2014 - assessment of soil-pile-structure interaction influencing

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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

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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

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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

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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

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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

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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

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Floating piles 3D Numerical Results

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Fig. 13. 3D numerical predictions of the maximum lateral deformation under the inuenHachinohe earthquake.6789

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(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

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(d)

ests for the xed base, shallow foundation, and end-bearing pile foundation casesrthquake; (d) Hachinohe earthquake.

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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.

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(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

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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.

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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 & Sons; 1975.[2] Priestley MJN. Performance based siesmic design. In: 12th World conference

on earthquake engineering (12WCEE). Auckland, New Zealand: New ZealandNational Society for Earthquake Engineering; 2000.

[3] Kramer SL. Geotechnical earthquake engineering. Prentice Hall; 1996.[4] Wolf JP. Dynamic soilstructure interaction. Englewood Cliffs, New

Jersey: Prentice-Hall; 1985.[5] Guin J, Banerjee PK. Coupled soilpilestructure interaction analysis undermodulus and damping ratio of the soil with the cyclic shear strain.Another advantage of the current numerical modelling techniqueis performing the SSPSI analysis in a fully-coupled manner inwhich main components of the interaction including subsoil, pilefoundation, and superstructure are modelled simultaneously with-out resorting to independent calculations of site or superstructureresponse, or application of the pile group interaction factors.

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 shaking table experimental tests have been conducted and alaminar soil container has been designed to simulate the free eldsoil response by minimising the boundary effects, and the super-structure has been simulated as a multi-storey frame representingmost of the dynamic properties of the prototype structure.Employing FLAC3D, a fully nonlinear three dimensional numericalmodel has been adopted to perform time-history analyses imple-menting hysteretic damping of the soil to represent the variationof the shear modulus and damping ratio of the soil with the cyclicshear strain. Free eld boundary conditions have been assigned tothe numerical model, and appropriate interface elements, capableof modelling sliding and separation between the pile and soil ele-ments, have been considered.

By comparing predicted and observed results, it can be con-cluded that the employed numerical model is appropriate for 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 of the soilstruc-ture system, in both cases with shallow or pile foundations, maylead to erroneous evaluations of structural displacements and in-ter-storey drifts affecting the predicted damage level of structuraland non-structural elements as well as the lateral load carryingmechanism of the soilstructure system.

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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

2014 - assessment of soil-pile-structure interaction influencing

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