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Pile Foundations during Earthquakes in Liquefiable Soils – Theory to Practice Deepankar Choudhury,1 Kaustav Chatterjee,2 Ashutosh Kumar3 and Reshma R. Phule4 Abstract Local soil sites and topography existing at a particular place play a significant role in amplification of ground motion in the event of an earthquake. As a result the frequency and amplitude of the seismic waves change when it travels from the bedrock to the ground surface and can have a detrimental effect on the overlying structure under resonance condition. Hence, construction of a stable foundation of high rise structures ensuring safety to the population in seismically active areas is a challenge faced by civil engineers. Liquefaction of soil subjected to earthquake loading is a major factor affecting pile foundations in seismically vulnerable areas. Ground motions, superstructure response, free field site response, soil – pile – superstructure interaction and pile – soil – pile interaction also have a significant impact on behavior of piles and pile-raft foundation installed in liquefied soil in seismic areas. In the present study, seismic equivalent linear ground response analysis at a particular soil site comprising of loose sand and subjected to four seismic ground motions is implemented using SHAKE2000. Fourier amplification ratio and response spectrum curves are generated which are useful for design engineers in evaluating the seismic hazard at a particular place. After carrying out ground response analysis and using the obtained results, static and dynamic analysis of single pile, passing respectively through homogenous and liquefied soil is implemented using finite difference based program FLAC3D (2009). The influence of dynamic loading on pile passing through liquefied soil is studied in terms of bending moment and deflection along pile depth. Finally, the present study is extended to combined pile-raft foundations (CPRF) under seismic conditions and their advantage over conventional pile foundation is proposed through practical case study. Professor, Department of Civil Engineering, IIT Bombay, Powai, Mumbai – 400076, India. Also, Adjunct Professor, Academy of Scientific and Innovative Research (AcSIR), New Delhi, India. Email:
[email protected] 2 PhD Research Scholar, Department of Civil Engineering, IIT Bombay, Powai, Mumbai – 400076, India. Email:
[email protected] 3 PhD Research Scholar, Department of Civil Engineering, IIT Bombay, Powai, Mumbai – 400076, India. Email:
[email protected] 4 PhD Research Scholar, Department of Civil Engineering, IIT Bombay, Powai, Mumbai – 400076, India. Email:
[email protected] 1
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Introduction A major factor governing the choice of suitable foundation types for high rise structures is the in-situ subsoil condition existing at that particular place.The infrastructural growth in urban cities having high population necessitates the construction of several high rise buildings and structures. These create a challenging task for civil engineers to ensure both safe and economic design for such structures. However, this task becomes more complex in seismically active areas. Indian design code IS 1893: Part 1 (2002) enlists the various criteria to be considered for seismic design of structures (mostly superstructures) in earthquake prone areas. But the use of various seismicity parameters for design and analysis of pile foundation and combined pile raft foundation for high rise constructions in seismically active areas are still comprehensively not available in Indian code. So, studying the response of ground to seismic motions, liquefaction vulnerability of the particular site and preparation of liquefaction hazard maps have become the basic topics of research among many civil engineers across the globe. If the top soil at a particular site is loose or soft or swelling in nature, pile foundations are widely used for transferring vertical and lateral loads from the superstructure to the deeper soil strata and hence have wide applications in both onshore and offshore engineering. The analysis of pile foundation under the impact of seismic forces can be categorized as a complicated soil – pile – superstructure interaction problem with soil – pile interaction having a significant impact on behavior of piles passing through liquefying soil. Hence, liquefaction induced lateral ground spreading is an important factor causing the failure of pile foundation in liquefaction prone areas [Tokimatsu et al. (1998), Ishihara (1997), Abdoun and Dobry (2002), Abdoun et al. (2003), Phanikanth et al. (2013)], as shown in Figs. 1 and 2. However, Bhattacharya (2003) and Dash et al. (2010) have mentioned aboutbuckling instability as another possible failure mechanism of pile foundation, especially when subjected to combined vertical load with lateral load on the pile top. Various researchers have studied and analyzed the behaviour of pile foundations in liquefiable soil using experimental, analytical and numerical procedures. Bhattacharya et al. (2004) proposed an alternative mechanism based on pile buckling for predicting the failure of pilesin liquefiable soil deposits during an earthquake. Liyanapathirana and Poulos (2005a,b) carried out numerical and analytical studies to obtain the response of pile foundation in liquefying soil based on finite element method and solving the partial differential equation of beams on Winkler foundation. Dynamic centrifuge tests for pile foundations in liquefiable soil have been carried out byAbdoun et al. (2003), Ghosh and Madabhushi (2007), Knappett andMadabhushi (2012) while shaking table tests have been conducted by Motamed et al. (2009, 2010) and Motamed and Towhata (2009) and few others. Ayothiraman and Boominathan(2013) carried out dynamic experiments on instrumented model aluminium single piles embedded in clay and studied its bending behaviour under the influence of lateral loading. The influence of axial loads on response of piles in liquefiable soil has been thoroughly studied by Knappett and Madabhushi (2009a, b), Dash et al. (2010). Bhattacharya et al. (2009) proposed an approximate method based on equivalent single degree of freedom model to obtain the fundamental frequency of a structure – pile – soil system while Shadlou and Bhattacharya (2014) proposed a set of formulas to obtain the dynamic stiffness of a pile for use in inertial interaction analysis based on elastodynamic solutions. Maheshwari et al. (2004) implemented numerical modelling of single pile and pile groups using a FORTRAN code 3dNDPILE to study the influence of non-linear behaviour of soil on the dynamic behaviour of piles while Maheshwari and Emani (2014) carried out seismic analysis of pile groups in three dimensional domain with non-linear behaviour of soil using FECIFECM coupling in Hybrid Domain and HiSS Plasticity model.Sarkar et al. (2014) proposed new
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(a) (b) Fig. 1: Liquefaction induced ground spreading and failure of (a) Kawagishi – Cho apartment and (b) Showa bridge due to lateral spreading during the 1964 Nigaata earthquake [ Kramer (2005)]
Fig. 2: Pile failure during earthquake as proposed by Tokimatsu et al.(1998) [see Choudhury et al. (2009)]
methodologies for implementing seismic requalification studies of pile foundations in liquefiable soils. If the soil layers under the structure have adequate bearing capacity, then raft can provide economical solution unless the serviceability of the structure is affected due to unacceptable settlement. In such cases conventional pile foundation is preferred which takes care of serviceability criteria but may leads to uneconomical design. Hence, introduction of few numbers of piles below raft foundation would reduce the settlement of raft foundation and make the foundation system economical. This composite foundation system which consist of spread foundation i.e. raft and comparatively few number of piles (friction or end bearing piles) is called Combined Pile-Raft Foundation (CPRF) [(Katzenbach et al. (2012)]. International Society of Soil Mechanics and Geotechnical Engineering provideinternational guideline for CPRF design,construction and practice, cover page of it is shown in Fig. 3(a) [Katzenbach and Choudhury (2013)]. Total resistance against the superstructure load depends upon sum of settlements of foundation and consists of pile resistance and base resistance, as shown in Fig. 3(b) and mentioned by Katzenbachet al. (2009).
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Rtotal,k= ΣRpile,k, j+ RRaft, k
(1)
Pile resistance: Rpile,k,j(s)=Rb,k,j(s) +Rs,k,j(s)
(2)
Raft resistance: Rraft,k(s)=∫∫ σ(s, x, y) dxdy
(3)
where, R total,kistotal resistance of CPRF, RRaft,kischaracteristics raft resistance, (ΣRpile,k,j) is characteristics pile resistance and s is settlement of spread foundation. The characteristic base resistance results from the integration of the settlement dependent contact pressure σ(s, x, y) in the ground plan area of the raft. The distribution of the super structure loadbetween the different bearing structures of a CPRF is described by the CPRF coefficient, αCPRF (Eq. 4) which defines the ratio between the amount of load carried by the piles Σ Rpilek,i and the total load of the building Rtotal. Co-efficient of one represents a freestanding pile group, ignoring the existence of the raft. Co-efficient of zero represents the load carried by raft foundation alone. CPRF coefficient: α CPRF =
∑ Rpile k,i Rtotal,k
(4)
The use of piles to reduce raft settlement may lead to economical design without compromising safety and performance of foundation system and the use of this technique have been increasing widely in recent years [Poulos et al. (2001)]. Detailed investigations of several high rise buildings like Messeturm Tower, Main Tower, Eurotheum etc. mainly in Frankfurt am Main, Germany were carried out which state that load sharing in CPRF depends upon pile-soil-raft interactions [(Katzenbach et al. (2005)]. Case study on twelve storey building in Japan revealed that CPRF shows the consistent behaviour under liquefiable soil deposits when Tohoku earthquake of Mw=9 struck the Japan on March 11, 2011 [Yamshita et al. (2012)].
(a) (b) Fig. 3: (a) Cover page of International Guideline for CPRF and (b) Schematic representation of load carrying and interaction mechanism of CPRF foundation[(Katzenbach and Choudhury (2013)]
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SEISMIC EQUIVALENT LINEAR GROUND RESPONSE ANALYSIS Site specific ground response analysis is an element of significant influence in the field of geotechnical earthquake engineering since it gives an insight to the profound impact of local soil conditions and geological features on the ground response, during an earthquake. Several factors like topography existing at a particular site, bedrock conditions and local soil layers amplifies the ground motion when it travels from the bedrock to the surface, which can have a detrimental influence on the overlying structures. Site specific ground response analysis can be implemented using linear, equivalent linear and non linear approach [Kramer (2005)]. In the present study, seismic equivalent linear ground response analysis is conducted at a soil site comprising of homogenous dry loose sand layer, having unit weight (γ) 14kN/m3, friction angle (ϕ) 300 and Poisson’s ratio (μ) 0.35, by using SHAKE2000 [Ordonez (2012)] computer software. Four different seismic ground motions are chosen for the analysis which includes 1989 Loma Gilroy, 1995 Kobe, 2001 Bhuj and 2011 Sikkim earthquake motions. It is observed from Fig. 4(a) that the maximum output acceleration at the ground surface is significantlyamplified when Kobe motion is the input motion, thereby indicating the ground acceleration at the bedrock level have been amplified by almost 1.72 times of that at the bedrock level. Similarly, for 1989 Loma Gilroy, 2001 Bhuj and 2011 Sikkim motions, the output acceleration got amplified at ground surface by 1.20, 2.39 and 1.58, respectively. The response spectrum curves at the ground surface for the various input motions are illustrated in Fig. 4(b). It is observed that the maximum spectral acceleration is 1.502g at a time period of 0.84sec for 1995 Kobe motion. The peak spectral acceleration for 1989 Loma Gilroy, 2001 Bhuj and 2011 Sikkimmotions are 0.56g at 0.4sec, 0.31g at 0.52sec and 0.46g at 0.21sec, respectively. LIQUEFACTION EVALUATION USING SIMPLIFIED PROCEDURES Liquefaction occurs generally at shallow depths in mostly loose to moderately dense granular saturated soils, due to rapid loading during seismic events due to insufficient time for dissipation of excess pore-water pressures through natural drainage. Urban structures were considered safer before the 2001 Bhuj earthquake, which shattered the myth of urban seismic safety through widespread
Fig. 4: (a) Variation of amplification of acceleration (amax(g)/MHA) along depth and (b) Response spectrum curves at 5% damping at ground surface
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damage to modern buildings [(Sinha et al. (2008)]. Prakash and Sandoval (1992) investigated the liquefaction behavior of low plasticity silts while Prakash and Guo (1998) carried liquefaction studies of silts with having clay content. Prakash and Puri (2010) conducted cyclic triaxial tests to study the liquefaction behavior of silts and silt clay mixers over a range of plasticity values. Measures to mitigate the damages caused by liquefaction require accurate evaluation of liquefaction potential of soils which has been carried out in the recent past using methods proposed by Robertson and Wride (1997), Youd et al. (2001), Seed et al.(2001, 2003), Cetin et al. (2004), Idriss and Boulanger (2004, 2008) and Boulanger and Idriss (2014). These methods, also known as deterministic methods, were based on standard penetration tests (SPT), cone penetration tests (CPT) and shear wave velocity (Vs) criteria and were developed from field liquefaction performance cases at sites that had been characterized with the corresponding in situ tests.In these procedures, liquefaction of a soil is predicted to occur if the factor of safety (FS), defined as the ratio of cyclic resistance ratio (CRR) of the soil required to cause initial liquefaction, divided by the cyclic stress ratio (CSR) induced at various depths within the soil by the earthquake, is less than or equal to 1. No soil liquefaction is predicted if FSis more than 1. All these procedures use common equation to determine the CSR based on the seismic (ground motion) parameters derived from the ground response analysis for specific site, but each equation use different relationship to determine the stress reduction coefficient ‘rd’[Choudhury et al. (2014)]. The CRR also varies significantly between these procedures because each procedure uses different factors to account for earthquake magnitude and overburden stresses. The variation of CRR for the three commonly used procedures of Youd et al. (2001), Cetin et al. (2004) and Idriss and Boulanger (2008) are shown in Fig. 5 andthe variation of the overburden stress factor, Kσ,proposed by Youd et al. (2001), Cetin et al. (2004) and Idriss and Boulanger (2004) are shown in Fig. 6.
Fig. 5: comparison of crr curves
Fig. 6: comparison of Kσ correction factors
AnALYSIS oF PILE FoundAtIonS In HoMoGEnouS And LIQuEFIEd SoIL After conducting site specific ground response analysis and liquefaction study at a seismically active area, the next important area of study is analysis of pile foundation subjected to input seismic motions and passing through liquefiable and non-liquefiable soil layers. Generally, piles are initially analyzed and designed for static loading and if found safe against deflection and bending, they are subjected to dynamic loadsand analysis is carried out, as has been done in the present study.
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Analysis of pile subjected to static loading As shown in Fig. 7, a single pile is modeled in finite difference based geotechnical software FLAC3D (2009) which is a 3-Dimensional explicit finite difference based computer program and performs a Lagrangian analyses for engineering mechanics computation.The numerical model developed in the present study is validated with Reese and Matlock (1956) theory for both free headed and fixed headed piles subjected to lateral load. A concrete pile of M30 grade concrete and loose sandy soil is chosen as the input parameters [Rao et al. (2013)], with the normal stiffness and shear stiffness at soil-pile interface being calculated as per Timoshenko and Goodier (2002). The results obtained are illustrated in Figs. 8 and 9 for fixed headed and free headed piles, respectively and maximum pile head deflection and pile bending momentarenoted. It is observed that the present results are in good agreement with the results obtained using Reese and Matlock (1956) theory for both free headed and fixed headed piles.
(a) (b) Fig. 7: Mesh for (a) soil grid and (b) soil-pile grid considered in the present analysis in FLAC3D
Fig. 8: comparison of (a) pile deflection and (b) pile bending moment along depth obtained in present study with that obtained from reese and Matlock (1956) theory for fixed headed pile under static condition
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Fig 9: Comparison of (a) pile deflection and (b) pile bending moment along depth obtained in present study with that obtained from Reese and Matlock (1956) theory for free headed pile under static condition
Analysis of piles subjected to dynamic loading in liquefied soil After conducting static analyses of the pile soil model generated in FLAC3D (2009), dynamic analyses of a free headed pile with floating tip boundary condition is implemented by subjecting the piles to 4 different input seismic motions. The ground motions chosen for the present analyses are 1989 Loma Gilroy motion, 1995 Kobe motion, 2001 Bhuj motion and 2011 Sikkim earthquake motion, having a wide variation of strong motion parameters like frequency content, bracketed duration and peak ground acceleration, as indicated by Phanikanth et al. (2011). The input motions are applied at the tip of the free headed pile having length (l) 8m and diameter (d) 600mm and embedded in a two-layered soil system comprising of a liquefiable layer of loose sand underlain by a non-liquefiable layer of stiff clay, as shown in Fig. 10. The water table is located at the ground leveland the same is simulated in the numerical model. The non-liquefiable clay layer and liquefiable loose sand layer are modeled using Mohr-Coulomb and Byrne (1991) constitutive model, respectively. The analysis is conducted for kinematic loading by varying the depth of liquefied layer (Lliq) and the response of the pile in terms of bending moment and deflection along the pile depth is obtained.
Fig. 10: (a) Dynamic centrifuge test model of Abdoun et al. (2003) and (b) model used in the present study for dynamic analysis of single pile in FLAC3D
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AnALYSIS And rESuLtS The behaviour of a free headed pile passing through liquefiable loose sand layer and embedded in non-liquefiable stiff clay layer under the action of various input seismic motions are illustrated in Figs. 11 (a) and (b). It is observed that the maximum bending moment occurs at a depth of 5m, which is the interface of two layers, for all the input ground motions. Further, the maximum bending moment of 175kNm is observed for 1995 Kobe motion, while for 1989 Loma Gilroy, 2001 Bhuj and 2011 Sikkim motions the bending moment magnitudes are 118.3kNm, 125kNm and 132kNm, respectively. The variation of pile deflection with pile depth is illustrated in Fig. 11(b). It is observed that the maximum pile head deflection is 48.2cm under the action of 1995 Kobe motion, while for 1989 Loma Gilroy, 2001 Bhuj and 2011 Sikkim motions, the deflection magnitudes are 40.6cm, 23.9cm and 32cm, respectively. The magnitude of pile deflection decreases with depth of the pile and reaches almost zero beyond a depth of 5m. This is due to the deeper soil layers are more densified owing to the high overburden stress transmitted from the top soil layers [Chatterjee and Choudhury (2013)]. Beyond a depth of 5m, the non-liquefiable stiff clay layer has both higher density and higher shear stiffness and normal stiffness compared to the liquefiable loose sand layer, and thereby resulting in almost negligible displacement. The influence of depth of liquefiable layer on soil – pile response and subjected to kinematic loading is also studied in the present analysis. The depth of liquefied layer Lliqis varied in terms of total pile length l as Lliq/l = 0.25, 0.5, 0.625, 0.75 and 1.0. The variation of bending moment and deflection with pile depth for 2001 Bhuj motion for different combinations of Lliq/l is shown in Figs. 12(a) and (b). It is observed for 2001 Bhuj motion that maximum pile head deflection is 10.4cm when Lliq/l is 0.25 and it increases to 23.9cm when Lliq/l increases to 0.6, thereby showing an increase of 56.5%. However, when the entire soil is liquefying, i.e., Lliq/l = 1.0, the deflection at the pile head is 14.6cm while at the pile tip it is -16cm. A similar trend of results is also observed for the other 3 different earthquake motions. The variation of bending moment with pile depth for different combinations of Lliq/l ratio is shown in Fig. 12(a). The maximum bending moment is observed to occur at the interface of the liquefied layer and non-liquefied layer and is maximum when Lliq/l ratio is approximately 0.6 and with a subsequent increase in depth of liquefying layer,
Fig. 11: Variation of (a) bending moment (knm) with pile depth (m) and (b) deflection (cm) with pile depth (m)for different earthquake motions considered in the present study
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Fig. 12: Variation of (a) bending moment (knm) with pile depth (m) and (b) deflection (cm) with pile depth (m)for different combinations of Lliq/l ratio and subjected to 2001 Bhuj motion
the magnitude of bending moment reduces. A maximum bending moment of 125kNm is observed for 2001 Bhuj earthquake motion and when Lliq/l ratio is 0.6. However, when Lliq/l ratio increases to 1.0, the bending moment reduces to 68kNm, thereby showing a reduction of approximately 45.6%. SEISMIc AnALYSIS oF coMBInEd PILE rAFt FoundAtIonS (cPrF) The analysis of CPRF has been carried out in the present study using FLAC3D (2009). Four concrete piles having length 9m and diameter 0.5m is modelled having 2m spacing centre to centre. The raft of size 4m x and 1m thickness is modelled. The Young’s modulus of raft and piles is taken as 41.7 GPa. Dimensions of raft and piles are taken from Horikoshi et al. (2003) and Eslami et al.(2011). Toyoura sand was considered in the present study. The modulus of elasticity of Toyourasand is taken as 4x104kPa, unit weight of Toyoura sand is 16.3 kN/m3 with Poisson’s ratio of 0.3, dilation angle of 1o, cohesion of 10 kPa. The properties of Toyoura sand is taken from Kimura et al. (1985). Modeling of cPrF Mohr-Coulomb yield criterion is chosen for Toyoura sand model and raft and piles are considered as elastic member in the analysis. Four square piles having 9m length and size of 0.5m is generated and connected rigidly to the raft. The three dimensional view of soil-pile-raft model generated in FLAC3D is shown in the Fig. 13. In order to simulate actual field condition, soil mass is fixed at base and roller boundaries are employed at the sides so that soil model can deform under the application of vertical load. The capacity of a pile is a function of the skin friction resistance along the pile shaft only is considered in the present analysis. The skin friction resistance is modelled by placing an interface along the pile surface. The interface properties have been selected as per experimental results of Horikoshi et al. (2003) and are provided between the soil and raft in order to prevent separation between raft and soil. The model was first brought to equilibrium state under gravity loads before the installation of piles. Then piles were modelled and again brought into equilibrium. Actual vertical stress (according to the equation = γh) at theinitial condition are -260 kPa at the bottom and 0 kPa at top. It is observed that the vertical stresses computed by FLAC3D are almost
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same values as specified.The load applied on to the raft is in theform of vertical pressure which is obtained from scaling the mass of raft i.e. 4.69 kg used in centrifugerotated at 50g to normal gravity level which comes around 366.4kPa of vertical pressure. The boundary conditions of the model is modified in dynamic analysis by applying free – field boundary condition which contain dampers in normal and shear directions to absorb the seismic wave reflections at the model boundaries, there by simulating the infield conditions. Analysis and results The model is analyzed under static loading condition which shows the maximum bending moment of 138 kNm and settlement of 2.5 cm. Total vertical loads taken by all the piles are around 2600kN which gives the CPRF co-efficient, i.e., total load carried by the pile foundation to the total load applied to the CPRF system is around 0.44.For the validation of model an input dynamic loading of 1 m/s2 in the prototype scale with a frequency of 1Hz was applied at the model base which is very similar to centrifuge test done by Nakai et al. (2004) has been applied to the base of soil in x- direction. Fig. 14 shows the maximum bending moment response of pile along its length under sinusoidal loading condition in FLAC3D. From Fig. 15 it can be inferred that bending moment response is in agreement with Eslami et al. (2011) when compared for sinusoidal loading case and the maximum bending moment observed on the pile top is 148 kNm. After validation of the present model, same model is analysed under real acceleration time history of 1979 El-Centro and 2001 Bhuj earthquake motions having PGA of 0.106g and 0.425g, respectively. Bending moment response of pile in CPRF under acceleration time history of El- Centro and Bhuj Earthquake is shown in Fig. 15. Under El-Centro loading condition maximum bending moment observed is 259 kNm which matches with the bending moment response shown by Eslami et al. (2011). Under 2001, Bhuj earthquake loading, maximum bending response is 196.9 kNm. Total settlement of the CPRF is found to be 5.7 cm under El-Centro loading condition which is nearly same as mentioned by Eslami et al. (2011) and in case of 2001Bhuj motion total settlement of nearly 3cm is observed.
Fig. 13: Three dimensional view of CPRF model in FLAC3D (2009)
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Fig. 14: Bending moment variation along pile length in sinusoidal loading condition in FLAc3d
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Fig. 15: Bending moment variation along pile length under different cases
cASE StudY oF cPrF undEr EArtHQuAKE condItIon A practical case study on a base isolated twelve storey building of Tokyo having height of 38.7m which was founded on loose silty sand underlain by soft cohesive soil having huge potential for liquefaction. CPRF with deep cement mixing was employed to increase the bearing capacity for raft foundation and cope of with the problem of liquefaction, shown in Fig. 16. The PGA of 1.748m/s2 was observed near the ground surface. The other building Hardon Experimental of Ibarakiin Japan having height of 19m is founded on subsoil having medium to dense sand underlain by cohesive layer. The building is supported by CPRF with 371 PHC piles. The PGA at the depth of 6m below the ground was recorded as 3.24m/s2. In both the buildings, settlement of the foundation, contact
(a) (b) Fig. 16: (a) tweleve storey building in tokyo and (b) Ground subsidence along the experimental hall [after Yamashita et al. (2012)]
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pressures and pore water pressures were monitored from the date of construction to the stuck of 2011, Tohuku Pacific Earthquake Mw=9. All monitored data were consistent before and after the earthquake hence showing excellent performance of CPRF during earthquake, showing excellent performance of CPRF during earthquake. Conclusions The major conclusions from the present study are highlighted below as follows: 1. Seismic equivalent linear ground response analysis shows that the amplification of ground motion is more for 2001 Bhuj and 2011 Sikkim motions, compared to 1995 Kobe motion, although the later had higher amplitude of motion at bedrock level. This is due to the higher frequency content and bracketed duration of the former two seismic motions. 2. The static analysis of pile foundations using FLAC3D (2009) is found to validate well with Reese and Matlock (1956) theory. Further, with an increase in vertical load on the pile top, the lateral load carrying capacity of both fixed headed and free headed pile is found to increase substantially. 3. Dynamic analysis of pile embedded in liquefiable soil shows that maximum bending moment occurs at the interface of the liquefiable soil layer and non-liquefiable soil layer and bending moment is maximum when the depth of the liquefiable layer is approximately 60% of the total pile length. 4. Present study shows that response of CPRF is quite well because of major contribution of raft in load sharing which is not the case in case of pile group foundation where raft does not have any contribution in load sharing. 5. Behavior of CPRF under sinusoidal loading, Bhuj and El-Centro loading conditions has been carried out using finite difference based software FLAC3D. The bending moment response of pile under dynamic loading matches with the results proposed by Eslami et al. (2011). The obtained results show that maximum bending moment along the pile length was observed to be 148kNm, 259 kNm and 196.9 kNm under sinusoidal, El- Centro and Bhuj earthquake respectively and there is 7%, 87% and 42% increase in bending moment, respectively when compared with static case. 6. It can be concluded that CPRF is taking load under combined interaction of soil- pile-raft. Major portion of load has been taken by raft as compared to piled foundation. Bending moment response shows that CPRF is showing consistent behavior under different earthquake loading conditions. Practical case study of CPRF during earthquake also validates the theoretical findings. REFERENCES 1. 2. 3. 4.
Abdoun, T. and Dobry, R. (2002).“Evaluation of pile foundation response to lateral spreading,” Soil Dynamics and Earthquake Engineering, 22(9–12), 1051–1058. Abdoun, T., Dobry, R., O’Rouke, T.D. and Goh, S.H. (2003). “Pile foundation response to lateral spreads: Centrifuge modelling,”Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 129 (10), 869878. Ayothiraman, R. and Boominathan, A. (2013). “Depth of fixity of piles in clay under dynamic lateral load,” Geotechnical and Geological Engineering, 31, 447-461. Bhattacharya, S. (2003). “Pile instability during earthquake liquefaction,” Ph.D. Thesis, University of Cambridge, U.K.
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