[seid] seismic liquefaction

Seismic liquefaction, lateral spreading, and flowslides: a numerical investigation into voidredistribution

Mahmood Seid-Karbasi and Peter M. Byrne

Abstract: Experience from past earthquakes indicates that seismically induced large lateral spreads and flow slides inalluvial sand deposits have taken place in coastal and river areas in many parts of the world. The ground slope in theseslides was often not very steep, gentler than a few percent. Recent research indicates that the presence of low-permeabilitysilt or clay sublayers within the sand deposits is responsible for this behaviour. Such layers form a barrier to upwardflow of water associated with earthquake-generated pore pressures. This causes an accumulation of pore water at thebase of the layers, resulting in greatly reduced strength and possible slope instability. This paper uses an effectivestress coupled stress-flow dynamic analyses procedure to demonstrate the effects of a low-permeability barrier layer onground deformations from an earthquake event. The analyses show that an expansion zone develops at the base of barrierlayers in stratified soil deposits under seismic loading which can greatly reduce shear strength and result in largedeformations and flow failure. Without such a layer or layers, the slope may undergo significant displacements, but nota flow slide. Slopes with a barrier layer can be stabilized by drains.

Key words: liquefaction, lateral spreads, stratification, flow failure, dynamic analysis, UBCSAND model, drain.

Résumé : L’expérience antérieure des tremblements de terre indique que de grands étalements latéraux et des couléesdans les dépôts de sable alluvionnaires se produisent dans des régions côtières et de rivières dans plusieurs parties dumonde. La pente du terrain dans ces glissements n’est souvent pas très abrupte, plus douce que quelques pour cent. Larecherche récente indique que la présence de sous-couches de silt ou d’argile de faible perméabilité dans les dépôts desable est responsable de ce comportement. De tels couches forment une barrière pour l’écoulement de l’eau vers lehaut associé avec les pressions interstitielles générées par le tremblement de terre. Ceci cause une accumulation de pressioninterstitielle à la base des couches, et résulte en une résistance grandement diminuée et une instabilité potentielle despentes. Cet article utilise une procédure d’analyses en contrainte effective couplée à la dynamique contrainte-écoulementpour montrer les effets d’une couche-barrière de faible perméabilité sur la déformation causée par un événement detremblement de terre. Les analyses montrent qu’une zone d’expansion se développe à la base des couches-barrières dansles dépôts de sol stratifié sous un chargement séismique qui peut réduire grandement la résistance au cisaillement etrésulter en de grandes déformations et en une rupture par écoulement. Sans une telle couche ou de telles couches, lapente peut subir des déplacements appréciables, mais pas un glissement par écoulement. Les pentes avec des couches-barrières peuvent être stabilisées par drains.

Mots-clés : liquéfaction, étalements latéraux, stratification, rupture par écoulement, analyse dynamique, modèleUBCSAND, drain.

[Traduit par la Rédaction] Seid-Karbasi and Byrne 890

Introduction

Experience from past earthquakes indicates that lateralspreads and flow slides have taken place in liquefied ground incoastal and river areas in many regions of the world, includingAlaska, Japan (Niigata), and Turkey. Movements may exceedseveral metres, even in gentle slopes of less than a few percent

(Kokusho 2003). Submarine slides have been seismically trig-gered in many regions as reported by Scott and Zuckerman(1972) and Hamada (1992). More interestingly, lateral spreadsor flow slides have occurred not only during earthquake shak-ing, but also after earthquake shaking. These large movementsare mainly driven by gravity, although they are initially trig-gered to liquefy by seismic stresses.

In the Hakusan District in Niigata City, Japan, an area250 m × 150 m bordering the Shinano River moved toward theriver during the 1964 Niigata earthquake (Kawakami andAsada 1966). Girders of the Showa Ohashi Bridge fell downdue to flow failure of the liquefied riverbed, which moved thesupporting pile foundations. Eyewitnesses reported that thegirders began to fall a few minutes after the earthquake motionhad ceased (Hamada 1992). At the lower San Fernando dam, ahydraulic fill in California, flow failure due to liquefaction inthe upstream slope occurred about a minute after the strong

Can. Geotech. J. 44: 873–890 (2007) doi:10.1139/T07-027 © 2007 NRC Canada

873

Received 3 November 2004. Accepted 29 January 2007.Published on the NRC Research Press Web site at cgj.nrc.caon 14 August 2007.

M. Seid-Karbasi1,2 and P.M. Byrne. Department of CivilEngineering, University of British Columbia, 6250 AppliedScience Lane, Vancouver, BC V6T 1Z4, Canada.

1Corresponding author (e-mail: [email protected]).2Present address: Golder Associates Ltd., 500-4260 StillCreek Drive, Burnaby, BC V5C 6C6, Canada.

shaking in the 1971 earthquake (Seed 1987). Mochikoshi tail-ings dams in Japan failed as a result of the 1978 Izu-Ohshim-Kinkai earthquake due to liquefaction-induced flow slides,causing release of the tailings (Ishihara 1984). One dam failedduring the shaking, and a second dam failed 24 h later. Figure 1shows a sand boil formed in the Mochikoshi disposal pondcomprising sandy silt and silt layers. More recently, Harder andStewart (1996) reported that in the 1994 Northridge, California,earthquake, liquefaction-induced flow failure of the 24 m highTapo Canyon tailings dam occurred about 10 min after themain shock. Yoshida et al. (2005) reported large lateraldisplacements in the order of several metres that occurred innear-level ground conditions in the 1983 Nihonkai–Chubuearthquake, with the movements obtained by comparing aerialphotographs taken before and after the earthquake.

Flow failures in liquefied soil deposits have also occurredin submarine slopes triggered by earthquakes. The slopes inthese slides were gentle, normally less than 5° and some-times less than 1° (Hampton and Lee 1996). During the1964 Alaska earthquake, the port cities of Valdez andSeward suffered great loss of life and property from large-scale submarine slides (Coulter and Migliaccio 1966; Lemke1967). In the 1999 Kocaeli earthquake, Turkey, widespreadlarge lateral spreads are reported by Kokusho (2003). Coastalareas along Izmit Bay suffered submarine slides triggered bythe earthquake. The 1995 Aegion earthquake of magnitude(Ms) 6.2 caused intense ground failures in Eratini, Greece(Bouckovalas et al. 1999). Geotechnical exploration revealedthat the soil profile at the failure sites is characterized by acontinuous interchange of silty sand and clay layers. Theground inclination in these cases was less than 7°. Duringthe 1980 Mammoth lakes earthquake in California, a 2 km ×20 km area of seafloor, consisting of sand and mud on aslope of 0.25°, suffered a flow slide. There was evidence ofliquefaction in the form of sand boils on the seafloor asdescribed by Field et al. (1982).

Although lateral flow failures have been reported in pastearthquakes causing damage to structures (e.g., Berrill et al.1997), the mechanism leading to large lateral displacementsis still poorly understood. Sand and silt deposits often com-

prise many sublayers as a result of the sedimentation pro-cess. A number of researchers have examined the effect oflayering on post-liquefaction sliding, including Scott andZuckerman (1972), Huishan and Taiping (1984), Liu andQiao (1984), Elgamal et al. (1989), Adalier and Elgamal(1992), Fiegel and Kutter (1992), Kokusho (1999, 2000),Kulasingam et al. (2001, 2004), Malvick et al. (2002, 2005,2006), Yang and Elgamal (2002), Kulasingam (2003), Seid-Karbasi and Byrne (2004a), Malvick (2005), Byrne et al.(2006), and Seid-Karbasi3. Based on physical model testsand site investigations, Kokusho (1999) and Kokusho andKojima (2002) concluded that failure can be caused by theformation of a water film at the base of a sublayer leading toa zone of essentially zero strength. The water-film effect isassociated with upward flow of water arising from liquefac-tion. When such flow is impeded by a lower permeabilitylayer, it results in a void redistribution and an accumulationof pore water that can lead to the formation of a water filmor a thin water-rich zone.

Kulasingam et al. (2004) using centrifuge testing demon-strated that water films are an extreme case of void redistri-bution but are not a requisite for large ground deformation tooccur. Under postseismic conditions, the void expansionmay cause the strength to drop significantly and becomesmaller than the driving stress, thereby resulting in largeground deformations. The permeability contrast to causesuch an effect can be associated with silt and (or) clay layerssandwiched within sand and (or) gravel layers. Layering andassociated void ratio change during and or after an earth-quake can explain why the steady-state strength approachbased on an undrained condition (Poulos et al. 1985) canlead to significantly higher values of residual strength thanthose estimated from back analysis of field case studies(e.g., Seed and Harder 1990; Stark and Mesri 1992; Olsonand Stark 2002; among others). This has already beenpointed out by a number of researchers (Seed 1987; Castroet al. 1989; Byrne and Beaty 1997; Kokusho 2003; Seid-Karbasi and Byrne 2004a; Byrne et al. 2006).

Our understanding of the behaviour of liquefiable soilshas increased dramatically in the past 30 years because of(i) observations from field case histories, (ii) extensive labo-ratory testing of soil elements under monotonic and cyclicloading, (iii) model testing of earth structures under simu-lated earthquake loading, and (iv) development of numericalmodeling procedures.

Laboratory model testing suggests that slopes comprisedof clean sands and having relative densities, Dr, as loose as20% are unlikely to suffer a flow slide (Kulasingam et al.2004). These slopes can be readily triggered to liquefy andmay undergo large displacements, but their strengths aregenerally adequate to prevent a flow slide. However, if thesands contain low-permeability silt layers that impede drain-age, expansion of the sand skeleton beneath the silt occursand can lead to the formation of a water film and completeloss of strength and a flow slide.

Traditionally, drainage conditions in laboratory investiga-tions to characterize liquefiable soils are considered as eitherundrained or (fully) drained, but these conditions do not rep-

© 2007 NRC Canada

874 Can. Geotech. J. Vol. 44, 2007

Fig. 1. Cross section of a sand volcano formed on the surface ofthe disposal pond (modified from Ishihara 1984).

3 Effects of void redistribution on liquefaction-induced ground deformations in earthquakes: a numerical investigation. Ph.D. dissertation. Inpreparation.

Administrator

Highlight

Administrator

Highlight

resent the real situation for stratified deposits in the field be-cause during and after shaking water migrates from zoneswith higher excess pressure towards zones with lower excesspressure.

In this paper a dynamic coupled stress-flow analysis pro-cedure is utilized to study the mechanisms involved in largedeformations leading to lateral spreads and flow failures ingentle slopes in liquefiable ground. Based on such analyses,implications for design of mitigation methods to resist seismicloading are examined.

Sand liquefaction and failure

Seismic liquefaction refers to a sudden loss in stiffnessand strength of soil due to cyclic loading effects of an earth-quake. The loss arises from a tendency for soil to contractunder cyclic loading, and if such contraction is prevented orcurtailed by the presence of water in the pores that cannotescape, it leads to a rise in pore-water pressure and a resultingdrop in effective stress (Silver and Seed 1971; Martin et al.1975). If the effective stress drops to zero (100% pore-waterpressure rise), the strength and stiffness also drop to zeroand the soil behaves as a heavy liquid. Unless the soil isvery loose (e.g., Dr ≤ 20%), however, it will dilate and regainsome stiffness and strength as it strains. If this strength issufficient, it will prevent a flow slide from occurring butmay still result in excessive displacements commonly re-ferred to as lateral spreading.

The excess pore water generated by seismic loading gen-erally drains upwards, and if no barrier layers are present,dissipation results in a decrease in void ratio and an increasein strength and stability. The potential for lateral spreadingand flow slides can greatly increase if low-permeability layers,e.g., silt layers within a soil deposit, impede drainage, form-ing a barrier to flow that can result in an expansion of thesand skeleton (void redistribution) and accumulation of porewater at the base of the layers.

The majority of the laboratory element tests conducted toassess liquefaction resistance of sands have been carried outunder undrained conditions. Pore-pressure rise in these ele-ments occurs under applied shear stress, causing a reductionin effective stresses and strength. In recent experimentalstudies, Vaid and Eliadorani (1998) and Eliadorani (2000) dem-onstrated that a small net flow of water into an element(injection) causing it to expand results in additional pore-pressure generation and further reduction in strength. Theyexamined this phenomenon by injecting or removing smallvolumes of water from the sample as it was being shearedand referred to this as a partially drained condition.

The results of inflow tests on Fraser River sand indicate apotential for triggering liquefaction at constant shear stress.Small imposed expansive volumetric strains resulted in effec-tive stress reduction and flow failure of samples of sand con-solidated to an initial state of stress corresponding toRc 1c 3c= ′ ′ =σ σ/ 2, as shown in Fig. 2a, where Rc is theconsolidation effective stress ratio, and ′σ1c and ′σ3c are themajor and minor principle effective stresses, respectively. Vaidand Eliadorani (1998) defined this condition as instability thatoccurs when a soil element subjected to small effective stressperturbation cannot sustain the current stress state and results inrunaway deformation as seen in Fig. 2b, or liquefaction flow.

In these tests, expansive volumetric strains (εv) wereimposed by injection of water into the samples at a constantrate d dv 1ε ε/ .= − 0 4, where ε1 is the axial strain. The sampleswere stable under the initial stress state. The stress paths fol-lowed during injection indicate a reduction in effectivestresses at a constant shear stress. At each initial confiningstress, the stress changes leading to instability occurred withlittle change in shear stress and void ratio and at very smallaxial strains of the order of 0.1%. Positive pore pressures con-tinued to develop even beyond the phase-transformation line.This occurs because the rate of imposed expansive volumetricstrain is greater than the tendency for the skeleton to expand.

Chu (1991), Bobei and Lo (2003), Sento et al. (2004), andYoshimine et al. (2006) have reported similar results for sandsand silty sand materials. As a result, soil elements can liquefydue to expansive volumetric strains that cannot be predictedfrom analyses based on the results of undrained tests.

To investigate the effects of low-permeability layers onground deformations due to earthquakes, it is necessary topredict the generation, redistribution, and dissipation ofexcess pore pressures during and after earthquake shaking.A fundamental approach requires a dynamic coupled stress-flow analysis. In such an analysis, the volumetric strains arecontrolled by the compressibility of the pore fluid and flowof water through the soil elements. To predict the instabilityand liquefaction flow, an effective stress approach based onan elastic–plastic constitutive model (UBCSAND) was used.The model is calibrated against laboratory element data andcentrifuge data and is described in the next section.

© 2007 NRC Canada

Seid-Karbasi and Byrne 875

Fig. 2. Partially drained instability of loose Fraser River sand(modified from Vaid and Eliadorani 1998): (a) stress paths;(b) strain paths. Drc, relative density at consolidation; ′σ3c, minorprincipal stress at consolidation.

Administrator

Highlight

Administrator

Highlight

Stress–strain model for sand

The UBCSAND constitutive model is based on the elastic–plastic stress–strain model proposed by Byrne et al. (1995) andhas been further developed by Beaty and Byrne (1998) andPuebla (1999). The model has been successfully used in ana-lyzing the CANLEX liquefaction embankments (Puebla et al.1997) and predicting the failure of the Mochikoshi tailings dam(Seid-Karbasi and Byrne 2004b). It has been used to examinedynamic centrifuge test data (e.g., Byrne et al. 2004; Seid-Karbasi et al. 2005) and also partially saturation effects (Seid-Karbasi and Byrne 2006). It is an incremental elastic–plasticmodel in which the yield loci are lines of constant stress ratio(η = τ / σ′, where η is the shear stress ratio, τ is the shear stress,and σ′ is the effective normal stress on the plane of maximumshear stress). The flow rule relating the plastic strain incrementdirections is nonassociated (see Fig. 3) and leads to a plasticpotential defined in terms of the dilation angle.

Elastic propertiesThe elastic component of response is assumed to be iso-

tropic and specified by a shear modulus, G e , and a bulkmodulus, Be , as follows:

[1] G K PP

G

ne e

aa

e

=′

σ

[2] B Ge e= α

where KGe is an elastic shear modulus number, Pa is the

atmospheric pressure, σ σ σ′ = ′ + ′( ) /x y 2, n e is an elastic expo-nent approximately equal to 0.5, and α ( ( ) / ( ))= + −2 1 3 2ν ν1depends on Poisson’s ratio, ν, and ranges from 2/3 to 4/3.

Plastic propertiesThe plastic shear strain increment d pγ is related to the

stress ratio, dη, where η = τ / σ′, as shown in Fig. 4 and canbe expressed as follows:

[3] ddpp

γ ησ

=′G /

where G p is the plastic shear modulus and is given by ahyperbolic function as follows:

[4] G G Rip p

ff= −

1

2η

η

where Gip is the plastic shear modulus at a low stress ratio

level (η = 0); η f is the stress ratio at failure and equals sinϕ f , where ϕ f is the peak friction angle; and Rf is the failureratio. Gi

p in turn is related to G e and the relative density ofthe sand. The associated increment of plastic volumetricstrain, d v

pε , is related to the increment of plastic shear strain,d pγ , through the flow rule as follows:

[5] d dvp

cvε γ ϕ η= −(sin )

where ϕcv is the friction angle at constant volume or phasetransformation. It can be seen from Fig. 3 that significantshear-induced plastic compaction is occurring (where plasticpotential vectors slope to the right) at low stress ratios,

whereas no compaction is predicted at stress ratios corre-sponding to ϕcv (plastic potential vector is vertical). Forstress ratios greater than ϕcv , shear-induced plastic expan-sion or dilation is predicted (where plastic potential vectorsslope to the left).

This simple flow rule is in close agreement with the char-acteristic behaviour of sand observed in drained laboratoryelement testing. The response of sand is controlled by theskeleton behaviour outlined previously. The presence of afluid (air–water mix) in the pores of the sand acts as a volu-metric constraint on the skeleton if drainage is fully or par-tially curtailed. It is this constraint that causes the porepressure rise that can lead to liquefaction. Provided the skel-eton or drained behaviour is appropriately modeled undermonotonic and cyclic loading conditions and the stiffness ofthe pore fluid and drainage are accounted for, the liquefac-tion response can be predicted.

This model was incorporated in the commercially avail-able computer code FLAC (fast lagrangian analysis of con-tinua), version 4.0 (Itasca Consulting Group Inc. 2005). Thisprogram models the soil mass as a collection of grid zonesor elements and solves the coupled stress flow problem usingan explicit time stepping approach. The program has a numberof built-in stress–strain models including an elastic–plastic

© 2007 NRC Canada


Fig. 3. Moving of yield loci with stress ratio and plastic strainincrement vectors. φcv, friction angle at constant volume; εv

p,plastic volumetric strain.

Fig. 4. Plastic shear strain increment and shear modulus.

Mohr–Coulomb model, and UBCSAND is a variation of thismodel in which friction and dilation angles are varied toincorporate the yield loci and flow rule described previ-ously. Pore fluid stiffness and Darcy hydraulic flow are ba-sic to the FLAC program, so only the skeleton stress–strainrelation is needed to simulate liquefaction, and that is whatthe UBCSAND model does. Drainage conditions are builtinto FLAC, and drained, undrained, or coupled stress flowconditions are specified by the user.

The key elastic and plastic parameters can be expressed interms of relative density, Dr , or normalized standard pene-tration test values, ( )N1 60 . Initial estimates of these parame-ters have been approximated from published data and modelcalibrations. The response of sand elements under mono-tonic and cyclic loading can then be predicted and the resultscompared with laboratory data. In this way, the model canbe made to match the observed response over the range ofrelative density or (N1)60 values. The model has also been

calibrated to reproduce the National Center for EarthquakeEngineering Research (NCEER) 1997 (Youd et al. 2001)triggering chart, which in turn is based on field experienceduring past earthquakes and is expressed in terms of thestandard penetration test resistance value, ( )N1 60 . The modelproperties to obtain such an agreement are therefore ex-pressed in terms of (N1)60.

Model simulation of laboratory element tests

The model was applied to simulate cyclic simple sheartests under undrained conditions. Figure 5 shows model pre-dictions along with test results on Fraser River sand. Thetest had an initial vertical consolidation stress ′ =σv 100 kPaand Dr = 40%. The results in terms of stress–strain andexcess pore pressure ratio, Ru , compare reasonably well withthe laboratory data. A comparison of model predictions withtests results in terms of required number of cycles to triggerliquefaction for different cyclic stress ratios, CSR, is shownin Fig. 5c and shows good agreement.

The model was also used to predict the effect of both theundrained condition and partial drainage as observed intriaxial monotonic tests. The partial drainage involved inject-ing the sample with water to expand its volume as it wassheared. The injection causes a drastic reduction in strength.In the numerical model, the same volumetric expansion was

© 2007 NRC Canada


Fig. 5. Comparison of predicted and measured response for FraserRiver sand: (a) stress–strain; (b) excess pore pressure ratio, Ru,versus number of cycles; (c) CSR versus number of cycles (testsdata from Sriskandakumar 2004).

Fig. 6. Predictions of element response in undrained and partiallydrained (inflow) triaxial tests for Fraser River sand: (a) stress–strain; (b) volumetric strains; (c) stress paths (modified fromAtigh and Byrne 2004).

applied and the results are shown in Figs. 6a, 6b, and 6c interms of stress–strain, volumetric strain versus mean stress,and stress path, respectively (model prediction with the solidline). The predictions are in remarkably good agreementwith the measured data.

The previous simulations illustrate that the model can gen-erate the appropriate pore pressures and stress–strain responseto undrained loading and can account for the effect of volu-metric expansion caused by inflow of water into an element.

Soil profile used in the analyses

The soil profile used in this study is a 10 m thick depositrepresenting a submerged ground condition as shown inFig. 7a. It comprises a loose sand deposit resting on animpermeable rigid foundation. The effect of a low-permeability layer within the loose sand at a depth of 4 m isexamined (see Fig. 7b). Fraser River sand with a relativedensity of Dr = 40% is considered to represent the loosesand. Materials properties are listed in Table 1, in which ρd ,n, and k are the material dry density, porosity, and perme-ability, respectively. The UBCSAND model was applied tothe loose sand layer with a corresponding equivalent

UBCSAND ( )N1 60 value. The low-permeability silt layerbarrier is simulated with a Mohr–Coulomb model having afriction angle ϕ = 30° and permeability k one thousand timeslower than that of the loose sand layer. Its stiffness in termsof bulk modulus and shear modulus was modeled as 1.0 ×104 and 0.5 × 104 kPa, respectively. It is not considered togenerate excess pore pressure.

Input base motion in terms of an acceleration–time historyis shown in Fig. 8. It is a harmonic (sinusoid) excitation

© 2007 NRC Canada


Firm impervious ground

Groundwater table

Element:(1,13)(1,10)(1,5)(1,3)

Material:Silt barrierLoose sandDrain column

(m)

0 1 2 3 4

-1

0

1

2

3

4

5

6

7

8

9

10

-1 (m)

(b) (c)(a)

Fig. 7. Ground conditions used in this study: (a) case I, uniform profile without low-permeability sublayer; (b) case II, profile withlow-permeability sublayer; (c) case III, profile treated with drain column.

MaterialDry density, ρd

(×1000 kg/m3)Porosity,n

UBCSAND(N1)60

Permeability,k (m/s)

Loose soil, Dr = 40% 1.50 0.448 6.2 8.81×10–4

Silt barrier 1.50 0.448 — 8.81×10–7

Table 1. Properties of the materials used in the analyses.

Fig. 8. Acceleration–time history for base input harmonic motion.

applied at the base of the soil layer. It ramps up to 2.5 m/s2

within 1 s and dies out in 2 s, lasting for 7 s in total.Analyses were conducted for three cases: (I) sloping

ground without a low-permeability layer, (II) sloping groundwith a low-permeability layer, and (III) sloping ground witha low-permeability layer treated with a drain column.

Analyses and results

To model the free-field condition, a mesh with 9 × 22zones as illustrated in Fig. 7 was used. Material types arerecognized with different permeability values as shown inthe figure. The nodes on the left and right boundaries werelinked to force the soil column to deform as a shear beam.The earthquake motion was applied as a time history ofacceleration at the base of the mesh.

Ground behaviour in cases I and IIA uniform sloping ground condition with 1° inclination

without a low-permeability layer was analyzed as a bench-mark condition (case I), and then the same soil profile with alow-permeability layer (case II) was analyzed. The barrier

layer was located at a depth of 4 m and had a permeability1000 times lower than that of the loose sand. The predictedbehaviour for these cases in terms of excess pore pressure,surface lateral displacement, and deformation pattern arepresented and compared in this section.

Figure 9 shows the time histories of excess pore pressureratio, Ru, for selected depths for the two cases (for positionsof the points refer to Fig. 7). The predicted patterns ofexcess pore pressure in case I (see Fig. 9a) indicate thatessentially high excess pore pressures build up within thesoil profile during the strong shaking. Small dilation spikesare predicted when Ru ≈ 1 and are less pronounced in theupper parts as a result of upward inflow. It may be seen thatexcess pore pressures dissipate somewhat more rapidly atdepth, e.g., Ru at 10 s is 55% and 75% for elements (1, 3)and (1, 13), respectively. A similar trend has been observedin centrifuge tests conducted for level and sloping ground(Taboada and Dobry 1993a, 1993b, 1998).

The corresponding Ru time histories for case II at differentdepths are shown in Fig. 9b. Again, pore pressures increasevery rapidly; however, beneath the barrier layer they remainhigh (Ru ≈ 100%) after the end of shaking, and dissipation

© 2007 NRC Canada


Fig. 9. Excess pore pressure ratio Ru versus time at selected points (elements (1, 13), (1, 10), and (1, 3), see Fig. 7) with increasingdepth: (a) case I; (b) case II.

occurs at greater depths. This indicates that water flowsfrom the greater depths towards the layer beneath the bar-rier, causing higher excess pore pressures to last for a signif-icantly longer time compared with the case without abarrier. The injected flow causes an expansion of the layerbeneath the barrier to occur at essentially zero effectivestress and leads to large deformation.

Figure 10 shows the predicted deformed meshes for thetwo cases. As seen from Fig. 10a, distortion in case I is pro-nounced at the base and tapers off towards the surface,resulting in a maximum displacement of 0.95 m at the topsurface. This pattern compares well with dynamic centrifugedata reported by Sharp et al. (2003), who modeled a 10 muniform liquefiable layer with a 5.2° inclination shaken with

22 cycles of harmonic motion with peak ground acceleration(PGA) = 0.23g at 2 Hz using viscose pore fluid. The soil usedin their tests was Nevada sand 120, deposited at Dr = 45%.

The deformation pattern after 30 s for case II is shown inFig. 10b. Comparing the pattern and magnitude of lateraldisplacements for these cases (with and without a barrierlayer shown in Figs. 10a and 10b), it can be seen that themagnitude has increased from 0.95 to 1.75 m and the patternis quite different, with a large slippage occurring at the baseof the barrier layer.

Time histories of horizontal displacement of the top sur-face for the two cases (with and without a barrier layer) arecompared in Fig. 11, which indicates that for case I, withoutthe barrier, displacements occur during shaking and ceaseshortly after the end of shaking. It can be seen that the sur-face displacements are much larger when a barrier layer ispresent. They are larger during shaking and continue toincrease after shaking ceases.

Figure 12 shows a profile of volumetric strain beneaththe barrier for case II and indicates that the lower twothirds of the soil profile contracts while expansion occursin the upper one third of the profile, with the highest rateof expansion directly beneath the barrier. Therefore, un-drained conditions do not exist, especially locally withinthe soil deposit beneath the barrier. This also suggests thatundisturbed samples taken prior to the earthquake will notbe representative of conditions during and shortly after theearthquake due to void redistribution resulting from the up-ward flow of water.

© 2007 NRC Canada


Fig. 10. Deformation pattern of soil profile: (a) without barrier, case I (with maximum lateral displacement of 0.95 m after 14 s);(b) with barrier (darker area), case II (with maximum lateral displacement of 1.75 m after 30 s).

Fig. 11. Surface lateral displacement (X-dis) versus time for pro-files with and without a barrier.

The increasing rate of expansion predicted in Fig. 12 asthe barrier is approached suggests that very high strains maybe occurring in the zone at or near the boundary. This isreferred to as strain localization.

Similar analyses were also carried out for the same soilprofile including a barrier layer with zero inclination (levelground). In this case, no significant lateral displacements arepredicted because of zero static driving force or shear stressbias for level ground conditions.

Flow failure

The results from the previous section revealed that delayedlarge deformations in liquefiable gently sloping grounds canoccur when a sublayer with low permeability is present.However, two questions arise in this regard. (1) In view ofthe predicted localization in the element beneath the barrier,to what extent is the predicted deformation pattern (i.e., con-traction in the lower section and expansion in the upper sec-tion) related to the model configuration (mesh size effects)?(2) What are the requirements for a flow failure?

The rest of this section discusses the results of analysescarried out to answer these questions.

The predicted strain localization implies that the com-puted results can be mesh-size dependent, as also noted byYang and Elgamal (2002) and Uzuoka et al. (2003). Toinvestigate scale effects on the predicted characteristicbehaviour of liquefiable grounds with a sublayer barrier, aseparate series of analyses was conducted for the same soilprofile with a barrier but modeled in two different ways:(i) as a 10 m soil profile in a 1g field (prototype scale), and

(ii) as a 0.1 m soil profile in a 100g field (model scale) fol-lowing appropriate conversion laws for modeling describedby Schofield (1981) and Kutter (1995). It is noted that themesh size used in the latter way of modeling was one-hundredth that used in the former.

The base input motion and mechanical properties of materialsare the same as those used previously. The displacement pat-terns were identical for both models when examined inprototype scale and are the same as those shown in Fig. 10b(the profiles of volumetric strain beneath the barrier layer forboth models were also identical and the same as thosedepicted in Fig. 12). This suggests that for a soil profilecomprising a barrier layer, contraction at the lower parts andexpansion at the upper parts are characteristic behavioursthat occur due to pore pressure migration and result in voidredistribution regardless of layer size. These findings indi-cate that the phenomenon of pore-water redistribution can becaptured in centrifuge tests and that the actual physical sizeof the layers is not important. What matters from a numericalmodeling point of view is the size of the mesh in relation tothe size of the liquefiable layer.

Further studies were undertaken to analyze the problemunder the same conditions but using meshes of decreasedsize for the element beneath the barrier. The results showedthat a decrease in element size results in greater volumetricexpansion for the very first element at the base of the barrierlayer. Figure 13a shows maximum volumetric strain of theelement beneath the barrier versus the element thicknessnormalized with respect to soil layer thickness beneath thebarrier. The figure suggests that void redistribution leads to avery thin, water-rich zone at the base of the barrier. Ifenough water flows into the soil element, it can expand untilthe steady or critical state at zero effective stress is reached.This corresponds approximately to the maximum void ratiostate. At this state the skeleton can undergo no furtherexpansion, and this is simulated by setting the dilation angleto zero. Additional inflow will result in the formation of awater film at the interface and zero shear strength.

The predicted increasing trend in volumetric strain of thebarrier base element with a decrease in its size (shown inFig. 13a) suggests that volumetric strain of the base elementbecomes infinity as the size of the base element approacheszero. This implies that for a liquefiable layer with a sublayerbarrier, flow failure occurs regardless of other involvedfactors, i.e., liquefied soil layer thickness, density, shakinglevel and duration, and ground slope.

Figure 13b shows time histories of volumetric strain forthe base element with various element thickness ratios(ETRs) for the analyzed sloping ground using theUBCSAND model. It can be seen that the amount of volu-metric expansion in the base element increases with timeand with a decrease in element thickness. More detailedanalyses indicate that the volumetric expansion does not goto infinity at the boundary. Further analyses to examine flowand expansion issues were carried out and are presented asfollows.

Figure 14 shows the analyses results in terms of the iso-chrones of Y-Flow, the vertical specific discharge velocity(flow rate per unit area) at different time intervals for themesh shown in Fig. 7b using the UBCSAND model. The

© 2007 NRC Canada


Fig. 12. Profile of volumetric strain beneath the barrier layer(after 30 s).

figure indicates that water flows upward from the lowerparts that have the highest excess pore pressures towards theupper low-permeability layer where the excess pore pressureis lowest. Maximum flow occurs near the mid-height of thesoil layer with a descending trend as it approaches theboundaries both top and bottom. Flows initially increasequickly with time and then drop off as pore pressures dissi-pate. Figure 15 shows the time history of volumetric strainrate for the base element of the mesh shown in Fig. 10b. Itcan be seen that the rate of expansion tends to essentiallyzero some time after shaking ceases. Similar results wereobtained for the meshes with smaller base zone thicknesses.Figure 16 shows the maximum expansion rate versus nor-malized zone thickness. It clearly shows no indication of aninfinite value for the expansion rate as the grid sizeapproaches zero. The results indicate that, although theexpansion increases with a reduction in element thickness,its maximum value is finite.

From a practical point of view, localized flow failure occursin a thin zone of soil when a shear band of limited thickness

forms within the soil mass. In the next section an approachfor numerical modeling of flow failures is presented.

Approach to modeling localized flow failureThe effect of mesh size can be approximately accounted

for in the analysis by specifying a dilation cut-off on skeletonexpansion that is related to the initial relative density andmesh size. Denser material would have more dilation capacityto reach its maximum void ratio, at which point dilationwould be set to zero. For a coarser mesh, less expansion ispredicted for the element and the amount of expansionrequired to trigger dilation cut-off would be less.

The mesh size should also be related to the particle size.Roscoe (1970) noted that a shear band may occur within athin zone with a thickness 10–20 times that of the meangrain size of the soil, D50 . Similar finding have been reportedfrom further experimental investigations (e.g., Alshibli andSture 1999; Desrues and Viggiani 2004; Lu et al. 2004) andnumerical investigations (e.g., district element method,DEM) by a number of researchers. A reasonable lower limit

© 2007 NRC Canada


Fig. 13. Volumetric strain of the base element with various thicknesses: (a) maximum volumetric strain versus normalized elementthickness; (b) time histories of volumetric strain.

Administrator

Highlight

on mesh size could be based on the size of the shear band.For a soil profile comprising Fraser River sand of Dr = 40%with D50 = 0.3 mm, a shear band could form within a zone3–6 mm in height. In Fig. 17 the data shown in Fig. 13a areplotted in semilogarithmic format with the initial portionextrapolated to reach a value corresponding to an elementthickness of 3 mm (ETR = 3/6000 = 0.0005). The maximumvolumetric strain predicted for that ETR is about 5.2%. Ifthe shear band zone can expand by 5.2% before reaching itscritical state, then it can retain some strength and prevent aflow slide. A key question then is as follows: What is themaximum volumetric expansion (threshold strain) that anelement can tolerate before a water film forms?

Potential maximum volumetric (threshold) strain can beestimated based on data from various soils available in theliterature. Table 2 lists a number of soils with their minimumand maximum void ratios compiled by Cubrinovski and Ishihara

(2002) and Olson and Stark (2003). The threshold volumetricstrain, ( )εv T , can be calculated as follows:

[6] ( ) maxεv T rmin

i

=−

+D

e e

e1

where Dr is the soil relative density, emax is the maximumvoid ratio, emin is the minimum void ratio, and ei is the initialvoid ratio. The numerator of eq. [6] should be e ess min−(where ess is the steady-state void ratio at zero effectivestress), but it can be approximated by e emax − min for the sakeof simplicity. A similar approach was adopted by Tsukamotoet al. (2004) and Sawada et al. (2006) to estimate potentialmaximum liquefaction-induced settlement. In Table 2, thepotential maximum expansion is also provided for a typicalmedium-dense state (Dr = 50%). Figure 18 shows the upperand lower bounds for potential maximum expansion for acomplete range of initial Dr calculated based on eq. [6]. Arange of relative densities from 25% to 80% is of interestfrom a practical point of view. For Fraser River sand at Drof 40%, (εv)T = 6.9%.

For the case involved in Fig. 17 with Fraser River sand,the predicted maximum volumetric strain in the shear bandzone was 5.2%. Since this is less than the threshold value,i.e., ( )εv T = 6.9%, a flow slide is not predicted. If the expan-sion limit is reached, the soil element deforms in a steady-state condition with essentially zero shear strength. Not allelements along a potential slide surface will necessarilyreach this state, and the low residual strengths back-calculated from field experience are a reflection of the aver-age strength on the failure surface.

When the element thickness is chosen equal to the shearband thickness, there is no correction for element size. It isgenerally not practical to make the element height this small,and a correction for element size is then needed. To examinethis approach, the sloping-ground problem was analyzed usingthe mesh shown in Fig. 7b. This mesh has an ETR of 0.042compared with a localized failure or shear band zone ratio of0.0005. Based on Fig. 13, the predicted volumetric strain in theelement beneath the barrier would be 0.79% compared with5.10% in the shear band. To account for this, the strain in theelement must be modified to ( )*ε εv v= (5.10/0.79) = 6.4εv ,where 6.4 is a mesh correction factor.

© 2007 NRC Canada


Fig. 14. Vertical specific discharge (Y-Flow) isochrones at differ-ent times.

Fig. 15. Time history of volumetric strain rate of base element.

Fig. 16. Maximum expansion rate versus base element thicknessratio. Isochrones at different time intervals for UBCSAND modelwith 0.5 m element height.

© 2007 NRC Canada


Fig. 17. Estimate of maximum volumetric strain, ( )maxεv , for a base element with 10D50 (0.003 m) thickness for Fraser River sandusing extended trend of ( )maxεv with element thickness ratio.

Void ratio

Soil emin emax

Expansion potential,(εv)50 (%) Reference

Fraser River sand 0.596 0.900 8.8 Vaid and Eliadorani 1998Cambria sand 0.538 0.767 7.0 Lade et al. 1998Nevada sand 50/80 0.581 0.858 8.0 Lade et al. 1998Nevada sand 80/200 0.617 0.940 9.1 Lade et al. 1998Nevada fines 0.754 1.178 10.8 Lade et al. 1998Ottawa sand 50/200 0.550 0.805 7.6 Lade and Yamamuro 1997Ottawa sand F-95 0.580 0.865 8.3 Lade and Yamamuro 1997Host sand A2 0.60 0.98 10.6 Thevanayagam 1998Toyoura sand 0.616 0.988 10.3 Zlatovic 1994Ottawa sand 0.48 0.78 9.2 Salgado et al. 2000Ottawa sand C-109 0.50 0.83 9.9 Pitman et al. 1994Quiou sand 0.78 1.20 10.6 Pestana and Whittle 1995Mine tailings sand 0.69 1.06 9.9 Vaid et al. 1985Brasted sand 0.48 0.79 9.5 Cornforth 1974Dune sand 0.54 0.91 10.7 Konrad 1990Well-rounded silica sand 0.67 1.06 10.5 Konrad 1990Nerlerk 0%–2% 0.62 0.94 9.0 Sladen et al. 1985Tottri sand 0.938 1.008 10.2 Takeshita et al. 1995Monterey No. 9 sand 0.53 0.86 9.7 Riemer et al. 1990Massey Tunnel sand 0.712 1.102 10.2 Konrad and Pouliot 1997Quebec sand 0.54 0.79 7.5 Konrad 1998Sand B 0.50 0.84 10.2 Castro 1969Sand C 0.66 0.99 9.0 Castro 1969Sand A 1.23 1.88 12.7 Castro 1969

Table 2. Minimum and maximum void ratio and potential maximum expansion of sands at Dr = 50%.

The analysis showed that the ( )*εv so computed was lessthan the threshold value of 6.9%, and the results wereidentical to those predicted in Figs. 11 and 12. When theduration of base motion was increased from 7 s to 12 s, thecomputed ( )*εv was greater than 6.9%, at which time dila-tion was suppressed and a flow slide was predicted.

Figure 19 shows the surface displacement – time historiesfor both cases. It can be seen that surface displacement for12 s motion is ever increasing, whereas displacements stopsome time after the end of shaking in the 7 s case. Groundsurface velocity is a good indicator of flow failure. Figure 20shows ground surface velocity versus time for the twoevents. For the 7 s event, the surface velocity decreases andthen becomes zero after the end of shaking, whereas for the12 s event it is essentially constant over that period. This re-veals that flow failure takes place in the second case whenthe dilation threshold is reached.

Countermeasure for flow failure

Drains have the potential to nullify the barrier effect andcurtail or prevent the occurrence of lateral spread or flowslides in the event of an earthquake. They can facilitate dis-sipation of excess pore pressure and mitigate the impedanceeffects of low-permeability layers. This is examined for thesame soil profile as that used in the previous analyses with

1.0° inclination treated with gravel drains penetratingthrough the low-permeability sublayer down to the loosesand layer at the bottom (case III).

For the study conducted in plane strain conditions, thedrains have been accounted for with a soil column having apermeability 100 times greater than that of the surroundingliquefiable soils. In a real prototype problem, the three-dimensional effects of the drain columns pattern can betreated in plane strain analysis using an equivalent drain areaapproach (e.g., Indraratna and Redana 1997, 2000). Fig-ure 21a shows the model (case III) with a drain. Figure 21bshows distributions of maximum excess pore pressure ratio,Rumax , within the model along with flow vectors after 3.5 sof shaking. It is seen that Rumax increases with an increase indistance from the drain column. The drain exhibits someexcess pore pressure during shaking. The maximum pre-dicted lateral displacement in this case was negligible, evencompared with that of the model without a low-permeabilitylayer (case I). This demonstrates that the drain column is apromising measure to mitigate liquefaction-induced defor-mations.

A comparison of predicted time histories of excess porepressure buildup for a point at mid-depth of loose sand forcases I, II, and III is shown in Fig. 22, which indicates thatdrains can reduce generated excess pore pressures and signi-ficantly speed up the dissipation of excess pore pressure.The effect of drains in decreasing excess pore pressure duringshaking is also reported by Chang et al. (2004) and Brennanand Madabhushi (2005) in field model tests and centrifugetests, respectively. Figure 23 shows excess pore pressure ratiosmeasured in field tests within a liquefiable soil for a depositwithout and with vertical drains subjected to a harmonicexcitation. These data show that the excess pore pressurebuild-up rate is significantly slower with drains and confirmsthe effectiveness of drains as a remediation measure.

Practical implications

It was demonstrated that the applied dynamic procedurecan reproduce the seismic response of liquefiable soil whena low-permeability sublayer is present. The presence of sucha sublayer impedes upward drainage that can lead to the for-mation of water films having essentially zero shear strengthand result in large ground deformation and flow failure. The

© 2007 NRC Canada


Fig. 18. Maximum expansion potential versus initial relative densityfor sands (based on Table 2 data and eq. [6]).

Fig. 19. Time histories of surface lateral displacement (X-dis) fortwo events.

Fig. 20. Time histories of surface lateral velocity (X-vel) for twoevents.

Administrator

Highlight

Administrator

Highlight

analyses also show that drains can be very effective in alle-viating the destructive effects of sublayer barriers.

Important aspects for design are as follows:(1) Appropriate site investigation techniques should be applied

to detect low-permeability, thin layers within sand and(or) gravel layers.

(2) If a low-permeability sublayer is present and significantliquefaction is predicted for the design earthquake, eitherlow residual shear strengths consistent with field backanalysis should be used or drainage–densification shouldbe considered. The design of drains–densification canbe optimized from a coupled stress-flow dynamic analysis.

(3) The perfect undisturbed sample obtained prior to anearthquake will not be representative of conditions duringand shortly after the earthquake due to expansion result-ing from the upward flow of water in stratified deposits.Dissipation of excess pore pressure some time after theearthquake will reconsolidate soil elements beneath thebarrier such that the perfect sample taken some timeafter the earthquake will also not be representative ofthe critical conditions.

Conclusions

Many failures of earth structures and submarine slideshave been reported during past earthquakes worldwide. Anumber of civil structures and soil deposits in coastal orriver areas have suffered large deformations during pastearthquakes as a result of soil liquefaction. The deformationsmay exceed several metres, even in gentle slopes of less thana few percent. Deformations occur not only during but alsoafter earthquake shaking.

Clean, loose (e.g., Dr ≥ 20%) sands are unlikely to suffera flow slide. Although they can be triggered to liquefy andundergo large strains and displacements, their undrainedstrengths are generally adequate for stability. However, if the

sands contain low-permeability layers that impede drainage,a void expansion resulting in a water-rich thin zone or waterfilm may cause a near complete loss of strength of the soildirectly beneath that layer and result in a flow slide. The lowresidual strengths based on back analyses of field case histo-ries and reported by Seed and Harder (1990) are likely a resultof void expansion related to the presence of barrier layers.

A numerical approach is used in this paper that captureselement sand behaviour in monotonic and cyclic loading underdifferent drainage conditions; undrained and inflow wereutilized to study the effects of low-permeability layers onthe sloping ground response during earthquake loading. Thefollowing conclusions are made based on the analysis results:

(1) The computed results show that for a homogeneoussand layer that is triggered to liquefy, upward flowresulting from excess pore pressure causes dissipationand a reduction in void ratio within the sand layer. How-ever, if a low-permeability layer(s) impedes drainage,then contraction occurs at the base of the sand layer butexpansion at the top where flow is impeded, and thisexpansion is responsible for flow failures. The occur-rence of contraction and expansion zones, respectively,at the lower and upper zones due to pore pressure redis-tribution is a characteristic behaviour of a liquefiabledeposit with sublayer barriers.

(2) Void redistribution results in a thin, water-rich zone atthe base of a barrier layer that could ultimately form awater film when enough water is available for injection.The thickness of this shear band zone is related to particlediameter; if and when the predicted expansion exceedsthe threshold expansion, a water film will form, and thisis simulated in the analyses by setting the dilation equalto zero.

(3) It is generally not practical to have the mesh size assmall as the shear band thickness, in which case a cor-rection to the computed expansion is required.

© 2007 NRC Canada


(a)(a)

Vertical drain

Rumax

0.65

0.700.75

0.80

0.850.90

0.95

1.00

(b)

0.65

0.700.75

0.80

0.850.90

0.95

1.00

(b)

k (m/s)8.81 x 10-7

8.81 x 10-4

8.81 x 10-2

Fig. 21. Treated model, case III: (a) central vertical drain; (b) Rumax and flow vectors at 3.5 s.

Administrator

Highlight

Administrator

Highlight

Administrator

Highlight

Administrator

Highlight

Administrator

Highlight

Administrator

Highlight

(4) Most of the relative movements occur at the base of thelow-permeability sublayer. Maximum displacements occurat or near the surface above the barrier layer.

(5) A large part of the deformations may occur some timeafter shaking has ceased, depending on the time neededfor migration of water from zones with higher excesspore pressures.

(6) Installation of vertical drains that penetrate the barrierlayers can mitigate the destructive effects of low-permeability layers. This has been demonstrated in theseanalyses and has also been observed from field modeltests.

(7) The design of these drains can be assessed fromdynamic coupled flow – effective stress analyses usingappropriate modeling parameters and design inputmotions. The dimensions and location of remediationmeasures can be optimized from dynamic analyses.

Acknowledgments

The authors acknowledge the financial support from BCHydro through the Professional Partnership program and thesupport of the Natural Sciences and Engineering ResearchCouncil of Canada (NSERC) through Strategic LiquefactionGrant NSERC 246394 and NSERC COSTA Grant 03608-CG068625. The authors are also grateful to ProfessorD.L. Anderson and to Ernest Naesgaard for helpful discus-sions during the course of this research. In addition, the au-thors express their appreciation to the reviewers for theirconstructive comments, which led to many changes from theoriginal manuscript.

References

Adalier, K., and Elgamal, A.W. 1992. Post-liquefaction behavior ofsoil systems. Report, Department of Civil Engineering,Rensselaer Polytechnic Institute, Troy, N.Y.

Alshibli, K., and Sture, S. 1999. Sand shear band thickness mea-surements by digital imaging techniques. Computing in CivilEngineering, ASCE, 13: 103–109.

Atigh, E., and Byrne, P.M. 2004. Liquefaction flow of submarineslopes under partially undrained conditions: an effective stressapproach. Canadian Geotechnical Journal, 41(1): 154–165.

Beaty, M.H., and Byrne, P.M. 1998. An effective stress model forpredicting liquefaction behavior of sand. In Geotechnical Earth-quake Engineering and Soil Dynamics III: Proceedings of aSpecialty Conference, Seattle, Wash., 3–6 August 1998. Editedby P. Dakoulas, M. Yegian, and R.D. Holtz. ASCE GeotechnicalSpecial Publication 75, Vol. 1, pp. 766–777.

Berrill, S.A., Christensen, S.A., Keenan, R.J., Okada, W., andPettinga, J.R. 1997. Lateral-spreading loads on a piled bridgefoundation. In Proceedings of the 14th International Conferenceon Soil Mechanics and Foundation Engineering, DiscussionSession, Seismic Behavior of Ground and Geotechnical Struc-tures, Hamburg, Germany, 6–12 September 1997. A.A.Balkema, Rotterdam. pp. 173–183.

Bobei, D.C., and Lo, S.-C. 2003. Strain path influence on thebehavior of sand with fines. In Soil and Rock America 2003:Proceedings of the 12th Panamerican Conference on SoilMechanics and Geotechnical Engineering, Cambridge, Mass., 22–26 June 3003. Edited by P.J. Culligan, H.H. Einstein, and A.J.Whittle. Verlag Gluckauf GMBH, Essen, Germany. pp. 583–588.

© 2007 NRC Canada


Fig. 22. Predicted time history of Ru at mid-depth of loose sand:(a) without barrier layer; (b) with barrier layer; (c) with barrierlayer treated with drain (Seid-Karbasi3).

Fig. 23. Measured Ru in field liquefaction test with and withouta drain (data from Chang et al. 2004).

Bouckovalas, G.D., Gazetas, G., and Papadimitriou, A.G. 1999.Geotechnical aspects of the 1995 Aegion (Greece) earthquake.In Proceedings of the 2nd International Conference on Earth-quake Geotechnical Engineering, Lisbon, 21–25 June. Edited byP. Seco e Pinto. A.A. Balkema, Rotterdam. pp. 739–748.

Brennan, A.J., and Madabhushi, S.P. 2005. Liquefaction and drainagein stratified soil. Journal of Geotechnical and GeoenvironmentalEngineering, ASCE, 131: 876–885.

Byrne, P.M., and Beaty, M.H. 1997. Post-liquefaction shearstrength of granular soils: theoretical/conceptual issues. In Pro-ceedings of the Workshop on Post-Liquefaction Shear Strengthof Granular Soils, Urbana–Champion, Ill., 17–18 April 1997.pp. 16–45.

Byrne, P.M., Roy, D., Campanella, R.G., and Hughes, J. 1995.Predicting liquefaction response of granular soils from pressure-meter tests. In Static and Dynamic Properties of Gravelly Soils:Proceedings of Sessions of the ASCE National Convention, SanDiego, Calif., 23–24 October 1995. Edited by M.D. Evans andR.J. Fragaszy. ASCE Geotechnical Special Publication 56,pp. 122–135.

Byrne, P.M., Park, S., Beaty, M., Sharp, M., Gonzalez, L., andAbdoun, T. 2004. Numerical modeling of liquefaction and com-parison with centrifuge tests. Canadian Geotechnical Journal,41(1): 193–211.

Byrne, P.M., Naesgaard, E., and Seid-Karbasi, M. 2006. Analysisand design of earth structures to resist seismic soil liquefaction.In Proceedings of the 59th Canadian Geotechnical EngineeringConference, Hardy Lecture, Vancouver, B.C., 1–4 October 2006.Canadian Geotechnical Society, Alliston, Ont. pp. 1–24.

Castro, G. 1969. Liquefaction of sands. Ph.D. thesis, Harvard Uni-versity, Cambridge, Mass.

Castro, G., Keller, T.O., and Boynton, S.S. 1989. Re-evaluation ofthe Lower San Fernando Dam. Contract Report GL-89-2, US ArmyWaterways Experiment Station, Vicksburg, Miss. Vols. 1 and 2.

Chang, W.J., Rathje, E., Stokoe, K.H., II, and Cox, B.R. 2004.Direct evaluation of effectiveness of prefabricated vertical drainin liquefiable sand. Soil Dynamics and Earthquake EngineeringJournal, 24: 723–731.

Chu, J. 1991. Strain softening behavior of granular soils understrain path testing. Ph.D. thesis, University of New South Wales,Kensington, NSW, Australia.

Cornforth, D.H. 1974. One-dimensional consolidation curves of amedium sand. Géotechnique, 24: 678–683.

Coulter, H.W., and Migliaccio, R.R. 1966. Effects of the earth-quake of March 27, 1964 at Valdez, Alaska. US GeologicalSurvey, Professional Paper 542-C. 36 pp.

Cubrinovski, M., and Ishihara, K. 2002. Maximum and minimumvoid ratio characteristics of sand. Soils and Foundations, 42: 65–78.

Desrues, J., and Viggiani, G. 2004. Strain localization in sand: anoverview of the experimental results obtained in Grenoble usingstereophotogrammetry. International Journal for Numerical andAnalytical Methods in Geomechanics, 28: 279–321.

Elgamal, A.W., Dobry, R., and Adalier, K. 1989. Small scaleshaking table tests of saturated layered sand–silt deposits. InProceedings of the 2nd US–Japan Workshop on Soil Liquefac-tion, Large Ground Deformation, and Their Effects on Life-lines, Grand Island and Ithaca, N.Y., 26–29 September 1989.Edited by T.D. O’Rourke and M. Hamada. NCEER TechnicalReport NCEER-89-0032, National Center for Earthquake Engi-neering Research (NCEER), Buffalo, N.Y. pp. 233–245.

Eliadorani, A.A. 2000. The response of sands under partiallydrained states with emphasis on liquefaction. Ph.D. thesis, CivilEngineering Department, The University of British Columbia,Vancouver, B.C.

Fiegel, G.L., and Kutter, B.L. 1992. Liquefaction mechanism forlayered soil. Journal of Geotechnical Engineering, ASCE, 120:737–755.

Field, M.E., Gardner, J.V., Jennings, A.E., and Edwards, B.D.1982. Earthquake induced sediment failures on a 0.258 slope,Klamath River Delta. Journal of California Geology, 10: 542–546.

Hamada, M. 1992. Large ground deformations and their effects onlifelines: 1964 Niigata Earthquake. In Case studies of liquefac-tion and lifeline performance during past earthquakes. Vol. 1.Japanese case studies. Edited by M. Hamada and T.D.O’Rourke. NCEER Technical Report NCEER-92-0001, NationalCenter for Earthquake Engineering Research (NCEER), Buffalo,N.Y. pp. 3/1–3/123.

Hampton, M.A., and Lee, H.J. 1996. Submarine landslides. Reviewsof Geophysics, 34: 33–59.

Harder, L., and Stewart, J. 1996. Failure of Tapo Canyon tailingsdam. Journal of Performance of Constructed Facilities, ASCE,10(3): 109–114.

Huishan, L., and Taiping, Q. 1984. Liquefaction potential of sat-urated sand deposits underlying foundation of structure. InProceedings of the 8th World Conference on Earthquake En-gineering, San Francisco, Calif., 21–28 July 1984. Prentice–Hall Inc., Englewood Cliffs, N.J. Vol. 3, pp. 199–206.

Indraratna, B., and Redana, I.W. 1997. Plane strain modelingsmear effects associated with vertical drains. Journal ofGeotechnical Engineering, ASCE, 119: 1321–1329.

Indraratna, B., and Redana, I.W. 2000. Numerical modeling ofvertical drains with smear and well resistance installed in softclay. Canadian Geotechnical Journal, 37(1): 132–145.

Ishihara, K. 1984. Post-earthquake failure of a tailings dam due toliquefaction of the pond deposit. In Proceedings of the Interna-tional Conference of Case Histories in Geotechnical Engi-neering, University of Missouri-Rolla, Rolla, Mo., May 1984.Vol. 3, pp. 1129–1143.

Itasca Consulting Group, Inc. 2005. Fast lagrangian analysis ofcontinua (FLAC), version 5, user’s guide. Itasca ConsultingGroup, Inc., Minneapolis, Minn.

Kawakami, F., and Asada, A. 1966. Damage to the ground andearth-structures by the Niigata earthquake of June 16, 1964.Soils and Foundations, 1: 14–30.

Kokusho, T. 1999. Water film in liquefied sand and its effect onlateral spread. Journal of. Geotechnical and GeoenvironmentalEngineering, ASCE, 125: 817–826.

Kokusho, T. 2000. Mechanism for water film generation and lat-eral flow in liquefied sand layer. Soils and Foundations, 40:99–111.

Kokusho, T. 2003. Current state of research on flow failure consid-ering void redistribution in liquefied deposits. Journal of SoilDynamics and Earthquake Engineering, 23: 585–603.

Kokusho, T., and Kojima, T. 2002. Mechanism for post-liquefaction water film generation in layered sand. Journal ofGeotechnical and Geoenvironmental Engineering, ASCE, 128:129–137.

Konrad, J.-M. 1990. Minimum undrained strength of two sands.Journal of Geotechnical Engineering, ASCE, 116: 932–947.

Konrad, J.-M. 1998. Sand state from cone penetration tests. Géo-technique, 48: 201–215.

Konrad, J.-M., and Pouliot, N. 1997. Ultimate state of reconstitutedand intact samples of deltaic sand. Canadian Geotechnical Journal,34(5): 737–748.

Kulasingam, R. 2003. Effects of void redistribution onliquefaction-induced deformations. Ph.D. thesis, Civil and En-vironmental Engineering Department, University of California,Davis, Calif.

© 2007 NRC Canada


Kulasingam, R., Malvick, E.J., Boulanger, R.W., and Kutter, B.L.2001. Void redistribution and localization of shear strains inmodel sand slopes with silt seams. Report on first year activities.In Proceedings of the US–Japan Cooperative Research in UrbanEarthquake Disaster Mitigation Workshop, Seattle, Wash., 16–18 August 2001. MCEER, N.Y. pp. 117–128.

Kulasingam, R., Malvick, E.J., Boulanger, R.W., and Kutter, B.L.2004. Strength loss and localization at silt interlayers in slopesof liquefied sand. Journal of Geotechnical and Geo-environmental Engineering, ASCE, 130(11): 1192–1202.

Kutter, B.L. 1995. Recent advances in centrifuge modeling of seis-mic shaking. In Proceedings of the 3rd International Conferenceof Recent Advances in Geotechnical Earthquake Engineering andSoil Dynamics, St. Louis, Mo., 2–7 April 1995. University ofMissouri-Rolla, Rolla. Vol. 2, pp. 927–942.

Lade, P.V., and Yamamuro, J.A. 1997. Effects of nonplastic fineson static liquefaction of sands. Canadian Geotechnical Journal,34(6): 918–928.

Lade, P.V., Liggio, C.D., and Yamamuro, J.A. 1998. Effects ofnon-plastic fines on minimum and maximum void ratios of sand.ASTM Journal of Geotechnical Testing, 21: 336–347.

Lemke, R.W. 1967. Effects of the earthquake of March 27, 1964 atSeward, Alaska. US Geological Survey, Professional Paper 542-E.

Liu, H., and Qiao, T. 1984. Liquefaction potential of saturated sanddeposits underlaying foundation of structure. In Proceedings ofthe 8th World Conference on Earthquake Engineering, San Fran-cisco, Calif., 21–28 July 1984. Prentice–Hall Inc., EnglewoodCliffs, N.J. pp. 199–206.

Lu, X., Shuyun, W., Yihua, W., and Cui, P. 2004. An approximatemethod for evaluating the shear band thickness in saturatedsand. International Journal of Numerical and Analytical Methodsin Geomechanics, 28: 1533–1541.

Malvick, E.J. 2005. Void redistribution-induced shear localizationand deformation in slopes. Ph.D. dissertation, University of Cal-ifornia at Davis, Davis, Calif. 285 pp.

Malvick, E.J., Kulasingam, R., Kutter, B.L., and Boulanger,R.W. 2002. Void redistribution and localized shear strains inslopes during liquefaction. In Physical Modeling inGeotechnics, ICPMG‘02: Proceedings of the InternationalConference, St. John’s, Nfld., 10–12 July 2002. Edited by R.Phillips, P.J. Guo, and R. Popescu. A.A. Balkema, Rotterdam.pp. 495–500.

Malvick, E.J., Kutter, B.L., Boulanger, R.W., Kabasawa, K., andKokusho, T. 2005. Void redistribution research with 1-g andcentrifuge modeling. In Proceedings of the 16th InternationalConference on Soil Mechanics and Geotechnical Engineering(ICSMGE), Osaka, 12–16 September 2005. A.A. Balkema, Rot-terdam. Vol. 4, pp. 2543–2546.

Malvick, E.J., Kutter, B.L., Boulanger, R.W., and Kulasingam, R.2006. Shear localization due to liquefaction-induced void redis-tribution in a layered infinite slope. Journal of Geotechnical andGeoenvironmental Engineering, ASCE, 132: 1293–1303.

Martin, G.R., Finn, W.D.L., and Seed, H.B. 1975. Fundamentals ofliquefaction under cyclic loading. Journal of the GeotechnicalEngineering Division, ASCE, 101: 423–438.

Olson, S.M., and Stark, T.D. 2002. Liquefied strength ratio fromliquefaction flow failure case histories. Canadian GeotechnicalJournal, 39(3): 629–647.

Olson, S.M., and Stark, T.D. 2003. Use of laboratory data to confirmyield and liquefied strength ratio concepts. CanadianGeotechnical Journal, 40(6): 1164–1184.

Pestana, J.M., and Whittle, A.J. 1995. Compression model forcohesionless soils. Géotechnique, 45: 611–631.

Pitman, T.D., Robertson, P.K., and Sego, D.C. 1994. Influence offines on the collapse of loose sands. Canadian GeotechnicalJournal, 31(5): 728–739.

Poulos, S.J., Castro, G., and France, J.W. 1985. Liquefaction evalu-ation procedure. Journal of Geotechnical Engineering., ASCE,111: 772–792.

Puebla, H. 1999. A constitutive model for sand analysis of theCANLEX embankment. Ph.D. thesis, Civil Engineering Depart-ment, University of British Columbia, Vancouver, B.C.

Puebla, H., Byrne, P.M., and Phillips, R. 1997. Analysis ofCANLEX liquefaction embankments: prototype and centrifugemodels. Canadian Geotechnical Journal, 34(5): 641–654.

Riemer, M.F., Seed, R.B., Nicholson, P.G., and Jong, H.L. 1990.Steady state testing of loose sands: limiting minimum density.Journal of Geotechnical Engineering, ASCE, 116: 332–337.

Roscoe, K.H. 1970. The influence of strains in soil mechanics.Géotechnique, 20: 129–170.

Salgado, R., Bandini, P., and Karim, A. 2000. Shear strength andstiffness of silty sand. Journal of Geotechnical and Geo-environmental Engineering, ASCE, 126: 451–462.

Sawada, S., Tsukamoto, Y., and Ishihara, K. 2006. Residual defor-mation characteristics of partially saturated sandy soils sub-jected to seismic excitation. Journal of Soil Dynamics andEarthquake Engineering, 26: 175–182.

Schofield, A.N. 1981. Dynamics and earthquake geotechnicalcentrifuge modeling. In Proceedings of the International Confer-ence on Recent Advances in Geotechnical Earthquake Engi-neering and Soil Dynamics, St. Louis, Mo., 26 April – 2 May1981. Edited by S. Prakash. University of Missouri-Rolla, Rolla.Vol. 3, pp. 1081–1100.

Scott, R.F., and Zuckerman, K.A. 1972. Sandblows and liquefac-tion. In The Great Alaskan earthquake of 1964. EngineeringPublication 1606, National Academy of Sciences, Washington,D.C., pp. 170–189.

Seed, H.B. 1987. Design problems in soil liquefaction. Journal ofGeotechnical Engineering, ASCE, 113: 827–845.

Seed, R.B., and Harder, L.F. 1990. SPT-based analysis of cyclicpore pressure generation and undrained residual strength. InProceedings of the H.B. Seed Memorial Symposium, May 1990.Edited by J.M. Duncan. BiTech Publishing Ltd., Richmond,B.C. Vol. 2, pp. 351–376.

Seid-Karbasi, M., and Byrne, P.M. 2004a. Liquefaction, lateralspreading and flow slides. In Proceedings of the 57th CanadianGeotechnical Engineering Conference, Québec City, Que., 24–27 October 2004. Canadian Geotechnical Society, Alliston, Ont.p. G13.529.

Seid-Karbasi, M., and Byrne, P.M. 2004b. Embankment dams andearthquakes. International Journal on Hydropower and Dams,11(2): 96–102.

Seid-Karbasi, M., and Byrne, M.P. 2006. Effects of partial satura-tion on liquefiable ground response. In GeoCongress 2006:Geotechnical Engineering in the Information Technology Age,Atlanta, Ga., 26 February – 2 March 2006. Edited by D.J.DeGroot, J.T. DeJong, D. Frost, and L.G. Baise. ASCE, Reston,Va. Paper 11803.

Seid-Karbasi, M., Byrne, P.M., Naesgaard, E., Park, S.,Wijewickreme, D., and Phillips, R. 2005. Response of slopingground with liquefiable materials during earthquake: a class Aprediction. In International Association of Computer Methodsand Advances in Geomechanics, IACMAG: Proceedings of the11th International Conference, Turin, Italy, 19–21 June 2005.Edited by G. Barla and M. Barla. Patron Editore, Bologna, Italy.Vol. 3, pp. 313–320.

© 2007 NRC Canada


Sento, N., Kazama, M., Uzuoka, R., Ohmur, H., and Ishimaru, M.2004. Possibility of post liquefaction flow failure due to seepage.Journal of Geotechnical and Geoenvironmental Engineering,ASCE, 130: 707–716.

Sharp, M., Dobry, R., and Abdoun, T. 2003. Liquefaction centri-fuge modeling of sands of different permeability. Journal ofGeotechnical and Geoenvironmental Engineering, ASCE, 129:1083–1091.

Silver, M.L., and Seed, H.B. 1971. Volume change in sands duringcyclic loading. Journal of the Soil Mechanics and FoundationsDivision, ASCE, 97: 1171–1182.

Sladen, J.A., D’Hollander, R.D., Krahn, J., and Mitchell, D.E.1985. Back analysis of the Nerlerk berm liquefaction slides.Canadian Geotechnical Journal, 22: 579–588.

Sriskandakumar, S. 2004. Cyclic loading response of Fraser RiverSand for numerical models simulating centrifuge tests. M.A.Sc.thesis, Civil Engineering Department, University of BritishColumbia, Vancouver, B.C.

Stark, T.D., and Mesri, G. 1992. Undrained shear strength of lique-fied sands for stability analysis. Journal of Geotechnical Engi-neering, ASCE, 118: 1727–1747.

Taboada, V., and Dobry, R. 1993a. Experimental results of modelNo. 1 at RPI. In Verification of Numerical Procedures for theAnalysis of Soil Liquefaction Problems (VELACS): Proceedingsof the International Conference, Davis, Calif., 17–20 October1993. Edited by K. Arulandan, R.F. Scott, and X. Zeng. A.A.Balkema, Rotterdam. Vol. 1, pp. 3–17.

Taboada, V., and Dobry, R. 1993b. Experimental results of modelNo. 2 at RPI. In Verification of Numerical Procedures for theAnalysis of Soil Liquefaction Problems (VELACS): Proceedingsof the International Conference, Davis, Calif., 17–20 October1993. Edited by K. Arulandan, R.F. Scott, and X. Zeng. A.A.Balkema, Rotterdam. pp. 277–294.

Taboada, V., and Dobry, R. 1998. Centrifuge modeling of earth-quake-induced lateral spreading in sand. Journal of GeotechnicalEngineering, ASCE, 124: 195–206.

Takeshita, S., Takeishi, M., and Tamada, K. 1995. Static liquefactionof sands and its liquefaction index. In Proceedings of the 1stInternational Conference on Earthquake Geotechnical Engi-neering, Tokyo, Japan, 14–16 November 1995. Edited by K.Ishihara. A.A. Balkema, Rotterdam. Vol. 1, pp. 177–182.

Thevanayagam, S. 1998. Effect of fines and confining stress onundrained shear strength of silty sands. Journal of Geotechnicaland Geoenvironmental Engineering, ASCE, 124: 479–491.

Tsukamoto, Y., Ishihara, K., and Sawada, S. 2004. Settlement ofsilty sand deposits following liquefaction during earthquakes.Soils and Foundations, 44: 135–148.

Uzuoka, R., Sento, N., and Kazama, M. 2003. Numerical analysison liquefaction-induced progressive deformation with pore waterpressure migration. In Proceedings of the 3rd International Sym-posium on Deformation Characteristics of Geomechanics, Lyon,22–24 September 2003. Edited by H. Di Benedetto, T. Doanh,H. Geoffroy, and C. Sauzét. A.A. Balkema, Rotterdam.pp. 1095–1101.

Vaid, Y.P., and Eliadorani, A. 1998. Instability and liquefaction ofgranular soils under undrained and partially drained states.Canadian Geotechnical Journal, 35(6): 1053–1062.

Vaid, Y.P., Chern, J.C., and Tumi, H. 1985. Confining pressure,grain angularity, and liquefaction. Journal of Geotechnical Engi-neering, ASCE, 111: 1229–1235.

Yang, Z., and Elgamal, A. 2002. Influence of permeability onliquefaction-induced shear deformation. Journal of EngineeringMechanics, ASCE, 128: 720–729.

Yoshida, N., Yasuda, S., and Ohya, Y. 2005. Two criteria forliquefaction-induced flow. In Proceedings of the 16th Interna-tional Conference on Soil Mechanics and Geotechnical Engi-neering (ICSMGE), Performance Based Design in EarthquakeGeotechnical Engineering: Concepts and Research SatelliteConference, Osaka, Japan, 10–15 September 2005. A.A. Balkema,Rotterdam. pp. 109–116.

Yoshimine, M., Nishizaki, H., Amano, K., and Hosono, Y. 2006.Flow deformation of liquefied sand under constant shear loadand its application to analysis of flow slide of infinite slope.Soil Dynamics and Earthquake Engineering Journal, 26: 253–264.

Youd, T.L., Idriss, I.M., Andrus, R., Arango, I., Castro, G., Christian,J., Dobry, J., Finn, L., Harder, L., Jr., Hynes, H.M., Ishihara, K.,Koester, J., Liao, S.S., Marcuson, W.F., III, Martin, G., Mitchell,J.K., Moriwaki, Y., Power, M.S., Robertson, P.K., Seed, R.B.,and Stokoe, K.H., II. 2001. Liquefaction resistance of soils:summary report from the 1996 NCEER and 1998 NCEER/NSFWorkshops on Evaluation of Liquefaction Resistance of Soils.Journal of Geotechnical and Geoenvironmental Engineering,ASCE, 127: 817–833.

Zlatovic, S. 1994. Residual strength of silty soils. D.Eng. thesis,University of Tokyo, Tokyo.

© 2007 NRC Canada


[seid] seismic liquefaction

Documents