value
Effective angle of shearing resistance at critical state, φ’ 22
o
Parameter controlling the slope of the isotropic normal
compression line in the ln(1+e) versus lnp plane, λ* [13]
0.11
Parameter controlling the slope of the isotropic normal
compression line in the ln(1+e) versus lnp plane, κ* [13]
0.026
Parameter controlling the position of the isotropic normal
compression line in the ln(1+e)–lnp plane, N
1.36
Parameter controlling the shear stiffness at medium- to
large- strain levels, r
0.65
Parameter controlling the initial shear modulus upon 180°
strain path reversal, mR
14
Parameter controlling the initial shear modulus upon 90°
strain path reversal, mT
11
Size of elastic range, R 1×10
-5
Parameter controlling the rate of degradation of the stiffness
with strain, βr
0.1
Parameter controlling the degradation rate of stiffness with
strain, χ
0.7
Initial void ratio, e 1.05
Dry density (kg/m
3
) 1136
Coefficient of permeability, k (m/s) 1×10
-9
TABLE III. ADOPTED CONCRETE PARAMETERS IN FEM
Description Value
Young's Modulus, E 35GPa
Poisson's ratio, νννν 0.3
Density, ρρρρ 2400kg/m
3
V. INTERPRETATION OF COMPUTED RESULTS
A. Induced Pile Settlement Due to Twin Tunnels
Figure 3 illustrates the pile settlement induced after the first
and th second tunnel at different depths relative to the pile
(zt/Lp). The long-term pile settlement (i.e. 15 years after twin
tunneling completion) is included in the figure. The measured
induced pile settlement tunneling at different depths in
centrifuge modeling reported by [9, 15] is also shown in the
figure for comparison. A linear increase in twin tunneling-
induced settlement was observed as the tunnel depth increased
from zt/Lp=0.67 to 1.33. However, as the depth increased
further (1.50≤zt/Lp≤2.0), the induced settlement decreased with
a similar trend. This computed result is consistent with the one
measured in centrifuge tests simulating tunneling at different
depths relative to piles in stiff clay [9] and in sand [8]. This
tunneling-induced settlement mechanism with zt/Lp can be
attributed to the influence zone of tunneling-induced ground
movement and stress-release-affected regions. In the case of
zt/Lp=1.33, the entire pile stood within the influence zone of
tunneling-induced ground movement, and the stress release
region was developed directly underneath the pile toe. When
zt/Lp=0.67, 0.83 and 1.00, the pile was located only partially
within the influence zones. Therefore, the pile experienced
larger settlement in the case of zt/Lp=1.33 than in these cases.
Although the entire pile in the cases of zt/Lp=1.50, 1.67, 1.83
and 2.00 was located inside the tunneling-induced ground
movement, the tunneling-induced pile settlement was less than
that of zt/Lp=1.33. This is because the location of the pile toe
was beyond the tunneling-induced stress-release-affected
region in these cases.
Fig. 3. Computed load settlement curve from the pile load test without
tunneling
Qualitatively, the second tunneling-induced settlement
trend with respect to zt/Lp was similar to the first tunneling-
induced settlement. However, the magnitude of the pile
settlement induced by the second tunnel was higher than that
induced by the first tunnel in each case. This can be attributed
to the degradation of clay stiffness around the pile due to stress
release and the development of shear strains as a result of the
first tunnel excavation. The largest settlement (175% of Sp due
to the first tunnel) was induced by the second tunnel when
zt/Lp=0.67. It can be seen that the settlement of the pile reduced
in the long term (i.e. 15 years after completion of the twin
tunnels) in all cases. This pile heave is attributed to the
dissipation of excessive negative pore pressure generated
0.00
0.50
1.00
1.50
2.00
2.50
3.00
3.50
4.00
4.50
5.00
5.50
6.00
6.50
7.00
7.50
8.00
0.50 0.67 0.83 1.00 1.17 1.34 1.50 1.67 1.84 2.00
N
o
rm
a
li
se
d
p
il
e
h
e
a
d
s
e
tt
le
m
e
n
t
(S
p
/d
p
):
%
Tunnel depth relative to pile (zt/Lp)
Loganathan et al., 2000 After first tunnel
Ng et al., 2013 After twin tunnels
Series6 sLong term settlement (15 years after twin tunnelling)
(Cummulative)
Computed (This study)Measured (Centrifuge)
1.33 1.83
Engineering, Technology & Applied Science Research Vol. 10, No. 2, 2020, 5361-5366 5364
www.etasr.com Mangi et al.: Parametric Study of Pile Response to Side-by-Side Twin Tunneling in Stiff Clay
around the pile due to the twin tunnels. The maximum
reduction of the pile settlement (17.2% of the settlement after
twin tunnels) was observed in the case zt/Lp=1.33.
B. Changes in Axial Load Distribution
Figure 4 illustrates the axial force distribution along the pile
with normalized depth (i.e., Z/Lp) below the ground surface
after twin tunneling when zt/Lp=0.67, 1.00 and 1.33. The axial
load distribution before tunneling (after applying the working
load) is also included in the Figure as a reference. Before
tunneling, the pile carried approximately 75% of the working
load (i.e. 1010kN) with its shaft resistance and the remainder
with its end-bearing resistance. Because tunneling was carried
out near the mid-depth of the pile shaft in the case zt/Lp=0.67,
tunnel-induced reduction in normal stresses to the pile shaft
and downward soil movement caused an increase in axial load
along the entire length of the pile after the first and second
tunnel excavations. At the end of the first tunneling, the
maximum increment in the axial force (63% of that at working
load) was computed at Z/Lp=0.7, which is above the tunnel
spring line. By inspecting the axial load distribution after the
first tunnel excavation, it is observed that along the upper half
of the pile (0≤Z/Lp≤0.6), the shaft resistance decreases to zero.
Consequently, the load was transferred to the lower half of the
pile. To maintain equilibrium, the pile had to settle to further
mobilize the end-bearing and shaft resistance along the lower
portion (Z/Lp>0.6). This led to increases of 73% and 24% in the
mobilized end-bearing and shaft resistance at the lower portion
(Z/Lp>0.5) respectively upon completion of the first tunnel. The
second tunneling in the case zt/Lp=0.67 caused a further
reduction of the normal stresses to the pile shaft. Consequently,
soil settled more than the pile, resulting in Negative Skin
Friction (NSF) along the upper half of the pile (0≤Z/Lp≤0.6).
This suggests that this portion of the pile is subjected to
dragload by the surrounding soil. This caused the pile to settle
more than that due to the first tunnel. To maintain vertical
equilibrium of the pile, the soil surrounding the lower part of
the pile (Z/Lp>0.6) resisted its settlement by mobilizing
Positive Skin Friction (PSF) at the pile–soil interface and end-
bearing resistance at the toe of the pile. Owing to the second
tunneling, the end-bearing and mobilized shaft resistance
increased to 190% and 18%, respectively.
Owing to twin tunneling adjacent to the pile toe
(zt/Lp=1.00), the axial load increased along the pile at depths
ranging from 0.42 to 1.0Lp. The increase in axial load resulted
from reduced shaft resistance, which was caused by stress
release from twin tunneling adjacent to the pile toe. The
reduction percentages in shaft resistance were 22% and 43%
after the first and twin tunneling (cumulative) respectively.
Consequently, the axial load borne by the shaft resistance along
the pile at depths ranging from 0.42 to 1.0Lp was transferred
downward to the pile toe, leading to a 75% and 156% increase
in mobilized end-bearing resistance after the first and twin
(cumulative) tunneling, respectively. In contrast to the
tunneling near the mid-depth of the pile shaft (zt/Lp=0.67) and
adjacent to the pile toe (zt/Lp=1.00), the axial load decreased
along the entire length of the pile when twin tunnels were
excavated below the pile toe (zt/Lp=1.33). The advancement of
the twin tunnels led to reductions in the end-bearing of the pile
as a result of stress release from 2% volume loss. To
compensate for the decrease in end-bearing resistance, the pile
had to settle substantially to mobilize the shaft resistance along
the entire pile length. This result is similar to those measured in
the field [16]. The end-bearing decreased by 12% and 28%
owing to the first and twin tunnels, respectively.
Fig. 4. Axial load distribution along the pile length
C. Twin-tunneling Induced Bending Moment along the Pile
Figure 5 illustrates the induced bending moment along the
pile after the first and second tunneling in the cases of
zt/Lp=0.67, 1.00 and 1.33. A positive bending moment means
that tensile stress was induced along the pile shaft facing the
first tunnel. The measured bending moment of a single pile
subjected to single tunneling in centrifuge model test [9] was
also included for comparison. Since there was no rigid
constraint at the pile head (pile head was free to move and
rotate), no bending moment was induced at/near the head of the
pile in all the cases. It can be observed that in the case of
zt/Lp=0.67, the maximum positive bending moment was
induced in the pile at Z/Lp=0.6 near spring line of the first
tunnel. Authors in [17] have also reported a measured
maximum bending moment occurring at the spring line of the
tunnel near the shaft of the pile group. This happened because
the pile was subjected to lateral soil movement towards the
tunnel resulting from significant stress release. The magnitude
of the maximum positive bending moment was 400kNm
(which is 50% of the pile BM capacity). On the other hand
negligible bending moment was induced near the pile toe
(below the tunnel spring line), because the tunneling induced
soil movement below the tunnel spring line was insignificant
[2]. Subsequently, the pile was subjected to stress release on
the opposite side of the pile as a result of the second tunnel’s
excavation near the mid-depth of the pile shaft. Consequently,
the induced-bending moment decreased along the pile after the
second tunneling. The induced-bending moment along the pile
was not returned to zero after the second tunneling as the soil
does not behave elastically. Finally, a positive bending moment
0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
100 200 300 400 500 600 700 800 900 1000 1100 1200 1300 1400
N
o
r
m
a
li
z
e
d
d
e
p
th
(
Z
/L
p
)
Axial force: kN
Before tunnelling C/D=1.5
1st tunnel 2nd tunnel
C/D=1.5 C/D=1.5
C/D=1.5 C/D=1.5
C/D=1.5 C/D=1.5
After 1st tunnelling After 2nd tunnelling
zt/Lp = 0.67 zt/Lp = 0.67
zt/Lp = 1.00
zt/Lp = 1.33 zt/Lp = 1.33
zt/Lp = 1.00
Engineering, Technology & Applied Science Research Vol. 10, No. 2, 2020, 5361-5366 5365
www.etasr.com Mangi et al.: Parametric Study of Pile Response to Side-by-Side Twin Tunneling in Stiff Clay
with magnitude of 50kNm was induced. Owing to the soil
inward movement towards the first tunnel when zt/Lp=1.00,
positive bending moment was induced at the lower part of the
pile (0.7≤Z/Lp≤1.0). To counter-balance this induced positive
bending moment, negative bending moment was developed
along the upper part of the pile (0≤Z/Lp≤0.7). The magnitudes
of maximum induced positive and negative bending moment
were 130kNm. The measured induced bending moment along
the pile due to single tunneling in centrifuge reported by [9]
was positive and smaller than the one computed in magnitude.
This may be caused by the tunneling-induced volume loss
which was modeled as 1% in the centrifuge test. The
subsequent tunneling at zt/Lp=1.00 on the opposite side of the
pile reduced the induced bending moment significantly. The
maximum induced bending moment was 37kNm at Z/Lp=0.6
after the twin tunnel excavation.
Fig. 5. Induced bending moment along the pile length
In contrast to the induced bending moment in the case
zt/Lp=0.67, negative bending moment was induced along the
entire pile length due to the first tunnel advancement in the case
of zt/Lp=1.33. However, the magnitude of maximum bending
(100kPa at Z/Lp=0.7) was less than the one in the case
zt/Lp=0.67. Also, no bending was induced at the pile toe and
head. It can be seen that the induced bending moment during
the first tunnel advancement in the case of tunneling near the
mid-depth of the pile (zt/Lp=0.67) and adjacent the pile toe
(zt/Lp=1.00) was less than that in the case of tunneling below
the pile toe (zt/Lp=1.33). Therefore, in the case of zt/Lp=1.33,
the most critical issue to be considered is the relatively large
settlement.
VI. DISCUSSION
It is well recognized that the stress-strain relationship of
soils is highly nonlinear even at very small strains. The
stiffness of most soils decreases as strain increases and depends
on the recent stress or strain history of the soil. Owing to non-
linear soil behavior, a tunnel excavation can cause reduction in
the stiffness of the soil. Therefore, it is vital to investigate the
pile responses not only to the first tunnel but also to the
subsequent tunnel in clay in a twin-tunneling transportation
system. Keeping these issues into consideration, effects of twin
side by side tunneling on a single pile in stiff clay were
investigated in this study. It was revealed that the second
tunneling caused larger settlement and smaller bending
moment in the pile. This finding is the major contribution of
this parametric study.
VII. CONCLUSIONS
Based on the modeled ground conditions, geometry, and
tunneling method, the following conclusions can be drawn:
• Due to the degradation of the stiffness of the clay
surrounding the pile as a result of tunneling-induced stress
release and shear strain, the second tunneling caused larger
settlement than the first in each case. When zt/Lp=0.67 the
second tunneling-induced settlement was the largest (i.e.
175% of Sp from the first tunnel) of all cases. The
cumulative settlements of the pile resulting from the
working load and twin tunneling in the cases zt/Lp=0.67,
1.00 and 1.33 were 25, 39, and 47mm (i.e. 3.1, 4.9, and 5.9
of the pile diameter) respectively.
• The first tunneling in the case zt/Lp=0.67 induced the largest
bending moment (50% of the pile BM capacity) at the
spring line of the tunnel. However, the induced bending
moment along the pile decreased significantly due to the
excavation of the second tunnel in each case because of the
side-by-side twin tunneling configuration in which the
second tunnel caused a stress release on the opposite side of
the pile.
• When of zt/Lp=0.67, the second tunneling-induced soil
movement due to stress release mobilized negative shaft
friction at the upper part of the pile. Consequently, a
downward load transfer was observed along the pile, further
mobilizing the pile end-bearing. Similarly, in the case
zt/Lp=1.00, the load borne by the pile shaft was transferred
to the pile toe as a result of the twin tunneling reduction in
the shaft resistance at the lower part of the pile
(0.3≤Z/Lp≤0.9).
ACKNOWLEDGEMENT
The authors would like to acknowledge the financial
support provided by the Quaid-e-Awam University of
Engineering, Science & Technology, Sindh and Pakistan.
REFERENCES
[1] G. Gudehus, “A comprehensive constitutive equation for granular
materials”, Soils and Foundations, Vol. 36, No. 1, pp. 1-12, 1996
[2] K. Ishihara, “Liquefaction and flow failure during earthquakes”,
Geotechnique, Vol. 43, No. 3, pp. 351-451, 1993
[3] C. W. W. Ng, “The state-of-the-art centrifuge modelling of geotechnical
problems at HKUST”, Journal of Zhejiang University Science A, Vol.
15, pp. 1-21, 2014
[4] J. Jaky, “The coefficient of earth pressure at rest”, Journal of the Society
of Hungarian Architects and Engineers, pp. 355-358, 1944
0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
-150 -100 -50 0 50 100 150 200 250 300 350 400 450
N
o
r
m
a
li
z
e
d
d
e
p
th
(
Z
/L
p
)
Induced bending moment: kNm
Measured Series3
Series11 Series12
1st tunnel 2nd tunnel
C/D=1.5 C/D=1.5
C/D=2.5 C/D=2.5
C/D=3.5 C/D=3.5
After 1st tunnelling After 2nd tunnelling
Meaured (Loganathan et al., 2000)
Computed (This study)
zt/Lp = 0.67 zt/Lp = 0.67
zt/Lp = 1.00
zt/Lp = 1.33 zt/Lp = 1.33
zt/Lp = 1.00
Engineering, Technology & Applied Science Research Vol. 10, No. 2, 2020, 5361-5366 5366
www.etasr.com Mangi et al.: Parametric Study of Pile Response to Side-by-Side Twin Tunneling in Stiff Clay
[5] K. H. Chiang, C. J. Lee, “Responses of single piles to tunneling-induced
soil movements in sandy ground”, Canadian Geotechnical Journal, Vol.
44, No. 10, pp. 1224-1241, 2007
[6] A. Niemunis, I. Herle, “Hypoplasic model for cohesionless soils with
elastic strain range”, Mechanics of Cohesive-frictional Materials, Vol. 2,
No. 4, pp. 279-299, 1997
[7] C. W. W. Ng, M. A. Soomro, Y. Hong, “Three-dimensional centrifuge
modelling of pile group responses to side-by-side twin tunnelling”,
Tunelling and Underground Space Technology, Vol. 43, pp. 350-361,
2014
[8] C. W. W. Ng, H. Lu, S. Y. Peng, “Three-dimensional centrifuge
modelling of the effects of twin tunnelling on an existiong pile”,
Tunneling and Underground Space Technology, Vol. 35, pp. 189-199,
2013
[9] N. Loganathan, H. G. Poulos, D. P. Stewart, “Centrifuge model testing
of tunnelling-induced ground and pile deformations”, Geotechnique,
Vol. 50, No. 3, pp. 283-294, 2000
[10] Abaqus user’s manual, version 6.8.2, Dassault Systemes, 2008
[11] D. Masin, “Hypoplastic cam-clay model”, Geotechnique, Vol. 62, No. 6,
pp. 549-553, 2012
[12] A. M. Marshall, R. J. Mair, “Tunneling beneath driven or jacked end-
bearing piles in sand”, Canadian Geotechnical Journal, Vol. 48, No. 12,
pp. 1757-1771, 2011
[13] S. W. Jacobsz, J. R. Standing, R. J. Mair, T. Hagiwara, T. Sugiyama,
“Centrifuge modelling of tunnelling near driven piles”, Soils and
Foundations, Vol. 44, No. 1, pp. 49-56, 2004
[14] P. W. Mayne, F. H. Kulhawy, “K0-OCR relationships in soils”, Journal
of the Geotechnical Engineering-ASCE, Vol. 108, No. 6, pp. 851-872,
1982
[15] D. A. Mangejo, M. A. Soomro, N. Mangi, I. A. Halepoto, I. A. Dahri,
“A parametric study of effect on single pile integrity due to an adjacent
excavation induced stress release in soft clay”, Engineering, Technology
& Applied Science Research, Vol. 8, No. 4, pp. 3189-3193, 2018
[16] M. A. Soomro, A. S. Brohi, D. K. Bangwar, S. A. Bhatti, “3D numerical
modelling of pile group responses to excavation-induced stress release in
Silty Clay”, Engineering, Technology & Applied Science Research, Vol.
8, No. 1, pp. 2577-2584, 2018
[17] D. R. Coutts, J. Wang, “Monitoring of reinforced concrete piles under
horizontal and vertical loads due to tunnelling”, International Conference
on Tunnels and Underground Structures, Singapore, November 26-29,
2000