1
Experimental Investigation of Longitudinal Bending of Buried
2
Steel Pipes Pulled through Dense Sand
3
Mohamed Almahakeri1; Amir Fam2; Ian D. Moore3
4
Affiliation:
5
1
Graduate Student
6
2
Professor and Canada Research Chair in Innovative and Retrofitted Structures
7
3
Professor and Canada Research Chair in Infrastructure Engineering
8
Dept. of Civil Engineering, Queen’s University, Kingston, ON K7L 3N6
9 10 11 12 13 14 15 16 17 18 19 20
Corresponding Author: Ian D. Moore Ellis Hall, Room 249 Queen’s University 58 University Ave. Kingston, On, Canada K7L 3N6 Phone: (613) 533-3160 Email:
[email protected]
21 22 23
Abstract
24
North America is traversed by many high pressure oil and gas transmission pipes, and the
25
stability of that essential buried infrastructure must be maintained under a variety of earth
26
loading conditions. In this study, a series of pipe bending experiments have been conducted on
27
105 mm (4.1 inch) outside diameter and 1830 mm (6 feet) long steel pipes buried in dense sand
28
placed in a 4000x2000x2000 mm (157.5x78.7x74.7 inch) test pit. The pipe ends were pulled by 1
29
two parallel cables attached to a spreader beam outside the test region, which was pulled by a
30
hydraulic actuator. The study investigated burial depth-to-diameter (H/D) ratios of 3, 5 and 7 as
31
well as two horizontal extents for the soil behind the pipe distances of 3D and 9.5D. Special
32
consideration was made to assess the influence of friction between the pulling cables and soil. It
33
was shown that this friction is significant and may contribute about 20% of the maximum pulling
34
load for the case of H/D = 3. Consistency of results was established using four test repetitions
35
for some cases. While the horizontal extent of soil behind the pipe tested in this study had an
36
insignificant influence on the pulling forces, the burial depths significantly influenced the
37
ultimate pulling forces for the system. The failure mechanism controlling the limiting pulling
38
force in these tests was consistently governed by soil, and the pipe remained elastic.
39 40
Key words: steel pipe, buried pipe, soil, bending, ground movement
41 42
INTRODUCTION
43
Buried energy pipelines cross zones of soil instability and may need to be designed to
44
resist lateral or other forces resulting from potential soil movements. Relative pipe-soil
45
movements can result from natural phenomena such as soil creep, slope failures, landslides, and
46
earthquakes-induced faults. Soil loading may lead to pipeline failures which could negatively
47
affect the surrounding environment, local or global economy, and may even be a source of
48
catastrophic safety hazard if flammable contents ignite in populated areas. Pipeline failures due
49
to geotechnical causes are one of the main categories identified by several pipeline regulatory
50
and operation-monitoring agencies around the globe. In Canada, the National Energy Board
51
(NEB, 2008) reported that 7% of pipeline failures are due to geotechnical causes. In the USA,
52
the Pipeline and Hazardous Materials Safety Administration (PHMSA, 2010) reports that 2.5% 2
53
of significant pipeline incidents are due to earth movements. For Europe, the European Gas
54
pipeline Incident data Group (EGIG, 2008) reported that 7% of all pipeline failures were caused
55
by ground movements. Further statistical analysis of the frequency of failure data presented by
56
(EGIG, 2008) on nine pipe size categories (covering pipe diameter sizes ranges from below 102
57
mm (4 inch), up to above 1219 mm (48 inch)) yields percentages of 21% and 26% of the failures
58
being for pipe diameters of 0-102 mm (0-4 inch), and 127-254 mm (5-10 inch), respectively.
59
Such percentages make these two pipe size categories at a greater risk of failure than those in the
60
seven higher remaining size categories surveyed. Flexural behavior of buried pipes due to lateral
61
earth movement has rarely been studied, with few full-scale experimental data sets available in
62
the public domain. Some design guidelines (e.g. the American Society of Civil Engineers
63
(ASCE, 1984), and the Pipeline Research Council International (PRCI, 2004)) provide suggested
64
relations to estimate the lateral loads imposed on a buried pipeline when subjected to lateral earth
65
movement. However, these formulations are based on plane strain conditions, which assume that
66
the pipe moves as a rigid body with no flexure of the pipe. While most of the previous research
67
work has been focusing on the force-displacement characteristics, little experimental work has
68
been conducted to examine the flexural behavior of buried steel pipes (e.g. Konuk et al., 1999).
69
The objective of the current study is to measure strains and deflections along pipes at various
70
burial depths, the progressive development of deflected pipe shape, and the manner in which
71
different lateral displacements at different positions along the pipe influence the development of
72
lateral force (and peak lateral force in particular). As a start, to achieve these goals, performance
73
of test setup and consistency of results are assessed, which involves explicit evaluation of
74
friction on the pulling mechanism, the boundary effects, and reproducibility of replicate tests.
75
The experimental testing program is conducted on small diameter (102 mm (4.0 in) nominal
76
diameter, D , and 105 mm (4.1 in) for outside diameter, Do) buried steel pipes of 2.7 mm wall 3
77
thickness, subjected to lateral earth loading, to understand their flexural behavior and the
78
resistance of the overall pipe-soil system. Nine tests on steel pipes buried in dense sand were
79
conducted using the geotechnical laboratory at Queen’s University. Steel pipes with length-to-
80
diameter ratio of 18 were pulled laterally at the pipe ends. Force-displacement curves, strains,
81
and deflected profiles of the pipes were monitored and recorded during the tests. Also, the initial
82
and final positions of the pipe and the soil surface deformations were surveyed before and after
83
each test.
84
LITERATURE REVIEW
85 86
Centrifuge Modeling
87
Centrifuge modeling has proven to be a very useful tool for studying pipe-soil interaction
88
simulating large diameter pipes and deep burial depths in a relatively small and cost-effective
89
experimental setting. Several studies have been conducted examining the relative movement
90
between pipe and soil, for both lateral and oblique movement directions. Paulin et al. (1995)
91
conducted centrifuge modeling tests investigating the effects of burial depth, displacement rate,
92
and trench width for rigid pipes buried in clay. O’Rourke et al. (2005) examined axial and
93
bending strains in buried aluminum pipes with different diameter to wall thickness (D/t) ratios
94
under lateral earth fault loading, simulating behavior of larger diameter steel pipes with
95
equivalent D/t ratios. Da et al. (2008) investigated the influence of pipe-fault orientation on pipe
96
behavior under earthquake faulting. Also, the orientation of relative pipe-soil movement has been
97
studied by Daiyan et al. (2010). One of the main outcomes of their testing work is that it
98
demonstrated that coupled pipe-soil interaction due to oblique loading can considerably increase
99
the soil restraint on the pipeline. However, this interaction is minimal at very low angles.
100 4
101
Model and Full Scale Testing
102
One of the first small scale test programs conducted on steel pipes (25 mm, 60 mm, and
103
114 mm in diameter) was performed by Audibert and Nyman (1977) where the soil behavior was
104
observed during lateral displacement of pipes buried in sand with different densities under a wide
105
range of embedment ratios (1.5 to 24.5). Audibert and Nyman (1977) suggested a rectangular
106
hyperbolic function for the dimensionless force-displacement relation which was in agreement
107
with the one suggested by Das and Seeley (1975) who studied vertical anchor resistance against
108
horizontal movement.
109
Trautmann and O'Rourke (1983) conducted full-scale experiments on pipes with
110
diameters of 102 mm and 324 mm subjected to lateral movements, within sand of three different
111
densities (corresponding to friction angles of 31°, 36°, and 44°). Those tests featured burial
112
depths ranging from 1.5 to 22 diameters. Their work was used as the basis for the ASCE (1984)
113
guidelines for calculating lateral soil loading on pipes. Nyman (1984) also proposed design
114
procedures to develop bilinear load-displacement relationships for soil loads on pipelines
115
subjected to horizontal and vertical movements.
116
Hsu (1993) conducted an extensive full-scale testing program (120 tests) examining
117
several preliminary variables such as burial depth, pulling rate, soil density, and pipe diameter.
118
His results showed that soil resistance and corresponding displacements exhibit a power law
119
relation with the pipe velocity. Hsu et al. (2001) presented an experimental study of the soil
120
friction loads for oblique pipeline movements in loose sand, and another on dense sand (Hsu et
121
al., 2006). Both of studies presented and discussed analytical models to predict the longitudinal
122
and transverse soil loads.
123
Paulin et al. (1998) conducted 24 tests during their large scale experimental program that
124
included five types of pipe-soil loading test conditions including upward movement, lateral 5
125
movement, downward movement, and axial movement. Since the results of that testing program
126
are proprietary, only relative pipe loading and displacement were presented. A number of other
127
studies examining the flexural behavior of buried pipelines have also been conducted at the
128
Centre for Cold Ocean Resources Engineering (C-CORE) by Konuk et al. (1999). Details of the
129
physical tests can be found in Hurley et al. (1998a, and 1998b). Konuk et al. (1999) conducted
130
two large-scale tests to assess the bending behavior of buried pipes in dense sand under lateral
131
loading. During each test, measurements were made of the deflected pipeline profile, pipeline
132
ovalization, pipeline forces, and the associated soil deformations. Plastic hinges developed in the
133
pipe for both of the tests. Most of this work conducted at C-CORE is proprietary and only partial
134
results are available (Konuk et al., 1999).
135
One of the most recent testing programs is the work conducted at the University of
136
British Columbia by Karimian (2006). The objective of that research was to investigate the
137
lateral and axial pipe-soil interaction of relatively large diameter rigid steel pipelines (324 and
138
457 mm). Special focus was given to study the effect on the interaction behavior of trenching
139
and geotextile lining of the trench.
140 141 142 143 144 145 146
6
147
EXPERIMENTAL PROGRAM
148
Nine full scale tests were conducted on small-diameter tubing of 105 mm (4.1 inch) outside
149
diameter size. These tubes (referred to as pipes throughout the remainder of the article) are
150
specified in accordance with the American Society for Testing and Materials specifications A513
151
(ASTM, 2008) for electric-resistance-welded carbon and alloy steel mechanical tubing. Each of
152
the test pipes were buried in dense Olivine sand, as shown in Table 1. Burial depths with
153
embedment ratios, defined as the ratio of soil cover height (measured from spring line of the pipe
154
to soil surface) to nominal pipe diameter (H/D), ranging from 3 to7, were examined. This range
155
covers shallow to deep pipe burial situations as examined in anchor plates by Rowe and Davis
156
(1982). Boundary limits of pipe position within the test pit were examined. This included two
157
distances to the rear wall from the back of the pipe, of 3D and 9.5D. The internal system friction
158
resulting from the interaction between the pulling cables and soil was examined by comparing
159
tests with pulling cables enclosed in hollow PVC tubes to ones with cables exposed to the soil.
160
Also, reproducibility of test data was examined by repeating some tests. As the key section of a
161
buried pipeline subjected to lateral earth movements is between points of inflection (zero-
162
moment points), this can be modelled in tests by pulling that section of the pipe at its ends, while
163
allowing them to rotate freely (Fig. 1). This analogy would not be applicable in situations where
164
non-uniform deformations develop in the soil moving past the pipe. For example, a parabolic soil
165
deformation distribution past the pipe would increase the loads at the pipe mid-section rather
166
than the pipe ends. However, in many situations, the movement of the soil relative to the buried
167
pipe is uniform along the pipe (where, for example, an intact block of soil is moving past the
168
pipe). In such situations, the testing approach is appropriate between the points of inflection
169
along the pipe.
170 7
171
Testing Facility and Setup
172
The test pit used for the current study has plan dimensions of 4000 x 2000 mm (157.5x74.7
173
inch), and is 2000 mm (74.7 inch) deep, with concrete walls (Fig.2). The pit was divided into
174
three main sections using retaining walls. The main (middle) section (section II in Fig. 2(a)),
175
with plan measurements of 2000 x 3010 mm (78.7x118.5 inch), was dedicated to the physical
176
testing of the pipe. The west end of the pit accommodated the assembly used to pull the pipe
177
laterally. It consists of a hydraulic actuator, a spreader beam, and a pair of turn buckles. The
178
actuator was situated outside the test pit, with the actuator piston extending to the spreader beam
179
through an access opening (Fig. 2(a)). On the east end of the pit a narrow, 406 mm (16 inch)
180
wide space (section III in Fig. 2(a)) running across the width of the pit was created to
181
accommodate the five string potentiometers, referred to as “string pots”, used to monitor the pipe
182
deflections during testing. The wall between Sections II and III needed to withstand passive
183
lateral soil pressure and was built using 50 x 100 mm (2x4 inch) lumber ribs bolted to the
184
concrete walls of the test pit, and then covered by 9.5 mm (3/8 inch) thick plywood sheets. On
185
the west end of section II, the retaining wall was designed to withstand the full active soil
186
loading during testing. It was built using a stack of 140 mm x 140 mm (5x5 inch) pressure
187
treated timbers. The timbers were cut to the total width of the test pit (2000 mm). A total of 14
188
pieces were used to erect a 1960 mm (77.2 inch) high wall. To transfer the reactions of this wall
189
to the concrete wall at the west end of the pit, both ends of each timber piece was supported by a
190
short timber piece of the same size, bearing directly against the concrete end-wall (Fig. 2(a)).
191
Numerical calculations for the expected pulling force at the deepest burial depth indicated that
192
negligible (0.35 mm) deflections develop at the mid-span position of the 2000 mm long wooden
193
pieces, and hence the end-wall was assumed to be rigid.
194 8
195
Side Wall Friction Treatment: Tognon et al. (1999) evaluated the effect of sidewall friction for a
196
buried pipe testing facility and showed that friction treatment on the side walls of a testing
197
chamber can significantly reduce the effect of friction between the soil and the side boundaries.
198
In the literature, several configurations have been used. Trautmann and O’Rourke (1983)
199
implemented friction treatment for their testing program by using glass on one side and Formica
200
on the other side to reduce friction. Konuk et al. (1999) used plywood lining for the sidewalls of
201
a test box. Karimian (2006) used stainless steel sides and demonstrated analytically that by
202
replacing the plywood side wall lining by stainless steel, the shearing resistance of the sidewalls
203
would reduce to half of its value. In the current study, the side wall friction treatment of Tognon
204
et al. (1999) was employed. This consists of two layers of polyethylene sheets, each 0.1 mm
205
thick, with the first layer fixed to the wall at the top edge using lumber strips (Fig. 2(c)). A thin
206
film, about 1 mm thick, of silicone-based bearing grease was then spread with a paint roller to
207
act as the lubricant. The second layer of polyethylene sheet was then rolled over the grease film,
208
which had sufficient cohesion to hold that second layer in place without any other physical
209
attachment. This arrangement provides a low friction angle between the sand and the side walls
210
(Tognon et al., 1999, report that the angle of friction is as low as 5°). Both Rowe and Davis
211
(1982) and Trautmann and O’Rourke (1983) examined the effect of aspect ratio, which is in this
212
case the pipe length-to-diameter (L/D) ratio, on the measured lateral forces, accounting for the
213
effect of boundary conditions. Trautmann and O’Rourke (1983) concluded that “a length-to-
214
diameter ratio in the range of five to ten will decrease end effects to relatively small levels”. In
215
the current study, the pipes have even higher L/D ratio of 18.
216
Soil Placement and Removal: Sand was poured into the test pit in several lifts, each about 163
217
mm (6.4 inch) thick, by dumping from a height of 1.0 m using a hopper with side chute. Each
218
pour was distributed evenly using shovels and levels, and then compacted using a hand tamper 9
219
with plate dimension of 254x254 mm (10x10 inch), and weighing 7.9 kg, by dropping it from a
220
constant height of about 300 mm (11.8 inch) to ensure consistent packing effort. The soil level
221
was first brought up to a height of 500 mm (19.7 inch) before placing the pipe to minimize the
222
effect of the bottom concrete boundary. The pipe, 1830 mm (6 feet) long, was then centered
223
between the side walls of the pit leaving a side clearance of about 85 mm (about 0.85D) on each
224
side of the pipe. After each test, sand was removed by means of hand shovels and placed in sand
225
bags, which were craned out of the pit.
226
Loading Mechanism: Some of the previous full-scale testing programs (Paulin et al. 1998,
227
Konuk et al. 1999, and Karimian 2006) used two independent loading points to pull the pipe, one
228
at each end. Other investigators (Trautmann and O'Rourke 1983, and Hsu 1993) used a single
229
loading point. In the present study, a single pulling point was employed. The actuator was
230
situated outside the test pit, and positioned on the centreline of the pit. The piston was extended
231
into the testing pit and connected to the middle of a stiff steel ‘spreader beam’ fabricated from
232
HSS (hollow square section) with cross-section 102x102x6.3 mm (5x5x3/8 inch). The beam was
233
connected to 6.3 mm (1/4 inch) diameter steel cables with a rated axial load capacity of 28 kN,
234
one on each end, to pull the pipe. Each end of the spreader beam rested on a 140x200 mm
235
(5.5x7.9 inch) low-friction, lubricated Teflon strips to minimize friction as the pipe is being
236
pulled. The Teflon strips were attached to the short timber pieces used to transfer retaining wall
237
loads to the test pit wall (Fig. 2(d)). Similar Teflon strips were fixed to the side walls near the
238
ends of spreader beam to minimize friction in case the beam rotated and touched the sidewall
239
during pulling. Turnbuckles were used to connect the spreader beam to the cables through slots
240
made into the timber retaining wall (Fig. 2(e)) and sealed using rubber plugs to prevent soil
241
leakage. The turnbuckles provided the ability to adjust length and remove any cable slack. The
242
cables were attached to the pipe ends by looping each steel cable around the pipe at a distance 30 10
243
mm (1.2 inch) from the pipe end (Fig. 2a). The free end of the cable was then crimped to the
244
main cable using a number of cable clips (saddles) with the number of saddles ranged from 1 to 3
245
saddles per cable end throughout the testing program. The cables (and the accompanying
246
saddles) then extended through the soil mass, using one of two configurations (Fig. 2(a, and b)).
247
In the first configuration, the cables were left exposed to the surrounding soil. In the second
248
configuration, the cables were encased in 89 mm (3.5 inch) diameter PVC tubes fixed to the
249
retaining wall, to minimize the friction between the cables-saddles system and the soil. The PVC
250
tubes were centered around the cables at their connection points at the front wall, but were
251
gradually bent upward (using soil support to maximize clearance above the cable as the PVC
252
tubes approach the test pipe). The ends of the PVC tubes near the attachments to the steel pipe
253
were elevated upwards by almost half their diameter, to provide maximum clearance above the
254
cables close to the steel pipe, to accommodate the upward movements of the pipe during testing.
255
The effect of the friction between the pulling cables and surrounding soil has been
256
addressed and/or mitigated in previous work. Trautmann and O’Rourke (1983) enclosed threaded
257
rods, used to pull the pipe buried through the sand, in PVC tubes to eliminate soil friction.
258
Karimian (2006) employed approaches recommended by the ASCE (1984), and the PRCI (2004)
259
to calculate the axial load in the cables arising from friction and used those results to modify
260
measured pulling force. In the present study, a more quantitative assessment was made, and an
261
approach to account for system friction is presented and discussed in a subsequent subsection.
262
To achieve this goal, auxiliary tests were conducted. A pulling test was carried out on just the
263
cables (i.e. without the test pipe) buried under the soil at a depth (cover) of 304 mm (12 inch),
264
representing the case of H/D = 3. Another test without the pipe was performed after enclosing
265
the cables in the 89 mm (3.5 inch) PVC tubes.
266 11
267
Test Materials
268
Backfill: Synthetic Olivine sand was used as the backfill material for all the experiments. This
269
sand can be described as medium-to-fine sand with generally angular soil particles. Synthetic
270
Olivine sand has the ability to reach a dense state with little compaction. Grain size analysis
271
indicates an average particle size (D50) of 0.67 mm (.026 inch), minimum particle size of 0.075
272
(2.9x10-3 inch) mm, and a coefficient of uniformity (Cu) of 1.84. Triaxial tests were performed to
273
assess the soil parameters, namely friction angle (), dilation angle (), and modulus of elasticity
274
(Es). The tests were performed at confining pressures of 18.3, 33, and 44 kPa. While the sand
275
was subjected to much lower initial confining stresses during pipe testing (0.8 to 3 kPa), it is
276
difficult to conduct triaxial tests at such low confining stresses, and friction angle was
277
extrapolated from the tests conducted at these higher pressures. Dry density and moisture content
278
of the sand after each lift during backfilling were checked using a nuclear densometer. Soil
279
parameters are summarized in Table 2.
280
To obtain a more accurate representation for the soil modulus of elasticity at low confining
281
stresses, the Janbu model (Janbu 1963) was used to represent the soil modulus distribution. The
282
model describes elastic soil modulus (Es) as a function of the confining stresses using the
283
following relation:
284
𝐸𝑠 𝑃𝑜
𝜎3 𝑛
=K( ) 𝑃𝑜
[1]
285
where Po is the atmospheric pressure at sea level (= 101.3 kPa); the Janbu model parameters n
286
and K are constants specific to each material; and 3 is the minimum effective principal stress.
287
Parameters n and K represent the power law parameters in the expression. They have been
288
determined for Olivine sand by plotting test data as (3/Po) versus (Es/Po) for the three triaxial
289
tests, which were then fitted with a linear relation on a log-log scale (Fig. 4). In this way, values
290
for K and n were found to be 326 and 0.86, respectively. 12
291
Steel Pipe: The steel pipe used was manufactured according to the American Society for Testing
292
and Materials Standard A513 (ASTM, 2008). Pipe dimensions and strength parameters are
293
provided in Table 3.
13
294
Instrumentation
295
Electric resistance strain gages and string pots were attached to the pipe to monitor bending
296
strains and deflections, respectively. String pot boxes were secured to the partitioning wall
297
separating sections II and III of the pit (Fig. 2). The strings were then passed through small holes
298
in the wall, into the soil, and enclosed in 50 mm (2 inch) diameter PVC tubes fixed to the rear
299
wall to avoid contact with the soil. Thin steel wire extensions were looped around and secured to
300
the test pipe and were connected to the hooks of the string pots. Displacements were measured
301
at five equidistant locations along the pipe (Fig. 3).
302
For the strain gages, an array of uniaxial strain gages was fixed to the front (passive), and
303
rear (active) spring lines of the pipe (Fig. 3). Results at the mid-span position are reported in a
304
subsequent section. A triaxial strain gage was also attached to the top (crown) of the pipe at one
305
diameter distance from the south end of the pipe. Extra slack was left in the strain gages wires by
306
embedding them with waves within the soil to allow for pipe movement.
307
To measure the pipe pulling load, a load cell (MTS System Corporation, Model 661.20A-
308
03) was used featuring an accuracy of ± 1 N. The load cell was connected between the spreader
309
beam and the actuator piston (Fig. 2). During testing, a data acquisition system (Vishay
310
Measurement Group-System 5000) was employed to collect force, strain and displacement data.
311
To check for pipe elevation and soil heave before and after testing, a laser level was used. Sand
312
density and water content were checked before each test using a nuclear densometer (by CPN
313
international Inc., model CPN MC-1DR-P) at the center of each of four quadrants of the soil for
314
each lift, as shown in the plan view, Fig 2(a).
315 316 317 14
TEST RESULTS AND DISCUSSION
318 319
Effect of Pulling System Friction
320
The pulling force on the cables buried under a soil cover of 304 mm, in the auxiliary test without
321
the pipe, was found to be 2.1 kN. This is of significant size given that it represents about 20% of
322
the total pulling force when it is attached to the pipe. In the other auxiliary test, when the cables
323
were enclosed in the 89 mm PVC tubes, a significant reduction of pulling force to 0.3 kN was
324
measured. These values are useful indicators of the expected change of force associated with
325
friction, but as will be seen in the following section, this force change varied from case to case
326
and the next section describes a unified adjustment technique developed for use with each
327
individual test.
328
Fig. 5 shows the actuator load-stroke curves for pipe specimen S3 (Table 1), which was
329
conducted with exposed cables, and the same pipe specimen for Test S4a, but with the cables
330
enclosed in PVC tubes. It can be seen that the difference in peak loads is consistent with the
331
difference between loads measured in the two auxiliary tests, 1.8 kN. It can also be seen that the
332
measured loads at any given stroke are different from the onset of loading, with the difference,
333
which represents friction between cable and soil, building up gradually to the 1.8 kN difference
334
at peak loads. In Test S3, the load measured at the instant of first sensing of load by the pipe,
335
indicated by first excitation of strain gage SG1 (Fig. 3) located at the pipe end, was about 1.145
336
kN, at an actuator displacement of 1.4 mm. On the other hand, the equivalent values in test S4a
337
were 0.23 kN and 0.2 mm, respectively. Table 4 summarizes the actuator load values recorded at
338
initial excitation of the outermost and middle strain gages SG1, SG4 and SG5 (Fig. 3). Based on
339
SG1, the point of initial load transfer to the pipe, the loads in all cases with enclosed cables
340
varied from 0.05 kN to 1.12 kN. Review of the actuator data retrieved proved that actuator stroke
341
is not relevant as it is sensitive to initial slack within the pulling cables before the start of the test. 15
342
These results have not therefore been included in this table. Data provided in Table 4 suggests
343
that variable, and unpredictable, amounts of load are needed to overcome friction. As such, an
344
approach for adjusting the actuator load-stroke curve was adopted for each test. First, the point
345
on the actuator load-stroke curve at which load was first transferred to pipe is established based
346
on first excitation of strain gage SG1. The actuator pulling force reading and stroke at this point
347
is then set as the new origin for the “Pipe Pulling Load” and “Pipe End Displacement”,
348
respectively. It is assumed that beyond this point, any further pulling would be transferred to the
349
pipe directly, except for the load used to overcome soil interaction with the short exposed portion
350
of cable. Furthermore, most of the slack in the cables would have been removed by then, except
351
for the small elongation due to the tensile loads applied during the remainder of the test. All
352
pulling loads and pipe displacements reported hereafter represent this approach of adjustment.
353 354
Effect of Distance to Rear Test Cell Boundary
355
Tests S1 and S2, both using exposed pulling cables, were compared to examine the effect of
356
front and rear boundary limits on the load-pipe end displacement curve (Fig. 6). Configurations
357
for both tests were the same, except for the position of the pipe relative to the rear and front end
358
walls. For test S1, the length of soil behind the pipe, measured from the centerline of the pipe to
359
the rear wall, was 305 mm (3D), and the length of soil in front of the pipe 2705 mm (26.6D). In
360
test S2, there was 965 mm (9.5D) of soil behind the pipe, and 2045 mm (20.1D) in front. Fig. 6
361
shows a difference of just 4% in the peak load. Also, the extent of soil surface deformation seen
362
behind the pipe after the test was concluded did not exceed distances of 105 mm and 94 mm
363
(equivalent to 0.92D and 1.02D) for tests S1 and S2, respectively. This indicates that in
364
situations where tight space is of a concern, a three pipe diameters to the rear boundary limit
365
should be sufficient. For the side boundary limits, a distance of 85 mm (0.85D) was left between 16
366
each end of the pipe and the side wall. While no physical testing was performed to assess the
367
effect of those side boundary limits, numerical calculations not included here show that side
368
clearance as low as 0.35D has a minimal effect on the load-deflection curvature of the pipe.
369 370
Test Repeatability
371
To assess the reproducibility of test results under the same burial conditions, tests S4a, S4b, S4c,
372
and S4d with embedment ratio of H/D = 3 were replicates. The only minor difference between
373
the tests is the number of clips used to grip the cable when looped around the pipe end, or when
374
threaded through the turnbuckle at the other end. Tests S4a and S4b used one cable clip for each
375
cable end, while tests S4c, and S4d used two clips instead after minor slippage (2 mm) of the
376
cable was noticed after test S4b. Fig. 7(a) shows the load-displacement curves, which depict
377
almost identical behavior in the elastic range, but with some variability in the peak load, which
378
ranged from 9.3 kN for test S4c to 10.4 kN for test S4a (mean of 9.92 kN and standard deviation
379
of 0.11 kN). Fig. 7(b) shows the pipe bending responses (i.e. mid-span deflection of the pipe
380
relative to a reference line connecting both ends). Some variability can be seen at the initial parts
381
of the curves. The maximum deflection varied from 2.5 mm in test S4b to 3.5 mm in test S4a.
382 383
Effect of Burial Depth
384
In the following section, the effect of embedment ratios of H/D = 3, 5 and 7 on several aspects of
385
the pipe-soil interaction are examined. Test S4a (Table 1) is selected to represent the burial depth
386
for H/D = 3.
17
387
Load-Displacement Responses: Fig. 8(a) and (b) show the load-displacement curves for the
388
three burial depths, for both pipe end, and pipe mid-span displacements, respectively. Clearly,
389
the peak pulling force increases with embedment ratio (10.4, 18.5, and 30.4 kN, for embedment
390
ratios 3, 5, and 7, respectively). It is important to note that failure in all cases was governed by
391
the soil and not the pipe as will be explained from the load-strain responses of the pipes at mid-
392
span. Fig. 8(c) shows the load-pipe bending deflection responses at mid-span, . The responses
393
are generally linear elastic and having similar slopes. Only the peak load and deflection increase
394
with increasing burial depth. Also, the three curves show some nonlinearity (concave down), i.e.,
395
there is less incremental load needed for the same incremental deflection of the pipe, even
396
though the pipe material is believed to still be behaving elastically (based on the measured
397
strains). This nonlinearity is not usually observed when conventional ‘in-air’ beam bending tests
398
are used (beam-bending experiments on the pipe featuring simple supports and traditional
399
characterization methods such as 3-point and 4-point bending, or uniformly distributed loads).
400
This difference between buried pipe response and ‘in-air’ response is consistent with earlier
401
studies that found that pipe bending response when buried in soil is different to simple beam-
402
bending behavior (e.g. Konuk et al., 1999; and Mahdavi, 2008).
403
Load-Strain Responses: Fig. 8(d) compares the load-bending strain responses at mid-span of the
404
pipes tested for embedment ratios 3 and 5 (difficulties were experienced with the strain gages for
405
the test at embedment ratio of 7, so the comparison is not available for that case). Strain
406
development with pipe loading for both tests are somewhat similar, except for the small shift in
407
initial load for the test with H/D = 5. It can be seen that beyond the peak load as the load drops,
408
the strain reduces suggesting that the curvature of the tube and hence its net deflection reduces.
409
The maximum axial strain reached at the mid-span of the pipe was for Test S5 to be 1100 micro
410
strain, which is far less than the yield strain of the pipe. It is important to note that these pipes 18
411
were not tested with internal fluid pressure, which would induce circumferential tensile strains
412
placing the pipe in a state of bi-axial stresses. Also, a number of geometrical factors are expected
413
to influence the longitudinal bending behavior of the pipe, such as its length to diameter ratio,
414
and its diameter to thickness ratio (the experimental results presented here are for one particular
415
choice of pipe diameter, length, and wall thickness). Numerical calculations can be employed to
416
investigate the effect of these parameters once the test data are used to establish the effectiveness
417
of the computer modeling (Fatemi et al., 2008, provide an example of this use of computer
418
analysis to establish the influence of pipe parameters).
419 420
Fig. 9 shows the development of pipe displacement (by measuring string pot extensions) at
421
different sections along the pipe versus pipe loading for embedment ratio H/D = 3 (Test S4a).
422
This specific test illustrates a situation where the pipe experiences slight rotation during pull. In
423
the ideal loading case, the extension of string pot pairs at pipe ends (curves for 1, and 5) should
424
be identical, as should displacements at the quarter positions (curves for 2, and 4). In this test,
425
the North end of the pipe advanced ahead of the South end (ahead by 7.4 mm at peak pipe
426
loading). Considering this pipe tilting or pipe rotation together with the bending deflection of the
427
pipe itself of about 3.5 mm, yields a total oblique orientation angle between the pipe span
428
relative to the pipe pulling direction of about 0.3 (from 90o). At post-peak loading, this value
429
increased to a maximum of 1.2. Phillips et al. (2004), and Daiyan et al. (2010) examined the
430
effect of combined axial and lateral pipe-soil interaction and developed interaction graphs for
431
combined loading with the calculated and measured horizontal and axial bearing capacity factors
432
(Nqh, and Nt, respectively) for the soil. One of the main outcomes from both studies is that
433
coupled pipe-soil interaction due to oblique loading is of minimal effect on the lateral soil
434
restraint on the pipeline at such low rotation angles. 19
435
Soil Surface Deformation: profile limits of the soil deformed surface were recorded after each
436
test. Typical features of deformed soil surface after lateral pipe pull reported in previous
437
experimental work (e.g. Rowe and Davis 1982, Trautmann and O’Rourke 1983, and Karimian
438
2006) were observed in the current study. The main characteristics are formation of a passive,
439
heaving, wedge in front of the pipe undergoing general shear failure, and an active, sinking,
440
wedge behind the pipe. Degree of surface deformation varied as pipe burial depth changed. Fig.
441
10 illustrates the soil surface deformation for the different burial depths tested. It can be seen that
442
for embedment ratio of H/D=3, soil dip behind the soil and the soil heave in front of the pipe is
443
very distinctive and of order of magnitude comparable to pipe diameter size, D, (1.1D below,
444
and 0.4D above initial soil surface, respectively). For embedment ratio of H/D=5, the soil
445
dipping behind the pipe is 0.9D, whereas soil heave in front of the pipe reduces considerably to
446
just 0.1D. For the deeper burial depth of H/D=7, soil surface movement was barely detectable
447
(0.1D, and 0.1D for soil dip and soil heaving, respectively). Such minimal soil surface
448
disturbance for embedment ratio of H/D=7 suggests that the failure mechanism here is close to
449
the deep, punching, failure mechanism observed by Audibert and Nyman (1977) with H/D
450
values ranging from 12-24, Rowe and Davis (1982) for deeply buried laterally loaded anchor
451
plates (with H/D values >3), and Trautmann and O’Rourke (1983) for deeply buried pipes with
452
H/D values ranging from 8-11.5. Ideally in this type of failure, only local failure of soil around
453
pipe takes place with no change occurring at the soil surface.
454
As for the pipe translation within the soil mass, Fig. 11 shows pipe initial and final positions at
455
pipe mid-span for the same tests under examination, which depicts a similar trend of change as
456
the one observed with soil surface heave in Fig. 10. At burial depth ratio of H/D=3, the pipe
457
elevated by 37 mm after a horizontal displacement of 102 mm, while for a similar pipe
458
horizontal displacement of 98 mm at burial depth ratio of H/D=5, the pipe elevated 17 mm only. 20
459
When the pipe was pulled for burial depth ratio of H/D=7 (which required longer horizontal
460
travel of 154 mm to reach peak loading), pipe also moved up by 17mm. Such upward movement
461
is expected since the pipe travels in the direction of least resistance. Since soil strength and
462
stiffness increases with depth (it is a function of the earth pressures), as the pipe moves forward,
463
it is easier to form a path which has an upward component, taking it into the weaker, less stiff
464
soil.
465
Other Pipe results: progress of the deflected shape of the pipe for all the tests was monitored and
466
recorded by plotting the instantaneous string pot readings along the pipe. Fig. 12 shows deflected
467
shape of the pipe at 50%, 75%, and 100% of peak pulling load, Pu, for pipe burial depth H/D=3
468
(Test S4a). The figure also illustrates the amount of tilt (exaggerated due to scale effect)
469
experienced by the pipe as discussed earlier.
470 471
Conclusion
472
A series of full-scale tests on high strength, steel pipes buried in dense sand under three burial
473
depth-to-diameter (H/D) ratios of 3, 5 and 7 has been conducted to study the pipe-soil
474
interaction, with focus on the flexural behaviour of the pipe under lateral earth loading with
475
idealizing earth lateral loading conditions on the pipe by pulling the pipe at its end instead. The
476
testing facility and experimental setup were discussed and demonstrated to meet the intended
477
purposes of the current study program with adequate accuracy. Other conclusions can also be
478
drawn from analyzing the test results:
479
a) Consistency of test setup configurations (geometrical setup, and material preparation)
480
was assessed with replicate tests, and four tests produced maximum pulling forces that
481
varied by less than 6%.
482
b) Friction within the pipe pulling system could lead to unpredictable, and potentially 21
483
significant (up to 20% of total pulling force), amount of load consumed in overcoming
484
cable-soil interaction. Mitigation of such factor is presented and discussed.
485
c) In similar testing configurations where tight space is a concern, a distance of three pipe
486
diameters to the rear boundary of the test box should be sufficient. Reduction to just three
487
diameters changed the measurement of pulling force by about 4%, which is less than the
488
amount seen in the replicate tests.
489
d) Bending deflection of the pipe is not linear with pulling force, even within the elastic
490
range of the steel material tested, which gives more significance to the study of bending
491
behavior of buried pipes in situ conditions than studying it at “in-air” loading
492
configurations.
493
e) By characterizing the load-bending deflection of flexible buried pipelines and
494
establishing bending tolerance estimations, better decisions about excavating,
495
straightening, and reburying pipes could be then conducted. The current two-dimensional
496
approach does not provide direct evidence of pipeline tolerance to these deflections).
497 498
Notation
499
The following symbols and abbreviations are used in this paper:
500
Cu
= coefficient of conformity of soil
501
D
= nominal pipe diameter
502
Do
= outside pipe diameter
503
D50
= average particle size of soil grains
504
Es
= soil modulus of elasticity
505
H
= burial depth to pipe spring line
506
K
= Janbu modulus multiplier at stress of 1 atmosphere (dimensionless)
507
L
= pipe length
508
n
= Janbu model index controlling level of stress dependence (dimensionless) 22
509
Nqh
= horizontal bearing capacity factor for sand
510
Pu
= peak soil resistance load to lateral pipe movement
511
Po
=atmospheric pressure at sea level
512
Pa
=actuator load
513
t
= pipe wall thickness
514
= bending deflection at mid span of the pipe
515
= sectional displacement of pipe
516
a
= actuator stroke
517
= friction angle of sand
518
= bulk unit weight of soil
519
= Poisson ratio
520
= minimum effective principal stress
521
= dilation angle of sand
522 523
References
524 525
ASTM A513. (2008).” Standard Specification for Electric-Resistance-Welded Carbon and Alloy Steel Mechanical Tubing.” American Society for Testing and Materials specification.
526 527 528 529
American Society of Civil Engineers. (1984). “Guidelines for the seismic design of oil and gas pipeline systems.” Committee on Gas and Liquid Fuel Lifelines, Technical Council on Lifeline Earthquake Engineering, ASCE, New York.
530 531 532 533
Audibert, J.M.E. and Nyman, K.J. (1977). “Soil Restraint against Horizontal Motion of Pipes”, Journal of the Geotechnical Division, Proceedings of the American Society of Civil Engineers, Vol. 103, No. GT10, pp. 1119-1142
534 535 536 537
Da, H., Abdoun, T. H., O'Rourke, M. J., Symans, M. D., O'Rourke, T. D., Palmer, M. C., & Stewart, H. E. (2008). “Centrifuge Modeling of Earthquake Effects on Buried High-Density Polyethylene (HDPE) Pipelines Crossing Fault Zones.” Journal of Geotechnical and Geoenvironmental Engineering, 134(10), 1501-1515.
538 539
Daiyan, N., Kenny, S., Phillips, R., and Popescu, R. (2010). “Numerical Investigation of Oblique Pipeline/soil Interaction in Sand.” Proceedings, International Pipeline Conference, Calgary.
540 541 542
Das, B. M., Seeley, G. R. (1975). “Load displacement relationship for vertical anchor plates.” Journal of Geotechnical Engineering Division, ASCE, 101(GT7), 711-715. 23
543 544
EGIG. (2008). “7th Report of the European Gas Pipeline Incident Data Group.” (http://www.egig.nl/downloads/7th_report_EGIG.pdf).
545 546 547 548
Fatemi, A., Kenny, S., Taheri, F., Duan, D., and Zhou, J. (2010). “End Boundary Effects on Local Buckling Response of High Strength Linepipe.” Proceedings, International Pipeline Conference, Calgary.
549 550 551
Hsu, T. W. (1993). “Rate effect on lateral soil restraint of pipeline.” Soils and Foundations, 33(4), 159-169.
552 553 554 555
Hsu, T.W., Chen, Y.J., and Hung, W.C. (2006). “Soil Friction Restraint of Oblique Pipelines in Dense Sand”, J. Transp. Eng. 132(2), pp. 175-181.
556 557
Hsu, T.W., Chen, Y.J., and Wu, C.Y. (2001). “Soil Friction Restraint of Oblique Pipelines in Loose Sand.”, J. Transp. Eng. 127(1), pp. 82-87.
558 559 560 561
Hurley, S., Zhu, F., Phillips, R. and Paulin, M. J. (1998a). “Large scale modeling of soil/pipe interaction under moment loading,” Test GSC 01 data report. Contract Report for Terrain Sciences Division. Geological Survey of Canada, C-CORE Publication 98 C-20.
562 563 564 565
Hurley, S., Zhu, F., Phillips, R. and Paulin, M. J. (1998b). “Large scale modeling of soil/pipe interaction under moment loading,” Test GSC 02 data report. Contract Report for Terrain Sciences Division. Geological Survey of Canada, C-CORE Publication 98 C-21.
566 567 568 569
Janbu, N. (1963). “Soil Compressibility as Determined by Odometer and Triaxial Test.” Proceedings of the European Conference on Soil Mechanics and Foundation Engineering, Vol. 1, pp 19-25, Weisbaden.
570 571 572 573
Karimian, S.A. (2006). “Response of Buried Steel Pipelines Subjected to Longitudinal and Transverse Ground Movement.”, PhD Thesis, University of British Columbia, Vancouver, Canada.
574 575 576
Konuk, I., Phillips, R., Hurley, S., and Paulin, M. J. (1999). “Preliminary Ovalisation of Buried Pipeline Subjected to Lateral Loading.” OMAE99/PIPE,-5023.
577 578 579 580
Mahdavi, H, Kenny, S., Phillips, R., and Popescu, R. (2008). “Influence of Geotechnical Loads on Local Buckling Behavior of Buried Pipelines.” Proceedings, International Pipeline Conference, Calgary. 24
581 582
National Energy Board. (2008, July). Focus on safety: A comparative analysis of pipeline safety performance 2000-2006.
583 584 585
Nyman, K.J. (1984), “Soil response against oblique motion of pipes”, Journal of Transportation Engineering, ASCE, Vol. 110, No. 2, pp190-202.
586 587 588 589
O’Rourke, M., Gadicherla, V., and Abdoun, T., (2005). “Centrifuge Modeling of PGD Response of Buried Pipe.” Earthquake Engineering and Engineering Vibration, Vol. 4, No. 1, pp. 6973
590 591 592 593
Paulin, M.J., Phillips, R., and Boivin, R. (1995). “Centrifuge modeling of lateral pipeline/soil interaction-Phase II.” OMAE 95, Proceedings of the 14th International Conference on Offshore Mechanics and Arctic Engineering, Copenhagen, Denmark, Volume 5, pp.107-123
594 595 596 597
Paulin, M.J., Phillips, R., Clark, J.I., Trigg, A., and Konuk, I. (1998). “A Full-scale Investigation into Pipeline/Soil Interaction.” Proceedings, International Pipeline Conference, Calgary, AB, Canada
598 599 600
Phillips, R., Nobahar, A., and Zhou, 1. (2004) "Combined Axial and Lateral Pipe-Soil Interaction Relationships." Proceedings, International Pipeline Conference, Calgary, AB, Canada.
601 602 603 604
PHMSA (2010). “Significant Pipeline Incidents by Cause.” (http://primis.phmsa.dot.gov/comm/reports/safety/SigPSIDet_1991_2010_US.html?nocache =1111).
605 606 607 608
Pipeline Research Council International, PRCI (2004) "Guidelines for the Seismic Design and Assessment of Natural Gas and Liquid Hydrocarbon Pipelines." Pipeline Research Council International, Inc. (PRCI), Catalogue No. L51927.
609 610 611
Rowe, R. K., and Davis, E.H. (1982). “The behaviour of anchor plates in sand.” Geotechnique, 32(1), 25-41.
612 613 614
Tognon, A.R., Rowe, R.K., and Brachman, R.W.I. (1999). “Evaluation of Sidewall Friction for a Buried Pipe Testing Facility.” Geotextiles and Geomembranes, 17: 193–212.
615 616
Trautmann, C.H., and O’Rourke, T.D. (1983). “Behaviour of pipe in dry sand under lateral and uplift loading”. Geotechnical Engineering Report, Cornell University, Ithaca, N.Y., 83–7.
617
25
618
619 620 621 622 623 624 625 626
627 628 629
631 632
Table 1. Test Matrix Distance to rear Test I.D. H/D wall Boundary** S1 3.0 D 3 S2 9.5 D 3 S3 9.5 D 3 S4a 9.5 D 3 S4b 9.5 D 3 S4c 9.5 D 3 S4d 9.5 D 3 S5 9.5 D 5 S6 9.5 D 7 * BL SF Rep BD
Pulling Cables Confinement Exposed* Exposed* Exposed Enclosed Enclosed Enclosed Enclosed. Enclosed Enclosed
Parameters Examined BL BL BD
SF SF
Rep Rep Rep Rep
BD BD
Uninstrumented Pipe ** Distance is in multiples of pipe diameter (D) Boundary Limit “System Friction” due to pulling cables confinement arrangement, by being either Exp. or Enc. Repeatability check test Burial Depth
Table 2. Soil Parameters Parameter Modulus of Elasticity (E), MPa Poisson Ratio () Bulk unit Weight (), kN/m3 Friction Angle () (Peak/Residual) Angle of Dilation ()
Value 16 0.3 15.1 53°/45° 16°
Table 3. Steel Pipe Parameters
Parameter Nominal Diameter (D), mm Outside Diameter (Do), mm Length (L), mm D/t Ratio L/D Ratio Wall Thickness (t), mm
630 Value 102 (4 in.) 105 (4.1 in.) 1829 (6 ft.) 37 18 2.7 (0.106 in)*
Modulus of Elasticity (GPa)
200
Yield Strength (MPa)
379
Ultimate Strength (MPa)
427
0.28 Poisson Ratio () Material Specifications ASTM A513 * measured (different from the 2.1mm (0.083 in.) value from ASTM A513 and the supplier) 26
633 634
635 636 637 638 639 640 641 642 643 644 645 646 647 648 649 650 651 652 653 654 655 656 657 658 659 660 661 662 663 664 665 666 667 668 669 670 671 672
Table 4. Load Cell Reading (kN) at 1st Excitation of Strain Gages 3 H/D Test I.D # S3 S4a S4b S4c S4d
5 S5
7 S6
SG1
1.145
0.232
1.074
0.345
0.052
0.446
1.117
SG5
1.471
0.232
1.373
0.534
0.052
0.476
1.520
SG4 (Mid-span) 1.264 0.485 2.191 0.736 0.723 Note: Shaded areas identify tests with enclosed pulling cables
0.446
2.130
27
673
Soil Loading
674
Inflection Points (Zero Moment)
Tested Portion 675 676 677 678 679 680 681 682 683 684 685
Fig. 1. Plan View of Lateral Earth Movement past a Buried Pipeline and the Experimental Test Section between Points of Inflection Being Modeled in this Study
686
Quadrant Quadrant Quadrant Quadrant
N
3 2
6
7
A
A
4
(b)
A
ua dra nt
(e)
4000 mm
3
H/D=7
4
694
696 697 698 699
5
584 mm
3010 mm
4
500 mm
Soil level @ H/D = 3
H
H/D=5
695
(d)
5
1
688 689 690 691 692 693
(c) 2000 mm
( a A)
Section III
Section II
(a) Section I
2000 mm
687
1: Actuator piston 2: Spreader beam 3: Retaining wall (front end wall) 4: Pulling cable (in exposed configuration) 5: Pulling cable with encasing PVC tube (in enclosed configuration) 6: Rear end wall 7: String potentiometer
406 mm
Fig. 2. Experimental Test Setup: a) Plan View, b) Elevation view (section A-A), c) View of Test Pit, d) Friction Treatment for Spreader Beam, and e) Spreader Beam-Turnbuckle Connection
700 701 28
702 703 704
SP5
SP3
SP4
String Potentiometer (SP1)
SP2
705
455 mm
706 707 708
SP Wires
Protecting Tubes
709 710 711 712
SG4R
SG5R
SG-Triaxial SG4F
SG5F
713
456 mm
714 715 716
SG2R Strain Gauge 1-Rear
SG3R
SG2F
SG3F
C.L .
Strain Gauge 1-Front (SG1F) 270 mm 102 mm
Fig. 3. Pipe Instrumentation: Top View Schematic Showing Strain Gauges and Stringpot Locations
717 718 719 720 721 722
Soil Modulus (Es)/ Pa
1000
y = 326.42x0.8603
100
10
Experimental 1
723 724 725
0.1
Confining Stress/ Pa
1
Fig. 4. Determination of Janbu Model Parameters 29
726
Actuator Pulling Load (kN)
14 Exposed Cables Setup (Test S3)
12 10 8
Enclosed Cables Setup (Test S4a)
6 4 Indicates point of 1st sensing of load
2 0 0
727
20 40 60 80 Actuator Stroke, a (mm)
728 729 730
Fig. 5. Effect of System Friction on Load-Displacement Curve Based on Actuator Stroke
731 732 733 734 10 9 8 7 6 5 4 3 2 1 0
Pipe Pulling Load (kN)
H/D =3 d= 9.5D (Test S1) d= 3D (Test S1)
d d
29.5 D 0
735 736 737
20
40
60
80
Pipe End Displacement (mm)
Fig. 6. Effect of Rear Soil Boundary on Load-Displacement Curve
738 739 740 30
741 12
H/D=3
(a)
Pipe Pulling Load (kN)
10 8 6 Test S4a Test S4d Test S4b Test S4c
4 2 0 0
742
12
20 40 60 Pipe End Displacement (mm)
H/D=3
80
(b)
Pipe Pulling Load (kN)
10 8 6 Test Test8S4c Test Test7S4b Test Test9S4d Test6S4a Test
4 2 0 0
743
1 2 3 Pipe Mid-span Deflection, (mm)
4
744 745 746 747
Fig. 7. Effect of Test Repeatability on (a) Load-Displacement Curves, and (b) Pipe Bending Deflection Curves at Mid-Span
748 749 750 751 752 753 754 755 31
756 757 758 40
30
missing data
H/D =7
25 20
H/D =5
15 10
H/D =3
5 0 0
759
20 40 60 80 100 Pipe End Displacement (mm)
120
35
Pipe Pulling Load (kN)
30
(b) missing data H/D = 7
20
10
H/D = 5 H/D = 3
5 0
760
(c)
30 25 missing data
20 15
H/D=7 (Test S6) Load
10
H/D=5 Load (Test S5)
(kN) (kN)
H/D=3 Load (Test S4a)
5
(kN)
0 -1
1
0 20 40 60 80 Mid-span Stringpot Extension () (mm)
100
3 5 7 9 Deflection, (mm)
11
20 Tension Compression 18 16 14 H/D = 5 12 10 8 6 H/D = 3 4 2 0 -1200 -800 -400 0 400 800 1200 Micro Strain (at Pipe Mid-span)
(d)
25
15
35
Pipe Pulling Load (kN)
(a)
Pipe Pulling Load (kN)
Pipe Pulling Load (kN)
35
761 762 763 764 765
Fig. 8. Effect of Burial Depth on (a) Load-Displacement Curves at Pipe Ends, (b) LoadDisplacement Curves at Mid-Span String Pot, (c) Load-Bending Deflection Curves at MidSpan, and (d) Load-Pipe Strain at Mid-Span Curves
766 767 768 769 770 771 32
772 773 774 12
Pipe Loading (kN)
10
(South End)
8 6 4 2 0 0
775 776 777
20 40 Stringpot Extension, , (mm)
60
Fig. 9. Sectional Displacements along the Pipe for H/D = 3
778 779 780
Soil Heave (mm)
60 40 20 0 -200 -20 0 -40 -60 -80 -100 -120 -140
781 782 783 784
H/D=3 (Test S4a) H/D=7 200
400
600
800
1000
H/D=5
Horizontal Distance from Initial Pipe Position (mm)
Fig. 10. Soil Surface Heave at Different Burial Depths (at Pipe Centerline) (Line markers indicate measured locations)
785 786 787 788 33
789 790 791
Burial Depth (mm)
-50
100 0 -100 0 -200 -300 -400 -500 -600 -700 -800
50
100
150
200
H/D=3 H/D=5 H/D=7
Horizontal Distance from Initial Pipe Position (mm)
Fig. 11. Initial and Final Pipe Position of Pipe at Different Burial Depths
792 793 794 795 796 797 798 799 800 801 802 803 804 805 806 807 808 809 810 811 812 34
813 814 815 816 817 818 819 820 821 822 823 824 825 826 827 828 829 830 831 832 833 834 835 836
35
