Journal of Rock Mechanics and Geotechnical Engineering xxx (2015) 1—8

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Journal of Rock Mechanics and Geotechnical Engineering

Full length article

Lateral response of pile foundations in liquefiable soils

Asskar Janalizadeh, Ali Zahmatkesh*

Babol University of Technology, Babol, Iran

ARTICLE INFO ABSTRACT

Article history: Liquefaction has been a main cause of damage to civil engineering structures in seismically active areas.

Received 6 IMarch 2015 The effects of damage of liquefaction on deep foundations are very destructive. Seismic behavior of pile

Received in revised form foundations is widely discussed by many researchers for safer and more economic design purposes. This

5 May 2()15 paper presents a pseudo-static method for analysis of piles in liquefiable soil under seismic loads. A free-

Accepted 6 May 2015 ~ ; . . , ■ j- , . .

Available online xxx field site response analysis using three-dimensional (3D) numerical modeling was performed to determine kinematic loads from lateral ground displacements and inertial loads from vibration of the su-

Ke words■ perstructure. The effects of various parameters, such as soil layering, kinematic and inertial forces,

Pile foundations boundary condition of pile head and ground slope, on pile response were studied. By comparing the

Lateral spreading numerical results with the centrifuge test results, it can be concluded that the use of the p-y curves with

Liquefaction various degradation factors in liquefiable sand gives reasonable results.

Pseudo-static method © 2015 Institute of Rock and Soil Mechanics, Chinese Academy of Sciences. Production and hosting by

Elsevier B.V. All rights reserved.

1. Introduction

The liquefaction is one of the challenging issues in geotechnical engineering and it damages structures and facilities during earthquakes. This phenomenon was reported as the main cause of damage to pile foundations during the major earthquakes (Kramer, 1996). In many earthquakes around the world, extensive damage to piles of bridges and other structures due to liquefaction and lateral spreading has been observed (Boulanger et al., 2003). Failures were observed in both sloping and level grounds and were often accompanied with settlement and tilting of the superstructure (Adhikari and Bhattacharya, 2008). The loss of soil strength and stiffness due to excess pore pressure in liquefiable soil may develop large bending moments and shear forces in the piles. If the residual strength of the liquefiable soil is less than the static shear stresses caused by a sloping site or a free surface such as a river bank, significant lateral spreading or downslope displacements may occur. The moving soil can exert damaging pressures against the piles, leading to failure (Finn and Fujita, 2002). The performance of structures above piles depends widely on the behavior of pile foundations under earthquake loading. During past earthquakes, because of inadequacy of the pile to sustain large shear forces and bending moments, the extensive damage in liquefiable soil has

* Corresponding author. Tel.: +98 9158342367. E-mail address: A.zahmatkesh@stu.nit.ac.ir (A. Zahmatkesh). Peer review under responsibility of Institute of Rock and Soil Mechanics, Chinese Academy of Sciences.

1674-7755 © 2015 Institute of Rock and Soil Mechanics, Chinese Academy of Sciences. Production and hosting by Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.jrmge.2015.05.001

been caused due to both lateral ground movement and inertial loads transmitted to piles. Under earthquake loading, the performance of piles in liquefied ground is a complex problem due to the effects of progressive buildup of pore water pressures and decrease of stiffness in the saturated soil (Liyanapathirana and Poulos, 2005). These effects involve inertial interaction between structure and pile foundation, significant changes in stiffness and strength of soils due to increase of pore water pressures, large lateral loads on piles, kinematic interaction between piles and soils, nonlinear response of soils to strong earthquake motions, kinematic loads from lateral ground displacements, and inertial loads from vibration of the superstructure (Bradley et al., 2009; Gao et al., 2011).

Various approaches including shaking table and centrifuge tests and also various numerical methods have been developed for the dynamic response analysis of single pile and pile group. The soil-pile—structure interaction has been investigated using the centrifuge test (e.g. Finn and Gohl, 1987; Chang and Kutter, 1989; Liu and Dobry, 1995; Hushmand et al., 1998; Wilson, 1998; Abdoun and Dobry, 2002; Su and Li, 2006) and shaking table test (e.g. Mizuno and Liba, 1982; Yao et al., 2004; Tamura and Tokimatsu, 2005; Han et al., 2007; Gao et al., 2011; Haeri et al., 2012). The obvious advantage of shaking table and centrifuge tests is the ability to obtain detailed measurements of response in a series of tests designed to physically evaluate the importance of varying earthquake characteristics (e.g. level of shaking, frequency content), soil profile characteristics, and/or pile—superstructure characteristics (Wilson, 1998). However, some limitations exist in centrifuge tests, for example, sand grains in centrifuge tests correspond to bigger gravel particles in prototype (Towhata, 2008).

To simulate the piles in liquefiable soil layers, Finn and Fujita (2002), Klar et al. (2004), Oka et al. (2004), Uzuoka et al. (2007),

Cheng and Jeremic (2009), Comodromos et al. (2009), and Rahmani and Pak (2012) used three-dimensional (3D) finite element method. The complexity and time-consuming nature of 3D nonlinear finite element method for dynamic analysis makes it useful only for very large practical projects or research and not feasible for engineering practice. However, it is possible to obtain reasonable solutions for nonlinear response of pile foundations with fewer computations by relaxing some of the boundary conditions in full 3D analysis (Finn and Fujita, 2002).

The simple approach for modeling and simulation of the piles in liquefied grounds is based on scaling of p-y springs, where p and y are the soil resistance per unit length of the pile and pile lateral displacement, respectively. Because of complexity and time-consuming of two-dimensional (2D) and 3D numerical modeling, most of the designers and researchers prefer to use one-dimensional (1D) Winkler method based on finite element or finite difference method for the seismic analysis of pile foundations. In pseudo-static method, a static analysis is carried out to obtain the maximum response (deflection, shear force and bending moment) developed in the pile due to seismic loading. In Winkler models, p-y curves are used to define the behavior of the nonlinear spring at any depth. These p-y curves can be obtained from the results of model tests or field (Liyanapathirana and Poulos, 2005). The Winkler assumption is that the soil—pile interaction resistance at any depth is related to the pile shaft displacement at that depth only, independent of the interaction resistances above and below (Wilson, 1998).

This pseudo-static method has been suggested early by Miura et al. (1989), Miura and O'Rourke (1991), Liu and Dobry (1995), JRA (1996), AIJ (1998) and recently by Liyanapathirana and Poulos (2005) and Elahi et al. (2010). This method for pile seismic analysis sometimes underestimates, and sometimes overestimates shears, moments and deflection of the piles. However, in many practical conditions, the results of pseudo-static method are reasonable (Tabesh, 1997).

In this paper, a pseudo-static method has been applied for estimation of the response of pile during dynamic loading. First, definition of the geometry and the soil modeling parameters are presented. Next, the numerical model is vertified by means of the centrifuge test. And then the effects of various parameters, including soil layering, kinematic and inertial forces, boundary condition of pile head and ground slope, on the behaviors of piles are studied.

2. Numerical analysis

All simulations were conducted using the open-source computational platform OpenSees (McKenna and Fenves, 2007). This platform allows for developing applications to simulate the performance of structural and geotechnical systems subjected to static and seismic loadings. In this paper, the steps for calculation of pile response are summarized as follows:

(1) A free-field site response analysis was performed during the dynamic loading using 3D numerical modeling. From this analysis, time history of ground surface acceleration and the maximum ground displacement along the length of the pile can be calculated.

(2) The dynamic analysis was performed using the time history of ground surface acceleration calculated in Step 1 for pile length above ground and superstructure with a fixed base. From this analysis, the maximum acceleration of superstructure can be calculated.

(3) In 1D Winkler analysis, the maximum soil displacement profile calculated in Step 1 and the maximum acceleration of superstructure in Step 2 were applied to the pile as shown in Fig. 1.

First, the time history of the ground surface acceleration and the maximum ground displacement at each depth were obtained from the free-field site response analysis. Taboada and Dobry (1993) and Gonzalez et al. (2002) showed that the pore pressure time histories recorded at the same elevation are identical, indicating the 1D behavior of the model. In free-field analysis, the model consists of a single column of 3D brick elements. The soil layers were modeled using cubic 8-noded elements with u-p formulation in which each node has four degrees of freedom: three for soil skeleton displacements and one for pore water pressure. To consider the effect of the laminar box in the numerical simulation, nodes at the same depths were constrained to have equal displacements in the horizontal and vertical directions. The pore water pressures were allowed to freely develop for all nodes except those at the surface and above the water table. The bottom boundary was assumed fixed in all directions.

The material model plays a key role in the numerical simulation of the dynamic behavior of liquefiable soils. The model in Dafalias and Manzari (2004), a critical state two-surface plasticity model, was used in this paper. This model requires fifteen material parameters and two state parameters to describe the behavior of sands and has been amply tested for simulating the behavior of granular soils subjected to monotonic and cyclic loadings (Jeremic et al., 2008; Taiebat et al., 2010; Rahmani and Pak, 2012). The key advantages of the model are that (1) it is relatively simple and (2) it has a unique calibration of input parameters. Thus, a single set of parameters independent of void ratio and effective consolidation stress level was used for the Dafalias and Manzari's material model. Table 1 presents the material parameters for Nevada sand. The additional parameters used for free-field analysis are presented in Table 2. It can be noted that at the onset of liquefaction, change of soil particles creates additional pathways for water. This leads to a significant increase in permeability coefficient (Rahmani et al., 2012). In this study, the permeability coefficient value was increased 10 times the initial value (suggested by Rahmani et al. (2012)).

For free-field analysis, the simulations were carried out in two loading stages. At the first stage, the soil skeleton and pore water weight were applied to soil elements. The values of stress and strain in this stage were used as initial values for the next stage of loading. At the second stage, dynamic analysis was performed by application of an input motion to the model base.

analysis.

Table 1

Material parameters for Nevada sand (Rahmani and Pak, 2012).

Elasticity Critical state Yield surface parameter, m Plastic modulus Dilatancy Dilatancy-fabric

Go n M c 1c e0 x h0 ch nb A0 nd zmax cz

150 0.05 1.14 0.78 0.027 0.83 0.45 0.02 9.7 1.02 2.56 0.81 1.05 5 800

Table 2

Additional parameters for pseudo-static and free-field analysis (Wilson, 1998; Rahmani and Pak, 2012).

Dr (%) Permeability Saturated unit Dry unit weight Friction Void

(m s-1) weight (kN m-3) (kN m-3) angle (°) ratio, e

35 7.05 x 10-5 19.11 14.9 30 0.743

80 3.7 x 10-5 19.91 16.2 39.5 0.594

In the second step, after free-field analysis, pile length above ground and superstructure were modeled. The pile was modeled as beam column elements with elastic section properties. The superstructure was modeled at the pile head. Generally, the superstructures above the pile foundations are multi-degree of freedom systems, but in the design of pile foundations, the superstructure was modeled as a single mass at the pile head to simplify the analysis. In this step, the base model was also fixed.

The model considered for the third step (pseudo-static analysis) is shown in Fig. 1. There are two versions of the pseudo-static BNWF method. These two methods are different in the way in which the lateral load on pile due to ground movement (kinematic load) is considered. The first BNWF requires free-field soil movements as an input. The free-field soil displacements are imposed on the free ends of the p-y springs due to lateral dilatation layers. In the second BNWF, the limit pressures over the depth of the lateral spreading soil were applied and the p-y springs were removed in this interval. In case of limit pressure, interaction of the soil and pile was not modeled, because the analysis is simple and can be done by hand calculation. Inertia forces from superstructure are represented as static forces applied simultaneously with lateral spreading demands. When limit pressures are applied directly to the pile nodes, bending moments and cap displacements depend on acceleration records and are greatly overpredicted for small to medium motions. This can be explained by the fact that the lateral spreading displacements were not large enough to mobilize limit pressures and actual pressures are smaller than limit pressures. However, for large motions, the pile cap displacements were considerably under-predicted (Brandenberg et al., 2007). In this paper, the free-field soil movement was used as an input. The cap mass, multiplied by the maximum acceleration of the superstructure obtained from Step 2 as a lateral force (F), was applied at the pile head. The material properties of p-y curves for non-liquefied sand were computed based on API (1987). These curves are defined by the following equation:

P = Putanh^y

where Pu is the ultimate bearing capacity at depth z, K is the initial modulus of subgrade reaction, and y is the lateral deflection. The initial tangent stiffness (Kjn), based on Eq. (1), is obtained as Kin = Kz. The p-y curves were modeled as zero-length elements with PySimple1 materials. Under dynamic loading, the piles are influenced by kinematic loads from lateral ground displacements and inertial loads from vibration of the superstructure. Fig. 1 shows an idealized schematic of the BNWF model for kinematic (F) and inertial (Ds) loads. The loss of bearing capacity for piles in loose

sandy soils (particularly vulnerable to liquefaction and lateral spreading during dynamic loading) also occurred. Therefore, the excessive forces imposed on the foundation due to ground displacement led to shearing of the piles and subsequent structural collapse of the superstructure. There are three methods for considering the influence of liquefaction on p-y curves in sand. In the first case, the lateral resistance of liquefiable sand is assumed to be zero. This method can lead to large design responses and high construction costs which may be very conservative (Rollins et al., 2005). Another approach is to treat liquefiable sand as undrained soft clay and use the p-y curves for soft clay. The undrained shear strength used in this case is obtained as a ratio of undrained shear strength to initial effective overburden stress using field data, and it is a function of overburden stress and relative density (Rollins et al., 2005; Varun, 2010). The third method for the simulation of pile response in liquefiable soils is the use of reduction factors, called p-multipliers. The p-y curves in liquefiable sand are multiplied by a factor usually between 0.01 and 0.3 to decrease the strength of sand due to liquefaction (Rollins et al., 2005; Brandenberg et al., 2007; Varun, 2010). In this paper, the third method (p-multipliers) was used. The free-field soil displacement and lateral force in head pile were imposed incrementally using a static load control integrator.

3. Validation of the proposed method

The performance and ability of the proposed approach to simulate pile behavior in liquefiable soil have been demonstrated by comparison between the numerical simulations and centrifuge tests performed by Wilson (1998). In these tests, the soil profile consisted of two horizontal layers of saturated uniformly graded Nevada sand (see Fig. 2). On prototype scale, the lower layer was 11.4 m thick with relative density of 80% (dense) and the upper was 9.1 m thick with relative density of 35% (loose). The single pile was a

Fig. 2. Layout of the model for centrifuge test by Wilson (1998).

steel pipe of 0.67 m in diameter, 18.8 m in length, and 19 mm in wall thickness. The pile tip was about 3.8 m above the container base. The superstructure mass (Ms) was 49.1 Mg. Properties of Nevada sand with Dr = 35% and 80% are presented in Table 2. The Kobe acceleration record (Fig. 3) was used as an input to shake model.

It is important to specify stiffness and lateral resistance of the p-y curves in liquefiable soil as explained in Section 2. Three cases were considered to evaluate the effects of variations in stiffness and lateral resistance of the p-y curves on the testing results. These cases include: (1) use of the p-y curves without the influence of liquefaction; (2) use of the p-y curves with a constant degradation factor in liquefied sand; and (3) use of the p-y curves with various degradation factors in liquefiable sand. In case (2), different degradation factors can be considered between 0.05 and 0.5 to reduce the strength of liquefiable soil. In this case, because of the small strength of liquefiable soil, especially that of surface ground, a degradation factor of 0.1 was considered. In case (3), variation in degradation factor with depth was taken from a small value (top of liquefiable layer) to 1.0 (bottom of liquefiable layer). An exponential decay function from bottom to top of liquefied layer is proposed as

R = R0eH ln(-R°) (2)

where R is a degradation factor, as a function of distance from the top of the layer (z); H is the liquefiable layer thickness; and R0 is the degradation factor at the top of the layer. Both the ultimate resistance and initial stiffness of the p-y curves in the liquefiable layer were taken to be R% of their unreduced magnitudes.

In free-field analysis under the Kobe acceleration record scaled to 0.04g, 0.12g and 0.22g, because of the ground level, lateral displacement (Ds) was less than about 5 cm and the effects of kinematic loads on seismic response are small. Therefore, in these tests, it is reasonable to only consider the inertial loads for calculation of the pile response. The superstructure displacement was calculated using acceleration of the superstructure obtained from the tests carried out by Wilson (1998). The comparison between the observed and simulated results of superstructure displacement is shown in Fig. 4. In cases (2) and (3), by considering the centrifuge test results, the best predictions were obtained with degradation factor of 0.1. These results clearly illustrate that the performance and accuracy of BNWF mainly depend on the accuracy in selection of the correct curves. As seen in Fig. 4, the deflections observed during the centrifuge test are much larger than those simulated without the influence of liquefaction. This is due to the loss of bearing capacity for the piles in loose and medium sandy soils during dynamic loading. It can be noted that use of constant degradation factors at various depths gives unreasonable results. The pile head displacements have a good agreement with the available experimental data in case (3).

-0.3 I-

Fig. 3. Acceleration record of Kobe earthquake scaled to 0.22g used in the centrifuge test by Wilson (1998).

0.1 0.15 0.2 0.25 0.3 0.35

Superstructure acceleration (g)

Fig. 4. Comparison of superstructure displacement in various cases with the centrifuge tests by Wilson (1998).

Fig. 5 compares the maximum bending moment recorded from the centrifuge tests with that obtained from present analysis. In this figure, variation of R is similar to case (3) and Fig. 5 shows that the results obtained from the numerical analysis agree with the values recorded during the centrifuge test. It can be said that increasing the value of R from a small value at the top of the layer to 1.0 at the bottom of the layer produces a reasonable response for the piles. Therefore, this method was used for subsequent analysis.

4. Results and discussion

In this section, the behavior of pile for various conditions is discussed. The soil profile is the same as the one used in the centrifuge test performed by Wilson (1998). The profile has two layers: the upper layer is liquefiable (relative density of 35%) while

Bending moment (kN m)

Fig. 5. Comparison of bending moment profiles with the centrifuge tests by Wilson (1998).

the lower layer (relative density of 80%) is not. Three different ground slopes of 1%, 2%, and 4% were considered. The water table was supposed to be 1 m, 2 m, 3 m, and 4 m below the ground. This means that the thicknesses of non-liquefiable surface crust are 1 m, 2 m, 3 m, and 4 m. The input motion for the model was a 20-cycles sinusoidal wave with a frequency of 2 Hz and the peak acceleration of 0.5g. It should be noted that the intensity, frequency content (e.g. predominant period) and the duration of strong shaking are important characteristics of an earthquake (Rathje et al., 1998). These characteristics affected the response of piles. The pile responses largely depend on the shaking amplitude. Increase in the shaking amplitude (because of more reduction of restraint on liquefied soil) resulted in a decrease in the restraint against bending under the lateral load, and the maximum bending moment in piles significantly increased (Gao et al., 2011). The frequency also had a significant effect on pile response.

The free-field analysis showed that displacement of level ground is significantly less than that of the sloping ground. Fig. 6 compares the displacement of sloping ground when the thickness of the non-liquefiable surface layer is 1 m and 2 m with ground slope of 2%. This figure highlights the importance of non-liquefiable surface layer as a key parameter on ground displacement. When liquefaction occurs in sloping ground, because of displacements developing up to several meters, large lateral forces may act on the pile. This phenomenon is commonly called lateral spreading (Klar et al., 2004). In lateral spreading, the driving forces only exceed the resisting forces during those portions of the earthquake that impose net inertial forces in the downslope direction. Each cycle of net inertial forces in the downslope direction causes the driving forces to exceed the resisting forces along the slip surface, resulting in progressively and incrementally lateral movement (Day, 2002). Based on the results of free-field analysis, the displacement profile can be matched with constant displacement across the upper soil layer, a linear variation across the liquefiable and non-liquefiable layers.

The variations in bending moment along piles in different ground slopes for various conditions are presented in Figs. 7—10. The results show that in sloping grounds, when a non-liquefiable soil layer overlies a liquefiable soil layer and piles are embedded in the non-liquefiable soil layer, the lateral spreading has more

Displacement (m)

Fig. 6. Displacement profile at two different thicknesses of non-liquefiable surface layer when the ground slope is 2%.

Bending moment (kN m)

Fig. 7. Variations in bending moment along the pile in different sloping grounds (free head, without superstructure).

influences on the damage of piles. An increase in bending moment occurred as the ground slope increased. After liquefaction, if the static shear stress caused by sloping ground is more than the shear strength of liquefiable soil, the non-liquefiable surface crust overlying a liquefied soil layer can slide with a considerable amount of displacement. In this condition (lateral spreading), the non-liquefiable surface layer was carried along with the underlying fully liquefiable soil and a large lateral force was imposed on the embedded piles (Ashour and Ardalan, 2011 ). This force due to the lateral movement of the non-liquefiable layer has the potential to induce large bending moments in the piles leading to failure.

The boundary condition of the pile head has an important effect on the pile responses (moments, shear and deflections). In layered

Bending moment (kN m)

Fig. 8. Variations in bending moment along the pile in different sloping grounds (fixed head, without superstructure).

Bending moment (kN m)

Fig. 9. Variations in bending moment along the pile in different sloping grounds (free head, with superstructure).

soil deposits, a liquefiable soil layer is overlain by a non-liquefiable layer; when the pile head is free, the maximum bending moment develops at a depth corresponding to the interface of liquefiable and non-liquefiable layers (see Figs. 7 and 9). When the pile head is fixed, the maximum bending moment develops at two locations: (1) at the pile head and (2) at the interface of the two layers (see Figs. 8 and 10).

The dynamic effects during earthquake on deep foundations are critically important. These effects include the kinematic forces applied by the soil to the pile foundation and the inertial forces of the superstructure due to earthquake. The combination of cyclic horizontal kinematic loads due to ground displacements and in-ertial loads from the superstructure determines the critical load for piles during the shaking phase (Cubrinovski et al., 2009). The

Bending moment (kN m)

Fig. 10. Variations in bending moment along the pile in different sloping grounds (fixed head, with superstructure).

kinematic loads depend on the magnitude of ground deformations and the stiffness of the soil during a given loading cycle. Due to the influence of liquefaction on free-field soil response and soil—pile-structure interaction, the magnitudes of inertial and kinematic loads change (Han et al., 2007). When the acceleration of ground surface or superstructure mass is large or lateral dynamic stiffness of pile group due to pile and/or soil stiffness is small, the inertial effects may become important (Elahi et al., 2010). Figs. 7 and 8 illustrate that in the absence of a superstructure, the maximum bending moment near the pile head decreases significantly, and a major difference is observed at that location between piles with and without superstructure, but the values are approximately unchanged at large depths (Figs. 9 and 10). In other words, the same kinematic forces were developed in the piles with and without superstructure. At greater depths, where the inertial effects from the superstructure are less significant, pile damage can occur due to the lateral loads arising from lateral spreading (the excessive ground movement). Both inertial and kinematic loads can cause damages at the pile head.

The inertia effects of the superstructure before development of the pore water pressures and liquefaction are important and the kinematic effects can often be neglected. Ishihara (1997) stated that inertial forces are the cause of development of the maximum bending moment near the pile head. These forces are predominant before liquefaction and are confirmed by the results of Figs. 7—10. If the shaking continues after liquefaction, the inertial forces are combined with kinematic forces on the pile foundation arising from large cyclic ground deformations. It can be said that the kinematic loading in the areas of lateral spreading with relatively strong non-liquefiable surface layers is important. Then, the pile failure near the bottom of the liquefiable layer is likely influenced by kinematic loads from the liquefiable layer, while failure near the pile head is likely influenced by inertial loads from the superstructure and kinematic loads from the non-liquefiable layer. It should be noted that the bending moments are underpredicted when the structural inertia forces are neglected.

The pile response is sensitive to the thickness of the non-liquefiable surface layer (H) and thickness of the liquefiable layer (L). Fig. 11 shows variations in pile head displacement when the ratio of thickness (H/L) for different sloping grounds is increased. The pile head displacement increases to a peak value and then

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----t—- 4%

0 I—_i_i_—I_i_i_i_—I_i_i_i_i_L__i_i_—I_i_i_i_—I_i_i_i_—

0 0.1 0.2 0.3 0.4 0.5 0.6

Fig. 11. Variations in pile head displacement against ratio of thickness (H/L).

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decreases with subsequent increase in thickness ratio. As the thickness of the non-liquefiable surface layer increases, horizontal kinematic loads due to ground displacements also increase and result in higher pile head displacement. However, when the non-liquefiable surface layer is thick, because of increasing effective stress on the liquefiable layer, ground displacements are decreased, resulting in smaller values of pile head displacement. In addition, when the non-liquefiable surface layer is thick, due to increase of static shear stress in ground with great slopes, ground displacement and pile head displacement also increase.

5. Conclusions

This paper presents a method for analysis of piles in liquefiable soil under seismic loads. Three steps for calculation of pile response were: (1) free-field response analysis using 3D numerical modeling for calculation of ground surface acceleration and the maximum ground displacement along the length of the pile, (2) the dynamic analysis of pile length above ground and superstructure for calculation of the maximum acceleration of superstructure, and (3) 1D Winkler analysis for calculation of pile response. All simulations in three steps were conducted using the open-source computational platform OpenSees. After verification of the numerical model using a centrifuge test, analyses were carried out for various conditions.

By comparing the numerical results with the centrifuge test, it can be concluded that using the p-y curves with various degradation factors in liquefiable sand produces reasonable results. In addition, the non-liquefiable surface layer especially in sloping ground plays a key role in ground displacement. When the pile head is free, the maximum bending moment develops at a depth corresponding to the interface of liquefiable and non-liquefiable layers. When the pile head is fixed, there are two locations for developing the maximum bending moment: (1) at the pile head and (2) at the interface of the two layers. Moreover, at greater depths, where inertial effects from the superstructure are less significant, pile damage may occur due to lateral loads arising from lateral spreading. Both inertial and kinematic loads can cause damages at the pile head.

Conflict of interest

The authors have declared no conflict of interest.

References

Abdoun T, Dobry R. Evaluation of pile foundation response to lateral spreading. Soil Dynamics and Earthquake Engineering 2002;22(9—12):1051—8.

Adhikari S, Bhattacharya S. Dynamic instability of pile-supported structures in liquefiable soils during earthquakes. Shock and Vibration 2008;15(6):665—85.

American Petroleum Institute (API). Recommended practice for planning, designing and constructing fixed offshore platforms (RP 2A-WSD). Washington, DC, USA: API; 1987.

Architectural Institutive of Japan (AIJ). Recommendation for design of building foundations. 1998.

Ashour M, Ardalan H. Piles in fully liquefied soils with lateral spread. Computers and Geotechnics 2011;38(6):821—33.

Boulanger RW, Kutter BL, Brandenberg SJ, Singh P, Chang D. Pile foundations in liquefied and laterally spreading ground during earthquakes: centrifuge experiments and analyses. Davis, USA: Center for Geotechnical Modeling, University of California; 2003. Report No. UCD/CGM-03/01.

Bradley BA, Cubrinovski M, Dhakal RP, MacRae GA. Intensity measures for the seismic response of pile foundations. Soil Dynamics and Earthquake Engineering 2009;29(6):1046—58.

Brandenberg J, Boulanger RW, Kutter BL, Chang D. Static pushover analyses of pile groups in liquefied and laterally spreading ground in centrifuge tests. Journal of Geotechnical and Geoenvironmental Engineering 2007;133(9):1055—66.

Chang GS, Kutter BL. Centrifugal modeling of soil-pile-structure interaction. In: Engineering Geology and Geotechnical Engineering: Proceedings of the 25th Symposium, Reno, Nevada. Rotterdam, the Netherlands: A.A. Balkema; 1989. p. 327—36.

Cheng Z, Jeremic B. Numerical modeling and simulation of pile in liquefiable soil. Soil Dynamics and Earthquake Engineering 2009;29(11—12):1405—16.

Comodromos EM, Papadopoulou MC, Rentzepris IK. Pile foundation analysis and design using experimental data and 3-D numerical analysis. Computers and Geotechnics 2009;36(5):819—36.

Cubrinovski M, Ishihara K, Poulos H. Pseudo-static analysis of piles subjected to lateral spreading. Bulletin of the New Zealand Society for Earthquake Engineering 2009;42(1):28—38.

Dafalias YF, Manzari MT. Simple plasticity sand model accounting for fabric change effects. Journal of Engineering Mechanics 2004;130(6):622—34.

Day RW. Geotechnical earthquake engineering handbook. New York, USA: McGraw-Hill; 2002.

Elahi H, Moradi M, Poulos HG, Ghalandarzadeh A. Pseudostatic approach for seismic analysis of pile group. Computers and Geotechnics 2010;37(1—2):25—39.

Finn WDL, Fujita N. Piles in liquefiable soils: seismic analysis and design issues. Soil Dynamics and Earthquake Engineering 2002;22(9—12):731—42.

Finn WDL, Gohl WB. Centrifuge model studies of piles under simulated earthquake loading from dynamic response of pile foundations—experiment, analysis and observation. Geotechnical Special Publication American Society of Civil Engineering (ASCE) 1987;11:21—38.

Gao X, Ling XZ, Tang L, Xu P. soil—pile—bridge structure interaction in liquefying ground using shake table testing. Soil Dynamics and Earthquake Engineering 2011;31(7):1009—17.

Gonzalez L, Abdoun T, Sharp MK. Modelling of seismically induced liquefaction under high confining stress. International Journal of Physical Modelling in Geotechnics 2002;2(3):1—15.

Haeri SM, Kavand A, Rahmani I, Torabi H. Response of a group of piles to liquefaction-induced lateral spreading by large scale shake table testing. Soil Dynamics and Earthquake Engineering 2012;38:25—45.

Han JT, Kim SR, Hwang JI, Kim MM. Evaluation of the dynamic characteristics of soil—pile system in liquefiable ground by shaking table tests. In: The 4th International Conference on Earthquake Geotechnical Engineering. Thessaloniki, Greece; 2007. Paper No. 1340.

Hushmand B, Scott RF, Crouse CB. Centrifuge liquefaction tests in a laminar box. Geotechnique 1998;38(2):253—62.

Ishihara K. Terzaghi oration: geotechnical aspects of the 1995 Kobe earthquake. In: Proceedings of the 14th International Conference on Soil Mechanics and Foundation Engineering. Rotterdam, the Netherlands: A.A. Balkema; 1997. p. 2047—73.

Japanese Road Association (JRA). Specification for highway bridges, part V: seismic design. Tokyo, Japan: JRA; 1996.

Jeremic B, Cheng Z, Taiebat M, Dafalias Y. Numerical simulation of fully saturated porous material. International Journal for Numerical and Analytical Methods in Geomechanics 2008;32(13):1636—60.

Klar A, Baker R, Frydman S. Seismic soil—pile interaction in liquefiable soil. Soil Dynamics and Earthquake Engineering 2004;24(8):551 —64.

Kramer SL. Geotechnical earthquake engineering. Upper Saddle River, USA: Prentice-Hall Inc.; 1996.

Liu L, Dobry R. Effect of liquefaction on lateral response of piles by centrifuge model tests. NCEER Bulletin 1995;9(1):7—11.

Liyanapathirana DS, Poulos HG. Seismic lateral response of piles in liquefying soil. Journal of Geotechnical and Geoenvironmental Engineering 2005;131(12): 1466—79.

McKenna F, Fenves GL. The OpenSees command language manual. Version 1.2. Berkeley, USA: Pacific Earthquake Engineering Research Center, University of California; 2007. http://opensees.berkeley.edu.

Miura F, O'Rourke TD. Lateral spreading effects on pile foundations. In: Proceedings of the 3rd US-Japan Workshop on Earthquake Resistant Design of Lifeline Facilities and Countermeasures for Soil Liquefaction. MCEER; 1991.

Miura F, Stewart HE, O'Rourke TD. Nonlinear analysis of piles subjected to liquefaction induced large ground deformation. In: Proceedings of the 2nd US-Japan Workshop on Liquefaction, Large Ground Deformations and Their Effect on Lifelines. MCEER; 1989.

Mizuno H, Liba M. Shaking table testing of seismic building-pile-soil interaction. In: Proceedings of the 5th Japan Earthquake Engineering Symposium, Tokyo, Japan; 1982. p. 1713—20.

Oka F, Lu CW, Uzuoka R, Zhang F. Numerical study of structure-soil-group pile foundations using an effective stress based liquefaction analysis method. In: Proceedings of the 13th World Conference on Earthquake Engineering, Vancouver, Canada; 2004. Paper No. 3338 (CD-ROM).

Rahmani A, Fare OG, Pak A. Investigation of the influence of permeability coefficient on the numerical modeling of the liquefaction phenomenon. Scientia Iranica 2012;19(2):179—87.

Rahmani A, Pak A. Dynamic behavior of pile foundations under cyclic loading in liquefiable soils. Computers and Geotechnics 2012;40:114—26.

Rathje EM, Abrahamson NA, Bray JD. Simplified frequency content estimates of earthquake ground motion. Journal of Geotechnical and Geoenvironmental Engineering 1998;124(2):150—9.

Rollins K, Gerber T, Lane J, Ashford S. Lateral resistance of a full-scale pile group in liquefied sand. Journal of Geotechnical and Geoenvironmental Engineering 2005;131(1):115—25.

Su D, Li XS. Effect of shaking intensity on seismic response of single-pile foundation in liquefiable soil. In: Ground modification and seismic mitigation. ASCE; 2006. p. 379—86.

Tabesh A. Lateral seismic analysis of piles. PhD Thesis. Sydney, Australia: Department of Civil Engineering, University of Sydney; 1997.

Taboada VM, Dobry R. Experimental results of model No. 1 at RPI. In: Arulanandan K, Scott RF, editors. Verification of numerical procedures for the analysis of soil liquefaction problems. Rotterdam, the Netherlands: A.A. Bal-kema; 1993. p. 3—18.

Taiebat M, Jeremic B, Dafalias YF, Kaynia AM, Cheng Z. Propagation of seismic waves through liquefied soils. Soil Dynamics and Earthquake Engineering 2010;30(4): 236—57.

Tamura S, Tokimatsu K. Seismic earth pressure acting on embedded footing based on large-scale shaking table tests. In: Seismic performance and simulation of pile foundations in liquefied and laterally spreading ground. Reston, USA: American Society of Civil Engineers; 2005. p. 83—96.

Towhata I. Geotechnical earthquake engineering. Berlin-Heidelberg, Germany: Springer-Verlag; 2008.

Uzuoka R, Sento N, Kazama M, Zhang F, Yashima A, Oka F. Three-dimensional numerical simulation of earthquake damage to group-piles in a liquefied ground. Soil Dynamics and Earthquake Engineering 2007;27(5):395—413.

Varun V. A non-linear dynamic macroelement for soil structure interaction analyses of piles in liquefiable soils. PhD Thesis. Atlanta, USA: School of Civil and Environmental Engineering, Georgia Institute of Technology; 2010.

Wilson DW. Soil-pile-superstructure interaction in liquefying sand and soft clay. PhD Thesis. Davis, USA: University of California; 1998.

Yao S, Kobayashi K, Yoshida N, Matsuo H. Interactive behavior of soil-pile-superstructure system in transient state to liquefaction by means of large shake table tests. Soil Dynamics and Earthquake Engineering 2004;24(5):397— 409.

Ali Zahmatkesh was born in 1984 in Ferdows, Khorasan, Iran. He received his M.Sc. degree in Geotechnical Engineering from Mazandaran University, Iran in 2010. Since 2012, he is a Ph.D. student at Babol University of Technology. His thesis topic is Analysis of Performance of Pile Foundations in Liquefied Soils. His research interests cover soil improvement, soil liquefaction and numerical modeling.