BURIED HDPE PIPELINES SUBJECTED TO NORMAL FAULTING-
A CENTRIFUGE INVESTIGATION
by
Tarek H. Abdoun, Da Ha, Michael J. O’Rourke, Michael D. Symans, Thomas D. O’Rourke;
Michael C. Palmer, and Harry E. Stewart
Affiliation:
Tarek H. Abdoun , Associate Professor, E-mail: abdout@rpi.edu
Da Ha, Doctoral Student, E-mail: had@rpi.edu
Michael J. O’Rourke, Professor, E-mail: orourm@rpi.edu
Michael D. Symans, Associate Professor, E-mail: symans@rpi.edu
Department of Civil and Environmental Engineering
Rensselaer Polytechnic Institute, Troy, NY 12180-3590, USA
Thomas D. O’Rourke, Professor, E-mail: tdo1@cornell.edu
Michael C. Palmer, Research Associate, E-mail: mcp5@cornell.edu
Harry E. Stewart, Associate Professor, E-mail: hes1@cornell.edu
School of Civil and Environmental Engineering
Cornell University, Ithaca, NY 14853-3501, USA
Corresponding Author:
Da Ha
JEC4049
Department of Civil and Environmental Engineering,
Rensselaer Polytechnic Institute
110 8th Street
Troy, NY 12180-3590, USA
Phone: 518 276 2064
Fax: 518 276 4833
Email: had@rpi.edu
1
Abstract
Permanent ground deformation (PGD) is arguably the most severe hazard for continuous buried pipelines. This paper presents results from four centrifuge tests designed to investigate the
differences in behaviors of buried HDPE pipelines subject to normal and strike-slip faulting.
Two centrifuge tests with different instrumentations but the same configurations were conducted
on each case of ground faulting (i.e. two tests on normal faulting and the other two tests on
strike-slip faulting). The tests results show that, as expected, the pipeline behavior is asymmetric
under normal faulting and symmetric under strike-slip faulting.
In the case of strike-slip fault-
ing the soil-pipe interacting pressure distribution is symmetric about the fault. While in the case
of normal faulting there is a pressure concentration close to the fault trace on the up-thrown side,
with much lower soil-pipe interacting pressures at other locations on the pipe.
The soil-pipe interacting force versus deformation relationship (i.e. p-y relationship) was obtained based on the experimental data. The p-y relationships for both the strike-slip and normal
faulting cases were also compared with the relationships suggested by the ASCE Guidelines
(1984). It was found that for the case of strike-slip faulting the experimental p-y relationship is
generally compatible to the ASCE Guidelines suggested values. In contrast, for the case of
normal faulting the experimental p-y relationship is much softer than the ASCE Guidelines suggested values. This is due to the fact that the “strip footing failure” assumption adopted by the
ASCE Guidelines does not reflect the real 3-D failure in the normal faulting case.
Keyword
Buried pipelines, Normal faulting, Strike-slip faulting, P-y relationship, Permanent ground deformation (PGD), Centrifuge testing
2
Introduction
There are three types of earthquake faults: strike-slip faults, normal faults and their combination- oblique faults. Damages to buried pipes at these abrupt ground offsets have frequently been
observed in the past earthquakes. For example, movement of the San Andreas Fault was responsible for extensive damage to water transmission lines during the 1906 San Francisco earthquake
(O’Rourke and Hamada, 1992). Landslides caused by the 1987 Ecuador earthquake were responsible for severe damage of the 0.66-m (26-in.) diameter Trans-Ecuadorian oil pipeline over a distance of 40 km, resulting in the loss of 60% of the export revenue of that country (O’Rourke et
al., 1987). During the 1999 Chi-Chi Earthquake in Taiwan, surface faulting with 4.0 m of vertical
movement caused severe damage to buried pipelines (EERI, 1999 a). An originally straight water
pipe was bent 90 degrees and buckled by the large force exerted by permanent ground distortion.
During the 1999 Izmit Earthquake in Turkey, a 2.2-m diameter steel pipe which was subjected to
about 3 m of strike slip fault offset, suffered leaks during the earthquake (EERI, 1999 b). The
orientation of the steel pipe with respect to the fault placed the line in compression during fault
offset and the resulting axial compression combined with flexure buckled the pipe at two locations.
In recognition of the severe consequence of this kind of earthquake hazard, pipe performances
under seismic faulting have been the subject of numerous research investigations. Experimental
work by Audibert and Nyman (1977) and Trautmann and O’Rourke (1983) and numerical/analytical work by Newmark and Hall (1975), Kennedy et al. (1977) and O’Rourke et al.
(1999) are examples. More recently, a centrifuge based experimental method was first implemented by O’Rourke et al. (2003) using the Rensselaer Geotechnical Centrifuge and a split container with inside dimension 1.0 m x 0.354 m x 0.203 m. These tests were successful in the sense
3
that the experimental equipment functioned well and the recorded strains were generally in good
agreement with those predicted by numerical simulations using finite element models.
In the centrifuge tests reported in this paper, the fault offset was simulated using a recently
updated split-box container (inside dimension: 1.14 m x 0.76 m x 0.20 m). The container is capable of simulating both vertical and horizontal offset in flight. Additional information about the
split container is presented by Ha et al. (2006).
Centrifuge Modeling of HDPE Pipe Response to PGD
All tests reported herein were conducted on HDPE pipe that complies with AWWA Standard
C901 (AWWA, 2003) for water service. The pipe has an outer diameter, D = 33.4 mm and a wall
thickness t = 1.96 mm, SDR (Standard Dimension Ratio) = D/t = 17. All the centrifuge tests were
carried out at a gravity level of 12.2 g, hence the model pipe geometry simulates a prototype pipe
with D = 407.5 mm and t = 24.0 mm.
Table 1 and Figure 1 summarize the four centrifuge experiments which were designed to
evaluate the behavior of buried pipelines subjected to normal versus strike-slip faulting. Note the
initial pipe-fault angle indicated in Table 1 is in plan for the strike-slip faulting case and is in elevation for the normal faulting case. Figure 1 presents the configurations of the Rensselaer splitbox container and the high density polyethylene (HDPE) experimental pipeline before and after
fault offset. Figure 1 (a) shows the configuration of strike-slip faulting, which corresponds to
models 1 and 2. Figure 1 (b) applies to models 3 and 4, which simulated normal faulting. The
nominal perpendicular initial orientation results in significant flexural strain as well as minor axial strain. As indicated in Table 1, during the test the movable portion of the container was offset
vertically 0.48 m for the normal faulting tests (tests 1 and 2) and was offset horizontally 1.06 m
4
for the strike-slip tests (tests 3 and 4) in prototype scale. It should be noted that the maximum
possible vertical offset in the split container is about 4 cm which corresponding to a prototype
offset of 0.48 m at 12.2 g.
To monitor pipeline response to fault offset, two types of instrumentation were used. In models 1, the HDPE pipe was instrumented with strain gages along the pipe springlines which measure the total longitudinal strain distribution on both the active and passive sides of the pipe. In
model 3, strain gages were instrumented at the top and bottom centerline of the pipe to capture
both the axial sand flexural strains due to offsets in the vertical plane. In this paper, the “longitudinal” strains are considered to be a combination of axial and flexural strains, where “axial”
strains are those due to direct axial tension or compression and “flexural” strains are due to
transverse bending in the horizontal plane for strike-slip faulting and bending in the vertical
plane for normal faulting. Axial strains were calculated as the average of the strain at opposite
springlines. Bending strains were calculated as one-half the difference between the longitudinal
strains at opposite sides. In models 2 and 4, the pipe was instrumented with two tactile pressure
sensor sheets manufactured by TEKSCAN Inc. The sensor sheets were wrapped around the test
pipes for a longitudinal distance of 0.25 m at model scale (3.0 m in prototype) on either side of
the fault. The pressure sensor sheet measures the normal pressure at the soil-pipe interface.
The tactile pressure sensors were calibrated using a sensor calibrator from TEKSCAN Inc.
(model # 5250). The tactile pressure sensor was calibrated at two pressure levels, 69 kPa (10 psi)
and 207 kPa (30 psi). The pressure sensor control software uses the two calibration points to get
calibration factors for other pressure levels, using nonlinear interpolation.
The pipeline was pin-connected to the split container end walls. As such, the centrifuge tests
simulate the case where a thrust or anchor block is located at the end wall. The tension force at
5
the end of the pipe was measured by a strain gage instrumented connecter. Using the material
(aluminum alloy) properties, cross-sectional area and the measured strain in the connector, the
force can be readily calculated.
Grain size effects on soil-structure interaction are an important issue in centrifuge modeling.
In this study the backfill material was processed so as to guarantee that no significant grain size
effect was expected. The soil used in the Rensselaer centrifuge tests were sieved from a glaciofluvial, well graded sand to produce sand suitable for centrifuge testing. Soil passing the #40
sieve (0.42 mm), but retained on the #200 sieve (0.075 mm) was used in the testing. The sieving
process resulted in very uniform sub-angular or sub-rounded quartz grains with an average grain
size diameter of D50 = 0.29 mm. Hence, a pipe outer diameter to average grain size ratio D/D50 =
115 was reached, which satisfied the criterion of D/D50 ≥ 48 recommended by the International
Technical Committee TC2 (2005) based on the centrifuge test data from Ovesen (1981) and
Dickin and Leung (1983). Table 2 lists the properties of the sand used in this series of centrifuge
tests.
The sieved sand was placed at a water content (wc) of about 4.0 ~ 4.5%. This moisture content
was chosen to replicate expected pipe burial condition. In addition, full-scale tests by researchers
at Cornell University (Turner, 2004; O’Rourke and Turner, 2006) have suggested that moist sand
may have a significantly higher lateral resistance than that of dry sand. However, more recent
work (Ha, 2007) has shown that there is little difference between the expected lateral resistance
for dry sand backfill and that with water content of about 4.0%.
The moist sand was placed and compacted in layers to a dry unit weight of 14.7 kN/m3 (internal peak friction angle = 40o, determined from direct shear test) and to a depth Hc of 1.12 m (in
the prototype scale) above the center line of the pipe (Hc/D = 2.8). The sand was compacted us-
6
ing an aluminum compacter (weight = 18.7 N, width = 0.157 m, length = 0.310 m). The thickness of each compacted lift is about 2.5 cm (1 inch). The compacting is equivalent to raising the
compacter to a height of 10 cm and dropping it 10 times. Hence the input energy is about
18.7*0.1*10/ (0.157*0.310) = 384.2 J/m2.
Post-Offset Observation
In the centrifuge tests, observations were made of the ground surface. Very different ground
surface deformation patterns were observed as shown in Figure 2. For strike-slip faulting, there
was soil disturbance within a narrow band at the fault. There were also large cracks nominally
parallel to the pipe on the passive side. That is, in Figure 2a there is a ground crack far to the left
of the pipe in the top portion of the photo and another crack far to the right of the pipe in the
lower portion of the photo (cracks indicted by thinner dotted lines). The cracks appear to be the
surface expressions of passive soil wedges. Additional cracks also appear between the outermost
surface expression of the passive soil block and the trace of pipeline, likely due to soil heaving
during the soil wedge formation process. Soil subsidence zones were observed on the active side
(i.e. soil in active condition) of the pipe.
For the normal faulting case, there is a big crack right at the fault trace due to the offset itself.
Soil cracks were also observed above the trace of the pipeline on the down-thrown side formed
as the pipeline plowing upward through the relatively weak over burden soil. Additional secondary soil cracks more or less parallel to the fault trace on the down-thrown side were observed
which were likely formed due to soil shear failure initiated by the fault offset itself. The secondary cracks nominally parallel to the fault result in a relatively wider zone of soil disturbance for
normal faulting, in comparison to the relatively narrow band of soil disturbance for strike-slip
faulting. Note the shape of the deformed pipe is indicated in Figure 2 as thicker dashed lines.
7
This was made possible by using the surface grid to locate the center and the two ends of the test
pipe.
Pipe Axial and Bending Strains
Figure 3 shows a comparison of the measured axial and bending strains for the normal
faulting and strike-slip faulting cases, respectively. Note due to limitations of the testing equipment, the normal faulting offset cannot exceed 0.48 m in prototype. As the axial strains are generally small, the pipe material is nearly linearly elastic. Note that the axial strains for strike-slip
faulting are nominally symmetric with respect to the fault. Also beyond about 2 m from the fault,
there is a linear decrease in axial strain with distance from the fault. This is consistent with a
constant longitudinal friction force per unit length at the soil-pipe interface. The peak axial strain
again for the strike-slip case is located at around ±1.0 m from the fault. In contrast, for the normal faulting case the axial strain distribution is not symmetric. Axial strains are substantially
larger on the up thrown side with the peak axial strain being located around -1.0 m from the fault.
For a given strike-slip offset, the bending strain are consistent with double curvature bending,
concave on one side of the fault and convex on the other. In contrast, the bending strain distribution is not symmetric for the case of normal faulting. Bending strains are substantially larger on
the up thrown side with the peak value located about -1.0 m from the fault, which is the approximate location of the peak axial strain.
The peak axial and bending strains were plotted versus the fault offsets (Figure 4). Note that
the peak axial strain vs. fault offset plots for both test 1 ( = -85o) and test 3 ( = 90o) are more
or less on the same curve. All the axial strains measured from test 1 ( = -85o) are less than 1.0%
and are roughly proportional to the offset. In test 3 ( = 90o) there is a slight deviation from linear
behavior (slight hardening) for the level of fault offset larger than 0.3 m.
8
On the other hand, the peak bending strain vs. fault offset plots for tests 1 and 3 are not the
same curve. For the normal faulting case (test # 3), the peak bending strains on the up-thrown
side are larger than those on the down-thrown side. For fault offsets larger than about 0.3 m, the
peak bending strain in the strike-slip case is roughly the average of the normal fault up-thrown
and down-thrown sides’ values. In both tests 1 and 3, the measured peak bending strains are located at about 1.2 m from the fault. For the normal faulting test (test 3), there was a slight deviation from linear behavior (slight softening) in the peak bending strain vs. fault offset curve at
fault offset levels used. However, for the strike-slip faulting test (test 1) the peak bending strain
reached a plateau with more or less a constant value when the fault offset exceeding 0.7 m. As
will be shown later, the passive soil failure wedge in the strike-slip faulting tests forms at an offset of roughly 0.7 m. That is, when the soil in the passive zone fails, the pipe continue to move
through the soil but the peak soil-pipeline interaction pressure as well as the pipe bending strain
remain nominally constant.
Soil-Pipe Interaction
In this section the measured pipe lateral force was calculated from the readings from the tactile
pressure sensor. The detailed information about tactile pressure sensor data interpretation technique was reported in a separate paper (Ha et al., 2007).
ASCE Guidelines (1984) suggests the following equation (see Eq. 2) for the calculation of
maximum lateral force on pipeline buried in sand in horizontal transverse movement:
Pu  HN qh D
(1)
where,  = effective unit weight of soil backfill; H = depth to springline of pipe; Nqh = horizontal bearing capacity factor for sand; D = pipe outer diameter.
9
In the current centrifuge testing case, = 14.7*(1+4%) = 15.3 kN/m, Nqh = 8.5 for  = 40o and
H/D= 2.8, H= 1.12 m, and D = 0.41 m. Hence,
Pu  HN qh D  15.3 *1.12 * 8.5 * 0.41  59.0 kN / m
The tactile pressure sensor data for the strike-slip faulting case were shown in Figure 5. A
symmetric pressure distribution (transverse horizontal) was developed in strike-slip faulting. The
peak pipe lateral force was concentrated at about ±0.3 m to ±0.6 m from the fault trace.
The observed peak value of the lateral force per unit length (near fault at the maximum offset)
for the strike-slip test (58.0 kN/m for friction coefficient  = 0) are reasonably close to the ASCE
(1984) value. It should be noted that the lateral pressure data shown in Figure 5 was directly calculated based on the tactile pressure sensor data. Since the tactile pressure sensor only measure
the normal pressure, the soil-pipe interface friction was not taken into account. A friction coefficient  = 0.4 is a reasonable value for soil-HDPE interface (O’Rourke et al, 1990). By taking
friction into consideration, the peak lateral force becomes 78.0 kN/m (Ha et al., 2007).
The tactile pressure sensor data for the normal faulting case are shown in Figure 6. In contrast
to the strike-slip faulting, it is clear that there is a concentration of transverse (vertical) pressure
on the pipe on the up-thrown side around -0.8 m from the fault. This is consistent with the bending strains in that the largest bending strains are on the up-thrown side and are located about 0.8
m from the fault. Also note that the axial strains are largest on the up-thrown side.
ASCE Guidelines (1984) suggests the following equation for the calculation of maximum lateral force on relative downward moving pipeline buried in sand:
1
qu  HN q D  D 2 N y
2
(2)
10
where,  = total unit weight of soil backfill; H = depth to springline of pipe; Nq and Ny = bearing
capacity factor for horizontal footing on sand loaded in the vertically downward direction; D =
pipe outer diameter.
In the current centrifuge testing case with H= 1.12 m and D = 0.41 m, for moderate dense sand
of 4% moisture content ( = 40o), = 15.3 kN/m, Nq = 65 and Ny = 80. Hence,
qu  15.3 *1.12 * 65 * 0.41  0.5 *15.3 * 0.412 * 80  560 kN / m
This value is much larger (about 10 times) than the measured maximum transverse force.
However, as shown in Figure 4, the measured peak vertical force has apparently not reached its
maximum value.
Figure 7 shows the peak pipe lateral force versus the fault offset relations. For both the normal
faulting and strike-slip faulting cases the relations are on the same curve for fault offset up to
about 0.4 m. For the case of strike-slip faulting, the peak lateral force started to decrease when
the fault offset reached 0.7 m. For the case of normal faulting, no decrease of peak lateral force
was observed for the fault offset range (i.e. up to 0.48 m) used in the current test.
The ASCE Guidelines (1984) also suggested the maximum elastic deformation yu and zu for
the horizontal and vertical transverse movement, respectively. For the current case of dense sand,
yu  (0.02 ~ 0.03)( H c 
D
0.41
)  0.025 * (1.12 
)  0.033
2
2
zu  (0.10 ~ 0.15) D  0.125 * 0.41  0.05
(3)
(4)
The equivalent elastic soil spring coefficient is simply the maximum resistance divided by one
half of the maximum elastic deformation, for example 2qu/yu for horizontal transverse case.
The following relation exists between bending strain and curvature:
11
b  c
d2y
dx 2
(5)
where b = extreme fiber bending strain, c = distance to the extreme fiber (outside radius for the
circular pipe specimen), y = pipe deflection and x = distance along the pipe. Note that Eq. 3 assumes that the transverse sections of the pipe remain plane after bending. Hence, the deflection
of the pipe perpendicular to the pipe longitudinal axis is obtained by double integration of the
bending strain:
 

y     b dx  dx
 c 
(6)
As the measured pipe strains are relatively small (less than 2.5% for both strike-slip and normal cases), it is reasonable to assume the pipe outer radius remains constant during offset. The
pipe lateral deformation at the locations of peak lateral resistance was calculated by using Equation 6 with the bending strain data fitted using cubic-spline interpolation. The resulting “p-y” relationship compared with the ASCE Guideline values are shown in Figure 8 and 9. Note the plotted centrifuge p-y relationships are obtained at the locations of peak lateral resistance.
Note the peak force values in Figures 8 and 9 have included the influence of friction, which
resulting in an increase of about 30% in the values of peak resistance. For the strike-slip case, the
measured peak resistance (78 kN/m) is higher than the ASCE value (59 kN/m) by 32% (see Figures 8). The measured p-y relationship is softer than the ASCE relationship. This can be explained by the fact that in the centrifuge model preparation the soil adjacent to the pipe cannot be
compacted very well due to the presence of sensors. Although this under compaction of sand
backfill surrounding testing pipe results in reduced initial stiffness of the p-y relationship, the ultimate soil resistance is compatible with the ASCE Guidelines (1984).
12
For the normal faulting case, the measured peak resistance is much lower (about 1/8 of the
value) than the ASCE value. However, due to the limitation of testing equipment on the normal
faulting magnitude, the peak resistance did not reach a plateau. The measured p-y relationship is
also much softer than the ASCE relationship. This is because the ASCE Guidelines equation is
based on the assumption that the maximum load on pipe is equal to the strip footing bearing capacity of the soil. However, in real normal faulting the soil deformation is not a 2-D case. The
soil on the up-thrown side is separated from the soil on the down-thrown side. This discontinuity
of the ground on the up-thrown side greatly reduces its bearing capacity since the soil close to the
fault on the up-thrown side is already in active mode.
Summary and conclusions
Four centrifuge tests were carried out to investigate the differences in behavior of buried pipelines subject to normal faulting and strike-slip faulting. For both fault faulting modes (i.e. normal
and strike-slip), the pipe model is pinned to the split container walls, simulating a prototype with
a thrust block somewhere near the fault. The axial and bending strains in the pipe were measured
during simulated fault offset. For the test geometry and boundary conditions, the pipe axial strain
is nominally a linear function of the fault offset. However, both the axial strain and the bending
strain distribution are not symmetric in the normal faulting case in contrast to the symmetric distribution in the strike-slip faulting case. For normal faulting, the peak axial and bending strains
are concentrated on the up-thrown side of the pipe. In the strike-slip faulting, the peak bending
strain has nominally an elasto-plastic type of behavior with respect to the fault offset.
Pipe lateral force was measured by using the tactile pressure sensor. The measured pipe lateral
force distributions are consistent with the strain gage measurements. That is, for the normal fault-
13
ing case the pipe lateral force are not symmetric and there is a concentration of lateral force on
the up-thrown side of the pipe. For the strike-slip faulting case the pipe lateral force are symmetric.
The peak lateral force are compared with the ASCE Guidelines (1984) suggested values. The
experimental data was also used to develop the soil-pipe p-y relationship. The results were compared with the ASCE Guidelines (1984) suggested relationships. For strike-slip faulting case, the
experimental p-y is softer than the ASCE value possibly due to the experimental model preparation. The ASCE suggested peak lateral force values are comparable to the measured ones. For the
case of normal faulting, both the ASCE suggested peak lateral force values and the stiffness of
the p-y relationship are much higher than the measured ones. This is due to the fact that in normal faulting, the soil deformation is in a 3-D condition instead of the 2-D assumption adopted in
the ASCE Guidelines (1984).
Acknowledgment
This work was supported primarily by the George E. Brown, Jr. Network for Earthquake Engineering Simulation (NEES) Program of the National Science Foundation under Grant Nos. CMS0421142, CMS-0086555, and CMS-0217366. Any opinions, findings and conclusions or recommendations expressed herein are those of the authors and do not necessarily reflect the views
of the National Science Foundation.
This project is part of a collaborative project involving
full-scale buried pipe tests at Cornell University and companion centrifuge tests at Rensselaer.
Dr. Yun Wook Choo from Korea Advanced Institute of Science and Technology (KAIST) provided assistance in the preparation of centrifuge models and in conducting the tests. The authors
14
also acknowledge lab support personnel at the Rensselaer Geotechnical Centrifuge Center and at
the Cornell University Lifeline Facility for their help in conducting tests presented in this paper.
References
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C901-02, January 2003.
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Ha D., Abdoun T., O’Rourke M., Symans M., O’Rourke T., Palmer M. & Stewart H. “Centrifuge
Modeling of Permanent Ground Deformation Effect on Buried HDPE Pipelines”, ASCE Journal of Geotechnical and Geoenvironmental Engineering (Accepted).
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16
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List of Symbols
c
= the outer radius of test pipe
Cc = coefficient of curvature
Cu = coefficient of uniformity
D
= pipe outer diameter
D50 = average particle size of sand backfill
H
= depth of soil from the surface to the center of the pipe
Nqh = horizontal bearing capacity factor for sand
p
= soil resistance in horizontal transverse direction
pu
= maximum soil resistance in the horizontal transverse direction
qu
= maximum soil resistance in the vertical transverse direction
t
= pipe wall thickness
yu
= maximum elastic deformation in the horizontal transverse direction
zu
= maximum elastic deformation in the vertical transverse direction
f
= maximum fault offset

= pipeline-fault orientation angle (strike-slip faulting)
  pipeline-fault orientation angle (normal faulting)

effective unit weight of soil backfill
′
= dry unit weight of soil backfill
b bending strain in test pipe
17
  friction coefficient
  friction angle of sand backfill
18
Table 1 Summary of Test Models Used in Centrifuge Testing
Model
Number
Initial Pipe-Fault Angle
Instrumentation
(degrees)
Offset Rate*
(m/min)
Peak Offset*
(m)
1
2
3
4
-85(SS)
-85(SS)
90(N)
90(N)
0.318
0.318
~0.318
~0.318
1.06
1.06
0.48
0.48
Note:
Strain Gage
Pressure Sensor
Strain Gage
Pressure Sensor
SS- Strike-Slip
N- Normal
*All dimensions in prototype scale
19
Table 2. Material Properties for Sand Backfill
Soil Properties
′Dry Unit Weight (kN/m3)
, Friction Angle (deg)
D50, average particle size (mm)
Cu, coefficient of uniformity
Cc, coefficient of curvature
Value
14.7
40
0.29
1.55
1.0
20
List of Figures:
Figure 1 Sketch of the centrifuge model before and after offset (dimensions in model scale)
Figure 2 Post-test surface observations of the two test setups
Figure 3 Axial and bending strain distributions in tests 1 and 3
Figure 4 Peak axial and bending strains versus fault offset
Figure 5 Measured lateral force distribution along the pipe for the strike-slip faulting caseassuming no friction (= 0)
Figure 6 Measured lateral force distribution along the pipe for the normal faulting caseassuming no friction (= 0)
Figure 7 Peak lateral forces versus fault offset
Figure 8 P-y relationship for the horizontal transverse (strike-slip faulting) case
Figure 9 P-y relationship for the vertical transverse (normal faulting) case
21
Split Container
0.04m
Fixed Portion
Fixed Portion
0.57m
85o
Movable
Portion
Test Pipe
Movable
Portion
0.04m
0.04m
Before Offset
Plan View
a)
Split Container
0.04m
Deformed Pipe
After Offset
-85o strike-slip faulting test setup, plan view
Test Pipe
Deformed Pipe
Down Thrown Side
0.08m
Fixed Side
0.57m
Fixed Side
Down Thrown Side
Before Offset
Elevation View
0.57m
0.04m
After Offset
b) 90o normal faulting test setup, elevation view
Figure 1 Sketch of the centrifuge model before and after offset (dimensions in model scale)
22
(a)  = -85o
(b)  = 90o
Figure 2 Post-test surface observations of the two test setups
23
1.5
 = 90o
Normal
 = -85o
Strike-Slip
Axial Strain (%)
1
0.5
0
0.12 m
0.24 m
0.48 m
0.73 m
1.06 m
-0.5
2
Bending Strain (%)
 = -85o
Strike-Slip
 = 90o
Normal
1
0
-1
-2
-6
-4
-2
0
2
4
6
-6
-4
Distance from Fault (m)
-2
0
2
4
6
Distance from Fault (m)
Figure 3 Axial and bending strain distributions in tests 1 and 3
24
1.2
3
Up- thrown Side, = 90o
Normal, = 90o
Peak Bending Strain (%)
Peak Axial Strain (%)
Strike-Slip, = -85o
0.9
0.6
0.3
0
Down- thrown Side, = 90o
2.4
Strike- slip, = -85o
1.8
1.2
0.6
0
0
0.4
0.8
1.2
0
Fault Offset (m)
0.4
0.8
1.2
Fault Offset (m)
Figure 4 Peak axial and bending strains versus fault offset
25
Force Distribution during Offset
80
1.06 m
0.73 m
0.49 m
0.24 m
0.12 m
0.00 m
Force Distribution (kN/m)
60
40
20
0
-20
Pipeline
-40
Fault
Strike-Slip
 = 85o
-60
-80
-4
-2
0
2
4
Distance from Fault (m)
Figure 5 Measured lateral force distribution along the pipe for the strike-slip faulting case- assuming no friction (= 0)
26
Force Distribution during Offset
20
Down-Thrown Side
Force Distribution (kN/m)
Up-Thrown Side
10
0
-10
Pipeline
Fault
-20
-30
0.48 m
0.24 m
0.12 m
0.00 m
Normal Faulting
 = 90o
-40
-50
-4
-2
0
2
4
Distance from Fault (m)
Figure 6 Measured lateral force distribution along the pipe for the normal faulting case- assuming
no friction (= 0)
27
80
Normal, = 90o
Peak Force (kN/m)
Strike-Slip, = -85o
60
40
H/D = 2.8
D = 0.41 m
20
0
0
0.4
0.8
1.2
Fault Offset (m)
Figure 7 Peak lateral forces versus fault offset
28
100
Horizontal Transverse
Peak Force (kN/m)
80
60
40
Centrifuge
ASCE
20
0
0
0.1
0.2
0.3
0.4
0.5
Relative Displacement (m)
Figure 8 P-y relationship for the horizontal transverse (strike-slip faulting) case
29
600
Peak Force (kN/m)
500
Vertical Transverse
400
300
Centrifuge
ASCE
200
100
0
0
0.02
0.04
0.06
0.08
0.1
Relative Displacement (m)
Figure 9 P-y relationship for the vertical transverse (normal faulting) case
30