6th International Conference on Earthquake Geotechnical Engineering 1-4 November 2015 Christchurch, New Zealand

Three Dimensional Numerical Analysis of Seismic Soil-Structure

Interaction Considering Soil Plasticity

R. Xu1, and B. Fatahi2

ABSTRACT Effects of soil Plasticity Index (PI) on the seismic response of mid-rise buildings considering

soil-structure interaction is investigated in this study. A 15-storey building resting on class E soil with different values of PI has been simulated utilising FLAC3D software. Fully nonlinear dynamic analysis under different earthquake recording including two near field and two far field earthquakes has been conducted and results in terms of the base shear, maximum lateral displacement and inter-storey drift are compared and discussed. Results indicate that as PI increases, the base shear of the structure resting on the soft soil, the maximum lateral displacement and inter-storey drift increase. It is concluded that practicing engineers should consider effects of soil plasticity on the seismic design of building frames constructed on soft soils accurately.

Introduction

During the recent decades, the importance of dynamic soil-structure interaction (SSI) has been well recognised in analysis of seismic response of superstructures. Several studies have shown that the superstructures are vulnerable to the effect of SSI, particularly when structures rest on soft soil deposits (e.g. Veletsos and Meek 1974; Hosseinzadeh and Nateghi 2004; Hokmabadi et al. 2014; Fatahi et al. 2014). When the soft soil deposit is subjected to strong seismic loading, significant soil damping is induced by soil modulus degradation and inertial interaction becomes predominant, causing excessive displacements near the ground surface. This behaviour of soft soil, which can be described in the forms of backbone curves, is nonlinear and hysteretic. The curves expressing shear modulus and damping ratio as a function of shear strain vary with soil characteristics, which significantly influence the structural response considering SSI. Moreover, Vucetic and Dobry (1991) concluded that plasticity of soils plays an important role in determining backbone curves in cohesive soils. Plasticity is an important characteristic of fine soils, which indicates the ability of a soil to deform irreversibly without cracking or crumbling. The plasticity of a soil can be described by Plasticity Index (PI) which indicates the range of water content where a soil exhibits plastic behaviour under stress. The lower and upper limits of the range of water content over which the soil behaves plastically are defined as Plastic Limit (PL) and Liquid Limit (LL), respectively, while the range of water content can be defined by Plasticity Index (PI) as follows: PI = LL − PL (1)

1PhD Candidate, School of Civil and Environmental Engineering, University of Technology Sydney (UTS), Sydney, Australia, Ruoshi.Xu@student.uts.edu.au 2Senior Lecturer of Geotechnical Engineering (PhD, CPEng), School of Civil and Environmental Engineering, University of Technology Sydney (UTS), Sydney, Australia, Behzad.fatahi@uts.edu.au

Dynamic Behaviour of Cohesive Soils Figure 1 shows the nonlinear stress-strain relationship of typical soils subjected to the cyclic loading. Two important characteristics of hysteresis loop, which are inclination and breath are used to describe the hysteresis response of soils. According to Kramer (1996), the inclination of the loop represents stiffness of the soil, which can be described at any point during the loading process by the tangent shear modulus 𝐺𝐺𝑡𝑡𝑡𝑡𝑡𝑡. It is clear that the tangent shear modulus varies throughout a cycle of loading, but its average value over the entire loop can be approximately represented by the slope of the line connecting the origin to the tip of the loop and the slope defines the secant shear modulus 𝐺𝐺𝑠𝑠𝑠𝑠𝑠𝑠. Gsec = τc

γc (2)

where, 𝜏𝜏𝑠𝑠 and 𝛾𝛾𝑠𝑠 are the shear stress and shear strain amplitudes at the defined point, respectively. Therefore, 𝐺𝐺𝑠𝑠𝑠𝑠𝑠𝑠 represents the inclination of the hysteresis loop. The breath of the hysteresis loop, which is related to the area within the loop, represents the absorbed energy, and damping ratio ξ can be defined as below: ξ = WD

4πWS= 1

2πAloopGγc2

(3) where, 𝑊𝑊𝐷𝐷 is the absorbed energy in one loop, 𝑊𝑊𝑆𝑆 is the maximum strain energy created by the loop, and 𝐴𝐴𝑙𝑙𝑙𝑙𝑙𝑙𝑙𝑙 is the area of the hysteresis loop which is equal to 𝑊𝑊𝐷𝐷. The parameters 𝐺𝐺𝑠𝑠𝑠𝑠𝑠𝑠 and ξ are used to describe the cyclic behaviour of the soil in the equivalent linear analysis and often referred to as the soil equivalent linear parameters.

Figure 1 (a) Hysteretic stress-strain relationship; (b) Backbone curve; and (c) Typical modulus reduction curve for soils (After Kramer, 1996)

The secant shear modulus of the soil element changes with the cyclic shear strain amplitude, which is large at low strain amplitudes and decreases as the shear strain amplitude increases. As shown in Figure 1b, the locus of tips of the hysteresis loops of different cyclic shear strain amplitudes form the backbone curve which matches the monotonic loading curve for the same type of soil and the slope of the backbone curve at the origin represents the maximum value of the shear modulus, 𝐺𝐺𝑚𝑚𝑡𝑡𝑚𝑚. At greater shear strain amplitudes, the modulus ratio, 𝐺𝐺𝑠𝑠𝑠𝑠𝑠𝑠 𝐺𝐺𝑚𝑚𝑡𝑡𝑚𝑚⁄ drops to values less than one. Therefore, to precisely represent the behaviour of the soil under the cyclic loading, both 𝐺𝐺𝑚𝑚𝑡𝑡𝑚𝑚 and modulus reduction curve are required.

Aloop=WD

γc γ

τ

τc

WsGtan

Gsec

(a) (c)

Gsec

γc

τ

γ

Gmax

Backbone curve

Modulus reduction curve

γc log γ

1

(b)

Vucetic and Dobry (1991) conducted a review on the available cyclic loading test results and concluded that the Plasticity Index (PI) is the main factor controlling the variations of the shear modulus reduction and damping ratio against the cyclic strain curve for a wide variety of cohesive soils. Solid lines in Figure 2a and 2b illustrate the ready-to-use charts provided by Vucetic and Dobry (1991) for modulus degradation and damping ratio, respectively. The aim of these two charts is to provide a design tool for practicing engineers since PI is readily available. As observed in Figure 2, when PI increases, 𝐺𝐺𝑠𝑠𝑠𝑠𝑠𝑠 𝐺𝐺𝑚𝑚𝑡𝑡𝑚𝑚⁄ increases while damping ratio decreases.

Figure 2 (a) Relation between G⁄Gmax versus cyclic shear strain for cohesive soils; (b) Relations between damping ratio versus cyclic shear strain for cohesive soils

Numerical Modeling

In this study, dynamic seismic soil-structure interaction analysis is conducted adopting direct method which evaluates the dynamic response in a single step, as it can perform fully nonlinear analysis. Furthermore, time domain analysis as recommended by Chu (2006) necessary to compute the nonlinear dynamic response is utilised. In order to have realistic fully nonlinear analysis, a three dimensional explicit finite difference based program FLAC3D (Fast Lagrangian Analysis of Continua) version 5.0 has been employed. In this program, behavior of different types of materials according to their prescribed constitutive models in response of applied loads and boundary conditions can be simulated. For a dynamic analysis, damping in the numerical simulation should be reproduced in magnitude and pattern related to the energy losses in the system when subjected to a dynamic loading. In soil and rock, natural damping is mainly hysteretic. Hysteretic damping algorithm is incorporated in FLAC3D dynamic analysis to simulate the realistic behavior of soils. Modulus degradation curves imply nonlinear stress-strain curves. In case of an ideal soil in which the stress depends only on the strain, an incremental constitutive relation from the degradation curve can be described by the strain-dependent normalised secant modulus (Ms) as follows: Ms = τ̅

γ (4)

where, τ� is the normalised shear stress which can be obtained through local shear stress divided by the initial shear modulus and γ is the shear strain. The normalised tangent modulus (Mt) is then obtained as follows:

0

0.2

0.4

0.6

0.8

1

0.0001 0.001 0.01 0.1 1 10

G / Gm

ax

Cyclic Shear Strain, γc %

PI = 200

PI = 100

PI = 50

PI = 0

PI = 15

PI = 30

After Vucetic and Dobry (1991) Applied SIG-III curves in this study

0

5

10

15

20

25

30

0.0001 0.001 0.01 0.1 1 10

Damp

ing Ra

tio, %

Cyclic Shear Strain, γc %

After Vucetic and Dobry (1991) Applied SIG-III curves in this study

PI = 200

PI = 100

PI = 50

PI = 30

PI = 15

PI = 0

(b)

(a)

Mt = dτ̅dγ

= Ms + γ dMsdγ

(5) The incremental shear modulus in a nonlinear simulation is then given by 𝐺𝐺𝑠𝑠𝑠𝑠𝑠𝑠, where 𝐺𝐺𝑠𝑠𝑠𝑠𝑠𝑠 is the given shear modulus obtained from Equation (2). The formulations described in Equations (4) and (5) are implemented in FLAC3D, by modifying the strain rate calculation so that the mean strain rate tensor (averaged over all subzones) is calculated before any calls are made to the constitutive model functions. At this stage, the hysteretic logic is invoked, returning a modulus multiplier, which is passed to any called constitutive model. The model then uses the multiplier Mt to adjust the apparent value of the tangent shear modulus of the full zone being processed. In this study, the tangent modulus model named SIG-III implemented in FLAC3D is employed to simulate hysteretic behavior in the soil deposit. The mathematical formulation of the model is defined as: Ms = a

1+exp (−log10(γ)−x0b )

(6) where, Ms is the secant modulus (G/Gmax), γ is the cyclic shear strain, and a, b and x0 are the model parameters. By adopting different model parameters, this model is able to generate different backbone curve for different types of material in dynamic analysis. Earthquake ground motions In order to perform a comprehensive study on the seismic response of the structural models, two near field seismic accelerations which are Northridge, 1994 and Kobe, 1995 and two far field seismic accelerations which are El-Centro, 1940 and Hachinohe, 1968 shown in Figure 3 are utilised in the time history analysis, which are selected by the International Association for Structural control and Monitoring for benchmark seismic studies (Karamodin and Kazemi 2008) and Figure 4 illustrates the spectral accelerations for each of earthquake input.

Figure 3 (a) Northridge Earthquake 1994; (b) Kobe Earthquake 1995; (c) El Centro Earthquake 1940; (d) Hachinohe Earthquake 1968

-1

-0.5

0

0.5

1

0 10 20 30

Accel

eratio

n (g)

Time (Sec)

(a)-1

-0.5

0

0.5

1

0 10 20 30 40 50

Accel

eratio

n (g)

Time (Sec)

(b)

-0.4

-0.2

0

0.2

0.4

0 10 20 30

Accel

eratio

n (g)

Time (Sec)

(c)-0.4

-0.2

0

0.2

0.4

0 10 20 30 40

Accel

eratio

n (g)

Time (Sec)

(d)

Figure 4 Spectral accelerations of four input earthquake with 5% damping ratio Structural models and soil deposit Figure 5 shows the dimensions of the structure adopted in this study, whose natural period is 1.28 seconds. SAP2000 V14 has been utilised for the structural design purpose. All the structural sections of the model have been designed based on the inelastic method assuming elastic - perfectly-plastic behavior. According to AS/NZS1170.1-2002 (Permanent, imposed and other actions), permanent and imposed loads are determined and applied to the structure model. It should be noted that cracked sections for the reinforced concrete sections are taken into account by multiplying second moment of area of the uncracked sections (Ig) by cracked section coefficients (0.35Ig for beams, 0.70Ig for columns and 0.25Ig for slabs) according to ACI318-08 (2008). The model foundation is a square shallow reinforced concrete foundation which is 14 meters in length and width, and 1 meter in depth. The entire numerical model has been illustrated in Figure 5. Then four earthquakes have been applied at the base. Finally, the structural members are designed in accordance with AS3600-2009 (Australian Standard for Concrete Structures) in a way that performance levels of the designed models stay in life safe level by limiting the maximum inelastic inter-storey drift to 1.5%.

Figure 5 (a) Fixed base model; (b) Flexible base model utilised in FLAC 3D simulation

0

1

2

3

0 0.5 1 1.5 2 2.5 3

Accel

eration

(g)

Period (s)

El-Centro

Hachinohe

Kobe

Northridge

Column: 700x700 mmBeam: 550x550 mmSlab thickness: 250 mm

Column: 800x800 mmBeam: 650x650 mmSlab thickness: 250 mm

Column: 750x750 mmBeam: 600x600 mmSlab thickness: 250 mm

14m60m

Quiet boundary coupled with free-field boundary

Beam structure elements

Shear wave velocity Vs= 150m/sMaximum shear modulus Gmax =

33100 kPaPoisson’s ratio n= 0.4

Soil density r= 1470 kg/m3

Interface elements

30m

(a) (b)

According to the classification of AS1170.4-2007 (Earthquake actions in Australia), a soft soil representing subsoil class E has been selected in this study. Figure 3b shows the characteristics of the soil and dimensions adopted in this study. Different sets of SIG-III model parameters a, b and x0 have been determined in order to regenerate backbone curves reported by Vucetic and Dobry (1991) to be used in the numerical simulation. Table 1 summarises the determined values of the model parameters.

Table 1 SIG-III model parameter for various Plasticity Indices

The interface elements are applied between the superstructure and the soil deposit to simulate the real behaviour between the structure and the subsoil. The interface between the foundation and soil is represented by normal (kn) and shear (ks) springs between two planes contacting each other and is modelled using linear spring system, with the interface shear strength defined by the Mohr–Coulomb failure criterion. Obviously, the relative interface movement is controlled by interface stiffness values in the normal and tangential directions. Normal and shear spring stiffness values for interface elements of the soil-structure model are set to ten times the equivalent stiffness of the neighbouring zone based on the recommendation of Itasca FLAC3D manual (2014).

ks = kn = 10 ��K+43G�

∆zmin� (7)

where, K and G are bulk and shear modulus of the neighbouring zone, respectively, and ∆zmin is the smallest width of an adjoining zone in the normal direction. Additionally, the interface elements have been applied in the way that the foundation can separate from the soil and gaping is allowed during the analysis. For side boundaries of the soil deposit, free field boundaries have been employed. Free-field boundaries have been simulated using a developed technique, involving one-dimensional free-field wave propagation in parallel with the main-grid analysis (Itasca, 2014). Thus, plane waves propagating upward undergo no distortion at the boundaries because the free-field grid supplies conditions identical to those in an infinite model. In addition, a rigid boundary as the bottom of the soil deposit has been adopted in this study.

Results and discussions

The results of dynamic analyses for 15 storey structural models in terms of the maximum base shear, lateral deflection and inter-storey drift under the influence of four earthquake records for fixed base condition and cases considering soil-structure-interaction with different Plasticity Indices are derived from FLAC3D history records and compared in Tables 2 - 3 and Figure 6.

According to Table 2, it is observed that the base shear of the structures modeled considering SSI is always less than the base shear of the corresponding fixed base cases. Comparing the base shear results, it is evident that as the PI of the subsoil increases from 0 to 200, the base

Plasticity Index(PI)) PI=0 PI=15 PI=30 PI=50 PI=100 PI=200

SIG-III model parameter

a 1.25 1.25 1.25 1.25 1.25 1.25b -0.7 -0.7 -0.7 -0.7 -0.7 -0.7x0 -1.75 -1.55 -1.2 -0.85 -0.5 -0.2

shear increases. As observed in Figure 2, by increasing the soil plasticity, the stiffness of the subsoil increases, while damping ratio decreases. Thus less distortion of the soil occurs as PI increases and consequently more energy transfers through the soil into the system. Thus, the increase of PI evidently leads to increase in the base shear of the building resting on the soft soil deposit.

Figure 6 Maximum lateral displacement (a) Northridge1994; (b) Kobe1995; (c) El-Centro1940; (d) Hachinohe1968

Table 2 Maximum base shear results obtained from different cases

Table 3 Maximum Inter-storey drift reported for different cases

Comparing the results in Figure 6 and Table 3 both taken when the maximum lateral displacement occurred at top floor, it becomes evident that when PI = 0 (e.g. low plasticity silt), the lowest maximum lateral displacement is observed (even lower than the

0

5

10

15

0 500 1000 1500

Stor

ey le

vel

Lateral displacement (mm)

PI0PI15PI30PI50PI100PI200Fixed 0

5

10

15

0 200 400 600 800 1000

Stor

ey le

vel

Lateral displacement (mm)

PI0PI15PI30PI50PI100PI200Fixed

(b)

0

5

10

15

0 100 200 300 400

Stor

ey le

vel

Lateral displacement (mm)

PI0PI15PI30PI50PI100PI200Fixed

(c)

0

5

10

15

0 100 200 300 400

Stor

ey le

vel

Lateral displacement (mm)

PI0PI15PI30PI50PI100PI200Fixed

(d)

(a)

EarthquakeBase shear kN

PI=0 PI=15 PI=30 PI=50 PI=100 PI=200 Fixed baseNorthridge 7068 12611 14952 25067 26855 27068 36653

Kobe 6966 9977 11765 14666 19607 23314 36127El-Centro 5148 8470 9200 11089 12062 12103 14345Hachinohe 4382 7267 10668 10640 11723 12474 13727

Earthquake Maximum inter-storey drift (%)PI=0 PI=15 PI=30 PI=50 PI=100 PI=200 Fixed base

Northridge 0.77 1.18 2.07 2.84 2.76 2.92 1.10

Kobe 0.65 1.20 1.84 2.00 1.83 1.84 1.42

El-Centro 0.48 0.60 0.76 0.84 0.91 0.92 0.48

Hachinohe 0.33 0.44 0.61 0.78 0.75 0.94 0.36

corresponding value for the fixed base case), while for PI = 200 (e.g. high plasticity clay), the highest maximum lateral displacement is achieved. The displacement increases as PI increases (except Kobe case, where due to the change in the system period, the second mode could be activated with extreme strong base shear at a certain input frequency). Although the shear wave velocity (or Gmax) is constant, the performance of the soil deposits is significantly influenced by PI, due to the combined effects of shear modulus degradation and damping ratio. Moreover, as the damping reduces due to the increase in the values of PI, the magnitude of acceleration reaching the structure and dominant frequency increase, which is less than dominant frequency of input acceleration for all cases. Thus, it is reasoned to conclude that due to the decrease in damping, the response of structure increases, which eventually indicates the variation of the damping ratio induced by PI has a dominant effect on the system behaviour. In addition, considering the conservation of energy, in relatively stiff soil deposits experiencing less soil distortion and damping during earthquake, more energy may transfer into the structure resulting in increased foundation rocking and structural displacement. Generally, for soils with low Plasticity Index (e.g. low plasticity silt), the maximum lateral displacement and inter-storey drift of the structure built on the soft soil are less than the corresponding values for the fixed base structure. In contrast, when soil has high plasticity (e.g. high plasticity clay), the maximum displacements and inter-storey drifts have been amplified significantly, especially, for the near field earthquakes (Northridge and Kobe). In fact, for both near field earthquakes, the maximum recorded inter-storey drifts considering SSI are well more than 1.5% (life save level), which is extremely dangerous and safety threatening.

Conclusions

In this study, the influence of Plasticity Index (PI) variations (PI=0 200) on the seismic response of mid-rise buildings has been numerically investigated. A 3D numerical soil-structure model has been developed and employed utilising FLAC3D adopting direct method of analysis. In order to capture soil nonlinearities, SIG-III model (Eqn. 6) has been used to simulate backbone curves of the shear modulus and the damping ratios versus the shear strain for soils with different Plasticity Indices. Numerical results show that as the Plasticity Index of subsoil increases (G/Gmax increases and damping ratio decreases), the base shear, the maximum lateral displacement and the maximum inter-story drift increase. The amplification of the lateral displacement and internal drifts could potentially change the performance level from life safe to total collapse which is safety threatening. It can be concluded that soil-structure interaction has considerable effects on the seismic response of mid-rise building frames resting on soft soil deposits and increase of the Plasticity Index could considerably amplify response of the structure. Thus, conventional design procedure excluding soil-structure interaction and soil plasticity may not be adequate for the safe design of mid-rise buildings resting on soft soils. Furthermore, in order to obtain reliable results, the influence of Plasticity Index should be taken into account while conducting soil-structure interaction analysis.

References

ACI318-08 (2008). Building Code Requirements for Structural Concrete and Commentary. American Concrete Institute.

AS/NZS1170.1. (2002). Structural Design Actions - Part 1: Permanent, Imposed and Other Actions, Standards Australia/Standards New Zealand, Sydney, Australia.

AS1170.4 (2007). Structural Design Actions - Part 4: Earthquake Actions in Australia, Standards Australian,

Sydney, Australia.

AS3600 (2009). Concrete Structures, Standards Australia Limited, Sydney, Australia.

Chu, D. (2006). Three-dimentional nonlinear dynamic analysis of soil-pile-structure interaction. PhD, Washington University.

Fatahi, B. & Tabatabaiefar, S. 2014, 'Fully Nonlinear versus Equivalent Linear Computation Method for Seismic Analysis of Midrise Buildings on Soft Soils', International Journal of Geomechanics, vol. 14, no. 4, pp. 1-15. DOI: 10.1061/(ASCE)GM.1943-5622.0000354

Hokmabadi, A.S., Fatahi, B. & Samali, B. 2014, 'Assessment of soil-pile-structure interaction influencing seismic response of mid-rise buildings sitting on floating pile foundations', Computers and Geotechnics, vol. 55, pp. 172-186.

Hokmabadi, A.S., Fatahi, B. & Samali, B. 2014, 'Physical modeling of seismic soil-pile-structure interaction for buildings on soft soils', International Journal of Geomechanics, vol. 15, no. 2, 04014046, DOI: 10.1061/(ASCE)GM.1943-5622.0000396

Hosseinzadeh, N.A. and Nateghi, F. (2004). “Shake table study of soil structure interaction effects on seismic response of single and adjacent buildings”, Proceedings of the13th World Conference on Earthquake Engineering, Vancouver, Canada, pp. 1918.

ITASCA (2014). FLAC3D version 5.01 Fast Lagrangian Analysis of Continua in three dimentions, User's Manual. Minneapolis, Minnesota, USA: Itasca Consulting Group, Inc.

Karamodin, A.K. and Kazemi, H.H. (2008). “Semi-active control of structures using neuro-predictive algorithm for MR dampers”, Structural Control and Health Monitoring, Vol. 17, No. 3, pp. 237–253.

Kramer, S. L. (1996). Geotechnical earthquake engineering, Prentice Hall.

Tabatabaiefar, H.R. & Fatahi, B. 2014, 'Idealisation of soil-structure system to determine inelastic seismic response of mid-rise building frames', Soil Dynamics And Earthquake Engineering, vol. 66, pp. 339-351.

Tabatabaiefar, S., Fatahi, B. & Samali, B. 2013, 'Lateral seismic response of building frames considering dynamic soil-structure interaction effects', Structural Engineering and Mechanics, vol. 45, no. 3, pp. 311-321.

Tabatabaiefar, S.H.R., Fatahi, B. & Samali, B. 2014, 'An empirical relationship to determine lateral seismic response of mid-rise building frames under influence of soil-structure interaction', Structural Design of Tall and Special Buildings, vol. 23, no. 7, pp. 526-548.

Veletsos, A.S. and Meek, J.W. (1974). “Dynamic behaviour of building-foundation system”, Earthquake Engineering & Structural Dynamics, Vol. 3, No. 2, pp. 121–38.

Vucetic, M. and Dobry, R. (1991). “Effects of soil plasticity on cyclic response”, Journal of Geotechnical Engineering, ASCE, Vol. 117, No. 1, pp. 89–107.

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WELCOME MESSAGEOn behalf of the Organising Committee I would like to extend a warm welcome to all participants of the 6th International Conference on Geotechnical Earthquake Engineering (6ICEGE), in Christchurch, New Zealand.

The 6ICEGE is organised by the New Zealand Geotechnical Society (NZGS) under the auspices of the Technical Committee TC203 (Technical Committee on Earthquake Geotechnical Engineering) of the International Society for Soil Mechanics and Geotechnical Engineering (ISSMGE). Following the highly successful conferences in Tokyo 1995, Lisbon 1999, Berkeley 2004, Thessaloniki 2007 and Santiago 2011, we are delighted to host the sixth event in this series of specialised conferences on earthquake geotechnical engineering, here in Christchurch.

The 6ICEGE offers an outstanding technical programme on a range of earthquake geotechnical engineering topics, presented in the published proceedings, and the oral and poster sessions. Following a rigorous review process nearly 400 papers have been accepted and included in the 6ICEGE proceedings. Over 200 papers will be presented in the oral sessions including 33 invited lectures and two special lectures, the 5th Ishihara Lecture and the 2nd Schofield Lecture.

Since the 5ICEGE in Santiago in January 2011, several major earthquakes have occurred throughout the world; the 2010-2011 Christchurch earthquakes, the 2011 Tohoku (Japan) earthquake and more recently the 2015 Nepal earthquake, to mention but a few. These earthquakes caused loss of life, extensive damage and tremendous impacts on people and communities. Christchurch was not an exception in this regard, and geotechnical problems and related damage were signature features of the Christchurch earthquakes. You will witness the tremendous impacts of the earthquakes on Christchurch, but also the impressive efforts in rebuilding the city. Indeed, Christchurch offers an exceptional context and venue for the 6ICEGE.

We believe the 6ICEGE provides an excellent opportunity for earthquake and geotechnical engineers, geologists and seismologists, consulting engineers, public and private contractors, regional and national authorities, and all those involved with engineering works and research related to earthquake geotechnical engineering, to exchange ideas and present their recent experience and developments. It also provides a great opportunity to enjoy New Zealand’s stunning natural and cultural beauty. Our Technical Tours and Accompanying Persons Programmes have been carefully tailored to facilitate your active participation, and we hope that you will find the 6ICEGE professionally rewarding, scientifically stimulating and personally enjoyable.

Thank you for your contributions, and enjoy the conference and New Zealand.

Misko CubrinovskiProfessor, University of Canterbury6ICEGE Chairman

ACKNOWLEDGEMENTSOn behalf of the Organising Committee I would like to acknowledge the seed funding for 6ICEGE provided by the conference partners, the New Zealand Geotechnical Society (NZGS), the Earthquake Commission (EQC), and the University of Canterbury Quake Centre (UCQC). Additionally, I would like to recognise the financial support of the 6ICEGE platinum sponsors, Tonkin and Taylor, and Golder Associates, as well as the numerous gold, silver and bronze sponsors. We appreciate your great support and contribution to the success of the 6ICEGE.

We are thankful to all authors for their high quality contributions. Special thanks are extended to the invited speakers for their excellent contributions and papers, which are the highlight of the 6ICEGE technical programme. The contribution of the numerous reviewers is also acknowledged and greatly appreciated.

Finally, I would like to express my warmest thanks to our postgraduate students for their assistance in various matters associated with this conference, and to the outstanding efforts of the Organising Committee of the 6ICEGE. I particularly appreciate the exceptional efforts and contributions of Charlie Price (NZGS Chair), Brendon Bradley (UC), Liam Wotherspoon (UA) and Mark Stringer (UC), and would like to extend my personal thanks to them.

Misko CubrinovskiProfessor, University of Canterbury6ICEGE Chairman

6TH INTERNATIONAL CONFERENCE ON EARTHQUAKE GEOTECHNICAL ENGINEERING

(6ICEGE)

ORGANISING COMMITTEEMisko Cubrinovski, Conference ChairmanBrendon Bradley, Programme ChairCharlie Price, Conference TreasurerCY ChinJennifer HaskellMarie-Claude HerbertTim McMorranKirsti MurahidyRolando OrenseGreg PrestonRichard SmithMark StringerMerrick TaylorLiam Wotherspoon

REVIEWERSGavin AlexanderJackee AllmondIoannis AnastasopolousClive AndersonRonald AndrusJawad ArefiM A L BakiChris BaxterGeorge BouckovalasRoss BoulangerHayden BowenBrendon BradleyAaron BradshawScott BrandenbergDarren ChianCY ChinUlas CilingirBrady CoxStephen CrawfordMisko CubrinovskiKarina DahlShideh DashtiJason De JongLijun DengJames DismukeAndy DoddsEce Eseller-BayatGeoffrey FarquharSebastiano FotiAndy FourieHideki FunaharaRussell GreenStuart HaighManouchehr HakhamaneshiJennifer HaskellRobert HillierRoman Hryciw

Chris HumphesonSeokho JeongRajibul KarimEdward KavazanjianRichard KellyTakashi KiyotaJonathan KnappettStavroula KontoeCarlo LaiTam LarkinChristian LedezmaWei LiRichard LittonWeian LiuRobert LoDomenico LombardiZiggy LubkowskiGopal MadabhushiTrevor MatuschkaBrett MaurerRory McCullyChris McGannMasoud MohajeriBrina MontoyaShinichiro MoriRamin MotamedTim MoteAlexei MurashevIslam NurulMitsu OkamuraGuney OlgunRolando OrenseJack PappinMichael PenderDimitris PitilakisZillur RabbiMizanur Rahman

Kelly RobinsonAbouzar SadrekarimiEsteban SaezToru SekiguchiJean Francois SemblatKostas SenetakisVlatko SheshovFrancesco SilvestriPaul SomervilleJohn St GeorgeLuke StorieMark StringerArmin StuedleinShuji TamuraMerrick TaylorTetsuo TobitaSelcuk ToprakHirofumi ToyotaGrigoris TsinidisYosimichi TsukamotoTsou-Shin UengRyosuke UzuokaRamon VerdugoAntonio Viana da FonscaManuela VillaniDariusz WanatoskiFrederick WentzDan WilsonClinton WoodLiam WotherspoonSuguru YamadaSusumu YasudaJeremy YeatsNozomu YoshidaKaterina ZiotopoulouJosh Zupan

The 6ICEGE Organising Committee would like to thank the following reviewers for undertaking the

reviews of the papers presented in these Proceedings.

Proceedings (Paper No. 372)

Conference Programme (Poster No. 85)

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Behzad Fatahi

From: 11125996@student.uts.edu.auSent: Tuesday, 1 December 2015 1:53 PMTo: Behzad FatahiSubject: FW: Reformatted Paper

FYI  From: Mark Stringer [mailto:mark.stringer@canterbury.ac.nz] Sent: Monday, 10 August 2015 7:01 AM To: 11125996@student.uts.edu.au; 6icege@tcc.co.nz Cc: behzad.fatahi@uts.edu.au Subject: RE: Reformatted Paper  Dear Ruoshi,   Thank you for the updated version of your paper, which I am pleased to accept.  Mark  From: 11125996@student.uts.edu.au [mailto:11125996@student.uts.edu.au] Sent: Sunday, 9 August 2015 5:37 p.m. To: 6icege@tcc.co.nz; Mark Stringer Cc: behzad.fatahi@uts.edu.au Subject: Reformatted Paper  Dear Mark,  Thank you for your email.  Please find the attachment for the reformatted paper.   Best Regards,  Ruoshi  This email may be confidential and subject to legal privilege, it may

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not reflect the views of the University of Canterbury, and it is not guaranteed to be virus free. If you are not an intended recipient, please notify the sender immediately and erase all copies of the message and any attachments. Please refer to http://www.canterbury.ac.nz/emaildisclaimer for more information.

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Behzad Fatahi

From: 11125996@student.uts.edu.auSent: Tuesday, 1 December 2015 1:50 PMTo: Behzad FatahiSubject: FW: 6ICEGE Review

Dear Behzad,  This following email is the review of the paper.  Ruoshi  From: Mark Stringer [mailto:mark.stringer@canterbury.ac.nz] Sent: Saturday, 20 June 2015 9:42 PM To: ruoshi.xu@student.uts.edu.au Subject: 6ICEGE Review

Dear Mr Xu,

We have recently concluded the review process for the 6ICEGE, and I am writing to you regarding your submission “Three dimensional numerical analysis of seismic soil-structure interaction considering soil plasticity.”

I am pleased to inform you that your paper is recommended for publication in the proceedings, subject to a small number of revisions.

- There are some errors in the table and figure numbers in the paper – please check all of the references.

- The reviewer highlighted that Figure 1 has been taken directly from Kramer 1996. If this is correct, then please include the reference in the figure caption.

- I think it would be useful to explain the normalization of shear stress in equation 4.

- On page 5, you refer to second moment of inertia (note that there is a spelling error on inertia) – do you actually mean second moment of area?

- On page 6, paragraph 3, when discussing the free field boundaries. Can you please include a reference for the technique.

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- Do the plots in figure 5 relate to a single time instant?

- Page 7 line 6: “mode” not “model.” J

- Page 7 line 9: there is a word missing between “variation” and “the”

- Page 7: you mention that the second mode of the structure might be being activated. However, this doesn’t seem consistent with figure 5, where you should presumably see the mode shape in some way?

- Page 7 end of line 10: please explain how the variation in damping is dominating the results.

- Please can you carry out these changes, and then resubmit your paper in PDF form to the secretariat at 6icege@tcc.co.nz, and copy me on the email, before 15th July.

When submitting the revised paper, please ensure that there are no page numbers on the manuscript as these will be added in later. Best wishes, Mark Stringer Post-Doctoral Fellow Department of Civil and Natural Resources University of Canterbury

This email may be confidential and subject to legal privilege, it may not reflect the views of the University of Canterbury, and it is not guaranteed to be virus free. If you are not an intended recipient, please notify the sender immediately and erase all copies of the message and any attachments. Please refer to http://www.canterbury.ac.nz/emaildisclaimer for more information.