Consolidation of a Soft Clay Composite ... a Soft ClayComposite 55 measuring approximately 15 cm in...

Copyright c© 2008 Tech Science Press CMES, vol.23, no.1, pp.53-73, 2008

Consolidation of a Soft Clay Composite: Experimental Results andComputational Estimates

A.P.S. Selvadurai1 and H. Ghiabi2

Abstract: This paper deals with the problemof the consolidation of a composite consisting ofalternate layers of soft clay and a granular ma-terial. A series of experiments were conductedon components to develop the constitutive mod-els that can be implemented in a computationalapproach. The constitutive response of the softclay is represented by a poro-elasto-plastic Camclay-based model and the granular medium by anelasto-plastic model with a Drucker-Prager typefailure criterion and a non-associated flow rule.The computational poro-elasto-plastic model isused to calibrate the experimental results derivedfrom the one-dimensional tests and to establishthe influence of the volume fractions and spa-tial arrangement of the soft clay and the granu-lar material on the consolidation response. Thecomputational developments are extended to ex-amine certain idealized situations involving clayinclusions that are distributed within a granularmedium in a regular arrangement.

Keyword: Soft clay, composite fill, poro-elasto-plastic geomaterials, clay lump, experi-mental study, computational modeling

1 Introduction

Although isotropy and homogeneity are dominantassumptions of geotechnical analysis there are sit-uations where these assumptions are not true rep-resentations of geotechnical practice (Booker andSmall, 1987; Seyrafian et al., 2006; Selvadurai,

1 Corresponding author. William Scott Professor and JamesMcGill Professor. Department of Civil Engineering andApplied Mechanics, McGill University, 817 SherbrookeStreet West, Montreal, QC, Canada H3A 2K6 E-mail:[email protected]; Tel: +1 514 398 6672; fax:+1 514 398 7361

2 Doctoral student

2007a). The paper by Eden (1955) discusses thepresence of varved features in clay specimens ob-tained from Steep Rock Lake, Ontario. Periodicdeposition of eroded material in lacustrine en-vironments can lead to geological deposits withanisotropy in strength and deformability charac-teristics. Examples of natural inhomogeneity inthe deformability characteristics of London Claywas reported by Ward et al. (1965). Inhomo-geneities can also be deliberately introduced toenhance either the load carrying capacity of softclays or to improve their rate of consolidation.Stone columns or lime-stabilized columns are in-stances where the stiffer inclusion provides the re-inforcement for strength and the reduction of set-tlements, while the incorporation of geosynthet-ics and sand layers are there to enhance the rateof consolidation of the construction. The theoret-ical, computational, laboratory experiments andfield studies in this area are far too numerous tooutline in detail. The reader is referred to thearticles, volumes and reviews by Gibson (1967),Gibson et al. (1971), Hughes and Withers (1974),Madhav and Vitkar (1978), Mitchell and Huber(1985), Schweiger and Pande (1986), Canetta andNova (1989), Tan et al. (1992), Lee and Pande(1998) and Selvadurai (2007a).

The major emphasis of this paper relates to a dif-ferent class of soil composite, which is encoun-tered in land reclamation practices and involvesthe use of dredged clay lumps that are interspersedwithin granular materials. A common construc-tion technique for these lumpy clay fills involvesplacing dredged clay lumps directly on top ofother layers, without filling the void space withanother material. As a result, the built up recla-mation fill experiences substantial compressionpurely due to the reduction of the void space, quiteapart from the consolidation of the clay lumps

54 Copyright c© 2008 Tech Science Press CMES, vol.23, no.1, pp.53-73, 2008

themselves, and this renders the technique inef-fective and unreliable unless account is taken ofthe large compression that can be associated withthe void closure. Examples of such large com-pression of fills constructed with clay lumps arereported by Leung et al. (2001) and Robinson etal. (2005). After clay lumps have been depositedin the reclamation fills, two different mechanismswill contribute to the settlement of the lumpy fills.The first mechanism corresponds to the expul-sion of fluid from the voids between lumps, inter-lump voids, while further settlement takes placedue to the decrease in the void ratio of the soillump, referred to as intra-lump voids. Callari andFederico (2000) implemented the double poros-ity approach, first used in the petroleum industryto model the flow in a fractured porous medium,to deal with the consolidation of structured clayfills. In this approach, the highly fissured porousmedium was modeled as a double porosity contin-uous medium, in which the first system of poresrepresents the net of fissures besides the poreswithin the intact lumps. To validate the doubleporosity model, the results of the proposed modelwas compared with that of finite element compu-tations. Furthermore, they conducted a number ofexperiments on structured clay samples consist-ing of an assembly of small cubical clayey blocks,each covered by a thin layer of geotextile. Thecomparison indicated that the theoretical doubleporosity curves reasonably agree with experimen-tal and computational results, especially for set-tlement evolution. Yang et al. (2002) also usedthe double porosity approach to present an ana-lytical solution for an idealized problem wherea linear elastic soil subjected to self-weight andsurcharge pressure. Subsequently, Yang and Tan(2005) implemented this formulation in finite el-ement analysis considering non-linear void ratioagainst permeability and effective stress relationsto ensure a more realistic prediction. This imple-mentation was then used to analyze results fromthe two one-dimensional consolidation tests ona lumpy clay layer. Nogami et al. (2004) alsodeveloped a numerical method for consolidationanalysis of lumpy clay fills, adopting the doubleporosity model and the meshless method basedon “radial point interpolation” approach. The re-

sults of the proposed computational approach in-dicated good agreement with the results of finiteelement analysis which has been previously for-mulated for consolidation of fissured clay. An at-tempt was then made to model the behavior of thelumpy samples tested in the national universityof Singapore (Wong, 1997); it was indicated thatthe double porosity model can reasonably predictthe consolidation behavior of lumpy fills providedthat proper material properties are given.

For a typical clay lump construction using, say,1.5 m diameter clay lumps in cubic packing, thevoid closure can result in a compression strain ofapproximately 57%, which would be regarded asunacceptable if the fill is directly used for sup-port of foundations. As a practical method to re-solve this problem, the constructed fills are usu-ally subjected to a surcharge load that acceleratesthe closure of the inter-lump voids (Karthikeyanet al., 2004). In order to achieve an effective com-pression from such a preloading process, the sur-charge should be placed over the fill after the dis-sipation of any negative pore pressure that mightbe generated in the lumps due to the removal ofoverburden pressure particularly during dredgingoperation (Robinson et al., 2004). The construc-tion time, therefore, could become longer depend-ing on different factors such as size of the dredgedlumps and the hydraulic conductivity of the fine-grained soil. An alternative to the aggregation ofclay lumps is to fill the void space with a granu-lar material that possesses a higher stiffness anda greater permeability thereby enhancing the loadcarrying capacity and minimizing the settlementsand rate of consolidation of the composite, due tothe reduction in the length of the drainage paths.

Reclaimed land that is constructed using claylumps in conjunction with an interspersed gran-ular fill is a soil composite and the mechanics ofsuch composites are influenced by the constitu-tive properties of the individual phases, the inter-faces between the phases and the associated vol-ume fractions. This paper investigates the con-solidation behavior of a composite consisting ofdisks of clay interspersed with layers of Ballotini.The experimental investigations are performed ona bench-scale oedometer with overall dimensions

Consolidation of a Soft Clay Composite 55

measuring approximately 15 cm in diameter and17 cm in height, with stacks of saturated clayinterspersed with Ballotini of varying thickness.The experimental results are evaluated using aporo-elasto-plastic model for the saturated clayand an elasto-plastic model for the granular lay-ers. The constitutive models are incorporated inthe ABAQUS code in order to validate the com-putational approach to be used in further compu-tational studies.

2 Materials and constitutive models

2.1 The silty clay

The fine-grained component of the compositespecimen was reconstituted from lumps of siltyclay obtained from an excavation located on Sher-brooke Street East (Montréal). A complete de-scription of the preparation of the silty clay usedin the bench-scale experiments is given by Ghiabiand Selvadurai (2006a).

The constitutive response of soft silty clays canbe modeled by appeal to a variety of constitu-tive models (Desai and Siriwardane, 1984; Darve,1990; Cambou and di Prisco, 2000). In this pa-per a Cam-clay model is chosen and the strengthand stiffness parameters of the silty clay were de-termined from the results of CU (ConsolidatedUndrained) triaxial and conventional consolida-tion tests (Ghiabi and Selvadurai, 2006a). Fromthe results of the two oedometric consolidationtests conducted on the reconstituted Montréalsilty clay, the slope of the hydrostatic compressionand swelling lines of the e vs. ln p curves wereestablished at λ =0.0564 and κ=0.0104, while thestress ratio parameter, M, estimated from resultsof CU Triaxial Tests, was 1.55. Furthermore,using the method presented by Casagrande, thehydraulic conductivity of the silty clay was esti-mated from the results of the oedometer test tobe in the range 1.3×10−8 cm/sec to 7.2 ×10−8

cm/sec.

2.2 The granular medium

Ballotini, an artificial particulate medium con-sisting of glass beads manufactured by Pot-ters Industries Inc. (La Prairie, Quebec;

www.pottersbeads.com) was used as the granu-lar component of the composite specimens. TheBallotini used was type A3 with a specific grav-ity estimated at GS=2.5. The grain size analysisconducted on the supplied material indicated that83% of this granular material ranged from 600μmto 850 μm, while approximately 17% had a parti-cle size between 300μm and 600μm.

The constitutive behavior of particulate materialssuch as ballotini can be represented by a varietyof models ranging from non-linear elastic mod-els to hypoplasticity (Hill and Selvadurai, 2005).The range of applicability of a particular modelwill ultimately depend on the mode of applicationof the constitutive model and the ease with whichthe parameters describing the model can be deter-mined through conventional tests. In this study,the mechanical behavior of the ballotini is mod-eled by appeal to a non-linearly elastic-perfectlyplastic constitutive model with a non-associatedflow rule. The specific model used to describefailure is a Drucker-Prager failure criterion de-fined by

f (σi j) =√

J2 +α1I1 −β (1)

and the plastic potential is chosen as

g(σi j) =√

J2 +α2I1 (2)

where I1, J2 are, respectively, the first and secondinvariants of the stress deviator tensor referredto effective stresses. The parameters of the con-stitutive model were determined from the resultsof a series of consolidated drained (CD) triaxial,bench-scale oedometer and isotropic compressiontests conducted on ballotini specimens. The re-sults indicated that the elastic stiffness of ballo-tini is highly dependant on the physical states ofthe material; a relation is presented as a bi-linearexpression that defines the modulus of elastic-ity as a function of the void ratio. To incorpo-rate the non-linear elastic approach with a plas-ticity concept, a computational procedure was de-veloped: The first modification involved the sub-incremental procedure in calculation of the elasticdeformation when the stress state lies within theyield surface. Hence, the elastic stiffness tensor[Ce] was updated in each sub-increment based on


the constitutive relations presented for the stiff-ness parameters. After the stress state reachedthe yield surface, further stress increments werecalculated considering the elastic-plastic stiffnesstensor [Cep]. In each sub-increment, the state ofstress, determined from the new stiffness tensorand strain increment, was scaled back to the yieldsurface. When the nonlinear elastic behavior isconsidered for the material, this tensor was recal-culated in each sub-increment based on the stateof stress and state of strain, which can change theconstitutive characteristics of the material. Thehydraulic conductivity is another important pa-rameter required in pore pressure diffusion- stressanalysis. This parameter is estimated at 1.6×10−3

cm/s from the results of the constant head perme-ability tests conducted on the ballotini specimens.The parameters for the constitutive models usedin the computational study are given in Table 1.

3 Experimental investigations

The experimental program described in this pa-per studies the behavior of a composite speci-men, consisting of a coarse-grained medium inter-spersed with disk-shape clay inclusions subjectedto bench-scale oedometer tests. The samples werefabricated in a consolidation cell with differentnumbers of silty clay disks placed within a granu-lar region. The results of the oedometric compres-sion test are used to assess the influence of the vol-ume proportion of the fine-grained region, in re-lation to the granular medium, on the mechanicalresponse of the entire sample. The experimentsalso provide useful data for comparison with thenumerical results to determine the predictive ca-pabilities of a computational approach.

3.1 Bench-scale experiments

The bench-scale experiments were designed fortesting the composite soil specimens in which acontrolled axial stress from static weights initi-ates the oedometric consolidation of the compos-ites (Figure 1). A weight rack was designed toaccommodate weight of up to 2.67 kN that couldbe added in loading increments. The initial stresson the sample was approximately 81 kPa and thiswas increased in subsequent loading sequences to

Table 1: Constitutive parameters used in compu-tations

Ballotini (non-linear elastic perfectly plastic)Elastic modu-lus

e > 0.605 27000

E(kPa) e < 0.605 (6.24-10.25e)×106

Poisson’s ratio 0.3α1 (Eq. 1) 0.298β (Eq. 1) 0α2(Eq. 2) 0.165Hydraulicconductivity k(cm/s)

1.6×10−3

Clay (Cam-clay model)Hydrostatic compression index λ 0.0564Swelling index κ 0.0104Poisson’s ratio 0.3Stress ratio M 1.55Hydraulic conductivity k (cm/s) 2×10−8

149, 219, 273, 383 and 493 kPa by adding weightsin five loading steps. To minimize the develop-ment of dynamic effects in the loading process,the load was applied gradually using a scissor jackinstalled in the loading frame. Each load steplasted for approximately 24 hours, after which noexcess pore pressure was expected to remain inthe system; this limit was arrived at through acomputational modeling of the experimental con-figuration.

The instrumentation used in the experimentalsetup included a load cell located on the top of theplunger and two LVDTs connected to the rectan-gular plate to measure the uniaxial compressionof the sample (Figure 1). The load cell used inthis experiment had a capacity of 10 kips (44.48kN) with an accuracy equal to ±0.04% FS. TwoLVDTs, with a displacement range of 2.5 cm,were attached to opposite sides of the plunger torecord the vertical compression of the specimen.

3.2 Sample preparation

Considering the consistency of the reconstitutedclay paste, a PVC mold (internal diameter 12.5cm, height 1.94 cm) with a PVC base plate


Figure 1: Loading setup and instrumentations used in the bench-scale test

(b) (a) (c) (d)

Figure 2: Preparation of the silty clay disks (a) PVC mold (b) Compacting clay paste with plunger (c)Trimming the clay disk using wire saw (d) Placing the molded clay in a sealed bag

was used to fabricate the clay discs (Figure 2a).To avoid air entrapment within the specimen, aplunger was used to slightly compact the soilpaste inside the mold (Figure 2b) and the topof the sample was trimmed using a wire cutter(Figure 2c). The mold containing the specimenswere then placed in a sealed bag (Figure 2d) andstored at a temperature of -10◦C for 24 hours,after which the frozen clay discs were removedfrom the mold. The shape of the discs and theirwater content was maintained by re-freezing themin sealed bags. The moisture content determinedfrom tests on frozen clay disks indicated that thespecimens were at, approximately, the liquid limitof the silty clay (25%). This value is lower thanthe range of 30% to 50% which is reported asmoisture content of the dredged lumps in differ-ent reclamation projects. Due to the high silt con-tent, the reconstituted silty clay has a low liquidlimit and plasticity index; therefore, the soil with

30% to 50% moisture content will be in form ofslurry. Since the clay lumps dredged from theseabed tend to have the consistency of a slurry,the soil paste with approximately 25% moisturecontent was used in the experimental study, whichshowed good consistency during sample prepara-tion.

The composite soil specimens were prepared inthe cylindrical cell of the bench-scale setup withan internal diameter of 15.2 cm. The samples (ap-proximate height 16.8 cm) contained either four,six or eight disks of clay with the volume fractionevaluated at 31%, 47% and 62.5%, respectively.The disks were spaced at equal distances and theballotini was dispersed and slightly compacted inlayers to construct the soil composite (Figure 3);the amount of ballotini used in the composites wasmeasured for estimating the initial physical stateof the granular part of the composite. The com-posite samples were saturated with deaired wa-


(b) (a) (c) (d)

Draining vent

Clay disks

ballotini

base plate

Water inlet

geotextile

Stainless steel

container

16.8 cm

6.25 cm

4.2 cm

1.94 cm

2.8 cm

1.94 cm

2.1 cm

1.94 cm

Plunger

O-ring

7.6 cm

Figure 3: (a) Cross sectional view of the 8-disk composite sample (b) Planar cross section (c) the 6-diskcomposite sample (d) the 4-disk composite sample

ter prior to application of the load, and the waterlevel was maintained at the upper level of the sam-ples for a period of two days. During applicationof the axial load, proper drainage was assured byplacing two layers of geotextile at the upper andlower surfaces of the samples, which were con-nected to the draining vent and water container,respectively.

3.3 Experimental results

The measurement of the axial compression of thecomposite specimens forms an important compo-nent of the experimental results that will be usedin the calibration of the computational predic-tions. A further bench-scale test was conductedon a pure ballotini sample to assess the influenceof the volume proportion of clay inclusions on themechanical properties of the composites. Figure 4shows a comparison of the time-dependent settle-ment response of the three samples of the compos-ites compared with that of a pure ballotini spec-imen. The tests indicate that an increase in thevolume proportion of clay results in a higher ax-ial compression of the composites; the axial set-tlement of the four-disk sample at the end of thefinal load step was 2.8 times that of a pure bal-lotini specimen. Furthermore, at the terminationof the last load step, the eight-disk sample ex-hibits approximately five times more axial com-pression than that experienced by a pure ballotinisample. The experimental results also show a pro-gressive axial compression at the end of each load

step: This phenomenon is mainly related to thetime-dependent compression of ballotini due tothe internal re-arrangement of the granular assem-bly. Such time-dependent behavior in granularmedia, particularly when they are used in low den-sity conditions in laboratory investigations, hasbeen observed, among others, by Murayama etal., (1984), Lade (1994), Di Prisco and Imposi-mato (1996) and Kuwano and Jardine (2002) andfurther discussed in the article by Ghiabi and Sel-vadurai (2006b). Comparing the results of dif-ferent tests, it was shown that this phenomenonwas diminished in the tests conducted on fabri-cated samples with a lower volume of ballotini;therefore, in the six- and eight-disk composites,an ultimate compression state was observed at theend of the load steps.

0

2

4

6

0 1 2 3 4 5

Time (day )

Set

tlem

ent

(mm

)

8-diskcomposite

6-diskcomposite

4-diskcomposite

Homogeneousballotini

Figure 4: Axial settlement of the samples in fivesuccessive load steps


During the axial loading of the soil composites,a mixture interface can develop at the boundaryof the two materials as a result of penetration ofthe ballotini particles into the soft clay. In thisstudy, the development of this mixture interfacewas investigated by dissecting the composite sam-ples at the termination of the tests. These inves-tigations indicated that there was no measurablepenetration of the ballotini grains into the soft claydisks at this stress level, although a layer of ballo-tini beads had adhered to the surface of the disks;these could be easily removed from the clay sur-face. The issue of particle penetration and its in-fluence on the consolidation response of the clayin terms of alteration of the fluid transport bound-ary conditions at the interface was recently in-vestigated by Selvadurai (2007b), using an ide-alized model of contact of discs at an interface.For both hexagonal and cubic arrangements of thecontacting discs, the consolidation rate exhibits amarginal change but the total consolidation defor-mations are uninfluenced.

4 Computational modeling of fluid-saturatedmedia

4.1 Constitutive equations

Many geotechnical engineering problems includetime-dependent mechanical responses of soil me-dia associated with the expulsion of the pore fluidpressure, which depends on hydraulic propertiesof the material and the associated boundary con-ditions. To account for this behavior it is neces-sary to couple the equations governing mass con-servation, the flow of pore fluid through the soilskeleton, the constitutive equations governing themechanical behavior of the soil skeleton and theequations of equilibrium.

In this section we examine the poro-elasto-plasticmodeling of the time-dependent response of theclay composite. The poro-elasto-plastic model-ing extends the classical theory of poroelasticitydeveloped by Biot (1941) to include effects offailure of the porous soil skeleton and the atten-dant alterations in the stiffness and fluid transportproperties of the soil skeleton. The extensionsof poroelasticity to include failure, yield, damage

and viscoplasticity of the soil skeleton has foundmany useful applications, particularly in the areaof geomechanics, biomechanics, energy resourcesexploration and environmental geotechnique. Theliterature in these areas is extensive and no at-tempt is made to provide a comprehensive surveyof the topics. The review articles and volumesby Cheng and Dusseault (1993), Coussy (1995),Selvadurai (1996, 2007b), Lewis and Schrefler(1998), Thimus et al. (1998), Auriault et al.(2002), Ehlers and Bluhm (2004) and Selvaduraiand Shirazi (2005) can be consulted for up-to-dateliterature.

Under the assumption of small-strain theory, theequation of equilibrium for quasi-static conditionscan be written as

∇.σσσ +ρg = 0 (3)

where σσσ is the total stress tensor and ρ is the av-erage density of the multiphase system defined as

ρ = (1−n)ρ s +nρw (4)

where ρ s and ρw are, respectively, densities of thesolid phase and water, n is the porosity and g isan acceleration vector usually related to the grav-itational effects. Considering the Darcy’s law forthe flow through the porous skeleton and the in-compressibility of the pore fluid, the mass conser-vation equation for the pore fluid in a saturatedmedium can be stated in the form

∇.vs +∇.

[k

μw (−∇Pw +ρwg)]

= 0 (5)

where vs is the velocity vector of soil skeleton, kis the permeability tensor and Pw and μw are theisotropic pressure and viscosity of the fluid phase,respectively. The constitutive relationship of theskeleton is the additional equation that has to beconsidered in the analysis of the porous medium.Using the principle of effective stress, the consti-tutive equation can be written in the incrementalform:

Δσσσ = D′Δεεε +Δσσσ f (6)

where D′ is the effective elasto-plastic consti-tutive tensor, Δσσσ and Δεεε are incremental to-tal stress and incremental strain tensors, ΔσσσT

f ={ΔPw,ΔPw,ΔPw,0,0,0} in which ΔPw is thechange in pore fluid pressure.


4.2 Finite element formulation for consolida-tion problem

Complete expositions of the general approachfor computational modeling of poro-elasto-plasticmedia are given by Lewis and Schrefler (1998),Potts and Zradkovic (2001) and in the manualsassociated with the general purpose finite elementcode ABAQUS (ABAQUS/Standard, 2006). Thesalient equations are briefly summarized for com-pleteness.

The equilibrium condition can be expressed byconsidering the principle of minimum potentialenergy. The essence of the finite element methodis to discretize the domain of interest into ele-ments; therefore, minimizing the sum of the po-tential energy of the elements with respect toincremental nodal displacements gives a set ofequations of the form:

KGΔdGn +LGΔPwG

n = ΔRG (7)

where KG is the global stiffness matrix, LG isthe off diagonal sub-matrix in the consolidationstiffness matrix, ΔRG is the incremental vector ofglobal nodal forces, and ΔdG

n and ΔPwGn contain

the list of nodal displacements and fluid pressureof all nodes of the domain, respectively. Using theprinciple of virtual work, the governing equationcan be written in finite element form as:

LGT(

ΔdGn

Δt

)−ΦGPwG

n = nG +Q (8)

where Q represents any sources or sinks, and ΦG

is the permeability sub-matrix in the consolida-tion stiffness matrix defined as:

ΦG =N

∑i=1

(ΦE)i =N

∑i=1

⎛⎝∫

V

ETkEγ f

dV

⎞⎠

i

(9)

and

nG =N

∑i=1

(nE)

i =N

∑i=1

⎛⎝∫

V

ETkiGdV

⎞⎠

i

(10)

where the matrix E contains derivatives of thepore fluid interpolation function, γ f is the unit

weight of the fluid, iG is the unit vector with op-posite direction to gravity and T denotes the trans-pose. In this study, the general purpose finite el-ement program ABAQUS/Standard was used toperform numerical simulations of the mechani-cal behavior of composite models composed oftwo different materials. This program has the ca-pability to characterize constitutive properties ofmaterials using a user-defined material subrou-tine (UMAT). ABAQUS implements the directapproach in coupling the effect of transient flowof the pore-fluid and the stress-strain behavior ofthe soil skeleton; thus, Equations (7) and (8) canbe written in the following incremental form:[

KG LG

LGT −β ΔtΦG

][ΔdG

n

ΔPwGn

]

=

{ΔRG(

nG +Q+ΦG(

PwGn

))Δt

}(11)

where β is the parameter used for the approxima-tion in the flow equation. In order to ensure stabil-ity of the time marching process, it is necessary tochoose β ≥ 0.5(see e.g. Selvadurai and Nguyen,1995; Lewis and Schrefler, 1998). ABAQUS con-siders a fully implicit scheme (β = 1) to ensurethe numerical stability of the analyses. Theseequations form the basis of the iterative solutionof a time step in a coupled flow deformation anal-ysis in ABAQUS/Standard.

4.3 Computational analysis of soil composites

The computational modeling of the oedometriccompression of the composite samples was con-ducted using the general-purpose finite elementcode, ABAQUS/ Standard. Composite modelsthat include both ballotini and clay disks are ax-isymmetric models whereas the UMAT subrou-tine used for ballotini was based on a three-dimensional formulation. In order to reduce thecomputational time for modeling the compositeregion, only a sector region with an angle of22.5◦was modeled. This region could be easilydiscretized into brick elements. The finite ele-ment mesh discretizations of the composite do-mains used in this analysis are shown in Figure 5.In this study a 20-node brick element (C3D20RP)


Figure 5: (a) Schematic figure of 4-disk sample (b) mesh configuration and boundary conditions- 4-disksample (c) mesh configuration- 6-disk sample (d) mesh configuration- 8-disk sample

with quadratic variations in the displacement andlinear variations in the pore pressure was selected,which employed a reduced integration scheme.The boundary conditions in the modeling, whichwere defined to represent the oedometric condi-tion, are shown in Figure 5. The lateral bound-aries were constrained horizontally with no pro-vision for drainage; since drainage was providedthrough the geotextile layers and drainage vents,the upper and lower surfaces of the specimenswere assumed to be free draining. Furthermore,the pore pressure at the upper and lower surfaceswas assigned as zero and 1.65 kPa, respectively,since during the experiments the water level in thecontainer was maintained at the same level as thetop surface of the specimens.

Another phenomenon that takes place in lumpyfills is swelling of the clay lumps due to the suc-tion that occurs when the clay lumps are sub-jected to a stress level after being consolidatedunder higher stress state (Robinson et al., 2004).In composite fills, it is assumed that the granularfiller material is placed immediately after place-ment of the clay lumps. In other words, the suc-tion process progresses while the lumps are en-closed within the granular phase. As a result, theexistence of the filler phase can constrain the freeswelling of the lumps due to confining stressesinduced by the interaction between the two re-

gions. The magnitude of the contact stress andthe extent of the swelling process are, therefore,related to the stiffness characteristic of both mate-rials. In this research, the material representingthe fine-grained component, is a normally con-solidated silty clay reconstituted in the laboratory(Ghiabi and Selvadurai, 2006a). The clay lumps,prepared from reconstituted clay paste, contain nonegative pore water pressure and, therefore, theswelling process was not considered to be impor-tant in the experimental investigations and com-putations conducted on composite specimens aswell as on composite lumpy fills.

The composite models were subjected to a uni-form pressure applied to the upper surface, whichincreased from 81 kPa (initial load step) to 149,219, 273, 383 and 493 kPa in six loading steps.The contact stresses between the plunger and theupper surface of the soil composite was mod-eled using the “rigid body” constraint option inABAQUS.

In coupled pore fluid flow and stress analysis, thedissipation of the fluid pressure, initially gener-ated within the pores of a porous media, gives riseto a gradual transfer of stress to the solid skeleton;this time-dependent response is required for fulldissipation of the excess pore pressures. Figure6 shows the variation of the pore fluid pressurein the last loading step for three points selected


from different locations of the four-disk compos-ite sample (Figure 5), and indicate that the porefluid pressure rises during the application of theaxial load (the first second of the load step). Thefluid pressure at point C, located inside the claydisk, increases by a greater magnitude than thatcomputed for points A and B.

0

10

20

30

40

50

60

1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05

Time (second )

Po

re p

ress

ure

(k

Pa

) Point A

Point B

Point C

Last load step

Figure 6: Pore fluid pressure change in differentlocations of four-disk composite model

After the axial pressure was applied in each loadstep, the axial load was maintained for 24 hours.The variation of pore fluid pressure at points Aand B shows that no measurable excess pore pres-sure remained in the ballotini region 10 secondsafter reaching the maximum axial load. Also, dis-sipation of the excess pore pressure in ballotiniresulted in compression of the ballotini layers be-tween the clay disks. The pore fluid pressure inthe clay region was therefore primarily decreaseddue to a stress relaxation encountered in the cen-tral core of the composite, which can be verifiedby considering the variation of contact pressurecomputed at the top surface of the sample duringany load step. Figure 7 shows the contact pres-sure change at points D and E (Figure 5) of thefour-disk composite model during the last loadstep; here the vertical pressure applied to the cen-tral core gradually decreased, resulting in a riseof stress level within the outer annular part of thespecimens. The computations also show that aconsiderable portion of the axial load was carriedby the outer annular region surrounding the cen-tral composite core. In Figure 6 the excess porefluid pressure generated within the clay parts was

completely dissipated within 10000 seconds af-ter the beginning of the load step, at which timethe ultimate axial compression of the compositemodel was achieved.

0

200

400

600

800

1000

1200

1400

1600

1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05

Time (second )

No

rma

l st

ress

(kP

a)

Point D

Point E

Figure 7: Variation of contact stress in four-diskmodel (last load step)

A progressive compression of the saturatedporous samples occurred as a result of thetime-dependent consolidation process in the con-stituents; the normalized axial compression in thelast load step is presented in Figure 8. The com-putations show two different phases in the com-pression response of the composite; the first phaseoccurred due to dissipation of the fluid pressuretrapped within the pores of the ballotini region,while the second phase was a result of the expul-sion of the pore fluid from the pores of the clayinclusions. In addition, computational results in-dicate that in each load step 70% to 80% of thetotal settlement occurred during the first phase ofthe consolidation, which is referred to as the ini-tial settlement.

4.4 Comparison between numerical predic-tions and experimental results

Computations were performed using the finite el-ement models shown in Figure 5. The constitu-tive models developed for the constituents wereimplemented using the initial physical state of thematerials, measured during the fabrication of thecomposite specimens. Figure 9 shows the com-parison of the displacement-time responses be-tween the computational predictions and the ex-


0

0.2

0.4

0.6

0.8

1

1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05

Time (second)

4-disk composite

6-disk composite

8-disk composite

∞U

Ut

Phase 1 Phase 2

Figure 8: Axial compression of the compositemodels in the oedometric test (last load step)

perimental results. For the purposes of com-parison, results are presented for three compos-ite specimens (four-disk, six-disk, and eight-disksamples) subjected to an axial load that was in-creased in five consecutive load steps. It canbe seen that the computational estimates predict,reasonably accurately, the experimental trends ofload-displacement response for the three compos-ite specimens tested. The axial settlement of thecomposite at the end of the last load step, pre-dicted from computations is approximately 5% to10% lower than the experimental responses. The

0

2

4

6

0 1 2 3 4 5

Time (day )

Set

tlem

ent

(mm

)

Experimental results

Computational results

6-disk

composite

4-disk

composite

8-disk

composite

Figure 9: Axial compression of composite

hydraulic conductivity of the silty clay was es-timated using the consolidation curves obtainedfrom conventional oedometer tests. The experi-mental results indicated that this parameter is de-

pendent on the void ratio of the material; thehydraulic conductivity evaluated from the firstload step, in which the void ratio of the speci-men reduced to about 0.57, was approximately7.2×10−8cm/s while it decreased to approxi-mately 2×10−8cm/s at the void ratio close to 0.46.Furthermore, in the central core of the compos-ite specimens, that contained both soft silty claydisks and ballotini, a high proportion of the axialcompression of the composite core occurs in theclay disks, and not in the ballotini phase placedbetween them, due to the relative stiffness proper-ties of these constituents. Although higher com-pression was obtained by placing more clay diskswithin the composite, increasing the number ofdisks in the core region results in lower compres-sion of each disk due to the increase in the volumeof the soft soil. Therefore, lower compression ofthe clay inclusions gave rise to higher void ratioand hydraulic conductivity of the material.

Therefore, the computation attempts to predictthe mechanical response of the composite sam-ples considering two different values for hydraulicconductivity. The lower value was selected at2×10−8 cm/s corresponding to void ratio of 0.46,and the higher limit was assigned a value 10 timeshigher than the lower one, which could be an es-timation for the hydraulic conductivity when thesoil has a void ratio higher than 0.57 (the ini-tial void ratio of the normally consolidated siltyclay is approximately 0.65). The results of thecomputational estimates and experimental inves-tigations for the settlement-time curves are pre-sented in Figure 10. Numerical estimations indi-cated that, after 10,000 sec, no excess pore pres-sure remained within the composites; therefore,it could be concluded that the ultimate consol-idation settlement was achieved. In the four-disk sample reasonable correlation was obtainedfor the consolidation curves, although using thevalue of kclay = 2 × 10−8 cm/s gave better pre-dictions. In the six-disk and eight-disk samples,the experimental settlement curves shifted towardthe numerical results obtained using the higherhydraulic conductivity for the silty clay material(kclay = 2×10−7cm/s). Furthermore, in the eight-disk composite, there was a lower rate of consol-


idation during the first phase of the process. Thisphenomenon can be attributed to the slow dissipa-tion of excess pore pressure generated in the bal-lotini layers located between the clay disks. Sincethe clay disks were placed approximately 1.6 mmapart, it is likely that contact between the disks re-sults in the entrapment of fluid within the thin lay-ers of ballotini. This results in a slower transfer ofthe load to the circumferential ballotini region andtherefore a delay in the progress of the first phaseof the consolidation process.

The experimental observations indicated that thetime-dependent compression of ballotini results infurther settlement of the sample after the termina-tion of the consolidation process, a phenomenonthat is more observable in the four-disk compos-ite due to its higher volume proportion of ballo-tini. The reduction in the amount of ballotini inthe six-disk and eight-disk specimens results innegligible time-dependent axial deformation afterthe termination of consolidation process.

5 The computational simulation of lumpy fillcomposites

In this computational study, different compositefills are simulated to investigate the influence ofvarious parameters, on the consolidation responseof lumpy fill composites. The dimensions of thecomposite domain and clay inclusions were se-lected by examining several typical projects re-ported in the literature (Hartlen and Ingers, 1981;Leung et al., 2001; Karthikeyan et al., 2004).

5.1 The composite lumpy fill

Using available information on the dimensions ofthe lumpy fills in different reclamation projects,cube-shaped lumps 1.5 m in width and 1.4 min height (volume≈3.15 m3) in a composite do-main with a depth of 5.6 m was considered in thecomputational modeling; this domain can enclosethree composite lumpy layers separated by thethin layers of ballotini. The composite mediumis loaded by a 6 m square rigid footing placed onthe surface of the fill. Due to symmetries associ-ated with the problem only one-quarter of the do-main is modeled, and to reduce the computational

0

0.2

0.4

0.6

0.8

1

1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05

Time (sec )

Norm

ala

ized

set

tlem

ent


Numerical results

Numerical results

(kclay=2×10-8

cm/s)

(kclay=2×10-7

cm/s)

4-disk sample

0

0.2

0.4

0.6

0.8

1

1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05

Time (sec )

No

rma

lize

d s

ettl

emen

t


Numerical results

Numerical results

(kclay=2×10-8

cm/s)

(kclay=2×10-7

cm/s)

6-disk sample

0

0.2

0.4

0.6

0.8

1

1.E-02 1.E-01 1.E+00 1.E+01 1.E+02 1.E+03 1.E+04 1.E+05

Time (sec )

Norm

ali

zed

set

tlem

ent


Numerical results

Numerical results

(kclay=2×10-8

cm/s)

(kclay=2×10-7

cm/s)

8-disk sample

Figure 10: Consolidation curves of the three com-posite samples during the last load step σv =493kPa

costs of performing the coupled pore fluid diffu-sion and stress analysis associated with an elasto-plastic skeletal response, a domain with specificdimensions was selected where the boundary in-terface condition would not contribute apprecia-ble influences on the results. In this numericalstudy, the size of the fill in both horizontal direc-tions was set to 30 m with the vertical boundarieslocated 15 m away from the center of the footing.


Also in this study, the layers containing both bal-lotini and clay material were designated as com-posite layers. The composite fill can accommo-date three composite layers; each can contain dif-ferent volume portions of clay and ballotini. Fig-ure 11 shows one of the composite layers that con-tains 25% volume proportion of the clay material,the remainder being the ballotini phase.

Figure 11: Composite layer with 25%volumefraction of clay

In land reclamation projects, a layer of surchargeis used to accelerate the consolidation process(Hartlen and Ingers 1981; Karthikeyan et al.2004). The surcharge pressure and the self-weightof the deposited material are the loads that inducethe initial state prior to the application of imposedstructural loads, which in this case corresponds toa rigid footing. This computational study exam-ines the mechanical responses of the compositemodels subjected to different loading steps. Thefirst loading step includes the self-weight and asurcharge pressure of 90 kPa, (approximately cor-responding to a layer of sand 5 m in height witha unit weight of 18 kN/m3) gradually applied tothe surface over a one day period. In the secondloading step, a uniform pressure is applied to thesquare rigid footing; this uniform pressure is in-creased to 1500 kPa over a period of 25 min (rateof 1 kPa /sec ) and then decreased to zero duringthe unloading step, which accounts for the irre-versible deformation that occurs in the compositefill. In the reloading step the pressure is increasedto 1500 kPa over a 25 min period, followed bythe consolidation step, where the pressure on thefooting is maintained at 1500 kPa, to allow for ex-

pulsion of excess pore pressures generated in theclay lumps and in the ballotini region. The dura-tion of this consolidation step is set equal to 200days after which there is virtually no excess porepressure in the composite layer.

5.2 Composite fill with one composite layerplaced at different locations

As the first part of the computational study, a sin-gle composite layer is placed at three different lev-els within the granular fill, denoted as one-layer-B, one-layer-M, and one-layer-T, as shown inFigure 12(a), (b) and (c). The boundary con-ditions chosen to model the displacement andpore pressure at the boundaries of the domainare shown in Figure 12 (d). The only perme-able boundary with zero excess pore pressure de-fined in the model is the surface of the fill andall other boundaries are considered to be imper-meable. The composite domains are discretizedinto 20-node brick elements with quadratic dis-placement and linear pore pressure formulationsusing a mesh discretization of 0.70 m in the ver-tical direction and 0.75 m in the horizontal direc-tion, with all the clay lumps divided into 8 ele-ments. The interface between the clay regions andthe granular medium exhibits complete continu-ity in terms of the displacement and pore pressurefields. The computations can be used to predictthe settlement of lumpy fills due to self weight andthe pressure from the surcharge is used to inducepre-consolidation. These results indicate that softlumps within the granular fill contribute to highersurface settlements, with the maximum occurringin the area over the clay lumps (Figure 13): Inone-layer-B and one-layer-M fills, the profiles ofthe surface are almost identical, whereas the max-imum settlement in one-layer-T fill shows an in-crease of approximately 28% due to the presenceof soft clay lumps close to the surface.

The computations were extended to simulate thesettlement of a 6m×6m rigid footing resting onthe composite domain. The presence of a compos-ite layer in the fill results in 20% to 25% greaterinitial settlement compared to that of a homoge-neous ballotini fill (Figure 14). The results in-dicate that during a loading-unloading-reloading


ballotini

clay lumps

5.6 m 1.5 m 1.4 m3 m

0.7 m

2.1 m

2.1 m

1.4 m

0.7 m

3.5 m

15 m

(a)

(b)

(c)

(d)

Figure 12: The location of composite layer in (a)1-layer-B (b) 1-layer-M (c) 1-layer-T compositefills (d) mesh configuration and boundary condi-tions

cycle, the initial settlement of one-layer-B andone-layer-M fills were about 21.5 cm while ap-proximately 5% less settlement was obtained forthe one-layer-T composite model.

Self weight effects, surcharge pressure and foot-ing loads results in the generation of excess porefluid pressure, particularly within the clay re-gions, and the dissipation of this excess pore fluidpressure induces further settlement of the footingduring the consolidation step (200 days), and des-ignated as the “consolidation settlement”. Fig-

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0 0.5 1 1.5 2 2.5 3 x(m )

sett

lem

ent

(cm

)

granular fill

1-layer-T

1-layer-M

1-layer-B

Region over lumps

Figure 13: Surface settlement due to the applica-tion of gravitational load and surcharge pressure

0

200

400

600

800

1000

1200

1400

1600

0 5 10 15 20 25

settlement (cm )

pre

ssu

re o

n t

he

foo

tin

g (

kP

a) 1-layer-B fill

1-layer-M fill

1-layer-T fill

granular fill

Figure 14: Load-settlement response of the rigidsquare footing resting on 1-layer composite fills

ure 15 compares the consolidation settlement forthe different one-layer models: The first phase iscomplete within the first day and corresponds tothe dissipation of the fluid pressure in the poresof the ballotini skeleton. This is followed by fur-ther settlement resulting from consolidation in theclay inclusions. The largest settlement in this stepis 3.3 mm, occurring in the one-layer-M fill whilethe lowest occurs when the composite layer is lo-cated at the base of the fill (one-layer-B). Theseresults indicate that the maximum total settlementis encountered when the composite layer is lo-cated at the half-depth of the fill.

5.3 Composite fills with different volume pro-portions of clay

To examine the effect of volume proportion ofclay on the mechanical response of the compositefill, two- and three-layer composite fills are mod-


0

0.5

1

1.5

2

2.5

3

3.5

0.00001 0.001 0.1 10 1000time (day )

set

tlem

ent

( mm

)

granular fill

1-layer-B

1-layer-T

1-layer-M

Figure 15: Consolidation settlement of a squarerigid footing resting on 1-layer lumpy fills

eled (Figure 16). In the two-layer composite fill,the distance between two composite layers is 1.4m, which is reduced to 35 cm in the three-layerfill. The mesh discretization and boundary condi-tions of two-layer composite models are identicalto the one-layer model (Figure 12d). The meshdiscretization used for the three-layer model isalso shown in Figure 16d. The computations sim-ulate the settlement of a rigid square footing onlumpy fills. The settlement of the rigid footingincreases with the increase in the number of com-posite layers (Figure 17). As a result, the foot-ing placed on the three-layer lumpy fill displaysthe highest initial settlement, approximately twicethat for a homogeneous fill. Figure 17 also showsirreversible deformation after a complete loading–unloading cycle due to plastic deformation in boththe clay lumps and granular filler: An increase inthe number of lumps gives a higher irreversiblesettlement of the footing. At complete unload-ing, the three-layer lumpy fill had an irreversibledisplacement of 12.2 cm whereas the irreversibledisplacement in the one-layer-B fill was 6.4 cm.The computational results indicate that increasingthe volume of the clay lumps gives a higher ver-tical displacement of the rigid footing during theconsolidation step; the ultimate settlement of therigid footing (in this study after 200 days) in athree-layer fill is approximately twice that of theone-layer-B fill. To investigate the influence ofthe volume proportion of clay on the mechanicalresponse of the fills, the settlement of the footingis presented in terms of degree of consolidationU

defined as:

U =Δ(t)−Δ(0)

Δ(∞)−Δ(0)(12)

where Δ(0) is the vertical settlement of the foot-ing at the beginning of the consolidation step andΔ(∞)is the ultimate settlement (after 200 days).Figure 18 presents two different phases in the set-tlement response of the composite fills: The firstphase corresponds to the quick consolidation ofthe ballotini region and the second phase is at-tributed to the consolidation process in the clayinclusions. By increasing the number of the lumpsin the composite fill, the portion of settlement thatoccurs due to the consolidation of the clay lumpsincreases; in one-layer-B fill this represents 21%of the settlement, which rises to 36% in the three-layer composite region. The results also indicatethat the consolidation settlement is delayed whenthe volume proportion of the clay lumps is in-creased; as a result, the time required for 50% and90% consolidation, t50 and t90 is approximatelythree times longer in the three-layer fill comparedto that of the one-layer-B model.

5.4 Composite fills with densely packed claylumps

The composite layer used in the previous sectionto model the composite lumpy fills contained 25%volume fraction of lumps arranged in a squaregrid pattern (Figure 11). In order to investigatethe influence of volume fraction and arrangementof clay lumps on the mechanical characteristics ofthe composite fills, a dense composite layer, con-taining 50% by volume of lumps arranged in adense packing was considered (Figure 19). Twocomposite models are assembled, each with threedense composite layers located at the same depthsused previously for the three-layer fill (Figure16); the configuration of the dense compositemodels are shown in Figure 20, with either a cubi-cal (three-layer-D1) or a hexagonal (three-layer-D2) packing. Thus, in the three-layer-D2 fill, thelumps in the middle composite layer are only sur-rounded by ballotini. A mesh discretization iden-tical to that shown in Figure 16d for the three-layer model is used for the new dense composite


(d)

ballotini

Clay lumps

5.6 m 3 m

0.7 m

0.7 m

0.7 m

1.4 m

0.7 m

0.35 m

15 m

(a)

(b)

(c)

1.5 m 1.4 m

Figure 16: Locations of the composite layers in(a) 1-layer-B (b) 2-layer (c) 3-layer composites(d) mesh discretization of 3-layer model

models. Other details of the simulations, such asboundary conditions and loading specifications,are identical to those implemented in the previ-ous simulations. Figure 21 presents a comparisonof the surface settlement profiles for three-layercomposite fills (three-layer, three-layer-D1 andthree-layer-D2) due to self-weight and a uni-form surcharge pressure, with the maximum ini-tial surface settlement occurring in the area overthe lumps of the top composite layer. The re-sults indicate that increasing the volume propor-tion of soft clay in the fill results in a higher sur-face settlement. In three-layer-D1 fill the maxi-

0

200

400

600

800

1000

1200

1400

1600

0 5 10 15 20 25 30

settlement (cm )

pre

ssu

re o

n t

he

foo

tin

g (

kP

a)

granular fill

1-layer-B fill

2-layer fill

3-layer fill

Figure 17: Load-settlement response of the rigidfooting resting on different composite configura-tions

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.00001 0.001 0.1 10 1000

time (day )

deg

ree

of

con

soli

da

tio

n U

granular fill

1-layer fill

2-layer fill

3-layer fill

Consolidation settlement

1.65 mm

2.67 mm

4.05 mm

5.03 mm

Phase 2 Phase 1

t50t90

Figure 18: Normalized consolidation settlementof a footing resting on different lumpy fills

mum settlement is 5.1 cm, which is approximately60% higher than that of three-layer compositefill. The maximum surface settlement computedin three-layer-D2 fill is, however, 7.36 cm, indi-cating that larger surface settlement due to gravi-tational loads is obtained from a clay lump pack-ing in a hexagonal form. Figure 22 shows the set-tlement responses of the footing placed on threedifferent composite layers and subjected to uni-form pressure up to a maximum of 1500 kPa. Theinitial settlement of the footing in dense packingcomposite layers is approximately 120% higherthan that estimated for the three-layer compositefill, and can be attributed to the higher volume ofclay present in the dense composite layer. Dur-ing the loading step, the computations estimatethe same magnitude of footing settlement for both


cubic and hexagonal and dense packing config-urations; therefore, the difference of settlementpredicted during the initial load step remains rela-tively unchanged until the termination of the load-ing step. After complete unloading, however, 40.8cm of irreversible settlement is induced in the cu-bic packing three-layer-D1 fill, which is 4.4 cmlower than that for the three-layer-D2 fill.

Figure 19: Locations of the clay lumps in densecomposite layer

The settlement of the footing continues in the 200-day consolidation step due to the dissipation ofthe excess pore fluid pressure generated in thecomposite fills. The results indicate that the con-solidation settlement of the footing in the three-layer-D1 and three-layer-D2 composite fills is33.6 mm and 28.6 mm, respectively, which is ap-proximately 6 times greater than that computa-tionally estimated for the three-layer fill (Figure23). It was also observed that in densely packedfills a higher portion of the consolidation settle-ment occurs in phase 2 due to consolidation ofthe fine-grained lumps. Figure 24 indicates that indensely packed lumpy fills, only 25% of the con-solidation settlement takes place in phase 1 com-pared to 65% in loose packing layers. As a resultthe time period needed for 50% consolidation tooccur in the densely packed fills increases to about2.4 days compared to the 9 minutes estimated forthree-layer fill.

Figure 20: (a) location of section A in compositemodel (b) section A in cubical packing (3-layer-D1) (c) section A in hexagonal packing (3-layer-D2)

6 Concluding remarks

This study considered idealized lumpy soil re-gions where the voids between the lumps werefilled with a type of granular material, whichcan enhance the mechanical characteristics of thelumpy material by eliminating the large void re-gions. The experimental program included bench-scale oedometer tests conducted on compositespecimens with different volume proportions ofclay and these showed that the presence of clayinclusions gives rise to higher axial compression


0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

8.0

0 0.5 1 1.5 2 2.5 3x(m )

sett

lem

ent

(cm

)

granular fill

3-layer

3-layer-D1

3-layer-D2

Region over the lumps of

top composite layer

Figure 21: Profile of the surface settlement due tothe application of self-weight and surcharge load

0

200

400

600

800

1000

1200

1400

1600

0 20 40 60 80

settlement (cm )

pres

sure

on

th

e fo

oti

ng

(k

Pa

)

3-layer fill

3-layer-D1 fill

3-layer-D2 fill

Figure 22: Load-settlement response of the rigidfooting resting on 3-layer lumpy fills

0

5

10

15

20

25

30

35

40

0.00001 0.001 0.1 10 1000

time (day )

set

tlem

ent

( mm

)

3-layer fill

3-layer-D2 fill

3-layer-D1 fill

Figure 23: Consolidation settlement of a squarerigid footing resting on composite lumpy fills

of the samples. A computational approach, val-idated with the experimental results, simulatedthe behavior of a large-scale composite fill of fi-nite thickness subjected to self-weight, surcharge

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0.00001 0.001 0.1 10 1000

time (day )

deg

ree

of

con

soli

dati

on

U

granular fill

3-layer fill

3-laye-D1 fill

3-laye-D2 fill

t50t50

Phase 1 Phase 2

Figure 24: Degree of consolidation for 3-layerlumpy fills subjected to a footing load

pressure, and loaded by a rigid footing. Computa-tions showed that the response of a composite fillsubjected to self-weight and surcharge pressure,is significantly influenced by the location of thelumps: The closer the lumps are to the surface themore uneven the surface settlement profile. Theresults indicted that the packing configuration ofthe lumps (hexagonal or cubical) can also inducedifferent magnitudes of surface settlement.

The load-displacement response of a square rigidfooting placed on composite finite-layer with var-ious volume proportions of clay in relation to thegranular phase was also examined. Increasing thevolume proportion of clay inclusions gives rise tohigher initial settlements and irreversible defor-mations take place during the loading and com-plete unloading cycles, respectively. After reload-ing and maintaining the pressure, the footing dis-plays further settlements as a result of consolida-tion in the constituent soils within the compos-ite fill. Increasing the volume proportion of fine-grained inclusions, both by placing more lumps ineach composite layer, or by increasing the num-ber of composite layers raises the settlement ofthe footing during the consolidation step. Further-more, in composite layers with a higher volumeof clay, the time-dependent settlement of the foot-ing is delayed, which is attributable to the lowerhydraulic conductivity of the clay material com-pared to that of the filler soil.

The experimental and computational results canbe used to investigate the serviceability of com-


posite subsoils subjected to different types ofloads. The most important case would be for re-claimed lumpy fills where the inter-lump voids arefilled with a granular phase. The proposed com-putational scheme can be implemented to predictthe mechanical behavior of the composite layersincluding the settlement and the time required toachieve ultimate deformation in the fills.

Acknowledgement: The work described inthis paper was supported jointly by the MaxPlanck Forschungspreis awarded by the MaxPlanck Gesellschaft, Germany and the NSERCDiscovery Grant awarded to the first author. Theauthors are grateful to a reviewer for the highlyconstructive comments and for drawing attentionto further papers on the subject.

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Consolidation of a Soft Clay Composite ... a Soft ClayComposite 55 measuring approximately 15 cm in...

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