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Full-Scale Seismic Testing of Piles in Improved and Unimproved Soft
ClayFebruary, 2016
Full-Scale Seismic Testing of Piles in Improved and Unimproved Soft
Clay Bradley Fleming, Iowa State University Sri Sritharan, Iowa
State University Gerald A. Miller, University of Oklahoma
Kanthasamy K. Muraleetharan, University of Oklahoma
Available at: https://works.bepress.com/sri_sritharan/34/
Bradley J. Fleming,a) M.EERI, Sri Sritharan,b) M.EERI, Gerald A.
Miller,c)
and Kanthasamy K. Muraleetharanc)
A full-scale field investigation was performed to determine the
effects of soil improvement on the seismic resistance of piles in
soft clay. A soil improvement method, called cement deep soil
mixing (CDSM), was used to improve soil supporting a standard 324
mm diameter steel pipe pile subjected to simulated earthquake
lateral loads. An identical pile in unimproved clay was also tested
to determine the effects of the soil improvement. Compared to the
unimproved pile, the CDSM technique showed a 42% increase in pile
lateral strength, a 600% increase in effective elastic stiffness,
and a 650% increase in average equivalent damping ratio. The pile
in improved soil reached its lateral capacity at a head
displacement of 0.1 m, at which point the critical region at the
base of the pile above the improved ground experienced buckling and
subsequent fracture due to low cycle fatigue. [DOI:
10.1193/012714EQS018M]
INTRODUCTION
Structures and their foundations are subjected to forces created by
earthquakes, wind, waves, water current, vessel impact, ice, and
gravity. All loads applied to the superstructure must be
transmitted to the foundation, which are then transferred to the
surrounding soil. With significant lateral loads, such as those
created by earthquakes, use of pile foundations is one of a few
options to transmit large structural loads to competent soils.
Piles supported by competent soils are relatively easy to design
cost-effectively. However, thick layers of weak soils such as soft
clays are widespread in high seismic areas (e.g., San Francisco,
southern Nevada, Washington, Eastern Missouri, and Arkansas),
exacerbating design challenges. Driven piles are the preferred
supports for many structures found on saturated soft soils due to
ease of installation and their effectiveness at penetrating deep
layers of competent soils as needed. However, soft clays reduce the
lateral resistance of the pile-soil system, making the pile
foundation less cost-effective. In this case, the current design
practice is to use an increased number of larger diameter piles
(see SDC, Caltrans 2010), which may not be feasible in deep
deposits of soft clay. An innovative and more cost-efficient
solution to this problem is to improve the soil surrounding the
pile in the immediate vicinity
a) Department of Civil, Construction, & Environmental
Engineering, Iowa State University, 394 Town Engineering, Ames, IA
50011
b) Department of Civil, Construction, & Environmental
Engineering, Iowa State University, 351 Town Engineering, Ames, IA
50011
c) School of Civil Engineering and Environmental Science,
University of Oklahoma, 202 West Boyd Street, Norman, OK
73019
Earthquake Spectra, Volume 32, No. 1, pages 239–265, February 2016;
© 2016, Earthquake Engineering Research Institute 239
over a short depth, thereby increasing its lateral stiffness and
allowing the pile lateral strength to be fully developed.
Some well-known methods for improving the soil conditions in the
field include deep soil mixing, jet grouting, stone columns, and
simple soft soil replacement. However, soil improvement techniques
are not often used in design practices due to limited understanding
of the behavior of improved soil and interactions between the pile,
improved soil, and unim- proved soil. In addition, this limited
understanding often results in overly conservative soil improvement
designs that include large improvement volumes and higher cost as
compared to an efficient design. Most experimental and analytical
studies focused on utilizing soil improvement to mitigate
liquefaction of loosely deposited sand but without the presence of
piles (Mitchell et al. 1998, Martin et al. 2001, Hatanaka et al.
1987, Adalier 1996, Adalier et al. 1998, Iai et al. 1988, Akiyoshi
et al. 1993, Liu and Dobry 1997, Kawakami 1996). Seismic behavior
of piles in liquefiable sands has also been extensively studied
(e.g., Ashford et al. 2000a, Ashford et al. 2000b, Weaver et al.
2005, Boulanger and Tokimatsu 2006, Ohtomo 1996), while a majority
of studies for piles in soft clay have been investigated by the
offshore community (Vucetic and Dobry 1988, Basack and Purkayastha
2007). Only a few studies have addressed the seismic behavior of
pile foundations constructed on soft clays (Brown et al. 2001,
Wilson 1998, Boulanger et al. 1999, Meymand 1998, Lok 1999, Mayoral
et al. 2005), despite the widespread presence of the soft clay soil
in high seismic regions and the frequent need to locate bridges and
buildings in this soil type. Only recently, a few investigations
have been carried out to determine the effectiveness of ground
improvement on increasing the lateral resistance of pile
foundations embedded in soft clay.
Rollins and Brown (2011) improved the quasi-static and dynamic
behavior of full-scale pile groups in the field by treating soft
clay in the vicinity of the pile and/or the pile cap using several
different soil improvement techniques including jet grouting, soil
mixing, flowable fill, soil replacement, and rammed aggregate
piers. Generally, soil improvement methods using cement as the
stabilizing agent (e.g., jet grouting, soil mixing, and flowable
fill) pro- duced the largest increases in lateral resistance
compared with other soil treatment methods. The jet grouting
technique was applied to the soil region surrounding the piles
below the pile cap in a block type configuration. Neglecting the
passive resistance of the pile cap against the soil, the lateral
resistance of the system increased by 220% compared to an
unimproved pile group. In addition, the initial stiffness of the
system increased by a factor of about 14 based on the results from
small displacement tests (<0.03m).
These experimental studies show that improvement methods using
cement to treat the soil are viable techniques for significantly
increasing the resistance of piles in soft soil. However, limited
experimental research has been performed to isolate the effects of
soil improvement surrounding single piles and to demonstrate the
impact that soil improvement has on enhan- cing the seismic
behavior of piles in soft clay. To address this issue, a full-scale
experiment using dynamic and quasi-static lateral loads applied at
the pile head was conducted on two piles having identical pile
properties and site conditions but one was embedded in cement deep
soil mixing (CDSM) improved soil and the other was not. Responses
of the improved pile are computed in terms of percent increases in
key response characteristics as compared to the unimproved pile to
demonstrate the improved seismic behavior of CDSM treated piles
in
240 FLEMING ET AL.
soft clay. To assist in estimating the response of CDSM improved
piles, recommendations are given to estimate the static responses
of the improved soil based on back-calculation of experimental
data. The method is verified using a computer program called LPILE
(Ensoft 2010) to compare computed soil responses to the observed
responses. Finally, the impact of soil improvement is demonstrated
using LPILE to determine the wall thickness and diameter of an
unimproved pile to achieve the same response of a standard pile in
improved soil that is taken from the experimental results.
TEST SITE
For the field experiment, a soft soil site in Miami, Oklahoma, was
chosen. The test site consisted of a 4.4 m layer of soft clay
overlying a 2.0 m layer of sandy gravel and limestone bedrock. The
top 1.0 m consists of lean clay with gravel and occasional
construction debris. The soft clay at the site was classified as
lean clay (CL) according to the Unified Soil Clas- sification
system (ASTM D2487). Soil characterization at the site included 11
piezocone soundings (CPTu), two soil borings, Shelby tube sampling,
and a stand pipe piezometer to measure the depth of the water table
(Taghavi et al. 2010).
Figure 1 shows the geotechnical characteristics of the soil profile
at the test site. Undrained shear strength (su) and effective
friction angle ( 0) of the soil were calculated from CPTu tip
resistance (qt) and pore-water-pressure (U) using Equations 1 and
2, respec- tively (NCHRP Synthesis 368):
0 50 100 150 200
S u (CPTu)
S u (Triaxial)
S u (kPa)
q t
q t (MPa)
Lean clay with gravel and occasional debris
Medium stiff to very soft silty clay (30 LL, 12 PI)
Sandy gravel
Ground surface
Figure 1. Average geotechnical characteristics of the Miami,
Oklahoma, test site.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 241
EQ-TARGET;temp:intralink-;e1;41;640su ¼ qt σvo Nkt
(1)
qtσatmffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
σvo
0σatm p
(2)
where σvo is equal to the total overburden pressure, Nkt is equal
to the dimensionless cone factor, and σatm is equal to the
atmospheric pressure. An average unit weight (γ) of 19 kNm3
for the clay and 20 kNm3 for the sand was used to calculate the
overburden pressure. The water table, measured by a stand pipe
installed at the site, was 2.62 m below the ground surface at the
time of pile testing. Piezocone soundings were performed when the
water table was about 0.80 m below the ground surface, as shown in
Figure 1. This depth was used to determine the effective overburden
pressure (σvo
0). Nkt was determined to be about 17 based on the average
undrained shear strength of soft clay samples taken from the field
and tested in the lab. Based on the CPTu measurements, the average
undrained shear strength of the soft clay was generally in the
range of 30 to 95 kPa between the depths of 1.1 m and 4.4 m below
the ground surface. Several Shelby tube samples were tested in the
lab for consolidated undrained shear strength of the soft clay soil
that are shown in Figure 1 and are detailed in the site
investigation report (Taghavi et al. 2010). Generally, the shear
strengths produced in these tests were between 35 kPa and 60 kPa
for depths between 1.5 m and 2.6 m. Also determined in the lab were
the overconsolidation ratio (OCR), the liquid limit (LL), and the
plastic index (PI) for the clay, which are also shown in Figure 1.
At depths greater than 4.4 m, the average friction angle of
granular soil was between 40 degrees and 45 degrees based on CPTu
soundings.
TEST UNIT DETAILS AND CONSTRUCTION
A three-dimensional (3-D) rendering of the full-scale test
configuration is shown in Figure 2. Two identical steel pipe piles
having AISC size HSS12.75 0.375 and satisfying ASTM A106B
specifications were installed in improved and unimproved soft clay.
The
Improved Soil
Reaction Frame
Soft Clay
(a) (b) Reaction Frame
soil not shown
Figure 2. (a) Soil profile and test configuration and (b) picture
of test site during dynamic testing of pile in improved soil.
242 FLEMING ET AL.
average Young’s modulus and yield strength for the steel in each
test pile, determined from four tensile tests performed on samples
cut from the sidewall of each pile, were 213 GPa and 372 MPa,
respectively. A yield plateau in the stress-strain curve initiated
at an average strain of 0.17% and terminated at an average strain
of 1.1%, at which point the steel exhibited strain hardening. The
average ultimate strength of the steel was 588 MPa, which occurred
at an average strain of 15%. The total length of each pile was 7.6
m. Approximately 5.2 m of each pile was embedded below the ground
surface and into the underlying sandy gravel layer to help support
the pile and reduce the amount of full length rotation under
lateral loading. A hole was dug around each test pile, near the
ground surface, having approximate dimen- sions of 1.2 m in
diameter and 1.0 m in depth, to remove the effects of the clay with
gravel layer and to assist in improving the uniformity of the clay
in contact with the pile.
A pile cap, consisting of two halves of a concrete block and acting
as a seismic mass needed for dynamic testing, was clamped to the
head of the pile to be tested. A set of threaded rods, sent through
the pile cap, avoiding the pile in the center, were used to clamp
the pile cap to the pile and to support a stiffened C15 40 channel
used for connecting the quasi-static and dynamic actuators. The
mass was kept constant for all tests to isolate the effect of
varying natural frequency of the system. Although observing changes
in the natural frequency of the system would benefit determining
the effects of superstructure mass, the volume of concrete
necessary to achieve the effective mass of a bridge deck or
building would be unfeasibly large to include in the experiment.
Therefore, a smaller mass was included to achieve a target natural
frequency within the testing range so that reasonable estimates of
system damping, which are best found at the natural frequency of
the system, could be determined.
Construction of the improved soil volumes occurred before the
installation of piles. The improved soil was constructed using
CDSM. In this process, an ABI Mobilram machine with an augur drive
attachment was used to revolve a mixing tool into the soft clay. It
consisted of a hollow shaft and mixing paddles, as shown in Figure
3a. Cement grout was pumped through the hollow shaft and ejected
laterally behind the lower mixing paddle, where it was mixed with
the native soil. While still revolving, the mixing tool was
advanced to a depth of 4.0 m and retracted. This process was
repeated once more to form a well-mixed column of soil and cement.
The top of the improved soil column after mixing is pictured
(a) (b)
Figure 3. Pictures of (a) Mobilram and auger with mixing paddles
and (b) top of soil improve- ment column.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 243
in Figure 3b. Concrete mixer trucks transported the grout from the
batch plant to the site and dumped the grout into a hopper where
the grout was continuously mixed and pumped through the supply line
for the mixer. Water from a nearby river was also pumped into the
hopper periodically to maintain cement and water concentration
close to the target value.
The improved volume was conservatively designed to test the full
strength of an im- proved pile relative to an unimproved pile. The
objective was to improve the soil such that the test pile would
yield and form a plastic hinge as opposed to rotating under the
lateral load. The nominal dimensions of the improved volume were
13D 13D 13D, where D is equal to the outer diameter of the test
pile. Because of the presence of the hole around the base of the
pile, a distance of approximately 9D along the pile length denotes
the contact length between the pile and improved soil.
To construct the large improvement block, a single column was
constructed, after which the mixing tool was offset slightly to
achieve an overlap of approximately 0.3 m between columns. The
process was repeated to form a block of improved soil with the CDSM
column arrangement shown in Figure 4. A total of 16 columns with
diameters of 1.2 m were con- structed, resulting in a plan
dimension of 4.0 4.0m. A test pile was then installed in the center
of the improved volume while the improved soil was still wet using
the same ABI Mobilram machine but with a vibrating hammer
attachment. This process allowed the pile to be installed with
ease.
In the case of retrofitting an existing foundation with CDSM,
although not as effective as the block type configuration chosen
for the experiment as demonstrated by Rollins and Brown (2010), the
improved soil could be constructed as a wall in close proximity to
the foundation to increase soil resistance to lateral loading of
the foundation elements includ- ing piles and pile cap. If
possible, jet grouting, which also uses cement grout to improve
soil resistance, can be applied to the soil underneath the pile cap
and adjacent to piles via holes drilled through the cap to allow
access to this region. These retrofit configurations were
3962 mm
3962 mm
Test Pile
Figure 4. Mobilram (left) and CDSM column arrangement and test pile
placement (right).
244 FLEMING ET AL.
studied by Rollins and Brown (2010) for pile groups and resulted in
significant increases in system stiffness compared to an unimproved
pile group. As a basis for future development of CDSM improved
piles, the purpose of the current research is to isolate the
effects of soil improvement surrounding a single pile in absence of
an underground pile cap.
The grout used to construct the CDSM columns consisted of a
water-to-Portland-cement (Type I) ratio of 11 by weight. A total of
1.893m3 of cement grout was added to each column, resulting in an
approximate concentration of 20% cement by dry weight of soil in
the CDSM column. The average undrained shear strength of the
improved soil, determined in the lab using samples collected from
the field, was 1517 kPa. This was computed using a
compressive-strength-to-shear-strength conversion factor of 3.2 as
opposed to the conven- tional value of 2, typically used for clay.
The revised value developed from testing of clay soil with high
cement concentrations in direct shear (Porbaha et al. 2000). These
authors estab- lished the correlation shown in Equation 3 as a
result of their research, which demonstrated a quadratic
relationship between the shear strength (τf 0) and compressive
strength (qu) in units of kgfcm2 for the improved soil. However, it
is important to note that loading conditions for simple shear are
difficult to achieve in the laboratory and that results from low
normal stress conditions used in the tests are more appropriate for
shallow soil-cement:
EQ-TARGET;temp:intralink-;e3;62;425τf 0 ¼ 0.53þ 0.37qu 0.0014q2u
for qu < 60 kgfcm2 (3)
Figure 5 shows the plan and profile views of the test specimens and
reaction frame. The reaction piles, having AISC size HP10 42 and
conforming to ASTM structural grade A572-50, were installed outside
the zone of influence of the test pile in unimproved soil (TPU) and
the test pile in improved soil (TPI) that, according to the
Canadian Geotechnical Society (1978), was taken as a radial
distance of 10D away from the center of the test pile. The two
reaction piles closest to TPI were driven into two separate col-
umns of improved soil to increase the strength of the reaction
frame. Approximately 6.2 m of each pile was embedded below the
ground surface and into the underlying sandy gravel layer.
A steel frame, consisting of HP10 42 steel members and mounted to
the four reaction piles as shown in Figure 5, was designed to
support the dynamic actuator while maintaining an appropriate
distance between the actuator and the pile cap. A removable
cantilever seg- ment, which provided a platform for mounting the
dynamic actuator to the frame, was designed to support the dynamic
actuator on both sides of the main reaction frame for testing of
both piles. After removing the cantilever segment, the actuator for
quasi-static testing was mounted to a cross beam supported by the
reaction piles of the frame (shown on the right side of Figure 2a).
All members and connections within the reaction frame were designed
to support a maximum 888 kN lateral force, after considering a
safety factor of 2, which was applied at the center of the clevis
for the dynamic actuator just above the can- tilever and at the
center of the quasi-static reaction beam, separately. Frame size
and con- figuration was based on soil properties determined from
multiple CPTu soundings at the site. The shear strength of the soil
was reduced by 50%, which conservatively accounts for the
variability of soil strength in the field and resulted in a
conservative reaction frame design.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 245
INSTRUMENTATION
Figure 5 shows the locations of instrumentation installed at the
site. Each test pile was instrumented with over 32 strain gages,
four displacement transducers, three tilt meters, and three
accelerometers. Steel angles, having a size L2 2 18, were welded to
the test piles to safeguard strain gages and cables during driving.
Generally, tack welds were placed every 305 mm along the length of
the test pile, in between the strain gage locations to reduce the
risk of damaging them. Displacements of each test pile were
measured at three different elevations along the freestanding
portion of the pile. A fourth displacement transducer was placed at
the same elevation as the top transducer but with a horizontal
separation of 300 mm to measure possible twisting of the pile cap
during testing. All displacement transducers were mounted to a
reference beam placed adjacent to the test pile and supported
5.26 m
Accelerometers
1.83 m
(b) Profile View
Figure 5. Plan and profile views of full-scale test
configuration.
246 FLEMING ET AL.
by hollow structural members, which were driven outside the zone of
influence, again, taken as a distance of 10D away from the pile
(Canadian Geotechnical Society 1978). Forces applied to the pile
cap were measured using the load cells integrated into the
actuators. Three accelerometers located on the ground surface,
which were placed adjacent to the test piles, were used to measure
the vibration transmitted from the pile to the soil. Subsurface
accelerometers were also installed but only near the unimproved
pile. Pore-water-pressure transducers were also installed below
ground near the unimproved pile.
LOADING PROTOCOL
Dynamic and quasi-static testing protocols were selected for each
pile to evaluate the cyclic behavior of the test piles at varying
levels of loading. Generally, test piles were sub- jected to at
least three cycles of the same amplitude and frequency, after which
the amplitude was increased and the next load pattern was applied.
Figure 6 shows the loading protocol used for both test piles. TPU
was first subjected to dynamic loading with force control enabled.
Displacement magnitudes were between 0.2mm and 115mm. Initially,
fre- quency of the excitation was varied between 0.25 Hz and 8 Hz
for displacement magnitudes of less than 2mm. However, due to the
velocity limitations of the actuator, excitation frequencies were
reduced with increasing displacement magnitudes, as noted in Figure
6. For instance, displacement magnitudes of 50mm were achieved with
2 Hz excitation fre- quency while the displacement cycles at 115mm
were applied at a frequency of 0.4 Hz. Undergoing this process also
helped with understanding the pile response as a function of
frequency.
Quasi-static loading was performed to test piles at larger lateral
displacements. Displace- ment magnitudes less than the maximum
displacement reached during dynamic testing were applied to TPU in
the early stages of quasi-static testing. Only one cycle of the
load pattern was applied for each interval of increasing
displacement up to the maximum displacement of 115mm, after which
three cycles of symmetric loading were applied at each interval as
the displacement magnitude increased to a maximum of 406mm. This
was the last symmetric load pattern, which was dictated by the
stroke limitations of the actuator. The remaining cycles displaced
the pile 406 mm in the push direction and 566 mm in the pull
direction from center.
Initially, TPI was subjected to small vibrations with force control
enabled. Force control was limited to magnitudes smaller than 2.5
kN. Subsequently, displacement control was enabled for displacement
magnitudes between1mm and100mm. Frequency of the exci- tation was
varied between 0.25 Hz and 8 Hz for displacement magnitudes of 1mm.
For displacement magnitudes larger than 1mm, a load pattern
sequence was applied to TPI with similar increasing displacement
magnitude increments and excitation frequencies as TPU. The target
quasi-static load pattern for TPI was also similar to the load
pattern chosen for TPU. However, the pile failed before completing
the entire load pattern sequence.
TEST OBSERVATIONS
Figure 7 shows some of the key test observations. Under cyclic
lateral loading, soft clay surrounding TPU experienced significant
plastic deformations, evident by soil losing contact with the pile
(i.e., gapping) after the pile head was brought back to zero
displacement.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 247
100 10
1 0.1
0.01 0.1
1 10
Cumulative Full-Cycle Sine Waves
0 2 4 6 8 10 12 14 16 21 23 26 28 41 44 47 50 53 74
D is
p la
ce m
en t
(m m
10 F
re q
u en
H z)
Additional cycles of same magnitude and frequency not shown for
clarity
Force Control
Displacement Control
Cumulative Full-Cycle Sine Waves
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20
D is
p la
ce m
en t
(m m
10
Cumulative Full-Cycle Sine Waves
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19
D is
p la
ce m
en t
(m m
*
(d)
Figure 6. Test sequences used in following load protocol: (a) TPU
during dynamic testing, (b) TPU during quasi-static testing, (c)
TPI during dynamic testing, and (d) TPI during quasi-static
testing.
Conversely, the improved soil surrounding TPI remained relatively
intact during all tests. Less than 3 mm total measured gap width
occurring at the ground surface and short hairline cracking, which
ran parallel to the direction of loading, was observed in the
improved soil directly in front and behind the pile. TPI reached
its lateral capacity at a displacement of about 10 cm, at which
point the critical region of the pile about 75 mm above the surface
of the improved ground experienced buckling and fracture due to low
cycle fatigue.
GAP OPENING
The soil surrounding TPU had undergone significant plastic
deformations, causing gap- ping between the pile and the soil
during unloading. Following each critical load cycle, the width of
gap was measured manually using a scale at the ground surface after
the pile was brought back to zero displacement. Figure 8 shows
average gap width versus average pile head displacement for
multiple cycles of varying head displacements of TPU, suggesting a
linear trend. This was also observed in Suleiman et al. (2006)
during lateral load tests of a drilled shaft in frozen and another
drilled shaft in unfrozen soil. A linear regression through
(a) (b) (c)Hairline Crack
Figure 7. Test observations including (a) gapping of soil adjacent
to TPU, (b) hairline cracks in improved soil adjacent to TPI, and
(c) fracture of TPI after 10 cm displacement.
Pile Head Displacement (mm)
G ap
W id
th (m
y = 0.54x - 11
Figure 8. Measured gap width at the ground surface for TPU.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 249
the data shows a y-intercept of −11mm, which suggests the soil
adjacent to the sidewall of the pile at the ground surface
generally displaced 11 mm in the same direction as the pile
immediately after reversing the load. The rebounding effect of the
soil occurs regardless of displacement magnitude. This was verified
by comparing the displacement of the pile near the ground surface
and the measured gap width, which, on average, was a difference of
about 11 mm. This evidence is also present in the back-calculated
responses of soil displacement. From these measurements, the soil
rebound could be observed in higher resolution. It was found that
soil rebound increased with increasing displacement of the pile up
to a pile dis- placement magnitude of about 37 mm, after which the
rebound remained relatively constant at 11 mm. Gapping for TPI was
negligibly less than 3 mm for all quasi-static tests.
FORCE-DISPLACEMENT RESPONSES
The graphs in Figures 9 show the dynamic and quasi-static
force-displacement responses obtained at the pile head for TPU and
TPI. These responses show distinctively different hysteresis
behaviors, which are largely caused by the introduction of improved
soil in TPI. In addition, varying excitation frequency had
negligible effect on the global force- displacement response of the
system, as can be seen in the response envelope, as no abrupt
change to the stiffness or force resistance is seen under
increasing excitation frequency. Equivalent viscous damping ratios
(ξeq) were calculated using Equation 4 (Chopra 1995) for each cycle
at varying pile head displacements to quantify the inelastic energy
dissipa- tion (i.e., system damping with the linearly elastic
damping component removed) in both systems:
EQ-TARGET;temp:intralink-;e4;41;367ξeq ¼ 1
-200
-150
-100
-50
0
50
100
150
200
250
-250
-200
-150
-100
-50
0
50
100
150
200
250
Py - = -152 kN(a) (b)
Figure 9. Global force-displacement response of (a) TPU and (b)
TPI.
250 FLEMING ET AL.
where the inelastic energy dissipation is equal to the area within
the force-displacement hys- teresis loop (ED) of the quasi-static
data and ES is equal to the maximum strain energy in the system or
ksecu2m2 with a system stiffness ksec and maximum displacement um.
Figure 10 shows ξeq for quasi-static tests in TPI and TPU except
for tests associated with the þ400mm 475mm cycle of TPU due to the
asymmetric loading. Global responses for each system are detailed
in the paragraphs below.
UNIMPROVED PILE (TPU)
The strength and stiffness of TPU, determined from the quasi-static
response of the sys- tem, were estimated by obtaining lateral
applied force and head displacement at the first occurrence of
yielding in the pile, which corresponds to a positive strain of
0.17% developing along the outer surface of the pile. The pile
first yielded 1.32 m below the ground surface at an applied lateral
head force of 129 kN, termed herein as system strength (Py), and a
lateral head displacement of 17 cm. Therefore, the secant stiffness
of the system was 759 kNm. At the beginning of the experiment, TPU
was loaded with equal displacements in the push and pull directions
of loading. However, the midpoint of the full extension and
retraction range of the actuator was about 9 cm closer to the
reaction frame compared to the centerline of the pile, which caused
the force-displacement plot of the last three cycles to be
unsymmetrical in full range of the actuator displacement.
Generally, ξeq increased with increasing head displacement except
for lateral head dis- placements less than 11 cm, where previous
dynamic testing had disturbed the surrounding soil. In this region,
ξeq decreased with increasing head displacement because ksec of the
sys- tem increased as lateral head displacement approached the
previous maximum head
Figure 10. Equivalent viscous damping ratio for independent
hysteresis loops during quasi-static testing.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 251
displacement of 11 cm, which subsequently increased the elastic
strain energy stored in the system. In addition, ξeq decreased with
increasing number of cycles for any given head dis- placement. This
is attributed to an increase in gap width between the pile and soil
with increasing number of cycles and hence less soil is engaged
during repeated cycles.
IMPROVED PILE (TPI)
The first occurrence of yielding in TPI was at a negative head
displacement (pull) of 2.6 cm, which is an 85% decrease in yield
displacement compared to TPU. Correspondingly, Py was 152 kN in the
pull direction, which is 18% higher than Py for TPU, which is a
result of the plastic hinge located 1.0 m closer to the ground
surface. The stiffness of TPI, estimated using the same method to
find the stiffness of TPU, was 5;846 kNm, which is over seven times
greater than the effective stiffness of TPU. The pile also yielded
in the push direction, which was at a lateral head displacement of
about 3.6 cm. The force applied at the top of the pile was 191 kN,
resulting in an effective stiffness of 5;306 kNm, taken from the
origin of the force-displacement curve to the yield point.
Compared to TPU, ξeq is larger for TPI and is more consistent with
an increasing number of cycles, which suggests that almost all
energy dissipation was through the plastic action of the pile. For
head displacements less than 5 cm, ξeq values were less than 8%. On
occasion, ED computed from testing TPU was larger compared to ED
values of TPI, which was observed in the areas within hysteresis
loops, especially at larger displacements. Although energy
dissipation was larger in the system for TPU, so was ES due to
large lateral displace- ments of the pile head, resulting in
smaller ξeq values by comparison.
RESPONSE PROFILES
Figure 11 shows the strain profiles for selected head displacements
in the push direction for TPU and TPI. Plastic zones developed in
TPU and TPI at the locations corresponding to strains larger than
the yield strain (εy), which is equal to 0.17%. For simplicity,
only the strain gage measurements on the extreme tension side of
the pile when the pile is pushed are dis- cussed in this section.
Generally, the strain profiles on the extreme compression side in
the push direction mirror the measured strains shown in Figure 11
when the pile remains elastic. However, strains larger than εy are
influenced by the loading history and the cyclic moment- curvature
response of the pile section. This is the reason for the observed
increases and decreases in strain measurements on the extreme
tension side of TPU and TPI for increasing head displacement beyond
the limit required to yield the pile. However, the strain measure-
ments on the extreme compression side (not shown) change
appropriately to follow the moment-curvature behavior of the
section. Important parameters including plastic hinge loca- tion
and length can be observed in the data provided and are noted for
each system in the following paragraphs.
Maximum strain in TPU occurred 1.95 m below the ground surface and
the plastic zone near this location was approximately 2.5 m long
when a lateral head displacement of 400 mm was applied in the push
direction. Strains were also sufficiently large to cause a plastic
zone to occur along a short segment near the ground surface. The
steel angles protecting the strain gages were terminated at this
location, which caused a change in the flexural rigidity (EI) of
the pile with respect to distance along the pile length. The EI of
the pile was not only affected
252 FLEMING ET AL.
by the presence of the angles but also by the location and distance
between welds connecting the angles to the pile. Based on the
back-calculation of EI from strain gage measurements and applied
lateral loads, the elastic EI was found to increase from 25;400
kN-m2 (EI for a pile cross section without angles) at the
termination of the angles to a depth of about 2.2 m below the
ground surface, at which point the maximum EI of 33;400 kN-m2 for a
cross section including angles was achieved. This increase in EI
also increased the moment required to yield the pile based on
theoretical moment-curvature analyses of two cross sections. One
pipe section without angles and another with angles were analyzed
for flexural strength based on properties of the steel determined
in the lab. Bending forces sufficient to cause a pipe section to
yield without and with angles were 258 kN-m and 285 kN-m,
respectively. The moment in TPU, at 10 cm below ground, was within
this range based on strain gage measurements observed in Figure
11a. Therefore, the pile underwent yielding at this location.
However, the rate of increase in EI with respect to depth was
greater than the increase in the ratio of yield moment over the
curvature in the pile. Therefore, the pile did not yield at loca-
tions where the product of curvature and EI were less than the
yield moment. Once EI reached its maximum value, yielding occurred
again at the maximum moment location and the plastic zone increased
in length from this location with increasing load.
The maximum strain in TPI was located just below the ground
surface. The sidewall of TPI experienced inelastic strains at head
displacements larger than 36 mm in the push direc- tion, even in
phase 1 testing, which occurred during the 50mm displacement cycle,
as can
0 2000 4000 6000 8000 10000 -2
-1
0
1
2
3
4
5
6
Strain ( )
)
+/- 25 mm - 1.0 Hz +/- 50 mm - 1.0 Hz +/- 100 mm - 0.4 Hz +/- 150
mm +/- 200 mm
0 2000 4000 6000 8000 10000 -2
-1
0
1
2
3
4
5
6
Strain ( )
)
+/- 50 mm - 4.0 Hz +/- 100 mm - 4.0 Hz +/- 200 mm +/- 400 mm +/-
475 mm
Ground Surface Ground Surface
Bottom of Pile
30 cm195 cm
Push of +400 mm after a pull of -475 mm
(a) (b)
Figure 11. Strain profile of (a) TPU and (b) TPI.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 253
be seen in Figure 11b. The plastic hinge that developed in TPI was
approximately 90 cm long after the 200mm displacement cycle and
extended 61 cm below the ground surface. Also note that strain
measurements were a constant zero below a depth of approximately
1.5 m, which is within the improvement zone. This observation
suggests that the depth of the improved soil could have been
reduced by about 50% along the region where the pile is in contact
with the improved soil.
BACK-CALCULATING SOIL RESPONSE
Soil response can be back-calculated analytically from experimental
data assuming that soil displacement (y) and soil resistance (p) at
discrete layers along the length of the pile are equal to pile
displacement and the load distribution along the length of the
pile, respectively. Theoretically, based on Euler–Bernoulli beam
theory, y and p are calculated by double inte- gration of the
pile’s curvature and double differentiation of the pile’s moment,
respectively, with respect to the distance along the length of the
pile. If the pile behaves elastically, the moment profile can be
calculated in terms of cross-sectional behavior of the pile as the
pro- duct of flexural rigidity and curvature of the pile. In this
study, the theoretical flexural rigidity of the pile was calculated
and applied to discrete values of curvature obtained from strain
gage measurements. Brandenberg et al. (2010) have shown that the
accuracy of p is greatly affected by the method used to
differentiate the data. These authors further found that a dif-
ferentiating technique of minimizing weighted residuals gives
comparable performance to the method of fitting cubic splines
(Segalman et al. 1979) to the data and outperforms the curve
fitting of the higher-order polynomials method (de Sousa Coutinho
2006). The weighted- residual method performed slightly better than
the cubic spline method when noisy data are sampled using small
intervals along the pile length. The higher-order polynomials
method produced accurate derivatives for uniform soil profiles but
was not suitable for layered soil. Alternatively, a generalized
constraint function can be used for soil response and optimized
based on pile response and cross-sectional behavior as demonstrated
by Tehrani et al. (2014). However, this method requires a
constraint function for soil behavior, which has not been
established for improved soil. For these reasons, the
weighted-residual method was used to back-calculate soils response
in the current research.
WEIGHTED-RESIDUALS METHOD
Suppose there is a desired function gðzÞ equal to the derivative or
integral of a function f ðzÞ with respect to the distance z along
the pile length. As mentioned previously, only dis- crete values f
i are known at locations of zi. Correspondingly, gi are discrete
values obtained from the weighted-residuals method. Complete
details of the differentiation method can be found in Brandenberg
et al. (2010). The integration method can be performed in a similar
manner, which has not been described in previous literature.
Therefore, both methods are presented below in a compact form
as
EQ-TARGET;temp:intralink-;e5;41;154g ¼ df dz
≈ ½G ¼ 3½Z1½Δ½F; (5)
EQ-TARGET;temp:intralink-;e6;41;111g ¼ ð L
3 ½Z½Δ1½F; (6)
254 FLEMING ET AL.
f n1
2 66666664
. .. . ..
. . . .
2 66666664
2ðz2 z1Þ ðz2 z1Þ 0 0 · · · ðz2 z1Þ 2ðz3 z1Þ ðz3 z2Þ 0 · · ·
0 ðz3 z2Þ 2ðz4 z2Þ ðz4 z3Þ · · ·
. . . ..
. .. . ..
. . . .
· · · 0 ðzn1 zn2Þ 2ðzn zn2Þ ðzn zn1Þ · · · 0 0 ðzn zn1Þ 2ðzn
zn1Þ
3 77777775 :
Equations 5 and 6 were applied twice to moment and curvature data
to calculate p and y, respectively. Appropriate boundary conditions
were required to determine both rotation and displacement of the
pile. For small pile head displacements, as observed during dynamic
testing, the displacement and rotation were zero at the bottom of
the pile based on strain profile measurements. However, for head
displacement magnitudes of 100 mm and larger, the strain profiles
penetrated almost to the bottom of the pile, indicating potential
for full pile rotation. In this case, boundary conditions were
obtained using tilt sensors and displacement transducers located
near the pile head for quasi-static test results. The boundary
conditions for shear in the pile and soil reactions were known at
the ground surface and provided an additional discrete measurement
in the system of equations. The shear at the ground surface was
taken from load cell measurements in the actuator and the soil
reaction along the unsup- ported length of the pile was taken as
zero.
FIELD P-Y CURVES
Figure 12 shows the average normalized soil resistance (ppu) for
the CDSM improved soil and soft clay soil, which were calculated
from strain gage measurements taken in the field, where pu is the
ultimate resistance of the soil and is dependent on the effective
unit weight of the soil (γ 0s), undrained shear strength of the
soil (su), depth of the soil layer (z), and diameter of the pile
(d). Normalizing the soil resistance by computing ppu removes the
effects of the soil properties and overburden pressure such that
selected curves can be plotted together and verified with models
found in the literature. An appropriate theoretical pu profile was
used for each pile. Using limit equilibrium methods, Matlock (1970)
proposed the pu profile for clay to be computed as
EQ-TARGET;temp:intralink-;e7;62;136pu ¼ min
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 255
The model chosen to represent the unimproved soil was derived by
Matlock (1970), which is a model for soft clay. To the authors’
knowledge, the derivation of a p-y model for improved soil has not
been attempted by other researchers. However, the response of CDSM
soil may be similar to that of stiff clay (without free water) or
weak rock, which have predicted p-y curves based on the work
performed by Reese and Welch (1975) and Reese (1997), respectively.
Guo (2013) developed a general-use model for sand and clay, which
represents soil behavior with a linear elastic, perfectly plastic
p-y curve. All four of the aforementioned models were compared to
the back-calculated curves shown in Figure 12.
The back-calculated resistance of the soft soil is largely
nonlinear, as shown in Figure 12a. The back-calculated curve fits
Matlock’s curve well for soil displacements less than 20mm,
especially in the positive displacement region of the curve.
According to the figure, at maximum soil displacements (jyj >
50mm), the normalized resistance of the back- calculated curve
reaches an apparent maximum value less than the maximum soil
resistance of the model, which is likely due to a reduction in
effective overburden stress caused by the excavation of soil around
the test pile. Both the model and the back-calculated curves were
normalized using the same theoretical value for pu, which was
computed using Equation 7, taking z as the depth from the top of
the excavation. In reality, z is approximately equal to zero at the
bottom of the excavation and equal to the theoretical value at a
depth where the exca- vated soil has minimal effect on the
overburden pressure. This effect is difficult to determine for
practical purposes and thus was not accounted for when constructing
Figure 12a.
Figure 12b shows the p-y behavior of the improved soil. It is seen
that the resistance of the improved soil is relatively linear for
small lateral displacements (jyj < 1.5mm). An attempt
Figure 12 Normalized p-y responses for (a) unimproved soil and (b)
improved soil.
256 FLEMING ET AL.
was made to match the improved soil with a model developed for
stiff clay (Reese and Welch 1975), as was done by Huang (2011).
However, the stiffness of the model was significantly less than the
stiffness back-calculated from field experiments for soil
displacements greater than 1.5mm. The model developed for weak rock
(Reese 1997) has an initial linear p-y stiffness (K) dependent on
the Young’s modulus of the soil (Es) and a dimensionless constant
kir that was derived from experimental data. However, recommended
kir values for weak rock are significantly higher than the
back-calculated values found in the experiment, as can be seen in
Figure 12b. Despite this difference, both the model and the
back-calculated curve reach pu at about the same displacement of
about þ1.5mm. Both the stiff clay and rock models show large
nonlinear behavior, which was not observed in the back-calculation
of p-y for improved soil. The model described by Guo (2013) is
linear and fits the back- calculated curve very well. According to
the author, K for this model can be estimated as
EQ-TARGET;temp:intralink-;e8;62;493K ¼
0.087þeð11.49þ50eÞ (8)
where e is the ratio of the unsupported length of the pile over the
length of the pile below the ground surface, Ep ¼ EIpðπd464Þ is the
Young’s modulus of an equivalent solid pile, EIp is the flexural
rigidity of the pile, Es is the Young’s modulus of the soil, and ν
is the Poisson’s ratio of the improved soil. Es was taken as 500
MPa, which was the average modulus of stress-strain curves produced
by unconfined compression tests of the improved soil from the
field. This resulted in a K equal to 984MNmm assuming a Poisson’s
ratio of 0.25. The pu for both the Guo (2013) model and
back-calculated curve were computed using the theoretical equation
shown in Equation 7, where z was taken as the depth from the bottom
of the excavation.
ANALYSIS
Figures 13 and 14 show the results of a model compared to the
responses of the piles observed in the field. In this analysis
method, the pile was represented as a series of nonlinear
beam-column elements. These elements were connected to a bed of
nonlinear springs repre- senting soil response. The springs’
resistance, p, and deflection, y, were considered to vary
continuously with depth. Empirical models developed by Matlock
(1970) and Reese et al. (1974) were utilized for the response of
unimproved clay and sand, respectively. A theore- tical model
developed by Guo (2013) was used for the improved soil. The
responses of each of these models can be found above and the
properties of the system can be found in the details of the
experiment. Recommendations for p-y curves were based on
homogeneous soil layers with constant properties horizontally to
infinite distances. This was not satisfied at depths where improved
soil was present. However, using the response of TPI, it is shown
for soil improvement of sufficient width that the model deployed in
this paper was satisfactory. A computer program called LPILE,
developed by Ensoft, was used to ana- lyze the lateral load
behavior of a pile using the Winkler spring bed model
Generally, there was a good comparison between the field and
analysis for both systems. The global responses for TPI and TPU in
the field were not symmetric in terms of the force magnitudes in
the push and pull directions (i.e., the magnitude of the force in
the pull direc- tion is larger than the force in the push
direction), whereas the analysis is symmetric. This was
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 257
-500 -400 -300 -200 -100 0 100 200 300 400 500 -250
-200
-150
-100
-50
0
50
100
150
200
250
)
Displacement (mm) -500 -400 -300 -200 -100 0 100 200 300 400
500
-250
-200
-150
-100
-50
0
50
100
150
200
250
Displacement (mm)
(a) (b)
Figure 13. Comparing lateral pile head responses in the field
(solid) and LPILE (dashed) for (a) TPU and (b) TPI.
-300 -200 -100 0 100 200 300
-2
-1
0
1
2
3
4
5
6
)
+/- 8 mm - 8.0 Hz +/- 13 mm - 4.0 Hz +/- 19 mm - 2.0 Hz +/- 25 mm -
1.0 Hz Model
-300 -200 -100 0 100 200 300
-2
-1
0
1
2
3
4
5
6
)
+/- 13 mm - 4.0 Hz +/- 19 mm - 4.0 Hz +/- 25 mm - 1.0 Hz +/- 100 mm
- 0.4 Hz Model
Ground Surface
(a) (b)
Figure 14. Comparing response profiles for (a) TPU and (b) TPI in
the field and LPILE.
258 FLEMING ET AL.
likely the result of the frame and reaction piles interacting with
the test pile. Since the analysis does not account for the
influence of the frame, the response of the analysis does not match
the response observed in the field in the pull direction. The
moment profiles of the analysis also show good agreement with the
field results.
EVALUATING EQUIVALENT RESPONSES
As previously described, soil improvement significantly increases
the lateral stiffness and strength of a pile in soft clay and may
be a cost-effective solution as opposed to increasing pile diameter
and number of piles in the foundation. The cross-sectional area of
the pile (Ap) is directly related to the amount of steel in the
pile and approximately correlated to the cost of the pile.
Increasing the diameter of the pile also increases the cost of
manufacturing and installing the pile. However, the costs of
materials and construction are controlled by fluc- tuating market
prices across different areas of the world, which makes a detailed
cost analysis difficult and outside the scope of this paper.
Therefore, the motive for this section is to describe the
effectiveness of CDSM soil improvement in terms of an equivalent
pile in unim- proved soil but with enhanced section properties that
provide the same equivalent system strength and stiffness as a pile
in CDSM improved soil but with a smaller cross section. The
responses of TPI were used to compare with the responses of a
theoretical pile developed in LPILE (Ensoft 2010) with the same
soft soil conditions observed in TPU. Relative increases in the
diameter and wall thickness of the unimproved pile were provided as
a result of this comparison.
The strength of the system was taken as the force applied at the
pile head required to yield the steel along the outer radius of the
pile. This allowed the full strength of the pile to develop while
the surrounding soil provided the resistance required to resist the
applied load. System stiffness was also determined at the first
occurrence of yielding in the pile and was calculated by dividing
the system strength by the lateral displacement at the pile head.
To find the equivalent strength and stiffness of the theoretical
pile in unimproved soil equal to that of TPI, the pile diameter and
wall thickness were varied in LPILE until the response of the
system met the following conditions at the pile head
simultaneously: (1) a maximum strain in the pile equal to 0.17%,
which was εy, found in testing of the steel in the lab, (2) a
lateral force applied at the pile head equal to 152 kN, which was
Py observed in testing of TPI, and (3) a lateral displacement of
the pile head equal to 2.6 cm, which was the pile head displacement
that caused yielding in TPI.
Figure 15 shows the effect of system strength and stiffness in
terms of Ap for indepen- dently varying diameter and wall thickness
of the unimproved pile in LPILE. As the diameter of the unimproved
pile was increased and the wall thickness was held constant, the
resistance of the soil and pile increased, which in turn increased
the strength of the system. Increasing the wall thickness of the
pile and keeping the diameter constant also increased system
strength due to the additional resistance provided by the pile at
yielding, but the resistance of the soil remained constant.
Therefore, increasing the diameter of the pile for a given Ap had a
greater effect on system strength than increasing wall thickness
due to increasing resistance provided by the soil.
System stiffness was observed to have an optimal section size
denoted by the peak in the arched curves shown in Figure 15b. It
was observed that increasing Ap increased the applied
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 259
force required to yield the pile. Therefore, the stiffness of the
soil must decrease with increas- ing Ap due to the nonlinearity of
its behavior for a given pile diameter. It was also shown that pile
stiffness increased with increasing Ap. The peak in the curve of
constant diameter denotes the condition where system stiffness
transitions from pile stiffness dominated behavior to soil
stiffness dominated behavior. Therefore, the stiffness of the
theoretical unimproved pile in LPILE could not achieve the
equivalent stiffness of TPI for a diameter that was less than
0 50 100 150 200 250 0
2000
4000
6000
8000
10000
12000
0
200
400
600
800
1000
100 cm
(a) (b)
Figure 15. System strength and the effects of varying pile diameter
and wall thickness for a theoretical pile in unimproved soil.
0 20 40 60 80 100 120 140 160 180 0
40
80
120
160
200
240
280
320
360
TPI (Field) Equivalent System (LPILE)
Figure 16. Force-displacement behavior of TPI and LPILE analysis of
a pile with a 254 cm outside diameter and 0.3 mm wall
thickness.
260 FLEMING ET AL.
sufficiently wide. A diameter of at least 120 cm resulted in soil
resistances large enough to achieve the equivalent stiffness
provided that the wall thickness was greater than 5 mm. However,
the system strength corresponding to this section size was too
large compared to the strength of TPI. Therefore, the wall
thickness of the pile was reduced to decrease the strength of the
system. However, this also resulted in the reduction of system
stiffness, which had to be compensated by increasing the diameter
of the pile. Through this optimiza- tion procedure, the pile
diameter and wall thickness that produced an equivalent system
response to that of TPI were determined to be 254 cm and 0.3 mm,
respectively.
Figure 16 shows the force-displacement response of the equivalent
pile in LPILE com- pared to the measured response of TPI in the
field. The stiffness and strength of both systems were comparable,
as expected. However, the capacity of the equivalent pile was
generally 36% higher than the measured capacity observed for TPI in
the field. This was attributed to a longer plastic hinge length and
larger pile diameter in the equivalent pile compared to TPI.
CONCLUSIONS
With the objective of understanding and improving the seismic
behavior of piles sup- porting foundations in soft soil, a
full-scale field investigation was conducted by applying dynamic
and quasi-static loads to piles in improved and unimproved soft
clay. Both test units had identical pile dimensions and site
profiles. One of the piles was embedded in a volume of improved
soil (TPI), applied using the CDSM technique, and the other was
left unimproved (TPU). From the experimental results presented in
this paper, the following conclusions have been drawn:
1. The secant stiffness of TPU was 759 kNm at the first occurrence
of yielding in the pile, which occurred at a lateral head
displacement of 17 cm. The pile had undergone significant lateral
displacement before the capacity of the pile was achieved. This was
due to limited resistance of soft clay and as a result reduces the
effectiveness of the pile at lateral displacements less than 5 cm,
which is typically taken as the limit in seismic foundation design.
In addition, a gap had formed between the pile and soft clay, which
caused the system to become more flexible with increasing number of
cycles and magnitude.
2. The impact of soft clay was minimized with the application of
CDSM and increased shear and moment in the pile of TPI, allowing
the pile to achieve its capacity at more reasonable lateral
displacements (i.e., <5 cm). The stiffness and strength of the
sys- tem were 5;846 kNm and 152 kN, respectively, at the first
occurrence of yielding in the pile, after which the pile had
buckled and fractured just above the ground surface due to low
cycle fatigue. In addition, negligible gapping and cracking to the
CDSM improved soil was observed, even after 200mm of displacement
was applied to the pile head. For displacements beyond yielding
condition, the energy dissipation of TPI was mostly contained
within the pile.
3. With respect to the response of TPU, the improved soil
introduced the following changes to the lateral load responses of
TPI:
• Increased the system strength by 42%. • Increased secant
stiffness at yielding condition of the pile by a factor of
approximately 7.
FULL-SCALE SEISMIC TESTINGOF PILES IN IMPROVED ANDUNIMPROVED SOFT
CLAY 261
• Increased the equivalent damping ratio by a factor of
approximately 7.5. • Shifted the locationof
themaximummomentupwardsbyapproximately130cm.
4. LPILE, a computer program that utilizes the Winkler spring bed
model for the non- linear response of the soil in terms of soil
resistance, p, and displacement, y, per- forms satisfactorily in
predicting the monotonic response of a pile in unimproved soil and
improved soil of sufficient width and depth. The p-y curves
utilized in LPILE for unimproved soil also compare to the p-y
relationship back-calculated from pile strain measurements in the
field. The p-y response of the improved soil may be estimated using
a linear elastic, perfectly plastic model with a stiffness (K) of
984MNmm.
5. Two criteria were met for achieving the same strength and
stiffness of TPI using larger section sizes for a pile in
unimproved soil. The results showed that a pile section having an
outside diameter of 254 cm and a wall thickness of 0.3 mm can
achieve a strength and stiffness similar to that of an improved
pile. Installing a pile with this section size is not possible with
the construction equipment avail- able for pile driving. However,
these results do show the diameter of the pile required to engage
enough soft clay to resist the same loads applied to TPI.
Therefore, applying CDSM is an effective means to achieve 100% of
the strength of the pile.
ACKNOWLEDGMENTS
This material presented in this paper is based upon work supported
by the National Science Foundation under NEESR-SG Grant No.
CMMI-0830328. Dr. Richard J. Fragaszy serves as the program manager
for this grant. Any opinions, findings, and conclusions or
recommendations expressed in this paper are those of the authors
and do not necessarily reflect the views of the National Science
Foundation.
The Oklahoma Department of Transportation, the City of Miami, and
the Grand River Dam Authority provided support in locating and
evaluating the soil at the test site. Juan Baez, president of
Advanced Geosolutions Inc., and Arul Arulmoli, principal of Earth
Mechanics Inc., provided valuable expertise and resources for the
construction of the improved soil. Charbel Khoury from the
University of Oklahoma and Owen Steffens and Douglas Wood from Iowa
State University provided valuable assistance in constructing the
test setup. Robert Nigbor, Steve Keowen, and Alberto Salamanca of
nees@UCLA provided valu- able expertise and resources instrumenting
and testing the piles in the field.
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(Received 27 January 2014; accepted 24 November 2014)
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Full-Scale Seismic Testing of Piles in Improved and Unimproved Soft
Clay
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