-
ti
eerin
a r t i c l e i n f o
Article history:Received 22 April 2013Received in revised form
10 July 2013Accepted 29 August 2013Available online 23 September
2013
Keywords:Soilpilestructure interactionSeismic response
a b s t r a c t
tied as the key point in the performance-based seismic design
[2],where the overall performance of the building is controlled
duringthe seismic design process.
For determining the seismic response of structures, it is a
com-mon practice to assume the structure is xed at the base. In
fact, ifthe ground is stiff enough (e.g. structure founded on solid
rock) it is
to be consideredimposeddue to th
mic deformare induced within the structure due to the underneath
soThe process, in which response of the soil inuences theof the
structure and response of the structure inuences the mo-tion of the
soil is referred to as soilstructure interaction [3].
The dynamic equation of the motion for the structure (Fig. 1)can
be written as:
Mfug Cf _ug Kfug M1ug Fv 1where [M], [C] and [K] are the mass,
damping, and stiffness matricesof the structure, respectively. {u},
{ _u}, and {} are the relative nodal
Corresponding author. Tel.: +61 411603532; fax: +61
295142633.E-mail addresses:
[email protected] (A.S. Hokmaba-
Computers and Geotechnics 55 (2014) 172186
Contents lists availab
d
lsedi), [email protected] (B. Fatahi),
[email protected] (B. Samali).strength to performance. The
development of capacity designprinciples in the 1970s [1] was an
expression of the realisation thatthe distribution of strength
through a building was more importantthan the absolute value of the
design base shear which can be iden-
structed on soft soils, two modications needfor determining the
seismic response. First, theto the structure differs from the free
eld motionence of the structure. Secondly, additional
dyna0266-352X/$ - see front matter 2013 Elsevier Ltd. All rights
reserved.http://dx.doi.org/10.1016/j.compgeo.2013.08.011motione
pres-ationsft soil.motion1. Introduction
The seismic design of buildings has been undergoing a
criticalreappraisal in recent years, with change of emphasis
from
reasonable to assume that the input motion of the structure due
toa design earthquake is essentially identical to the motion of
thefree eld, which is dened as the motion experienced at the
samepoint before the structure is built. However, for structures
con-Shaking table testFLAC3DLaminar soil containerFrictional
pilesFloating pilesThe role of the seismic soilpilestructure
interaction (SSPSI) is usually considered benecial to the
struc-tural system under seismic loading since it lengthens the
lateral fundamental period and leads to higherdamping of the system
in comparison with the xed-base assumption. Lessons learned from
recentearthquakes show that xed-base assumption could be
misleading, and neglecting the inuence of SSPSIcould lead to unsafe
design particularly for structures founded on soft soils. In this
study, in order to bet-ter understand the SSPSI phenomena, a series
of shaking table tests have been conducted for three differ-ent
cases, namely: (i) xed-base structure representing the situation
excluding the soilstructureinteraction; (ii) structure supported by
shallow foundation on soft soil; and (iii) structure supportedby
oating (frictional) pile foundation in soft soil. A laminar soil
container has been designed and con-structed to simulate the free
eld soil response by minimising boundary effects during shaking
tabletests. In addition, a fully nonlinear three dimensional
numerical model employing FLAC3D has beenadopted to perform
time-history analysis on the mentioned three cases. The numerical
model adopts hys-teretic damping algorithm representing the
variation of the shear modulus and damping ratio of the soilwith
the cyclic shear strain capturing the energy absorbing
characteristics of the soil. Results are pre-sented in terms of the
structural response parameters most signicant for the damage such
as foundationrocking, base shear, oor deformation, and inter-storey
drifts. Comparison of the numerical predictionsand the experimental
data shows a good agreement conrming the reliability of the
numerical model.Both experimental and numerical results indicate
that soilstructure interaction amplies the lateraldeections and
inter-storey drifts of the structures supported by oating pile
foundations in comparisonto the xed base structures. However, the
oating pile foundations contribute to the reduction in the lat-eral
displacements in comparison to the shallow foundation case, due to
the reduced rockingcomponents.
2013 Elsevier Ltd. All rights reserved.Assessment of
soilpilestructure interacresponse of mid-rise buildings sitting
on
Aslan S. Hokmabadi , Behzad Fatahi, Bijan SamaliCentre for Built
Infrastructure Research (CBIR), School of Civil and Environmental
Engin
Computers an
journal homepage: www.eon inuencing seismicoating pile
foundations
g, University of Technology Sydney (UTS), P.O. Box 123, Sydney,
Australia
le at ScienceDirect
Geotechnics
vier .com/ locate/compgeo
-
truclate
anddisplacements, velocities and accelerations of the structure
with re-
(a)
Fig. 1. Schematic modelling of the multi degree freedom
structure considering: (a) sdeformation and rocking of the
structure supported by oating pile foundation; (c)
A.S. Hokmabadi et al. / Computersspect to ground, respectively.
{g} is ground acceleration, and {Fv} isthe force vector
corresponding to the viscous boundaries. This vec-tor is nonzero
only when there is a difference between the motionon the near side
of the articial boundary and the motion in the freeeld [4]. The
role of the seismic soilpilestructure interaction(SSPSI) is usually
considered benecial to the structural system un-der seismic loading
since it elongates the period of the structure andincreases the
damping of the structural system, so the considerationof SSPSI
tends to reduce the base shear and in turn structural de-mand of
the superstructure in comparison to the xed-base condi-tion. In
contrast, as shown in Fig. 1, SSPSI may increase the
overalldisplacement of the superstructure in comparison to the
xed-basecondition due to translation and rotation of the foundation
(e.g.Guin and Banerjee [5]; Yingcai [6]). The rocking stiffness is
devel-oped due to the resistance of the piles to vertical movement
[7],as shown particularly in Fig. 1b. Ma et al. [8] showed that
rockingmay be the most critical mode of vibration for a foundation
becauseof the very low hysteretic (material) damping, which will
lead tohigh motion amplitude when the excitation frequencies are
nearthe resonance state. The increase in the lateral deformation of
thebuilding can change the performance level of the structure and
isespecially important for tall, slender structures or for closely
spacedstructures that can be subjected to pounding when relative
dis-placements become large [3]. Moreover, increase in the total
defor-mation of the structure and in turn secondary P D
effectinuences the total stability of the structure. The lessons
learnedfrom post seismic observations of the past earthquakes such
as1985 Mexico City, 1994 Northridge, and 1995 Kobe
earthquakesprovided sufcient reason to believe that the SSPSI
effects shouldbe investigated with greater rigour and precision
(e.g. Mendozaand Romo [9]; Mizuno et al. [10]).
Pile foundations are usually employed to transmit
foundationloads through soil strata of low bearing capacity to
deeper soil orrock strata having a high bearing capacity and
stiffness. End bear-ing piles terminate in hard, relatively
impenetrable materials such
(b)
(c)
ture supported by oating pile foundation employing foundation
springs; (b) lateralral deformation of the xed-base structure.
Geotechnics 55 (2014) 172186 173as rock or very dense sand and
gravel, while oating piles obtain agreater part of their capacity
by skin friction or adhesion and aremostly employed in situations
where the bedrock is deep. Determi-nation of the pile foundation
seismic response is a complex processinvolving inertial interaction
between the structure and the pilefoundation, kinematic interaction
between piles and soils, andthe non-linear response of soils to
strong earthquake motions [7].However, simple methods such as
Winkler computational modelare often used in engineering practice
in which soilpile interac-tion is modelled using either linear or
non-linear springs. The reli-ability of these constitutive models
has been questioned by manydue to the simplifying assumptions
regularly used [11,12]. At rst,the applied earthquake motion in the
time history analysis is de-rived from the free eld motion ignoring
the presence of super-structure and pile elements. Secondly,
Winkler springs whichhave been developed initially to model single
pilesoil interaction,are not directly applicable to simulate pile
groups due to the over-lapping displacement elds of piles affecting
the individual pilestiffness [13]. The limitations of Winkler
methods and availabilityof advanced computational tools lead the
researcher to conductfully-nonlinear analysis to study the seismic
response of pile foun-dations. As mentioned by Chu [14], for
systems with strong nonlin-ear behaviour, coupled soilpilestructure
response analysis ishighly desirable which can explicitly express
the relationship be-tween the soil and the structural responses,
especially when thestiffness of the pile foundation signicantly
affect the overall soilresponse.
Although a number of works dealing with the SSPSI effects onthe
seismic response of structures are available in the literature,most
of them adopt simplied models (e.g. single degree of free-dom
system for superstructure or linear analysis) [1519]. Thepresent
research aims to study the effects of SSPSI on the seismicresponse
of the superstructure by employing the fully nonlinearmethod in
which main components of the interaction including
-
ously conducted shaking table tests (e.g. Chau et al. [22];
953 tonnes. The soil medium beneath the structure is a clayey
soil
andwith the shear wave velocity of 200 m/s and density of 1470
kg/m3.The horizontal distance of the soil lateral boundaries and
bedrockdepth has been selected to be 60 m and 30 m, respectively.
Thebuilding is resting on a footing which is 1 m thick and 15 m
wide.For the pile foundation case (Fig. 2c), a 4 4 reinforced
concretepile group with pile diameter and length of 1.2 m and 20
m,respectively, and equal spacing of four time the diameter (4d)
areconsidered. The piles are closed-end and have rigid
connectionwith the pile cap representing typical oating pile
foundations.Ishimura et al. [23]; Jakrapiyanun [24]; Pitilakis et
al. [25]; Mey-mand [26]) the superstructure is simplied as a single
degree offreedom oscillator in which the behaviour of the
soilstructure sys-tem may not be completely conforming to reality
and the effect ofhigher modes would not be captured. In the current
model tests,unlike the previous efforts, a multi-storey frame for
the superstruc-ture is adopted, representing the dynamic properties
of the proto-type structure such as natural frequency of the rst
and highermodes, number of stories, and density. Moreover, an
advancedlaminar soil container has been designed to simulate the
free eldsoil response by minimising boundary effects. Consequently,
thecurrent shaking table tests which simulate two common types
offoundations (shallow foundation and oating pile foundation) onthe
soft soil together with the xed-base model (excluding
thesoilstructure interaction) provide unique and valuable data
toinvestigate the inuence of the soilpilestructure interaction
onthe seismic response of buildings.
The experimental model tests have been carried out utilisingthe
shaking table facilities located at the structures laboratory ofthe
University of Technology Sydney (UTS). The size of the shakingtable
is 3 m 3 m, with maximum payload of 10 tonnes and over-turning
moment of 100 kN m. Furthermore, the shaking table canapply maximum
acceleration of 2.5 g with testing frequencyrange of 0.1100 Hz.
2.1. Prototype characteristics and scaling factors
A fteen-storey concrete moment resisting building frame withthe
total height of 45 m and width of 12 m consisting of threespans,
representing the conventional types of mid-rise momentresisting
buildings, is selected for this study as shown in Fig. 2.The
spacing between the frames into the page is 4 m. Natural fre-quency
of the prototype building is 0.384 Hz and its total mass issubsoil,
oating pile foundation, and superstructure, are
modelledsimultaneously. For this purpose, a three-dimensional
explicit -nite-difference program, FLAC3D [20], has been used to
numeri-cally model and examine the inuence of the
soilstructureinteraction on the seismic response of a 15-storey
moment resitingbuilding. Two types of foundations, shallow
foundations and oat-ing pile foundations, have been considered. The
proposed numeri-cal soilstructure model has been veried and
validated againstexperimental shaking table results.
2. Shaking table experimental tests
Model tests in geotechnical engineering offers the advantage
ofsimulating complex systems under controlled conditions
providingthe opportunity of better understanding the fundamental
mecha-nisms of these systems. Such tests are often used as
calibrationbenchmarks for numerical or analytical methods, or to
make quan-titative predictions of the prototype response [21]. In
most previ-
174 A.S. Hokmabadi et al. / ComputersIn order to achieve a
reasonable scale model, a dynamic simili-tude between the model and
the prototype should be applied asdescribed in the literature (e.g.
Harris and Sabnis [27]; Langhaar[28]; Meymand [26]). Dynamic
similitude governs a conditionwhere homologous parts of the model
and prototype experiencehomologous net forces. The scaling
relations for the variables con-tributing to the primary modes of
system response are presentedin Table 1.
Adopting an appropriate geometric scaling factor (k) is one
ofthe important steps in scale modelling on shaking table.
Althoughsmall scale models could save cost, the precision of the
resultscould be substantially reduced. Considering the mentioned
speci-cations of UTS shaking table, scaling factor of 1:30 provides
thelargest achievable scale model with rational scales, maximum
pay-load, and overturning moment meeting the facility
limitations.Thus, geometric scaling factor (k) of 1:30 is adopted
for experimen-tal shaking table tests on the scale model in this
study. Accordingto Table 1, apart from the geometric scaling which
should be im-posed to all the components, the required scaled
natural frequencyfor the structural model and the required scaled
shear wave veloc-ity and density of the soil mix should be 2.11 Hz,
36 m/s and1470 kg/m3, respectively. Moreover, the required scaled
naturalfrequency of the soil mix inside the soil container needs to
be10 Hz which is used as a benchmark to design the laminar
soilcontainer.
2.2. Model components of shaking table tests
The developed soilstructure model for shaking table tests
pos-sesses four main components including the model structure,
themodel pile foundations, the laminar soil container, and the
soilmix together with the imposed shaking events. Details and
charac-teristics of these components are explained below.
2.2.1. Model structureEmploying geometric scaling factor of
1:30, height, length, and
width of the structural model are determined to be 1.50 m, 0.40
m,and 0.40 m, respectively. In addition, according to the scaling
rela-tionship as shown in Table 1, the required natural frequency
of thestructural model is 2.11 Hz. In addition, the density of the
modeland prototype should be equal. Thus, the total mass of 106 kg
forthe model structure is obtained.
In order to simulate the prototype structure more accurately
onthe shaking table, the model structure has been designed
employ-ing SAP2000 [29] software considering the required
characteristicsof the model structure. The 3D numerical model
consists of fteenhorizontal steel plates as the oors and four
vertical steel plates asthe columns. Steel plate grade 250,
according to Australian stan-dards [30], with the minimum yield
stress of 280 MPa and the min-imum tensile strength of 410 MPa, has
been adopted in the design.The thickness of the steel plates has
been determined in the designprocess after several cycles of trial
and error in order to t the re-quired natural frequency and mass of
the model structure. Thenalised base plate is a 500 500 10 mm steel
plate while theoors consist of 400 400 5 mm plates and four500 40 2
mm steel plates are used for the columns. The con-nections between
the columns and oors are provided using stain-less steel metal
screws with 2.5 mm diameter and 15 mm length.After the numerical
modelling and design, the structural modelwas constructed in house.
The completed structural model isshown in Fig. 3.
2.2.2. Pile foundationSimilar to the model structure, the model
pile should be sub-
jected to the competing scale model criteria. In order to
achievea successful model pile design, the principal governing
factors of
Geotechnics 55 (2014) 172186pile response such as slenderness
ratio L/d, moment curvature rela-tionship, exural stiffness EI,
relative soil/pile stiffness, yieldingbehaviour/mechanism, and
natural frequency of vibration should
-
and Geotechnics 55 (2014) 172186 175A.S. Hokmabadi et al. /
Computersbe addressed [26]. By adopting the geometric similitude,
theoverall pile slenderness and relative contact surface area
wouldbe preserved in the model. This also guarantees that pile
grouprelative spacing and consequent group interaction would be
(a)
(b)Fig. 2. (a) Prototype xed-base structure; (b) prototype
structure supported by shallow f
Table 1Scaling relations in terms of geometric scaling factor
(k).
Mass density 1 Acceleration 1 Length kForce k3 Shear wave
velocity k1/2 Stress kStiffness k2 Time k1/2 Strain 1Modulus k
Frequency k-1/2 EI k5
Fig. 3. The completed model structure for shaking table
tests.replicated at the model scale. Thus, by considering the
geometricscaling factor (k) of 1:30 in this study, the model piles
should havea diameter of 40 mm with L/d ratio of 16.6.
The moment-curvature relation criterion represents pile
re-sponse to the lateral loading which is a function of the
exuralrigidity and yielding behaviour. Since in the present study
pilesare intended to respond in the elastic range (this assumption
isconrmed numerically), this criterion is achieved by scaling
theexural rigidity (EI) of the piles according to Table 1 (k5, k =
1/30)in addition to ensuring that the yielding point of the model
pileis equal to or greater than the scaled prototype. Furthermore,
byscaling the stiffness of the soil and pile consistently, the
relativesoil/pile stiffness parameter will be satised inevitably.
Therefore,the soilpile interaction should then be accurately
reproduced inthe model.
(c)oundation; (c) prototype structure supported by oating
(frictional) pile foundation.Previous researchers (e.g. Bao et al.
[31]; Chau et al. [22]; Taoet al. [32]) have used different types
of materials like aluminiumtubes, steel bars, and reinforced
concrete to build a model pile.Considering the selected scaling
factor in this study (k = 1/30)and in turn the required stiffness
and yielding stress for the modelpiles, a commercial Polyethylene
high pressure pipe with StandardDimension Ratio (SDR) of 7.4
according to the Australian Standard[33], is the selected candidate
which falls in the range of acceptablecriteria with 5% deviation
from the target value for EI. Moreover,Polyethylene pipes can
tolerate large deformation prior to theyielding point without any
brittle failure. Characteristics of themodel pile used in this
study are summarised in Table 2.
2.2.3. Soil mixA synthetic clay mixture was designed to provide
soil medium
for the shaking table testing. Previous researchers
(Meymand[26]; Turan et al. [34]; and Moss et al., [35]) reported
that areconstituted soil would not be able to satisfy the competing
scale
Table 2Characteristics of the model pile built from polyethylene
pressure pipe.
Outer diameter (mm) 40 Youngs modulus (MPa) 1.16E+3Wall
thickness (mm) 5.5 Density (kg/m3) 955Cross-sectional area
(mm2)5.78E+2 Poissons ratio 0.4
Moment of inertia (mm4) 8.33E+4 Flexural yield stress (MPa)
32
-
modelling criterion of shear wave velocity with enough
bearingcapacity for the foundation in shaking table tests while
syntheticclay mix can provide adequate undrained shear strength to
mobi-lise the required bearing capacity underneath the structural
modelmeeting the scale modelling criterion of the shear wave
velocity. Itshould be noted that, without providing enough bearing
capacityfor the structural model foundation, the underneath soil
may expe-rience failure or excessive settlements while testing
process is
176 A.S. Hokmabadi et al. / Computers andbeing undertaken. In
order to nd out the most appropriate mixfor the test program, three
different mixes (A, B, and C) were pro-duced and examined in the
UTS soils laboratory. The proportion ofdifferent mix components for
the three mixes are summarised inTable 3. Mix A, which is the
closest mix to what proposed by Mey-mand [26], has higher
percentages of kaolinite and bentinite, lowerpercentage of class F
y ash and lime, and the same percentage ofwater content comparing
to Mix B. Mix B and C have the same drycomponent percentages, but
the water content was increased by20% in Mix C in comparison to Mix
B in order to achieve better mix-ibility and workability for the
mix.
Each proposed mix was prepared three times to control
repeat-ability of the test and each time three cylindrical test
specimens ofsize D = 50 mm and h = 100 mmwere taken for the bender
elementtest which was performed to measure the shear wave velocity
ofthe soil over the curing age. The elapsed time from specimen
prep-aration to testing is termed curing age. To carry out bender
ele-ment tests, the soil specimens were placed between
benderelements as shown in Fig. 4a; and the shear wave velocity of
eachsoil specimen was obtained by measuring the time required for
thewave to travel between two bender elements using PC runningGDS
bender element control software. The adopted system has adata
acquisition speed of 2 MSamples/s, 16 bit resolution of
dataacquisition and the connection to the control box through USB
link.In this study, the propagated shear wave type has been sine
waveswith amplitude of 10 V and a period of 1 s. Fig. 4b shows the
sche-matic graphical signal processing to measure the shear wave
traveltime at the bender element test.
The extracted average shear wave velocities versus curing
agesfor the three different mixes over the period of two weeks
havebeen illustrated in Fig. 5. As shown in Fig. 5, the examined
soilmixes gain stiffness, and consequently shear wave velocity
in-creases with the curing age as expected (e.g. Wartman, [36];
Rie-mer, et al. [37]; Moss et al. [35]). However, only Mix C
producesthe required shear wave velocity of 36 m/s for the soil
model onthe second day of its curing age while the other two mixes
are un-able to produce such a low shear wave velocity as required.
Inaddition, in order to ensure that the undrained shear strength
ofthe proposed soil mix is adequate to satisfy the required
founda-tion bearing capacity underneath the structural model,
UnconnedCompression Tests were performed on three soil specimens
inaccordance with AS5101.4-2008 [38]. Eventually, desired soil
mixconsisting of 60% Q38 kaolinite clay, 20% Active-bond 23
Bentonite,20% class F y ash and lime, and 120% water (% of the dry
mix) hasbeen adopted for the shaking table tests in this study.
Table 4 sum-marised the soil mix properties at the second day of
its curing age.Accordingly, soil density on the second day was
determined to be
Table 3Proportion of different components for the examined
mixtures.
Mix Components Mix A (%) Mix B (%) Mix C (%)
Q38 kaolinite clay 67.5 60 60ActiveBond 23 bentonite 22.5 20
20Class F y ash + Lime 10 20 20
Watera 100 100 120
a % of the dry mix.1450 kg/m3 being almost equal to the
prototype soil density(1470 kg/m3) as required. Therefore, the
designed soil mixpossesses the required dynamic similitude
characteristics. Itshould be noted that the prototype soil acquires
the properties ofcemented soil that can be found in nature or
treated soil.
2.2.4. Shaking eventsThe input ground motions in this study are
represented by a set
of real earthquakes dened at the outcropping bedrock. Each
testmodel was subjected to two near eld shaking events
includingKobe, 1995, Northridge, 1994, two far eld earthquakes
includingEl Centro, 1940, and Hachinohe, 1968, and Sine Sweep test.
It iswell known that the intensity of shaking decreases as the
distanceincreases from the seismic fault where the earthquake
shaking isgenerated [39]. In addition, high frequency components
lose en-ergy more quickly than low frequency components while
travellingthrough the ground. As a result, near eld earthquakes
generatehigher ground peak acceleration and frequency component in
com-parison with the far led earthquakes. The characteristics of
thementioned earthquakes suggested by the International
Associationfor Structural Control and Monitoring for benchmark
seismic stud-ies [40] are summarised in Table 5. Referring to Table
2, althoughthe model earthquake magnitude remains the same as the
proto-type, time intervals of the original records should be
reduced bythe factor of 5.48 (k1/2, k = 1/30) whichmeans that the
scaled earth-quakes contain higher frequencies and shorter
durations. Thescaled acceleration records of the four adopted
earthquakestogether with the relevant frequency content obtained
from FastFourier Transform are illustrated in Fig. 6ad. In
addition, exponen-tial sine sweep wave with amplitude of 0.05 g,
exponentialincrease rate of 0.5 Hz, and frequency range of 150 Hz
has beenapplied to the test models in order to identify the dynamic
charac-teristics of the systems. Fig. 6e displays a schematic view
of theapplied exponential sine sweep waves.
2.2.5. Laminar soil containerA soil container is required to
hold the soil in place during shak-
ing table tests and provide connement. The ideal soil
containershould simulate the free eld soil response by minimising
bound-ary effects. Since the seismic behaviour of the soil
container affectsthe interaction between the soil and the
structure, the performanceof the soil container is of key
importance for conducting seismicsoilstructure interaction model
tests successfully [25]. Well-de-signed laminar soil container as
gured out by many researchers(e.g. Chau et al. [22]; Taylor, [42])
has advantage among othertypes in which lateral motion of the
entire depth of the laminar soilcontainer follows a sinusoidal
shape representing authentic condi-tions of the free eld ground
motion. Therefore, in order to performrigorous and reliable
experimental shaking table tests, a laminarsoil container has been
employed in this study.
Considering the adopted geometric scaling factor (1:30)
andallowing a further 10 mm on each side for construction
purposessimilar to Prasad et al. [43], the nal length, width, and
depth ofthe laminar soil container are selected to be 2.10 m, 1.30
m, and1.10 m, respectively. Same as the model structure, the
laminar soilcontainer is initially designed employing a 3D
numerical model.The key parameter in the design of the soil
container is the naturalfrequency of the container itself which
should be close to the nat-ural frequency of the soil deposit
inside the container (approxi-mately 10 Hz for this study) in order
to minimise the interactionbetween the soil and the container
during the shaking table test.
The employed laminar soil container consists of a
rectangularlaminar box made up of aluminium rectangular hollow
section
Geotechnics 55 (2014) 172186frames separated by rubber layers.
The aluminium frames providelateral connement of the soil, while
the rubber layers allow thecontainer to deform in a shear beam
manner. The soil container
-
eas
A.S. Hokmabadi et al. / Computers and Geotechnics 55 (2014)
172186 177(a)Fig. 4. (a) Bender element test setup; (b) schematic
graphical signal processing to m
city
(m/s
)
Bender Element Test(Amplitude= 10 V, period= 1 sec.)was xed and
secured on the shaking table using eight M38 boltspassing through
the provided holes. Then, the internal surface ofthe soil container
was covered and sealed with two layers of blackplastic sheeting.
According to Gohl and Finn [44] and Valsangkaret al. [45], 25 mm
thick absorbing layers of Polystyrene foamsheets have been
installed at the end walls of the soil containerto simulate viscous
boundaries in the free eld condition and min-imise the reection of
the outward propagating waves back intothe model. In addition, a
layer of well graded gravel was glued tothe bottom of the soil
container to create a rough interface be-tween the soil and the
base during the test. This layer provides fric-tion between the
timber base plate (as a bedrock) and the in-situ
involves a logarithmic frequency sweep holding a specied
accel-
Shea
r Wav
e Ve
lo
Curing Age (Days)
Mix A (67.5% kaolinite, 22.5% Bentonite, 10% fly ash and lime,
and 100% water)
Mix B (60% kaolinite, 20% Bentonite, 20% fly ash and lime, and
100% water)
Mix C (60% kaolinite, 20% Bentonite, 20% fly ash and lime, and
120% water)
Fig. 5. Average shear wave velocity for three mixes obtained
from Bender ElementTest.
Table 4Properties of the soil mix on the second day of
curing.
Soil properties Value
Mass density (kg/m3) 1450Shear wave velocity (m/s) 36Maximum
shear modulus, Gmax (kPa) 1776Undrained shear strength, Su (kPa)
3.1Plasticity Index, PI (%) 42
Table 5Utilised earthquake base motions.
Earthquake Country Year PGA (g) Mw (R)
Northridge USA 1994 0.843 6.7Kobe Japan 1995 0.833 6.8El Centro
USA 1940 0.349 6.9Hachinohe Japan 1968 0.229 7.5
a Obtained from PEER [41].eration constant at the base of the
structure. For the current SineSweep test, by increasing the
frequency of the shaking table from0.1 Hz to 50 Hz, the rst
resonance between the shaking tableand the structural model
frequencies showed the fundamental nat-soil mix ensuring negligible
relative slip between the soil and thebottom surface of the
container and to justify the xed-baseassumption in the computer
model. Fig. 7 shows the laminar soilcontainer adopted in this
study.
2.3. Shaking table tests program
The shaking table tests have been carried out under three
con-ditions: xed-base condition, with shallow foundation, and
withoating pile foundation. In the rst case, a xed base model
(con-structed structure directly xed on top of the shaking table)
hasbeen tested in order to ensure the structural model possesses
thetarget natural frequency and determine the damping ratio of
thestructural model. To achieve the above, constructed
structuralmodel was xed and secured on the shaking table as shown
inFig. 1. Displacement transducers (levels 3, 5, 7, 11, 13, and
15)and accelerometers (at levels 3, 5, 7, 9, 11, 13, and 15) were
in-stalled on the structure in order to monitor the dynamic
responseof the structure and to primarily measure the structural
lateral dis-placements. The recorded accelerations can be used to
check theconsistency and accuracy of the obtained displacements
througha double integration in time domain. In addition, by
recordingthe accelerometers which are installed on two edges of the
topoor, any possible torsion of the structure during the seismic
exci-tations could be monitored.
Initially, Sine Sweep test was performed on the structural
mod-el to determine the natural frequency of the model. Sine Sweep
test
(b)ure the shear wave travel time between the sender and
receiver bender elements.ural frequency of the model. The test was
repeated three times toensure the determined natural frequency is
adequately accurate.The resulting natural frequency of the
constructed structural mod-el obtained from Sine Sweep test results
was 2.19 Hz which is in avery good agreement with the desired
natural frequency of thestructural model (2.11 Hz). Therefore, the
constructed structuralmodel, with the natural frequency of 2.19 Hz
and the total mass
Duration (s) Type Hypocentral distancea (km)
30.0 Near eld 9.256.0 Near eld 7.456.5 Far eld 15.6936.0 Far eld
14.1
-
and-0.20.00.20.40.60.81.0
lera
tion
(g)
Scaled Kobe EarthquakeScaling factor = 1/5.48
-1.0-0.8-0.6-0.4-0.20.00.20.40.60.81.0
0 1 2 3 4 5 6
Acc
eler
atio
n (g
)Scaled Northridge EarthquakeScaling factor = 1/5.48
Time (sec)
178 A.S. Hokmabadi et al. / Computersof 104 kg, possesses the
required characteristics to meet the dy-namic similitude criteria.
The estimated value of the structuraldamping ratio of the
constructed structural model was determinedto be equal to 1.1%,
obtained from the free vibration lateral dis-placement records of
the structural model using the Taylor seriesexpansion [46].
After ensuring the adequacy of the structural model
character-istics, shaking table tests were performed by applying
scaled earth-quake acceleration records of Northridge, 1994 (Fig.
6a), Kobe,1995 (Fig. 6b), El Centro, 1940 (Fig. 6c), and Hachinohe,
1968(Fig. 6d) to the xed base structural model and the results in
termsof maximum lateral deections are presented in Section 4.
The second case of the shaking table tests was to study the
ef-fects of the soilstructure interaction under the shallow
foundationcase. After securing the laminar soil container on the
shaking table,
-1.0-0.8-0.6-0.4
0 2 4 6 8 10
Acc
e
-0.4-0.3-0.2-0.10.00.10.20.30.4
0 2 4 6 8 10 12
Acc
eler
atio
n (g
)
Scaled El Centro EarthquakeScaling factor = 1/5.48
-0.4-0.3-0.2-0.10.00.10.20.30.4
0 2 4 6 8 10
Acc
eler
atio
n (g
)
Time (sec)
Scaled Hachinohe EarthquakeScaling factor = 1/5.48
Time (sec)
Time (sec)
Fig. 6. Adopted shaking events in this study: (a) scaled
Northridge earthquake; (b) scaled(e) exponential sine sweep
wave.4.05.06.07.08.0
(g/H
z10
-4)
0.0
1.0
2.0
3.0
4.0
5.0
6.0
Am
plitu
de (g
/Hz
10-4
)
Frequency (Hz)0 10 20 30 40 50
Geotechnics 55 (2014) 1721862 cubic meters of the designed soil
mix (Mix C: 60% Q38 kaoliniteclay, 20% Active-bond 23 Bentonite,
20% class F y ash and lime,and water, 120% of the dry mix) was
produced and placed intothe laminar soil container. As explained in
Section 2.2.3, the de-sired soil mix acquires the required
stiffness and consequentlythe shear wave velocity after two days of
curing. As a result, thetime frame for the testing process was very
tight and time sensi-tive. Therefore, soil mixing and placement
needed to be carriedout in one day in order to produce a homogenous
soil mix, andafter two days of curing, the nal tests had to be
performed.
During the soil mixing process, ten cylindrical soil samples ofD
= 50 mm and h = 100 mm were taken from the soil mix for qual-ity
control of the mix. The entire mixing process and lling thelaminar
soil container were completed in one day. Then, the soilmix inside
the container was left to be cured for two days while
0.01.02.03.0
Am
plitu
de
Frequency (Hz)
0.0
0.5
1.0
1.5
2.0
0 10 20 30 40 50Am
plitu
de (g
/Hz
10-4
)
Frequency (Hz)
0.0
0.1
0.2
0.3
0.4
0.5
0.6
Am
plitu
de (g
/Hz
10-4
)
Frequency (Hz)0 10 20 30 40 50
0 10 20 30 40 50
Kobe earthquake; (c) scaled El Centro earthquake; (d) scaled
Hachinohe earthquake;
-
same age as for the shallow foundation case in order to make
theresults comparable, without any variation of the dynamic
soilproperties.
As mentioned in Section 2.2.2, commercial Polyethylene pres-sure
pipe was employed to build the model piles. The length ofthe model
oating piles is 66 cm, leaving 34 cm distance betweenthe piles toe
and the base. Wooden tips were tted to the modelpiles to provide a
closed end condition. The model piles were dri-ven into the soil
through a 150 mm tall wooden template to ensurelocation and
verticality. Moreover, employing template during theinstallation
process helps to achieve full connection between thepiles and the
surrounding soil without generating any gap due tothe installation
process. The template was constructed with specialcut outs to
accommodate a few millimetres of extra room for pileswith external
strain gages aiming to prevent any possible damage
A.S. Hokmabadi et al. / Computers and Geotechnics 55 (2014)
172186 179the surface of the soil container was covered and sealed.
On thesecond day, the structural model was lifted up and placed on
thedesignated location, without observing any excessive
settlement
Fig. 7. Final setup of the shaking table tests for the structure
with oating(frictional) pile foundation.or failure underneath the
base plate as predicted. In addition tothe instrumentation used on
the structure, vertical displacementtransducers were placed on the
level of base plate of the structure(simulating the foundation) to
determine the vertical displace-ments of the structure during the
testing process as shown inFig. 8a. Similar shaking events
including Sine Sweep test and fourscaled earthquake records (Fig.
6) have been applied to the system.The natural frequency of the
soilstructure model from the per-formed Sine Sweep test was
measured to be 1.60 Hz.
The last case of the shaking table tests was to consider the
oat-ing pile foundation and investigate the inuence of
soilpilestruc-ture interaction on the seismic response of the
superstructure bycomparing this case with the previously mentioned
xed-baseand shallow foundation cases. Since the properties of the
designedsoil mix is time dependent, this stage should be carried
out at the
(a)Fig. 8. Shaking table tests setup and connections for: (a)
mto the strain gages during installation.After installation of the
model piles the template was removed
and the steel plate (simulating the foundation) with
prefabricatedholes was tted over the group. Sixteen M12 bolts were
used toprovide xed connection between the piles head and the
steelplate as shown in Fig. 8b. The required nuts were xed to the
piletop with strong glue and steel rings before the test and the
strengthand capability of this connection technique was examined
success-fully. Then, the model structure was suspended from the
overheadcrane and connected to the steel plate from the pre-located
con-nections (Fig. 9b) similar to the xed-base and shallow
foundationcases.
Consequently, all the components of the system including
thecontainer, soil, piles, and superstructure were installed. The
samearrangement of displacement transducers and accelerometers
hasbeen used on the structure and the steel plate (simulating the
foun-dation). In addition, twelve strain gauges were installed on
the pilesand four 3D accelerometers were embedded inside the soil
body.Since the inuence of the soilstructure interaction on the
responseof the superstructure is the main objective of this
research, just thedata obtained from the instrumentation on the
structure itself, notincluding the soil and piles sensors, are
reported in this paper. Sim-ilar shaking events including Sine
Sweep test and four scaled earth-quake records have been applied to
the oating pile foundationsystem. The natural frequency of the
soilpilestructure modelfrom the performed Sine Sweep test was
measured to be 1.8 Hz.The results of the conducted shaking table
tests under the inuenceof four scaled earthquake acceleration
records in terms of the max-imum lateral deections of various
stories of the structure are pre-sented and discussed in Section 4.
The nal setup of the testsincluding the displacement transducers
and accelerometers at dif-ferent levels of the structural model for
the oating pile foundationsystem on the shaking table are shown in
Fig. 7.(b)odel shallow foundation; (b) model pile foundation.
-
and180 A.S. Hokmabadi et al. / Computers3. Development of 3d
numerical model
In order to conduct a fully coupled analysis of the entire
soilpilestructure system, a three dimensional numerical
soilstruc-ture model has been developed which treats the behaviours
ofthe soil and the structurewith equal rigour. Adopting
directmethodof analysis, the numerical model can perform fully
nonlinear time
(a)
(c)
(b)
Fig. 9. Numerical grid and model components in FLAC3D for: (a)
xed-basestructure; (b) structure supported by shallow foundation;
(c) structure supportedby oating (frictional) pile
foundation.history dynamic analysis to simulate realistic dynamic
behaviourof soil and structure under seismic excitations. According
to Chu[14], time domain analysis in necessary to compute the
nonlineardynamic responses of soilpilestructure systems as the
frequencydomain analysis can deal only with linear responses
without con-sidering any nonlinearities. In this study,
three-dimensional expli-cit nite difference based program FLAC3D,
Fast LagrangianAnalysis of Continua, version 4.0 [20] has been
employed followingthe other researchers experience (e.g. Comodromos
and Papado-poulou [47]; Rayhani and El Naggar, [48]). This program
can simu-late behaviour of different types of structures and
materials byelements which can be adjusted to t the geometry of the
model.Each element behaves according to a prescribed constitutive
modelin response to the applied forces or boundary restraints. The
pro-gram offers a wide range of capabilities to solve complex
problemsin mechanics such as inelastic analysis including plastic
momentand simulation of hinges for structural systems.
Three cases including: xed-base condition, the structure
sup-ported by the shallow foundation, and the structure supportedby
the oating pile foundation have been modelled separatelyand the
results are compared. The dimensions of the numericalmodels were
chosen to be similar to the experimental tests. Thereason for
choosing the soil deposit thickness of 30 m for the pro-totype is
that most amplication occurs within the rst 30 m of thesoil prole,
which is in agreement with most modern seismic codescalculating
local site effects based on the properties of the top 30 mof the
soil prole [48].
Experience gained from the parametric studies helped to nal-ise
the adopted mesh size and the maximum unbalanced force atthe grid
points to optimise the accuracy and the computationspeed,
simultaneously. For the oating pile foundation model, thegenerated
mesh comprised 10,868 zones and 16,356 grid points.Fast computation
facilities at University of Technology Sydneywere employed to
conduct the time-history analysis, and the com-putation took
approximately 20 h for a single analysis. The numer-ical grid and
model components in FLAC3D are shown in Fig. 9.
Adjusting the boundary conditions for the static analysis,
inwhich the system is under the gravity loads only, the bottom
faceof the model is xed in all directions, while the side
boundaries arexed in the horizontal directions. During the dynamic
time-historyanalysis, in order to avoid reection of outward
propagating wavesback into the model, quiet (viscous) boundaries
comprising inde-pendent dashpots in the normal and shear directions
are placedat the lateral boundaries of the soil medium. The lateral
boundariesof the main grid are coupled to the free-eld grids by
viscous dash-pots of quiet boundaries at the sides of the model, as
shown inFig. 2, to simulate the free eld motion which would exist
in theabsence of the structure and pile foundation. Rigid boundary
con-ditions is adopted to simulate the bedrock in the seismic
soilstructure interaction analysis as suggested by other
researchers(e.g. Dutta and Roy [49]; Spyrakos et al. [50]), and the
earthquakeinput motions are applied at the bedrock assuming
horizontallypolarised shear waves propagating vertically. Lu et al.
[51] empha-sised on the inuence of the gravity load on contact
state of thesoilstructure interface mentioning that signicant error
in theanalysis may occur if gravity is not taken into account in
the dy-namic analysis.
Solid elements are used to model the soil deposits, and
Mohr-Coulomb failure criterion is adopted. In addition, the
built-in tan-gent modulus function developed by Hardin and Drnevich
[52] isadopted to implement hysteretic damping of the soil
representingthe variation of the shear modulus reduction factor and
dampingratio with cyclic shear strain of the soil. This model is
dened as
Geotechnics 55 (2014) 172186follows:
Ms 1=1 c=cref 2
-
whereMs is the secant modulus (G/Gmax), c is the cyclic shear
strain,and cref is Hardin/Drnevich constant. In this study, cref =
0.234 whichgives the best numerical t to the backbone curves
suggested bySun et al. [53] for the ne grained soils (cohesive
soils) are adopted,and the comparison has been presented in Fig.
10.
There are two main analytical procedures for dynamic analysisof
soilstructure systems under seismic loads, equivalent-linearand
fully nonlinear method. The equivalent-linear method (e.g.Seed and
Idriss [54]) cannot capture directly any nonlinearity ef-fects
during the solution process and uses linear properties foreach
element that remain constant throughout the history of shak-ing,
and are estimated from the mean level of dynamic motion.Therefore,
the above mentioned strain-dependent modulus anddamping functions
for the soil are only taken into account in anaverage sense, in
order to approximate some effects of nonlinear-
loads were applied on the free end of the cantilever pile.
Therecorded deection from the FLAC3D model showed less than
2%difference from the existing analytical predictions, conrming
theaccuracy of the model.
Because of the different characteristics of the soil and the
super-structure/piles, sliding and separation may occur at the
soilstruc-ture interfaces [58]. Two sets of interface elements are
modelled inthis study. For the shallow foundation case, the
interface elementsare placed between the foundation and the soil
surface (Fig. 11a).However, for the pile foundation case, the
interface elements wereattached to the outer perimeter and bottom
of the piles as shownin Fig. 11b. It should be noted that in the
pile foundation case thereis no interface or attachment between the
foundation and the sur-face soil as some gap in the shaking table
tests was considered toavoid any pile-raft behaviour. Therefore,
there is no direct stress
A.S. Hokmabadi et al. / Computers and Geotechnics 55 (2014)
172186 181ity. As a result, during the low amplitude shakings in
the excitationhistory, soil elements will be modelled overdamped
and too soft,and during the strong shaking soil elements will be
modelledunderdamped and too stiff. In contrast, employing fully
nonlinearmethod, nonlinearity in the stress-strain law is followed
directlyby each element and the dependence of damping and
apparentmodulus on strain level is automatically modelled. Byrne et
al.[55] and Beaty [56] provided some overviews of the above
men-tioned methods and concluded that the most appropriate
methodfor a dynamic analysis of soilstructure system is a fully
nonlinearmethod. In addition, Lu et al. [57] studies illustrated
the potentialfor further reliance on the computer simulation in the
assessmentof the nonlinear seismic ground response using the
nonlinear dy-namic analysis. Consequently, fully nonlinear method
for dynamicanalysis of soilstructure systems has been employed in
this study.
The common soil tests such as bender element and density
testwere conducted on the obtained samples during the mixing
pro-cess on the second day of curing age as described in Section
2.3.The results are in good conformity with the initial laboratory
testresults and adopted in the numerical model as summarised
inTable 4.
The pile elements and superstructure are modelled with
solidelements considering elastic-perfectly plastic behaviour
withyielding criteria for the elements to control the possibly of
inelasticbehaviour in both superstructure and piles. The
formulationadopted in this study to simulate the inelastic
behaviour of pileand structural elements assumes that the material
behaves linearelastically until reaching the dened yield stress and
after reachingthis yield stress, the material deforms without
inducing additionalresistance. The yield stress for the material
used in the modelstructure is 280 MPa. As a calibration, a FLAC3D
analysis was rstconducted on a cantilever pile while the pile was
xed at one endinto the ground without the surrounding soil and
different lateral
Cyclic Shear Strain (%)
G /
G m
ax
Sun et al. (1988) (Utilised test: Resonant column
test)----Adopted in this study for ref =0.234
Backbone curve for cohesive soils
(a)
Fig. 10. Adopted tting curve for ne grained soil in this study
(after Sun et al. [57]): (a)ratio versus cyclic shear
strain.transfer between the foundation slab and the subsoil in the
pilefoundation case. According to Fig. 11c, the normal and shear
forcesthat describe the interface response in the elastic range are
deter-mined at calculation time (t + Dt) using the following
relations[20]:
FtDtn knunA rnA 3
FtDtsi Ftsi ksDut1=2Dtsi A rsiA 4where FtDtn and FtDtsi are the
normal and shear force vector attime (t + Dt), respectively. un is
the absolute normal penetrationof the interface node into the
target face.Dusi is the incremental rel-ative shear displacement
vector. rn and rsi are the additional nor-mal and shear stresses
added due to interface stress initialisation,respectively. kn and
ks are the normal and shear stiffnesses, respec-tively, and A is
the representative area associated with the interfacenode. The
lateral and axial stiffnesses of the interface elements areset to
ten times the equivalent stiffness of the neighbouring zone,based
on the recommended relationship by Rayhani and El Naggar[48] and
Itasca Consulting Group [20] for the isotropic soil medium,as
follows:
ks kn 10K 43GDzmin
5
where K and G are bulk and shear modulus of neighbouring
zone,respectively, and Dzmin is the smallest width of an adjoining
zonein the normal direction. This is a simplifying assumption that
hasbeen used to ensure that the interface stiffness has minimal
inu-ence on system compliance by avoiding the intrusion of
adjacentzones (a numerical effect) and preventing excessive
computationtime [48]. In addition, shear strength of the interfaces
was denedby MohrCoulomb failure criterion and the tensile strength
of the
Cyclic Shear Strain (%)
Dam
ping
Rat
io
Sun et al. (1988) (Utilised test: Resonant column
test)----Adopted in this study for ref =0.234
Backbone curve for cohesive soils
(b)
Relations between G/Gmax versus cyclic shear strain; (b)
relations between damping
-
piles and the supporting soil in the pile foundation case and
upliftin the shallow foundation case.
foun
andFinally, fully nonlinear time-history analysis is conducted
underthe inuence of the mentioned shaking events (Fig. 6), and
resultsin terms of the maximum inelastic lateral deections
determinedfor three mentioned cases are presented and discussed in
the nextsection.
4. Results and discussioninterfaces are set to zero in order to
allow gapping between the
A
B
kn
ks S SsTs
D
S = sliderTs= tensile strengthSs= shear strengthD = dilation
(assumed zero)Ks= shear stiffnessKn= normal stiffness
(c)
Fig. 11. Interface elements adopted in this study: (a)
interfaces between the shallowand surrounding soil; (c) components
of the interface constitutive model.
182 A.S. Hokmabadi et al. / ComputersResults of the conducted
shaking table tests and the 3D numer-ical predictions for the
maximum lateral displacements of thexed-base, shallow foundations,
and oating pile foundations arepresented in Figs. 12 and 13,
respectively. To determine the lateraldeections, the movement of
the shaking table has been subtractedfrom the storey movements.
Therefore, all the records are relativeto the base movements. It
should be noted that the presented dataare based on the lateral
deformation of each storey when the max-imumdeection at the top
level occurred. This approach givesmorereasonable pattern of the
structural deformation in comparisonwith the approach
thatmaximumabsolute storey deformation irre-spective of occurrence
time are recorded [59]. Fig. 14 illustrates asample of time-history
deformation records used to obtain the lat-eral deformations
reported in Figs. 12 and 13.
Comparing the results of the conducted shaking table tests(Fig.
12) and the 3D numerical predictions (Fig. 13) for the maxi-mum
lateral displacements of the xed-base, shallow foundations,and
oating pile foundations, it is observed that the trend and
thevalues of the 3D numerical predictions are in a good
agreementand consistent with the experimental shaking table test
results.Therefore, the developed 3D numerical model can replicate
thebehaviour of the soilpilestructure system with acceptable
accu-racy and is a rational and appropriate tool for further
studies of thesoilpilestructure interaction effects. The observed
disparity be-tween FLAC3D predictions and experimental measurements
inthe lower levels of the shallow foundation and oating pile
casescan be due to the nature of the numerical method, adopting
elas-tic-perfectly plastic MohrCoulomb model for the soil,
assumingideal rigid connection between the foundation and the pile
caps,and unavoidable experimental uncertainties. Moreover, as
anexample, Fig. 15 presents the time-history acceleration records
atthe top of the 15-storey model structure for the xed-base,
shallowfoundations, and oating pile foundations under the inuence
of1940 El Centro earthquake. Comparison of the measurementsand the
predictions indicates that the horizontal acceleration time curves
obtained from the 3D numerical analysis and the lab-oratory
experiments are in a reasonable agreement.
In order to draw a general conclusion to be used by
practicingengineers, the average values of the 3D numerical
predictions
(b)
(a)dation and the soil; (b) interfaces at the outer perimeter
and tip of the oating piles
Geotechnics 55 (2014) 172186and experimental values of
deformations for each case were deter-mined and compared in Fig.
16. In comparison to the xed basestructure, the maximum lateral
deection of the structure sup-ported by oating pile foundations
increases by 34%, and 27%based on the experimental measurements and
the 3D numericalpredictions, respectively. Moreover, the maximum
lateral deec-tion of the structure supported by the shallow
foundation is in-creased by 55% based on the experimental values
and 59% basedon the 3D numerical predictions in comparison to the
results ob-tained from the xed base structure. The natural
frequency ofthe system reduces due to the soilstructure interaction
(2.19 Hzand 2.11 Hz for the xed-base condition, 1.8 Hz and 1.88 Hz
forthe oating pile foundation, and 1.60 Hz and 1.64 Hz for the
shal-low foundation case based on the experimental results and
numer-ical predictions, respectively). Therefore, such decreases in
thenatural frequency (increases in the natural period) considerably
al-ter the response of the building frames under the seismic
excita-tion. This is due to the fact that the natural period of the
systemlies in the long period region of the response spectrum
curve,and the displacement response tends to increase. The pile
founda-tions reduce the lateral displacements in comparison to the
shal-low foundation case since the presence of stiff pile elements
inthe soft soil increases the equivalent stiffness of the ground
andinuences the dynamic properties of the whole system such asthe
natural frequency and damping.
Rocking component plays an important role in the lateral
defor-mation of the superstructure. According to Kramer [3],
relativelateral structural displacements under the inuence of the
soilstructure interaction consist of rocking and distortion
components.
-
and6789
101112131415
rey
Num
ber
Scaled Northridge Earthquake (1994)Near field Eearthquake,
Scaled factor = 1/ 30Mw = 6.7 (R), PGA = 0.843 (g)
(a)
A.S. Hokmabadi et al. / ComputersThe maximum vertical
displacement and the rocking angles of thefoundation in the instant
of the maximum deformation at the topof the structure are
summarised in Table 6. Accordingly, for theoating pile foundation
cases, approximately 27% of the maximumlateral deections were due
to the rocking component, while 73%took place due to the distortion
component. These values for theshallow foundation cases are 37% and
63%, respectively. For exam-ple, under the inuence of 1940 El
Centro earthquake, the maxi-mum lateral deection at the top of the
xed base model was
0123456789
101112131415
0 10 20 30 40
Stor
ey N
um
ber
Maximum Lateral Deflection (mm)
Scaled El Centro Earthquake (1940)Far field EarthquakeScaled
factor = 1/ 30
Mw = 6.9 (R), PGA = 0.349 (g)
Fixed base Experimental Results
Shallow foundation Experimental Results
Floating piles Experimental Results
012345
403020100
Sto
Maximum Lateral Deflection (mm)
Fixed base Experimental Results
Shallow foundation Experimental Results
Floating piles Experimental Results
(c)
Fig. 12. Recorded maximum lateral deection of the structure from
the shaking table tunder the inuence of: (a) Northridge earthquake;
(b) Kobe earthquake; (c) El Centro ea
0123456789
101112131415
403020100
Stor
ey N
umbe
r
Maximum Lateral Deflection (mm)
Scaled Northridge Earthquake (1994)Near field Eearthquake,
Scaled factor = 1/ 30Mw = 6.7 (R), PGA = 0.843 (g)
Fixed base 3D Numerical Results
Shallow foundation 3D Numerical Results
Floating piles 3D Numerical Results
0123456789
101112131415
Stor
ey N
umbe
r
Maximum Lateral Deflection (mm)
Scaled El Centro Earthquake (1940)Far field EarthquakeScaled
factor = 1/ 30
Mw = 6.9 (R), PGA = 0.349 (g)
Fixed base 3D Numerical Results
Shallow foundation 3D Numerical Results
Floating piles 3D numerical Results
0 10 20 30 40
(a)
(c)
Fig. 13. 3D numerical predictions of the maximum lateral
deformation under the inuenHachinohe earthquake.6789
101112131415
rey
Num
ber
Scaled Kobe Earthquake (1995)Near field EarthquakeScaled factor
= 1/ 30
Mw = 6.8 (R), PGA = 0.833 (g)
(b)
Geotechnics 55 (2014) 172186 183measured to be 13.63 mm due to
distortion component, whilethe maximum lateral deection at the top
of the structure sup-ported by oating pile foundation was 20.40 mm
with 7.62 mmof that value being due to rocking component and 12.78
mm tookplace due to distortion component. In the oating pile
foundationcases, rocking occurs due to the axial deformation of the
pileelements together with the deformation of the surrounding
andbeneath soil elements. The area replacement ratio of the pile
groupis 8% in this study and as a result piles attract signicant
axial
012345
403020100
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Maximum Lateral Deflection (mm)
Fixed base Experimental Results
Shallow foundation Experimental Results
Floating piles Experimental Results
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0 10 20 30 40
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ey N
umbe
rMaximum Lateral Deflection (mm)
Scaled Hachinohe Earthquake (1968)Far field EarthquakeScaled
factor = 1/ 30
Mw = 7.2 (R), PGA = 0.229 (g)
Fixed base Experimental Results
Shallow foundation Experimental Results
Floating piles Experimental Results
(d)
ests for the xed base, shallow foundation, and end-bearing pile
foundation casesrthquake; (d) Hachinohe earthquake.
0123456789
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ey N
umbe
r
Maximum Lateral Deflection (mm)
Scaled Hachinohe Earthquake (1968)Far field EarthquakeScaled
factor = 1/ 30
Mw = 7.2 (R), PGA = 0.229 (g)
Fixed base 3D Numerical Results
Shallow foundation 3D Numerical Results
Floating piles 3D Numerical Results
0123456789
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0 10 20 30 40
Stor
ey N
umbe
r
Maximum Lateral Deflection (mm)
Scaled Kobe Earthquake (1995)Near field EarthquakeScaled factor
= 1/ 30
Mw = 6.8 (R), PGA = 0.833 (g)
Fixed base 3D Numerical Results
Shallow foundation 3D Numerical Results
Floating pile 3D Numerical Results
0 10 20 30 40
(b)
(d)
ce of: (a) Northridge earthquake; (b) Kobe earthquake; (c) El
Centro earthquake; (d)
-
and184 A.S. Hokmabadi et al. / Computersforces. However, clearly
the rocking of the structure in the shallowfoundation case, without
pile elements, is muchmore than the casewith pile foundations
resulting in further amplication of the lat-eral deection.
Fig. 17 compares the 3D numerical predictions of the
structuraldemand in terms of the base shear for the xed-base
structure andthe structure supported by two types of foundations,
shallow foun-dations and oating pile foundations. In general, the
ratio of thebase shear for cases including soilstructure
interaction to that ofxed-base is less than one, demonstrating the
effect of soilstruc-ture interaction in reducing the base shear of
the structure. Thebase shear of the structure supported by the
oating pile founda-tion and shallow foundation is on average 78%
and 70% of the xed
Fig. 14. Sample experimental time-history displacement results
for the xed basemodel under the inuence of El Centro
earthquake.
-1.5
-1.0
-0.5
0.0
0.5
1.0
1.5
Acce
lara
tion
(g)
Time (sec)
Numerical Results Exp. Results
Fixed base
-3.0
-2.0
-1.0
0.0
1.0
2.0
3.0
0.0 0.5 1.0 1.5 2.0 2.5
Acce
lara
tion
(g)
Time (sec)
Numerical Results Exp. Results
Shallow foundation
-3.0
-2.0
-1.0
0.0
1.0
2.0
3.0
Acce
lara
tion
(g)
Time (sec)
Numerical Results Exp. Results
Floating piles
0.0 0.5 1.0 1.5 2.0 2.5
0.0 0.5 1.0 1.5 2.0 2.5
(a)
(b)
(c)
Fig. 15. Time-history acceleration records at top of the
15-storey model structureunder the inuence of El Centro earthquake
for: (a) xed-base structure; (b)structure supported by shallow
foundation; (c) structure supported by oating(frictional) pile
foundation.0123456789
101112131415
3020100
Stor
ey N
umbe
r
Maximum Lateral Deflection (mm)
Fixed base Numerical ResultsFixed base Exp. ResultsShallow
foundation Numerical ResultsShallow foundation Exp. ResultsFloating
piles Numerical ResultsFloating piles Exp. Results
Fig. 16. Average values of maximum lateral displacements base on
shaking tableexperimental measurements versus 3D numerical
predictions.
Table 6Maximum vertical displacement and rocking angle of the
base plate.
Scaledearthquakeaccelerationrecord
Maximum vertical displacement (rocking angle) of the
baseplate
Fixedbase
Shallow foundation Floating pile foundation
Experimental Numerical Experimental Numerical
Northridge 0 2.54 mm(0.58)
2.68 mm(0.61)
1.9 mm(0.43)
1.98 mm(0.45)
Kobe 0 1.32 mm(0.30)
1.45 mm(0.33)
0.43 mm(0.1)
0.52 mm(0.12)
El Centro 0 1.98 mm(0.45)
2.06 mm(0.47)
1.27 mm(0.29)
1.34 mm(0.31)
Hachinohe 0 1.47 mm 1.58 mm 0.93 mm 1.02 mm
Geotechnics 55 (2014) 172186base structure excluding
soilstructure interaction. Therefore,presence of oating pile
foundation increases the base shear andin turn demands of the
superstructure, in comparison with the casesupported by shallow
foundation.
Since the adopted model is a Multi Degree of Freedom
(MDOF)structure, inter-storey drifts can be estimated and employed
toinvestigate the performance levels of the building structures
underthe inuence of dynamic soilstructure interaction. The
corre-sponding inter-storey drifts of the average values of 3D
numericalmodel (Fig. 18) have been calculated using the following
equationbased on the Australian standard [60]:
Drift di1 di=h 6
where di+1 is deection at (i + 1) level, di is deection at (i)
level, andh is the storey height. In the performance-based seismic
design, theseismic performance (performance level) is described by
consider-ing the maximum allowable damage state (damage
performance)for an identied seismic hazard (hazard level).
Performance levelsdescribe the state of structures after being
subjected to a certainhazard level, and based on FEMA273/274 [61]
are classied as: fullyoperational, operational, life safe, near
collapse, or collapse. Overalllateral deection, ductility demand,
and inter-storey drifts are themost commonly used damage
parameters. The above mentionedve qualitative levels are related to
the corresponding quantitativemaximum inter-storey drifts (as a
damage parameter) of:
-
and0
100
200
300
400
500
600
700
Northridge Kobe El Centro Hachinohe
Bas
e Sh
ear (
N)
Earthquake
Fixed base
Shallow foundation
Floating pile foundation
Fig. 17. Base shear of the model structure obtained from 3D
numerical analysis for:xed-base structure; structure supported by
shallow foundation; and structuresupported by oating (frictional)
pile foundation.
56789
101112131415
tore
y N
umbe
rA.S. Hokmabadi et al. / Computersseismic soilstructure
interaction tends to increase the inter-storeydrifts of the
superstructure. The inter-storey drifts of the structuresupported
by the oating pile foundation are more than thexed-base conditions
excluding soilstructure interaction. How-ever, the structure
supported by oating pile foundation experi-ences less inter-storey
drifts in comparison to the structuresupported by the shallow
foundation. For example, the maximumrecorded inter-storey drift of
the xed base structure is measuredto be 1.48%, while the
corresponding value for the oating pilefoundation and shallow
foundation cases are 1.83% and 2.25%,respectively. In other words,
effects of soilpilestructure interac-tion (pile foundation) and
soilstructure interaction (shallowfoundation) induces 23% and 52%
increase in the recorded inter-sto-rey drifts, respectively. As a
result, the soilstructure interactionmay affect the performance
level of the structure and shift theperformance level of the
structure from life safe zone to nearcollapse or even collapse
levels.
Therefore, although SSPSI reduces the base shear of the
struc-ture leading to the reduction in the structural distortion in
compar-ison with xed base structure, considering the effect of
SSPSIincreases the overall lateral deformation and consequently
inter-storey drifts of the structure mainly due to the rocking
component.Moreover, in the seismic response of pile groups, rocking
andtranslation components are coupled and the response of the
under-neath soils to strong seismic shaking is strongly
nonlinear.
Practicing engineers can adopt this veried numerical model-ling
procedure in the design considering the effect of SSPSI withrespect
to the interface elements, boundary conditions, and hyster-etic
damping of the soil representing the variation of the shear
01234
0 0.5 1 1.5 2 2.5
S
Inter-storey Drift (%)
Fixed baseShallow foundationFloating pile foundationLife safe
limit (1.5%)
Fig. 18. Average 3D numerical inter-storey drifts for: xed-base
structure; struc-ture supported by shallow foundation; structure
supported by oating (frictional)pile foundation.References
[1] Park R. Reinforced concrete structures. New York: John Wiley
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5. Conclusions
The pile foundations are usually considered to be rigid enoughto
guarantee the restraint against rocking motions and conven-tional
xed-base models are used to predict the seismic responseof these
systems although the behaviour of system can be signi-cantly
affected by seismic soilpilestructure interaction. A seriesof
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By comparing predicted and observed results, it can be
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thesimulation of the soilpilestructure interaction under
strongground motions. In addition, based on the shaking table
resultsand 3D numerical investigations, it is observed that the
lateraldeections of structures siting on the oating pile
foundationsamplied in comparison to the xed base model (34% based
onthe experimental measurements and 27% based on the 3D numer-ical
predictions). This amplication for the structure siting on
theshallow foundation is more severe (55% based on the
experimentalmeasurements and 59% based on the 3D numerical
predictions).Therefore, the oating pile foundations increase the
lateraldisplacements of the superstructure in comparison with
thexed-base assumption, and reduce the lateral displacements
incomparison to the shallow foundation case due to the
rockingcomponents.
Consequently, seismic soilpilestructure interaction affectsthe
performance level of structures sitting on the soft soil
byincreasing the inter-storey drifts, which may shift the
performancelevel of the structure from life safe to near collapse
or even collapselevels. Therefore, ignoring the real deformability
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foundations, maylead to erroneous evaluations of structural
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Assessment of soilpilestructure interaction influencing seismic
response of mid-rise buildings sitting on floating pile
foundations1 Introduction2 Shaking table experimental tests2.1
Prototype characteristics and scaling factors2.2 Model components
of shaking table tests2.2.1 Model structure2.2.2 Pile
foundation2.2.3 Soil mix2.2.4 Shaking events2.2.5 Laminar soil
container
2.3 Shaking table tests program
3 Development of 3d numerical model4 Results and discussion5
ConclusionsReferences