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Soil Dynamics and Earthquake Engineering 27 (2007) 324–332

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Shaking table testing of geofoam seismic buffers

Richard J. Bathursta,�, Saman Zarnanib, Andrew Gaskinc

aGeoEngineering Centre at Queen’s-RMC, Department of Civil Engineering, 13 General Crerar, Sawyer Building, Room 2085,

Royal Military College of Canada, Kingston, Ont., Canada K7K 7B4bGeoEngineering Centre at Queen’s-RMC, Department of Civil Engineering, Queen’s University, Kingston, Ont., Canada K7L 3N6

cOntario Power Generation Inc., Bowmanville, Ont., Canada L1C 3Z8

Received 25 August 2006; accepted 31 August 2006

Abstract

The paper describes the experimental design and results of tests used to investigate the use of compressible EPS (geofoam) seismic

buffers to attenuate dynamic loads against rigid retaining wall structures. The tests were carried out using 1-m-high models mounted on a

large shaking table. Three different geofoam buffer materials retaining a sand soil were tested under idealized dynamic loading

conditions. The results of these tests are compared to a nominal identical structure without a seismic buffer. The test results demonstrate

that the reduction in dynamic load increased with decreasing seismic buffer density. For the best case reported here, the maximum

dynamic force reduction was 31% at a peak base acceleration of 0.7g.

r 2006 Elsevier Ltd. All rights reserved.

Keywords: Shaking table; Geofoam; Seismic; Buffer; Retaining wall; Expanded polystyrene (EPS)

1. Introduction

The use of vertical compressible layers placed againstrigid soil retaining wall structures to reduce lateral staticearth pressures has been reported in the literature bydifferent researchers [1–3]. The product of choice toperform this function is expanded polystyrene (EPS),which is called geofoam in modern geosynthetics nomen-clature [4]. A logical extension of this application is to usevertical compressible inclusions of EPS geofoam as seismicbuffers to attenuate earthquake-induced dynamic earthpressures against rigid walls [5]. Nevertheless, proof ofconcept through experimental testing is lacking. This paperdescribes the experimental design and results of tests usedto investigate the use of compressible expanded polystyrene(geofoam) seismic buffers to attenuate dynamic loadsagainst rigid retaining wall structures. The tests werecarried out using 1-m high models mounted on a largeshaking table. Three different geofoam buffer materials

e front matter r 2006 Elsevier Ltd. All rights reserved.

ildyn.2006.08.003

ing author. Tel.: +1613 541 6000x6479/6347/6391;

8336.

ess: bathurst-r@rmc.ca (R.J. Bathurst).

retaining a sand soil were tested under idealized dynamicloading conditions. The results of these tests are comparedto a nominal identical structure without a seismic buffer.The test results demonstrate that the reduction in dynamicload increased with decreasing seismic buffer density. Forthe best case reported here, the maximum force reductionwas 31% at a peak base acceleration of 0.7g. Somepreliminary results of this program have been reported byZarnani et al. [6]. However, in the current paper, details ofthe experimental design and examples of test results notpreviously reported are described.

2. Shaking table

The physical tests described in this paper were carriedout on a 2.7m� 2.7m shaking table located at the RoyalMilitary College of Canada (RMC). The table has a 5-tonne payload capacity and is driven by a 100-kN capacityMTS hydraulic actuator with ancillary controller and PCsoftware. The table is driven in the horizontal directiononly. However, the horizontal component of seismic-induced dynamic earth loading is typically the most

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Fig. 1. Cross-section view of general test arrangement and instrumenta-

tion.

Fig. 2. Particle size distribution for sand backfill.

Fig. 3. Results of direct shear box tests on sand backfill.

R.J. Bathurst et al. / Soil Dynamics and Earthquake Engineering 27 (2007) 324–332 325

important loading direction for the application underinvestigation. The table can excite the rated payload atfrequencies up to 13Hz and 72g. The maximum displace-ment of the actuator is 775mm.

The wall models were built in a stiff strong box (2.5mlong� 1.4m wide and 1.3m high) bolted to the steelplatform of the shaking table. A cross-section view of thegeneral test arrangement is illustrated in Fig. 1. The backfillextended 2m beyond the model wall at the front of thestrong box. This dimension was selected to preventintersection with a potential active failure mechanismcomputed using a Mononobe–Okabe wedge analysis. Thesidewalls of the strong box are comprised of 13-mm-thickPlexiglas sheets that are laterally braced by steel frames.The inside surface of the Plexiglas is covered by twolubricated layers of clear 6-mil polyethylene film. Thecombination of friction-reducing membrane and rigidlateral bracing was adopted to ensure that the test modelswere subjected to plane strain boundary conditions. Aplywood sheet was bolted to the top of the shaking table. Alayer of sand was epoxied to the plywood surface to createa rough surface with a friction angle equal to that of thesand backfill. The back of the box was rigidly braced. Thisis an energy-reflecting boundary that was adopted forsimplicity.

3. Materials

3.1. Backfill

A dry synthetic olivine sand was used as the backfill inthe test configurations. This material was selected becauseit is silica free and thus avoided the health danger of silicadust generation during material handling in an enclosedlaboratory environment. The sand is composed of angularto sub-angular particles with a specific gravity of 2.88. Thesoil is uniformly graded with a maximum particle size of2mm, coefficient of curvature of 1.27, coefficient ofuniformity of 2.5 and fines content less than 3% by weight.The particle size distribution is shown in Fig. 2. Conven-tional direct shear box tests [7] were carried out on

10 cm� 10 cm specimens of the sand at the same density asthe sand in the shaking table models. The results of thesetests are shown in Fig. 3. Sand properties are summarizedin Table 1.

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Table 1

Soil properties

Property Value

Density 15.5Mg/m3

Peak angle of friction 511

Residual friction angle 461

Cohesion 0

Relative density 86%

Dilation angle 151

Note: Strength parameters from 10 cm� 10cm in plan area direct shear

tests.

Table 2

Initial tangent elastic modulus for non-elasticized EPS based on

correlations with density

Reference Initial tangent elastic modulus, Ei (MPa)

Type I Type XI

r ¼ 16 kg/m3 r ¼ 12 kg/m3

Hazarika [9] 3.76 2.12

Missirlis et al. [10] 4.55 2.98

Negussey [11] 3.30 1.80

Negussey and Anasthas [12] 8.22 4.94

Obrien [13] 9.17 6.02

Anasthas et al. [14] 4.48 2.98

Duskov [15] 5.73 5.48

Duskov [16] 5.7 3.84

Negussey and Sun [17] 3.00 1.29

Horvath [4] 4.20 2.40

Eriksson and Trank [18] 4.79 2.87

Magnan and Serratrice [19] 4.06 3.03

Average 5.08 3.31

Standard deviation 1.89 1.48

Note: r ¼ density.

Table 3

Unmodified EPS geofoam properties

Property EPS type and density

Type I Elasticized Type XI

16 kg/m3 14 kg/m3 12 kg/m3

Elastic modulus, Eia(MPa) 4.7 0.27 3.1 (1.6)b

Poissions ratioc 0.09 0.08 0.07

Yield strength (kPa)c 30.4 23.6 16.8

Note:afrom manufacturers’ literature.bWith 50% of material removed.cAssumed using correlations for non-elasticized EPS by Horvath [4].

R.J. Bathurst et al. / Soil Dynamics and Earthquake Engineering 27 (2007) 324–332326

3.2. Geofoam

Three commercially available EPS materials wereselected in this investigation. The first product was a TypeI material according to the ASTM classification system [8]with a density of 16 kg/m3. The second product was anelasticized material with a density of 14 kg/m3. Thisproduct is produced from solid EPS block that is subjectedto a cycle of compression load–unloading in order toincrease the linear elastic range of the material behavior.The third material was a modified Type XI materialaccording to the ASTM classification system [8]. Thegeofoam inclusions were constructed from three 50-mm-thick sheets of EPS glued together. The Type XI materialwas modified by removing 57% of the EPS foam from twoof the sheets by circular coring using a 80-mm-diameterhole saw. The coring was done on a hexagonal close–pack-ing pattern. The third sheet was only partially cored fromone side so that the surface of the seismic buffer wascontinuous at the soil–geofoam interface in the testconfiguration. Coring reduced the density of the geofoamfrom 12 to 6 kg/m3. This material was modified in order toinvestigate the influence on wall response of geofoamdensity and modulus values that are significantly lowerthan those of the two unmodified geofoam materials.

The density of EPS is used as a characteristic property ofdifferent commercially available products [8]. However, themechanical property of interest when these products areused as compressible inclusions to attenuate lateral earthpressures is the elastic modulus. A large number ofcorrelations are reported in the literature to equate theinitial elastic modulus of EPS to product density. Theresults of these correlations for the Type I and XI materialsused in this investigation are summarized in Table 2. Therange of values is due to different methods of testing,specimen size, shape and rate of loading. Nevertheless,non-elasticized EPS products are linear elastic up to 1%strain. Elasticized EPS materials have a linear elastic rangeup to �40% strain but have a lower elastic modulus [4].Properties of the three materials used in this investigationare summarized in Table 3. The elastic moduli reported inthis table are based on the manufacturers’ literature. Thevalues for Poisson’s ratio and yield strength are based oncorrelations with density for non-elasticized EPS reported

by Horvath [4]. These values were required duringexperimental design to ensure that the combination ofelastic properties and thickness of the compressibleinclusion for the geofoam configurations could be expectedto reduce lateral earth pressures in the physical experi-ments. This experimental design stage was carried outusing a numerical model based on a FLAC code [20].Examples of FLAC numerical modeling of the testsdescribed in the current paper have been reported byZarnani and Bathurst [21]. The back-calculated values ofelastic modulus for the three products in Table 4 arediscussed later in the paper.

4. Test methodology

4.1. General arrangement

A total of seven test configurations were investigated.Four tests that capture the range of response in this

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Table 4

Wall model configurations

Wall Geofoam

Thickness Type Density Initial tangent elastic modulus,

(mm) (ASTM-C578-06) (kg/m3) Ei (MPa)a

1 0 N/A

(control)

2 150 I 16 4.8

4 150 Elasticized 14 1.3

6 150 XI 6b (50% removed by coring) 0.6

Note:aBack-calculated values from stress-strain load cycles.bOriginal density of EPS geofoam was 12 kg/m3.

Fig. 4. Placement of 1-m-high by 1.4-m-wide geofoam seismic buffer

against rigid aluminum bulkhead.

Fig. 5. Final placement of sand backfill against geofoam seismic buffer.

R.J. Bathurst et al. / Soil Dynamics and Earthquake Engineering 27 (2007) 324–332 327

experimental program are summarized in Table 4. Wall 1was the control test configuration with no geofoam seismicbuffer. This wall was constructed using a 1-m-high by 1.4-m-wide rigid aluminum bulkhead. The three walls with ageofoam seismic buffer were constructed with the config-uration illustrated in Fig. 1. The geofoam seismic bufferwas placed between the aluminum wall and the sandbackfill. Each wall was 1m in height and each seismicbuffer was 150mm thick. Hence, the only variable betweenthe seismic buffer tests was the geofoam properties.

The aluminum bulkhead was supported laterally by fiveload cells rigidly braced to the shaking table platform. Thebulkhead (and seismic buffer) was seated on an instru-mented footing supported by three frictionless linearbearings. The linear bearings were seated on five loadcells. The footing arrangement allowed the vertical andhorizontal load measurements at the wall boundaries to befully decoupled.

4.2. Construction

The seismic buffer was held in place temporarily usingduct tape (Fig. 4). The sand was placed in five 200-mm-thick lifts and manually leveled. Each lift was vibro-compacted by gently shaking the table for 90 s at 9Hz(peak base acceleration ¼ 0.1g). The final placement of thesand backfill is shown in the photograph in Fig. 5. The finaldensity of the tests was 15.5Mg/m3 as measured using anuclear density meter.

4.3. Instrumentation

The test models were instrumented with load cells tomeasure the boundary loads on the aluminum bulkhead asmentioned earlier. The load cells were button-type (Hon-eywell Sensotec, compression load cell model 53). Fourdisplacement potentiometers (Data Instruments) were usedto measure lateral deformations at the geofoam–soilinterface. The potentiometers were mounted directly onthe aluminum bulkhead and the cores passed through a 20-mm-diameter opening through the bulkhead and thegeofoam buffer to aluminum plates inserted flush with

the buffer surface (Fig. 4). An additional potentiometerwas mounted against the shaking table to measure thedisplacement-time history of the model base. An accel-erometer was also mounted directly on the shaking table to

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record the input base acceleration–time excitation history.Four other accelerometers (Schaevitz S05E Silicon modeltype) were embedded in the backfill soil at the locationsshown in Fig. 1. The accelerometers were anchored withina 100-mm-diameter by 70-mm-deep section of PVC pipe toensure that the devices remained level and moved in phasewith the surrounding sand during shaking.

The instruments were monitored by a separate high-speed data acquisition system (SERIES-7000 from Sym-metric Instruments Inc. and using SYSTEM 200 software).Data from a total of 19 instruments were recorded at aspeed of about 100Hz in order to prevent aliasing and tocapture peak response values.

Fig. 7. Two-second accelerogram window at amplitude step.

Fig. 8. Unfiltered measured base input accelerogram frequency distribu-

tion.

4.4. Model excitation

Following construction, the models were excited using adisplacement–time history selected to match a targetstepped-amplitude sinusoidal accelerogram with a fre-quency of 5Hz (Fig. 6). A 2-s window of the targetaccelerogram is illustrated in Fig. 7. The accelerationrecord was stepped in 0.05g increments and each amplitudeincrement was held for 5 s. The maximum base accelerationwas 0.8g.

A 5Hz frequency (i.e. 0.2 s period) at 1/6 model scalecorresponds to 2Hz (i.e. 0.5 s period) at prototype scaleaccording to the scaling laws proposed by Iai [22].Frequencies of 2–3Hz are representative of typicalpredominant frequencies of medium-to high-frequencyearthquakes [23] and fall within the expected earthquakeparameters for North American seismic design [24,25]. Thissimple base excitation record is more aggressive than anequivalent true earthquake record with the same predomi-nant frequency and amplitude [23,26]. However, it allowedall walls to be excited in the same controlled manner andthis allowed valid quantitative comparisons to be madebetween different wall configurations. As noted earlier inthe paper, the models were only excited in the horizontalcross-plane direction to be consistent with the criticalorientation typically assumed for seismic design of earthretaining walls [25].

Fig. 6. Target base excit

As a check on the input base acceleration, frequencyanalysis of the entire measured accelerogram was carriedout and this analysis is shown in Fig. 8. The measuredpredominant frequency of the unfiltered input accelero-gram can be seen to be close to a target value of 5Hz.

ation accelerogram.

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Fig. 9. Measured dynamic compressive stress–strain load cycles.

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5. Test results

The dynamic compressive stress–strain responses of thethree geofoam layers used in the models are shown inFig. 9. The hysteretic response curves have been calculatedusing the forces recorded by the lateral load cellssupporting the aluminum bulkhead and strains computedfrom the displacement potentiometers attached to thebuffer surface. The measurements are taken with respect toend of static construction. The strain amplitude for Tests 2and 4 is within the linear elastic range cited in the literature(1% strain). Test 6 with 50% material removed recorded astrain amplitude of about 2.2%. Hence, this material islikely to have exceeded the elastic limit of the material.

The relative dynamic compressibility of the threegeofoam materials can be compared by taking the slopeof the lines passing through the origin of the plots andparallel to the long axis of the hysteresis loops. The initialcompressibility of the three geofoam buffers in thisinvestigation can be seen to vary through about one orderof magnitude (e.g. Ei ¼ 4.8, 1.3 and 0.6MPa).

For an ideal linear elastic material, the load–unloadresponse curves would fall on a line with slope equal to thelinear elastic modulus of the medium. For Tests 2 and 4,the slope can be taken as an approximation to the linearinitial tangent elastic modulus and the values for these linestaken from Fig. 9 are reported in Table 4. The back-calculated elastic modulus for Test 2 falls within the rangeof values reported in the literature (Table 2). For Test 6,the linear elastic modulus is taken from the axis slope of theloops for compressive strains over the range of 0–0.5%(Ei ¼ 0.6MPa). The linear approximation to the hysteresiscurves with larger amplitudes can be approximated by areduced modulus value (e.g. Ep ¼ 0:4MPa). The discre-pancies between manufacturers’ literature and back-calcu-lated values for the initial elastic modulus of the EPSmaterials used in Tests 4 and 6 are likely due to theinfluence of specimen size and rate of loading noted earlier.Finally, there are more advanced models available tocharacterize hysteretic stress–strain loops for polymericmaterials [27,28]. However, the simple descriptorsused here are adequate to differentiate between thestiffness properties of the geofoam materials used in thisinvestigation.

Four accelerometers were placed in the backfill soil asdescribed earlier. The frequency response from thesedevices is shown in Fig. 10 for Test 2. Comparison withFig. 8 shows that the predominant frequency of the backfillwas constant in this experiment but that there wasacceleration amplification through the backfill. For exam-ple the maximum ratio of the Fast Fourier Transform(FFT) power values is 1.6.

The experimental design included the measurement ofvertical toe forces (Fig. 1). Fig. 11 shows the peak verticalforce–time response of the four walls in this study.The horizontal line in the figure is the self-weight of thealuminum (rigid) panel wall. The vertical load for the

control test is greater than the wall self-weight duringexcitation which may be expected assuming some interfacefriction between the retained sand soil and the panel face.The difference in self-weight of the walls with and withouta geofoam buffer is negligibly small due to the low densityof the geofoam material. The data in Fig. 11 show that asthe stiffness of the geofoam buffer material decreases, the

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Fig. 10. Acceleration frequency content and amplification in soil backfill during base excitation (Test 2).

Fig. 11. Peak vertical wall force versus time.

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vertical toe forces generally increase. This may be due togreater penetration of the sand particles into the less stiffpolystyrene materials.Fig. 12 shows the dynamic compression–time response

of the three walls with a geofoam seismic buffer. The datahave been plotted to show the average, minimumand maximum horizontal deformations taken at thesame time. The data in the figures shows that as themodulus of the geofoam decreases, the magnitude ofaverage horizontal compression increases. The lowerbound on the shaded response curves for Walls 2 and 4corresponds to measurements recorded by the lowestelevation displacement potentiometer, and the upperbound to the highest elevation displacement potentiometer.Hence, the lateral deformations over the height of the wallwere in phase during dynamic excitation. For Wall 6,however, the reverse was true. Hence the deformations

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Fig. 12. Dynamic compression–time response of seismic buffer models.

Fig. 13. Peak horizontal wall force versus peak base acceleration.

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were not in phase due to the reduced stiffness of thegeofoam.

A primary objective of this study was to demonstrateproof of concept by quantifying the reduction of dynamicearth forces against a rigid wall due to the inclusion of ageofoam seismic buffer. Fig. 13 shows the peak horizontalforce recorded at the end of construction and duringdynamic loading versus the measured peak base accelera-

tion value during base shaking. The structures with ageofoam inclusion generated less horizontal load than therigid structure during base shaking. The greatest reductionoccurred for Test 6, which was constructed with the lowestdensity geofoam inclusion and the lowest stiffness. Thereduction in lateral earth forces from the rigid wall case at acommon peak base acceleration of 0.7g was 16, 20 and31% for Tests 2, 4 and 6, respectively.

6. Conclusions

This paper describes the details of an experimentalprogram that was undertaken to examine quantitatively theconcept of geofoam compressible inclusions to reduce themagnitude of earthquake-induced dynamic forces againstrigid earth retaining wall structures. The experimentaldesign and test methodology using a shaking table showsthat large detectable differences in dynamic force reductionwere observed between rigid walls with and without ageofoam seismic buffer. The results presented heredemonstrate proof of concept and are a valuable set ofresults that can be used to verify numerical codes [21,29].Numerical codes can then be used to investigate a widerrange of problem geometry, soil conditions and geofoamthickness and stiffness with a view to optimizing thesesystems for site-specific earthquake design.

Acknowledgements

The authors are grateful for funding provided by theNatural Sciences and Engineering Research Council ofCanada, the Academic Research Program at RMC, andgrants from the Department of National Defence (Cana-da). The many discussions with our colleagues K. Hatamiand M. El-Emam during the preparation of this paper arealso gratefully acknowledged. Finally, the authors thankD. Van Wagoner of Geotech Systems Corporation forsupplying the EPS material and associated technical data.

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References

[1] Partos AM, Kazaniwsky PM. Geoboard reduces lateral earth

pressures. In: Proceedings of Geosynthetics’87, Industrial Fabrics

Association International. New Orleans, LA, USA, 1987. p. 628–39.

[2] Horvath JS. Compressible inclusion function of EPS geofoam.

Geotextiles and Geomembranes 1997;15(1–3):77–120.

[3] Karpurapu R, Bathurst RJ. Numerical investigation of controlled

yielding of soil-retaining wall structures. Geotextiles and Geomem-

branes 1992;11:115–31.

[4] Horvath JS. Geofoam Geosynthetic. Scarsdale: Horvath Engineer-

ing, P.C.; 1995.

[5] Inglis D, Macleod G, Naesgaard E, Zergoun M. Basement wall with

seismic earth pressures and novel expanded polystyrene foam buffer

layer. In: Proceedings of the 10th annual symposium of the

Vancouver geotechnical society. Vancouver, BC, Canada, 1996, 18pp.

[6] Zarnani S, Bathurst RJ, Gaskin A. Experimental investigation of

geofoam seismic buffer using a shaking table. In: Proceedings of the

North American geosynthetics society (NAGS)/GRI19 conference,

Las Vegas, NV, USA, 2005, 11pp.

[7] ASTM D 3080-04. Standard Test Method for Direct Shear Test of

Soils Under Consolidated Drained Conditions. American Society for

Testing and Materials, Philadelphia, Pennsylvania, USA, 2004.

[8] ASTM C 578-06. Standard Specification for Rigid Cellular Poly-

styrene Thermal Insulation. American Society for Testing and

Materials, West Conshohocken, Pennsylvania, USA, 2006.

[9] Hazarika H. Stress-strain modeling of EPS geofoam for large-strain

applications. Geotextiles and Geomembranes 2006;24(2):79–90.

[10] Missirlis EG, Atmatzidis DK, Chrysikos DA. Compressive creep

behavior of EPS geofoam. In: Proceedings of the 3rd European

Geosynthetics Conference. Munich, Germany, 2004, vol. 2, pp. 749–54.

[11] Negussey D. Design Parameters for EPS Geofoam. (Keynote paper).

In: Proceedings of the international workshop on lightweight Geo-

materials, Tokyo, Japan, 2002, 19pp.

[12] Negussey D, Anasthas N. Young’s modulus of EPS geofoam by

simple bending test. In: Proceedings of the 3rd international

conference of EPS Geofoam, Salt Lake City, Utah, USA, 2001, 14pp.

[13] O’Brien AS. EPS behavior during static and cyclic loading from

0.05% strain to failure. In: Proceedings of the 3rd international

conference of EPS Geofoam, Salt Lake City, Utah, USA, 2001, 11pp.

[14] Anasthas N, Negussey D, Srirajan, S. Effect of confining stress on

compressive strength of EPS geofoam. In: Proceedings of the 3rd

international conference of EPS Geofoam, Salt Lake City, Utah,

USA, 2001. 14pp.

[15] Duskov M. Materials research on EPS20 and EPS15 under

representative conditions in pavement structures. Geotextiles and

Geomembranes 1997;15(1–3):147–81.

[16] Duskov M. EPS as a light weight sub-base material in pavement

structures. PhD. thesis, Delft University of Technology, Delft, the

Netherlands, 1997.

[17] Negussey D, Sun MC. Reducing lateral pressure by geofoam (EPS)

substitution. In: Proceedings of the international symposium on EPS

construction method (EPS). Tokyo, Japan, 1996, pp. 202–11.

[18] Eriksson L, Trank R. Properties of expanded polystyrene—labora-

tory experiments, expended polystyrene as light fill material; technical

visit around Stockholm June 19, 1991, Swedish Geotechnical

Institute, Linkoping, Sweden, 1991.

[19] Magnan JP, Serratrice JF. Proprietes mecaniques du polystyrene

expanse pour ses applications en remblai routier, Bulletin Liaison

Laboratoire Ponts et Chaussees, LCPC, No. 164, 1989, pp. 25–31.

[20] Itasca Consulting Group. FLAC: Fast Lagrangian Analysis of

Continua, version 5. Itasca Consulting Group, Inc., Minneapolis,

Minnesota, USA, 2005.

[21] Zarnani S, Bathurst RJ. Numerical investigation of geofoam seismic

buffers using FLAC. In: Proceedings of the North American

geosynthetics society (NAGS)/GRI19 conference, Las Vegas, NV,

USA, 2005, 8pp.

[22] Iai S. Similitude for shaking table tests on soil-structure-fluid model

in 1 g gravitational field. Soils and Foundations 1989;29:105–18.

[23] Bathurst RJ, Hatami K. Seismic response analysis of a geosynthetic

reinforced soil retaining wall. Geosynthetics Int 1998;5(1&2):127–66.

[24] NBCC. National Building Code of Canada, National Research

Council of Canada, Ottawa, Ontario, Canada, 1990.

[25] AASHTO. Standard Specifications for Highway Bridges, American

Association of State Highway and Transportation Officials, Seven-

teenth Edition, Washington, DC, USA, 2002

[26] Matsuo O, Tsutsumi T, Yokoyama K, Saito Y. Shaking table tests

and analysis of geosynthetic-reinforced soil retaining walls. Geosyn-

thetics Int 1998;5(1&2):97–126.

[27] Bathurst RJ, Hatami K, Alfaro MC. Geosynthetics and Their

Applications, Geosynthetic-reinforced soil walls and slopes—seismic

aspects, Chapter 14. Thomas Telford, 2002.

[28] Yogendrakumar M, Bathurst RJ. Numerical simulation of reinforced

soil structures during blast loadings. Transport Res Rec

1992;1336:1–8.

[29] Bathurst RJ, Keshavarz A, Zarnani S, Take A. A simple displace-

ment model for response analysis of EPS geofoam seismic buffers.

Soil Dy Earthq Eng, in press.