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Geotextiles and Geomembranes 24 (2006) 156–168

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Particle-scale surface interactions of non-dilative interface systems

J.E. Dovea,�, D.D. Bentsb, J. Wanga, B. Gaoc

aVia Department of Civil and Environmental Engineering, 200 Patton Hall, Virginia Tech Blacksburg, VA 24061, USAbS&ME, 9751 Southern Pine Boulevard, Charlotte, NC 28273, USA

cGolder Associates, 3730 Chamblee Tucker Road, Atlanta, GA 30341, USA

Received 9 September 2005; accepted 6 January 2006

Available online 17 February 2006

Abstract

Millimeter and micron length scale particle–surface interactions are of fundamental importance in the modeling and actual field

behavior of geotechnical composites incorporating non-dilative interface systems. The results of a basic experimental study to explain

underlying causes of steady state strength behavior of these systems are presented. Single soda lime glass beads and natural subrounded

quartz sand were placed in unlubricated sliding against manufactured materials consisting of high-density polyethylene,

polyvinylchloride, fiber reinforced plastic, stainless steel and heat-treated steel. For high-density polyethylene and polyvinylchloride

sheet materials steady state is reached after the rate of wear diminishes. In contrast, fiber reinforced plastic and steels exhibit a higher

friction coefficient at large displacement than at peak due to interactions of the particle with the surface at the grain level. Initial peak and

steady state friction coefficients depend primarily on the initial surface microtopography, grain shape, relative hardness of the particle

and material pairing, normal stress and manufactured material morphology. The results give engineers an improved framework for

assessing laboratory results and potential influence of wear on mechanical and hydraulic properties of construction materials.

r 2006 Elsevier Ltd. All rights reserved.

Keywords: Interface friction; Interface behavior; Wear; Granular media; Interface design

1. Introduction

Non-dilative interface systems occur when soil particlesare large with respect to the asperities comprising thesurface topography of the opposing continuous construc-tion material (Dove and Jarrett, 2002). Little internal soilresistance is mobilized during shear and it is not possible todescribe the behavior of these systems using traditional soilmechanics principles. In contrast, dilative interface systemsoccur when the asperities engage the soil particles andinduce fabric and contact force anisotropy (Bathurst andRothenburg, 1990) within the interphase region and withinthe soil mass.

Non-dilative interface system behavior is important inmany practical geotechnical engineering problems. Some ofthese problems include interface strength behavior ofuntextured geomembrane interfaces with granular media,side resistance in piles (Heerema, 1979, 1980), cyclic

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

otexmem.2006.01.002

ing author. Tel.: +1540 231 2307; fax: +1 540 231 7532.

ess: jedove@vt.edu (J.E. Dove).

degradation of pile capacity, pipe friction forces (Tognonet al., 1999), earth and rock anchor capacity and soil-machine interaction. In addition, the further developmentof discrete and alternative continuum numerical methodsto allow modeling of plastic damage of a surface in contactwith soil particles requires an understanding of basicinteraction mechanisms, and experimental data for valida-tion. While the tribology literature is rich in studies onfriction and wear, little of that information is applicable togeotechnical systems.Dove and Frost (1999) derived a semi empirical

expression to describe peak state behavior of non-dilativeinterfaces in contact with assemblages of subroundedOttawa sand, spherical glass microbeads, and angularblasting sand. It was found that for interface systemscomposed of high density polyethylene (HDPE) geomem-brane and Ottawa sand, the peak friction coefficientdecreased with increasing normal stress until a criticalnormal stress between 50 and 100 kPa was reached. Izginand Wasti (1998) and O’Rourke and Druschel (1989) madesimilar observations in this stress range. However, when

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Nomenclature

a wear groove half-width (mm)F shear load (N)HCF Heywood circularity factorHs

v Vickers hardness of the manufactured materialsurface (GPa)

Hpv Vickers hardness of the particle (GPa)

Hvr Vickers hardness ratioh average wear groove depth (mm)

N number of particle passages over a givenvertical section of manufactured material sur-face

Np number of particle passages required to reach asteady state strength condition

W normal load (N)w wear groove width (mm)m coefficient of frictionmss steady state coefficient of friction

Fig. 1. Particle contacts at non-dilative interfaces. (a) Two-dimensional

discrete element model of a direct interface shear device illustrating the

possible distribution of contact forces for well graded circular disks

against a smooth, rigid surface (shaded region at bottom). Contact stress is

directly proportional to the thickness of the black lines. (b) Ottawa 20/30

sand contact areas on HDPE sheet at 100 kPa normal stress (Dove 1996).

Field of view is 8mm square. Particle displacement over the surface (inset).

J.E. Dove et al. / Geotextiles and Geomembranes 24 (2006) 156–168 157

normal stresses increased above critical, Dove and Frost(1999) reported an increasing peak friction coefficient.Luellen et al. (1999) also experienced this behavior whileevaluating materials for a composite landfill liner systemand observed a critical normal stress of approximately150 kPa.

Above the critical stress, the increase in peak frictioncoefficient is attributed to contact normal stresses sufficientto cause plastic deformation of the manufactured materialduring consolidation and wear during shear. Plasticdeformation and wear result in surface damage in theform of grooves. In contrast, steady state behavior is noteasily explained. In some cases, it is observed that steadystate is not attained in laboratory test. Instead, shearresistance monotonically decreases within the full travelrange of the shear device. Observations of polymer surfacesand surface topography profiles made after standardinterface shear tests indicate that groove depth and spacingvary widely. Typically there are only several deep grooveswhile depths of more shallow grooves are distributed over awide range. This is due to the redistribution of an appliednormal load within the particulate mass at the consolida-tion phase of an interface shear test. Fig. 1a illustrates howcontact forces might vary for an interface system composedof poly dispersed assemblages of circular disks. Contactforce on the rigid boundary is directly proportional to thewidth of the black contact force chains. Several contactshave relatively high loads as indicated by the thickness ofthe contact force chain lines. Resulting contact stresses canexceed the crushing strength of the soil grain or themanufactured material yield stress. It is postulated thatthose contacts with stresses exceeding the yield stressdominate the overall shear behavior. Friction behavior oflower stress contacts is governed by mechanisms proposedby Dove and Frost (1999).

Experimental evidence for variation in contact forces isshown in Fig. 1b. This image is of the actual contact areasfor uniformly graded Ottawa sand assemblage consoli-dated against a smooth HDPE geomembrane at a normalstress of 100 kPa (Dove, 1996). The size of the contact isproportional to the normal load transmitted to the polymersurface. Particle contact areas are the subrounded darkregions, which are formed as the normal load is applied.These contacts then traverse the opposing manufactured

material surface during the shear process. It is the particleand surface mechanical interaction that determines thestress-displacement response, not the strength or relativecompaction of the soil. Post-shear observations of manymanufactured material surfaces by the authors, and bymany other engineers (O’Rourke and Druschel, 1989;Negussey et al., 1989; Luellen et al., 1999; Stark andPoeppel, 1994; Frost et al., 2002; DeJong and Westgate,2005), indicate that particle movement, and the orientationof wear grooves, are in parallel paths as would be expectedin plane strain environments.The above discussion implies that a relationship exists

between the stress-displacement and strength behavior of

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Fig. 2. Stress-displacement behavior of some non-dilative interface

systems. (a) C-33 concrete sand/fiber reinforced plastic (FRP); (b) angular

blasting sand/high density polyethylene (HDPE) sheet; (c) subrounded

Ottawa sand/HDPE sheet; (d) spherical glass beads/HDPE sheet; and (e)

subrounded sand/A-4 steel. Tests at 100 kPa normal stress, 80% relative

density.

Fig. 3. Instrumented tribometer arm with particle contacting surface.

Stylus profilometer tip and arm are used to profile surface before and

during shear displacement.

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non-dilative interface systems and: (1) manufacturedmaterial properties, (2) soil particle shape, (3) particle tosurface contact stress, and (4) amount of shear displace-ment. Understanding this behavior is important to practi-cing geotechnical engineers because measured interfacestrengths can have large variation depending on materialpairing, as indicated on the laboratory interface shear testcurves of Fig. 2. In addition, the atypical stress-displace-ment responses and inability to reach steady state for somematerial pairs can lead to misinterpretations of laboratorydata.

This experimental research investigation was conductedto determine possible mechanisms for steady state frictioncoefficient evolution. The interaction of wear grooves, theinfluence of wear on mechanical properties of themanufactured materials and modeling of the frictionbehavior are beyond the scope of this paper.

2. Experimental methods and procedure

2.1. Experimental apparatus

A pin-on-disk tribometer was used to conduct interfaceshear tests with a single soil particle in contact with amanufactured material. This device allowed a single grainto repeatedly travel in the same path over the manufac-tured material surface. It consisted of a precision micro-stepping rotary positioner (turntable), an instrumentedcantilevered arm and a loading system, as shown in Fig. 3.The arm freely rotated in a vertical plane. A screw locatedat the end of the arm served as the ‘‘pin’’ to which a soilparticle was mounted, thus particles were not permitted torotate. This configuration prevented particle rolling. Forparticle rolling to occur in an assemblage, the resistanceoffered by mineral friction and mechanical interlock

between adjacent soil particles must be smaller that thefriction coefficient between the particles and manufacturedmaterial. In an actual laboratory test, some particles willrotate as the fabric formed at the interface will containparticles that are not frustrated.Normal load was applied to the arm with a precision air

cylinder and independently measured with a miniature loadcell. The rotation rate was a constant 0.028 radians persecond in all tests. Shear forces developed by the particlewere measured with strain gauges mounted on the arm in afull bridge configuration. Calibration of the arm wasaccomplished using precision weights and a miniature loadcell. The stylus profilometer (Fig. 3) measured the initialand evolving surface topography in a direction perpendi-cular to the particle path.The tribometer laboratory model provides insight into

processes occurring in shear devices and field applications.A ring shear device is illustrated on Fig. 4a where particleswith diameter D, are arranged on the surface at a constantradial distance, r, from the center. When rotation of theshear device begins, the particles move over the manufac-tured material and pass a typical section A-A. Assumingtrailing particles follow the lead particle in the samepath, the number of particle passages N, in one revolutionof the ring shear device at section A-A is 2pr/D. Eachparticle passage causes an incremental amount ofsurface wear. All other sections on the surface are subjectto the same process. A standard direct interface shear testdevice is shown in Fig. 4b. In these devices, thebottom surface is longer than the top shear box so theshear area remains constant during a test. Displacement ofthe top box occurs from left to right relative to the lowersurface as indicated by the right most portion of the dashedrectangle. The portion of the surface that experiences thefull number of particle passages is shown by the centrallylocated shaded region. Thus, if this were a 300mm shearbox with 100mm of displacement, only a 200mm longsection will experience the maximum number of particlepassages.

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Fig. 5. Particles used in this study: (a) soda lime glass beads; (b) sub-

rounded silica sand. Both images made at 6.7� magnification.

Fig. 4. Models of particle movement in shear tests. (a) Ring shear. Each

particle passage causes surface and/or particle wear as shown in Section A-

A depending on the relative hardness of the materials; (b) direct interface

shear. Only the shaded area experiences the maximum number of particle

passages.

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2.2. Materials

Particles consisted of 1.5mm diameter spherical sodalime glass beads and subrounded silica filter sand (Figs. 5aand b). Glass beads serve as a reference material becausethe spherical shape is well-supported theoretically andminimizes effects of particle shape and roughness. Thesubrounded sand appeared to be free of microfractureswhen examined under a microscope. Grain shape wascharacterized from images made of particles under amicroscope using the Heywood circularity factor (HCF),which is defined as (National Instruments, 1999):

HCF ¼Particle perimeter

2�ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi

p� Particle areap . (1)

The perimeter and area are based on the two-dimensionalprojection of the individual particles onto the plane of theimage. A circle has an HCF value of 1.0 and increases asangularity increases. HCF is independent of particle size.Grain roughness effects were minimized by using variablemagnification to maintain a constant field of view in themicroscope for particles of different sizes. Table 1 providesthe properties of the grains.

Manufactured materials were chosen for their applic-ability in geotechnical engineering construction and toobtain a range of relative hardness. Vickers Hardness, HV,(ASTM E384, 1989) was used in this study because of itswell-developed theoretical basis and because it represents astress. For steel, Hv is nominally three times the yieldstrength (Johnson, 1985), however this relation has notbeen proven for polymers. Relative hardness, Hvr is used inthis study to provide an index of the hardness contrast

between the particle and substrate:

Hvr ¼Hs

v

Hpv

, (2)

where Hsv is the Vickers hardness of the manufactured

material surface and Hpv is the Vickers hardness of the

particle. Tables 1 and 2 provide hardness values for theparticles and the manufactured material substrates used inthe experiments. Substrate material hardness was deter-mined from the average of 30 measurements using alaboratory Vickers device. Hardness of the soda lime glassbeads was determined by nanoindentation hardness testsconducted by the Nano Instruments Innovation Center ofMTS Corporation, Oak Ridge, Tennessee, USA. Nanoin-dentation tests could not be performed on the sub-roundedsand due to excessive surface roughness in the 50 to 60 nmlength scale range. Therefore, sub-rounded sand wasassigned a hardness of 7.5GPa, which is approximatelyin the middle of the range of 5.4–9.7GPa determined byRichardson (1967).The polymer materials include HDPE sheet, polyvi-

nylchoride (PVC) sheet and fiber reinforced plastic (FRP)sheet. HDPE is used extensively in geosynthetic applica-tions and its interface behavior with coarse-grained soil iswell established. PVC is available with virtually anyhardness and is widely used for underground pipes and

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Table 1

Properties of particles

Particle Vickers hardness, Hpv

(GPa)aYoung’s modulus (GPa)a Equivalent diameter

(mm)a,bHeywood circularity

factora,b

Soda lime glass bead 6.2c 70c 1.50 1.003

(0.6) (4) (0.01) (0.006)

Subrounded sand 5.4–9.7 – 1.9 1.016

(0.1) (0.007)

aStandard deviation values are given in parenthesis.bAverage of 24 particles.cDetermined by nanoindentation tests.

Table 2

Properties of manufactured materials

Substrate Substrate Vickers

hardness

Vickers hardness

coefficient of variation

Substrate thickness Tensile yield strengtha Initial average

roughness PRab

Hsv (Gpa) (%) (mm) (MPa) (mm)

High density

polyethylene0.037 1.7 1.5 17 8.48

Polyvinyl chloride 0.14 0.6 3.1 52 1.03

Fiber reinforced

polymer0.22 1.5 6.3 138 10.52

Steel 1.3 1.0 3.3 – 0.34

Heat treated steel 7.2 0.4 6.7 – 1.0

aManufacturers values.bUnfiltered (primary) value.

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casings. The sheet material used in this project was a Type IPVC that is typically used for underground piping.

The FRP material was manufactured by a pultrusionmethod with the fibers bound by a thermosetting polyesterresin matrix. A woven nylon veil layer was located justbelow the surface resin coating. Below the veil layer, thelongitudinal reinforcing consisted of continuous strandfiberglass filaments; a randomly oriented continuousfiberglass strand provided transverse reinforcing. FRP isincreasingly used for foundation elements and retainingstructures, especially in geochemically aggressive environ-ments (Pando et al., 2002).

Steel materials consisting of 308 stainless steel (SS) and aheat-treated American Iron and Steel Institute Type A2steel (HTS) were. Coupons of HTS were cut from A4 sheetstock and subsequently heat treated at 400 1. The surfacesof the steel coupons were cleaned on a polishing wheelprior to testing. Properties of the manufactured materialsare presented in Table 2.

A survey of construction materials used in civilengineering indicates that hardness values are typicallyless than about 1.9GPa. By comparing values in Tables 1and 2, it is seen that soil particles are much harder than thepolymer materials used herein and are typically harderthan construction materials in general. Therefore, it can beexpected that manufactured materials will undergo wearwith little or no degradation of the soil particles.

2.3. Test procedure

Coupons of the manufactured material were cut from abulk sample and epoxy-glued to an aluminum backingplate mounted on the turntable. Surfaces are cleaned withwater and allowed to completely dry. Vickers hardnesstests are conducted using the softest material (HDPE sheet)to assess the influence of the backing plate as the Vickerstest is a more intense point loading than the grains. Nodifference in Hv, was observed with or without the backingplate.A particle was then mounted at the end of a rigid set

screw using a small amount of epoxy glue. The screw wasplaced into a threaded hole at the end of the instrumentedarm and allowed to contact the construction material. Thelength of the instrumented arm resulted in a particle tracecircumference on the manufactured material of 57.5mm.Results presented in this study are for a constant normalload on the particles of 10.7N. This load produced plasticparticle–surface contact conditions during the initialloading stage and in the early part of shearing. A newmanufactured material sample was used for each test.Peak coefficients of friction were determined from the

single highest value of shear force occurring just afterinitiating disk rotation. This shear force and friction ishereafter referred to as the ‘pull-off value.’ After peak, testswere paused after each revolution for the first 10

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Fig. 6. Wear groove geometry.

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revolutions, then approximately every 10 revolutionsthereafter until completion of the test to profile the surfaceand measure the contact area. Post-peak friction coeffi-cients were determined at each pause by averaging theshear loads measured since the preceding pause. Eachtribometer test was allowed to continue until the averagefriction force becomes constant.

2.4. Material surface measurements

Profiles of manufactured material surface topographywere used to relate the degree of surface wear to thevariation in friction coefficient. Profiles were made with aTaylor–Hobson S3F stylus profilometer. A profile of eachconstruction material surface is made before each test thenat each subsequent pause. The profilometer was positionedsuch that the initial and all subsequent profiles wereoriented perpendicular to the particle path and were madein the same location in each revolution. Average groovedepth, h ¼ ðh1 þ h2Þ=2, and the groove half-width,a ¼ w=2, (Fig. 6), were determined from the profilometertraces.

The size of the contact area on the glass bead particleswas determined by analyzing digital images acquiredthrough a microscope at each pause in shear. Imaging thecontact area required that the set screw and particle beremoved from the instrumented arm. All mounting screwswere indexed to the arm so that the particle could bereplaced in the same orientation. Contact area measure-ments of subrounded sand particles could not be directlymade from images because of their irregular shape andsurface microroughness. Therefore, contact areas weredetermined using the method described by Shooter andTabor (1952).

3. Experimental results—polymer materials

3.1. HDPE sheet

Fig. 7 presents averages of friction coefficient, contactarea, groove depth and groove half width from four singleparticle tests using glass beads and three tests usingsubrounded sand. The friction coefficient for glass beadson HDPE decreases from the pull-off (peak) value of 0.22to a steady state value of about 0.14 after about 10 particlerevolutions. Peak friction coefficient for the sub-roundedsand is 0.53, which decreases after one revolution to about0.31 and then to a steady state value of 0.22 after about 7revolutions. These friction coefficients agree with results

from standard direct interface shear tests on the samematerials conducted by Dove and Jarrett (2002), and withthe results shown on Figs. 2c and d. However, it should benoted that the subrounded sand peak friction coefficient inFig. 2c is approximately 0.32, which is in good agreementwith the friction coefficient measured after the firstrevolution in the tribometer rather than the pull-off value.Fig. 7b shows the variation of particle to surface contact

area during shear. The normal load, W, is sufficient tocreate plastic contact conditions for the first few revolu-tions. An initial area, A, of 0.28mm2 is computed using therelation A ¼W=Hv, where Hv is the Vickers hardness ofthe softest material of the pair. This estimated area agreeswell with the measured contact area after the firstrevolution of 0.24mm2. Fig. 7b also indicates that particlecontact area grows as surface wear increases. Surface wearis most severe in the first four revolutions where plasticcontact conditions prevail. After about four revolutions,the contact area becomes greater than 0.28mm2, whichcauses a reduction in contact stress to below the HDPEhardness. At the same time, the friction coefficientdecreases dramatically in the first four revolutions andthen becomes essentially constant after about sevenrevolutions. Figs. 7c and d show that growth in groovedimensions drastically decrease after the initial 10revolutions but groove depth continues to increase at asmall rate. These data indicate that for HDPE sheet,friction coefficient is directly related to growth of the weargroove.Fig. 8 is a series of surface profiles that record wear

groove evolution in an HDPE sheet. The actual profiles are20mm in length but only the segment of interest in thevicinity of the groove is shown. Measurements made on thesurface profiles using an image analyzer indicates the crosssectional area of the groove is larger than the area of theshoulders on either side of the groove, indicating materialis lost from the surface. This conclusion is supported bydirect observation with a microscope the presence ofHDPE debris clinging to the particles, and surface filmsof transferred HDPE could be seen covering the particlecontact surfaces. In the ideal plowing mode of abrasivewear, material is not lost from the surface but is plasticallyredistributed (Kato and Adachi, 2001) to form the grooveshoulders. Since material is being lost, plowing is not theonly abrasive wear mechanism operating. The materialtransfer to the bead and the loss of material from thesurface provides evidence that the adhesive wear mode issignificant. Similar observations are made for the sub-rounded sand as shown in Fig. 7b. The worn profiles areobserved to be more angular in shape than the glass beadprofiles.

3.2. Polyvinyl chloride (PVC) sheet

Fig. 9 presents results for average friction coefficient,contact area, and groove depth and groove radius valuefrom three single particle tests using glass beads and three

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Fig. 8. Wear groove development for glass bead and HDPE sheet

interface system. Note vertical exaggeration. (a) Initial profile; (b) one

revolution; (c) 10 revolutions; and (d) 55 revolutions.

Fig. 7. Results of single particle tests for glass bead, subrounded sand and HDPE sheet interface systems. (a) Friction coefficient; (b) contact area; (c)

groove depth; and (d) groove half-width.

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tests using subrounded sand. Steady state is reached afterabout six revolutions of the glass bead and about eightrevolutions of the subrounded sand. The friction coefficient

of the PVC polymer decreases with increasing number ofparticle passages in a manner similar to the HDPEpolymer. However, unlike HDPE, there is little differ-ence between the steady state friction values of the twoparticles.Contact area corresponding to fully plastic contact

conditions of the PVC is 0.077mm2. Both particles reachthis contact area in about four revolutions, after which areagrows more slowly. Groove measurements indicate that thedepth continues to increase throughout the test with themost rapid increase in the first six to eight revolutions. Asanticipated, both the degree and rate of wear is muchhigher with the subrounded particle than the glass beadparticle.

3.3. Fiber reinforced polymer (FRP) sheet

Fig. 10 shows the changes in friction coefficient for thefiber reinforced polymer substrate. It was observed that theglass bead did not wear the surface consistently amongspecimens. In one case there was discernable wear only onportions of the particle trace after approximately 30revolutions. This caused considerable difficulty in inter-preting the profilometer traces; thus contact area, groove

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Fig. 9. Results of single particle tests for glass beads, subrounded sand and PVC sheet interface systems. (a) Friction coefficient; (b) contact area; (c)

groove depth; and (d) groove half-width.

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half width, and groove depth measurements for theglass beads were not considered reliable and are notreported.

The FRP material used in this study has a VickersHardness of about 0.22GPa making it considerablyharder than the HDPE or PVC. Unlike the other polymersubstrates, the FRP has steady state friction valuesthat are greater than, or nearly equal to, the initialpull off values for both particles. Initial friction co-efficient for the glass beads decreases to 0.11 from a peakvalue of 0.14 after the first revolution and increasesthereafter. The steady state friction coefficient of 0.28 isreached at about 40 revolutions and is much greater thanthe steady state values for the softer HDPE and PVCpolymers.

The peak friction coefficient for the subrounded sand is0.34. A steady state friction coefficient of 0.29 is reached atapproximately four revolutions, which is also greater thansteady state friction coefficient of the other polymermaterials. The subrounded sand was effective in wearinga groove in the specimens. There is a rapid initialpenetration of the surface, which decreases substantiallyafter 10 revolutions.

4. Experimental results—metals

4.1. Stainless steel

Two results are immediately apparent from Fig. 11showing average results for 5 tests with glass bead andthree tests using subrounded sand particles. First, a truesteady state is not reached. Secondly, the friction coeffi-cients initially decrease from the initial pull-off value, thenincrease. The initial decrease is more pronounced for thesubrounded sand but the rate of subsequent increase isgreater for the glass beads. Wear curves (Figs. 11c and d)are similar in shape but much lower in magnitude than thepolymers.

4.2. Heat treated steel

Fig. 12 indicates that glass beads and heat treated steelinterface systems have complex load-displacement beha-vior. For this system, the hardness of the particle is lessthan the substrate ðHvr ¼ 1:16Þ, thus wear is experienced bythe particle. The friction coefficient decreases from theinitial pull-off value then increases with a slight change in

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Fig. 10. Results of single-particle tests for glass beads, subrounded sand and FRP interface systems. (a) Friction coefficient; (b) contact area; (c) wear

groove depth; and (d) wear groove half-width.

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the slope. Observations made under the microscopeindicate the steel specimen was only ‘‘polished’’ in theparticle path. No measurable groove in the steel surfacewas formed during these experiments.

5. Discussion

This paper examines millimeter to micrometer lengthscale interactions between particles and surfaces andprovides explanations for the stress-displacement relation-ships and large displacement friction observed for selectednon-dilative interface systems. Grain shape, relative hard-ness of the mating materials and normal load on the graincontrol this behavior, however there are distinct differencesbetween polymer and metal manufactured materials.

Stress-displacement behavior of polymer systems iscontrolled by removal of an incremental amount of thesofter material during each revolution. For the HDPEpolymer, the rate of wear is most rapid in the initial five toten volutions or particle passages. As contact stressdecreases due to contact area growth, each increment ofmaterial removed or plastically deformed is less than the

preceding increment. These data indicate that shear forcedecreases as wear rate decreases.This process requires sufficient displacement to permit

contact area growth that leads to a reduction incontact stress. Grain shape plays a key role in this processas shown in Fig. 13a. When glass bead assemblages aredisplaced against HDPE in a shear box, it is typicallyobserved that a steady state is not reached within thetravel distance of the device. For example, in Fig. 2d it isseen that steady state is not reached within the maximumtravel of the shear box. For the subrounded sand, sevenrevolutions of the single particle device are required toreach steady state conditions corresponding to 11.5mm ofshear displacement. Likewise, direct interface shear testswith a subrounded Ottawa sand-HDPE interface system(Fig. 2c) reach a steady state at about 5–6mm. Similarly,Fig. 2b indicates the direct interface shear test using anangular blasting sand-HDPE interface system reachessteady state at about 4–5mm. The above resultssuggest that the displacement required to reach steadystate decreases as particle angularity increases. This isattributed to more efficient removal of the manufacturedmaterial.

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Fig. 11. Results of single particle tests for glass bead, subrounded sand and stainless steel interface systems. (a) Friction coefficient; (b) contact area; (c)

wear groove depth; and (d) wear groove half-width.

Fig. 12. Results of single particle tests for glass beads and heat treated steel interface systems. (a) Friction coefficient, and (b) contact area.

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Similar results are observed for PVC. Fig. 13a indicatesthat the number of particle passes to reach steady statedecreases because of the greater Hvr. This is to be expectedas the higher hardness of the PVC will support a greatercontact stress elastically, resulting in less wear.

Shear behavior of FRP is directly related to itsmorphology. A thin veneer coating of resin covers thereinforcing fibers. Examination of the grooves by micro-scope revealed extensive damage to the resin coating andformation of wear particles consisting of chips of resin.

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Fig. 13. Influence of hardness ratio on stress-displacement and strength

behavior. Open symbols—glass bead; solid symbols—subrounded sand.

(a) Number of particle passes, Np, required to reach steady state, and (b)

steady state friction coefficient.

Fig. 14. Images of particles and wear grooves for each system. (a) Glass

bead/HDPE, Hvr ¼ 0.006; (b) glass bead/PVC, Hvr ¼ 0.022; (c) glass

bead/FRP, Hvr ¼ 0.035; (d) subrounded sand/stainless steel, Hvr ¼ 0.24;

and (e) glass bead/heat treated steel, Hvr ¼ 1.16.

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After about 16 mm of particle penetration, the veneer layeris sufficiently removed for the particle to begin encounter-ing the veil layer. The fibers, or interphase region, presentincreased surface roughness to the particles resulting in anincrease in friction coefficient at steady state. Therefore,the friction behavior of fiber-reinforced composites can beexpected to depend not only on material hardness but alsoon fabrication methods.

Generation of surface wear debris by removal of thesofter material by the harder material also is an importantfactor controlling friction behavior (Bowden et al., 1943;Suh and Sin, 1981). Fig. 14 presents images of the particlesand manufactured material surfaces taken during theexperiments. Wear debris consisting of manufacturedmaterial chips are evident on the surfaces of both matingmaterials. In Fig. 14e, the manufactured material is harderthan the particle therefore the particle wears, as may beseen by the circular contact area on the tip of theparticle.

Wear debris interact with the moving particle and cancause an increase in friction force depending on theirrelative strength and hardness. In general, if the wearparticle is soft, as in the case of HDPE and PVC polymers,there is little or no additional shear resistance due to the

presence of the wear debris. Therefore, adhesive forcesbetween the polymers and particles must contribute tostrength. A definitive conclusion cannot be reached on theFRP material since the particle engaged the veil layer atrelatively small displacement. However, it might beexpected that the epoxy matrix would behave in a mannersimilar to the other polymers.In contrast, there is a significant difference in the

behavior of metals. When a hard particle slides against ametal surface, wear particles are formed that havegeometric roughness and relatively high strength. Wearmaterial is entrapped between the particle and surface,causing the friction coefficient to increase with slidingdistance as more debris are generated. A steady stateshould be reached as the generation of wear particles is

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balanced by the decrease in wear rate and the loss of wearparticles from the groove (Suh, 1986).

The results also demonstrate that particle–surfaceinteractions control the observed shear behavior. Thefriction coefficients measured in the single particle deviceand in laboratory tests using the same HDPE material arepractically the same (Figs. 2c, d and 7a). Results for FRPin Fig. 10 are in good agreement with the direct interfaceshear results of Fig. 2a and offer an explanation for theanomalous stress–displacement behavior. The interactionof wear debris from metal counterfaces on large displace-ment strength is likely responsible for the absence of a post-peak reduction in interface strength shown in Fig. 2e.

These results have practical implications. Engineersusing FRP for driven foundation elements and sheet pilesshould account for wear damage and any effects on long-term durability of the FRP. By working with manufac-turers, exciting opportunities for producing FRP materialsto meet project-specific requirements exist. The formationof wear grooves represents damage to a manufacturedmaterial and potential for degradation of physical andhydraulic properties, especially for geomembranes. DeJongand Westgate (2005) describe this in detail. The con-sequences of damage are project specific and should beassessed on a case-by-case basis using project materials.Data contained in the figures provide guidance on groovedimensions for round and subrounded particles on HDPEsheet for computing material stress and effective thickness.Where rounded or subrounded particles are used ininterface systems with polymers, judgment should be usedin selecting large-displacement strengths as the laboratorydevice may not provide sufficient travel to reach a steadystate condition. Finally, the results provide laboratory datafor engineers seeking to validate numerical models of non-dilative interface systems.

6. Conclusions

The results of this experimental study provide rationalexplanations for the observed behavior of non-dilativeinterface systems incorporating manufactured materials.The friction behavior in these unlubricated experiments isdominated by the shear force required to form the groovein the surface, the resistance due to wear particlegeneration and interaction, and the force required toovercome adhesive forces acting on the contact surfacearea. The results suggest that those particles with contactstresses greater than the hardness of the manufacturedmaterial will dominate the shear resistance measured in alaboratory test. These results provide engineers with animproved basis for interpreting laboratory results of givensystems, assessing manufactured material damage andpotential influence on mechanical and hydraulic properties,and a basis for selecting material pairs to be included ingeotechnical composite systems.

Acknowledgements

This material is based upon work supported by theNational Science Foundation under Grant nos. CMS-0196087 and CMS-0200949. Any opinions, findings, andconclusions or recommendations expressed in this materialare those of the authors and do not necessarily reflect theviews of the National Science Foundation. The manuscriptwas greatly improved by the thoughtful comments of theassessors.

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