A new method for remediation of sandy slopes susceptible to seepage flow using EPS-block geofoam

Geotextiles and Geomembranes 42 (2014) 166e180 Contents lists available at ScienceDirect Geotextiles and Geomembranes journal homepage: www.elsevier.com/locate/geotexmem A new method for remediation of sandy slopes susceptible to seepage flow using EPS-block geofoam A Tolga Özer a, *, Onur Akay a, Garey A Fox b, Steven F Bartlett c, David Arellano d a Department of Civil Engineering, Okan University, Tuzla Campus, Istanbul 34959, Turkey Department of Biosystems and Agricultural Engineering, Oklahoma State University, 120 Ag Hall Stillwater, OK 74078-6016, USA c Department of Civil and Environmental Engineering, University of Utah, 110 Central Campus Dr., Salt Lake City, UT 84112, USA d Department of Civil Engineering, The University of Memphis, 104 Engineering Science Building, Memphis, TN 38152-3180, USA b a r t i c l e i n f o a b s t r a c t Article history: Received 31 July 2013 Received in revised form December 2013 Accepted 24 January 2014 Available online March 2014 Using expanded polystyrene (EPS) geofoam (geofoam block) in slope remediation projects has drawn interest from the civil engineering sector for its ease of application and budget saving features According to design precedence, all slope remediation applications that use geofoam blocks should incorporate permanent drainage systems to prevent instability of the lightweight geofoam blocks due to hydrostatic and seepage pressures In this study, a new method for slope remediation using geofoam blocks was tested through physical laboratory experiments For this purpose, a total of 24 lysimeter (dimensions of 60 cm height, 20 cm width, and 200 cm length) experiments (including duplicates) were conducted in which seepage through a geofoam block slope system were generated with three different constant water levels in the water reservoir of the lysimeter Geofoam blocks (dimensions of 2.5 cm height, cm width, and 15 cm length) were assembled to form embankment type configuration at the toe section of the sandy slopes This study also included coupled numerical model simulations that were comprised of variably saturated flow modeling and slope stability modeling which could be implemented successfully for the global static failure analysis of the geofoam block slope system comprised of two mediums with different geotechnical characteristics In addition to global static stability failure analysis, which involved conventional limit equilibrium analysis for the geofoam block slope system, hydrostatic sliding mechanism was investigated which provided insight into using an overburden concept to increase the resistance against horizontal driving forces Experimental and numerical modeling results showed that the geofoam block slope system was stable even though the phreatic surface was above the bottom of the geofoam block assemblage For this reason, the embankment type configuration tested in this study can be considered a viable remediation technique where seepage induced deep-seated global stability and hydrostatic sliding failures are a concern Ó 2014 Elsevier Ltd All rights reserved Keywords: EPS-block geofoam Slope stability Slope remediation Hydrostatic sliding Seepage Introduction There are several factors that can trigger slopes to fail Steep slopes, low strength slope materials, weak foundation conditions, and earthquakes are major factors affecting slope instability Seepage is another primary cause of slope instability for both manmade and natural slopes (Fox and Wilson, 2010) Leaky pipes, irrigation, snowmelt, thawing ice lenses, runoff from uphill sources, the clogging of a drain, or shutting off a near-surface well might produce mounding of the phreatic surface within the slope above its steady-state position (Schmertmann, 2006) When this * Corresponding author Tel.: ỵ90 216 6771630x1978; fax: þ90 216 6771486 E-mail addresses: tolga.ozer@okan.edu.tr (A.T Özer), onur.akay@okan.edu.tr (O Akay), garey.fox@okstate.edu (G.A Fox), bartlett@civil.utah.edu (S.F Bartlett), darellan@memphis.edu (D Arellano) 0266-1144/$ e see front matter Ó 2014 Elsevier Ltd All rights reserved http://dx.doi.org/10.1016/j.geotexmem.2014.01.003 infiltrated water enters a slope faster than the excess pore-water pressures can dissipate, stability will be significantly reduced Pore-water pressure accretion is the most prevalent of failures on natural hillslopes (Sidle and Ochiai, 2006) The cause and nature of a slope failure must be understood before designing slope remediation systems (Duncan and Wright, 2005) Fay et al (2012) listed the essential elements of slope stabilization as proper planning and site investigation, understanding the soil, and knowing the surface and subsurface water conditions Since every slope repair project has unique causes, numerous types of remediation techniques have been developed (Dronamraju, 2008; Shah, 2008; Fay et al., 2012) These remediation techniques can be categorized in four different groups: mechanical stabilization techniques, earthwork techniques, erosion control techniques, and bioengineering techniques (Fay et al., 2012) A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 In addition to these listed slope stabilization categories, reducing the driving force is a viable alternative (Elragi, 2000) However, since the resisting forces along the failure surface are also dependent on the weight of the slope (resisting forces are proportional to normal stresses), the factor of safety against global stability failures (instability along rotational failure surfaces) can only be increased if the reduction of the driving forces is greater than the reduction in the resisting forces In order to reduce the driving forces in slopes, engineers have used several lightweight fill solutions (e.g., pumice, shredded tires, expanded polystyrene (EPS) geofoam (geofoam block), and tyre bales) Among these lightweight fill materials, a geofoam block has high strength to density ratio (Elragi, 2000; Stark et al., 2004) Due to this property, durability and ease of installation in the field, geofoam block has been gaining popularity since it was first used as a lightweight embankment fill by Norwegian Public Roads Authorities in 1972 (Aabøe, 2011) In addition to its application as a lightweight embankment fills for roadways, geofoam blocks were used for slope stabilization projects in Japan largely in the mid-1980’s to the mid-1990’s (Tsukamoto, 1996) Geofoam blocks have been used by Reuter and Rutz (2000), Reuter (2001), Mann and Stark (2007) in slope remediation projects in United States Even though geofoam blocks have experienced wide-spread use in slope stabilization and rehabilitation projects, there were no formal design guidelines or procedures until 2011 Arellano et al (2011) developed a design guideline, which was funded by the National Cooperative Highway Research Program (NCHRP), to use geofoam blocks for slope stabilization and repair projects In this design guideline, Arellano et al (2011) presented design procedure algorithms which were based on conceptual failure modes This design guideline is based on the recommendation that all geofoam block slope systems incorporate a drainage system since many of the geofoam block slope case histories evaluated as part of the NCHRP 24-11(02) research included use of underdrain systems below geofoam blocks to prevent water from accumulating above the bottom of the geofoam block assemblage Also, in some cases, drainage systems were incorporated between the adjacent upper slope material and geofoam blocks to collect and divert seepage water and thereby alleviate seepage pressures Even though the design procedure recommends permanent drainage systems, the groundwater table may rise in the long-term due to clogging of the drainage pipe as a result of improper design and/or poor construction in the field As a result, the groundwater table may rise above the bottom of the geofoam blocks which may cause global stability failure of the slope and/or hydrostatic sliding failure of geofoam block assemblage The behavior of geofoam block slope systems for remediation of sandy slopes with seepage was first studied by Akay et al (2012, 2013) using scaled physical slope experiments for marginally stable sand slopes Based on an extensive laboratory testing program, Akay et al (2013) concluded that in comparison with the results obtained from the nonremediated slope (“Matrix” configuration), the geofoam block configurations (“One Row” and “One Row Partial Bottom”) could be considered as a viable alternative remediation technique for shallow-seated failures; however, they seemed to be ineffective to prevent deep-seated global stability failures of a marginally stable steep sandy slope under seepage Therefore, Akay et al (2013) recommended that various geofoam block configurations be investigated to evaluate the use of geofoam block for remediation of sandy slopes that experience deep-seated global stability failures under seepage The overall objective of this study was to evaluate a geofoam block configuration in order to remediate a 1:1 sandy slope with a deep-seated slip surface with seepage In general, geosynthetic reinforcements are used to remediate/construct 1:1 or even steeper 167 sandy slopes (Benjamim et al., 2007; Portelinha et al., 2013) In this study, the possible use of geofoam blocks as a remedial geosynthetic alternative for steep sandy slopes subjected to seepage was investigated For this purpose, a small scale (1:20) laboratory, physical-slope modeling techniques (1-g model test) were utilized This laboratory method has been successfully performed to model not only the behavior of geofoam block slope systems with seepage forces (Akay et al., 2012, 2013), but also to model various geotechnical systems such as stone columns (Deb et al., 2011), geogrid reinforced foundations (Latha and Somwanshi, 2009), footing on geogrid reinforced clay slope (El Sawwaf, 2007a), geogrid and geotextile reinforced sand slopes (Lee and Manjunath, 2000; Yoo, 2001), geogrid reinforced flyash slope (Choudhary et al., 2010), horizontal anchor plates (El Sawwaf, 2007b), and geocell reinforced foundations (Dash et al., 2003) When compared to the field prototype, the main drawback of the 1-g small scale laboratory model is the differences in the stress levels between the 1-g model and field prototype (Akay et al., 2013; Choudhary et al., 2010; Latha and Somwanshi, 2009) However, the results of this research are relevant to revealing insights of using the proposed geofoam block configuration for remediation of sandy slopes susceptible to seepage forces at 1:20 scale This study also included numerical model simulations that were comprised of variably saturated flow modeling and slope stability modeling The model results were utilized in the determination of the factor of safety against prevailing failure mechanisms observed during laboratory lysimeter experiments Therefore, the factor of safety against global stability failure (FSGL) of the slope and the factor of safety against hydrostatic sliding of the geofoam block assemblage along the interface of the bedding level and the bottom of the embankment (FSSL) were calculated for the quantification of the performance of the geofoam block configuration Materials and methods 2.1 Laboratory lysimeter studies A total of 24 lysimeter experiments (including duplicates) were performed in this study Following Fox et al (2006), Wilson et al (2007), and Akay et al (2013), the lysimeter was constructed using 1-cm-thick Plexiglas and had the dimensions of 200 cm length, 20 cm width, and 60 cm height (Fig 1a) In addition to the soil compartment, the lysimeter had a water reservoir located at one end that generated the necessary hydraulic gradient for seepage to occur through the constructed slope The constant water level in the reservoir was adjusted to be higher than the base of the slope (25 cm, 38 cm, and 50 cm water pressure head) A stainless steel mesh having an opening size of 0.063 mm (equivalent to No 230 sieve size) and a perforated 1-cm-thick Plexiglas plate with 8-mmdiameter holes was placed between the reservoir and the soil compartment of the lysimeter The back-slope was uniformly compacted into the soil compartment of the lysimeter in 2.5 cm lifts to obtain a homogeneous domain with a dry density of 14 kN/m3 The constructed slope had a side-hill with a 45 angle (1:1 horizontal:vertical) In order to mimic field conditions in which the failed mass of the slope displaced at the toe provides resistance to subsequent failures, the slope was packed only to a length of 100 cm The slope height and width was 55 cm and 20 cm, respectively (Fig 1a) Data collection during an experiment included the pore-water pressures (h) developed inside the slope that were measured by 22 pencil-size tensiometers (Soil Measurement Systems, Tucson, AZ, USA) which were found to be successful at monitoring water pressure dynamics during previous soil column and lysimeter studies (Akay and Fox, 2007; Akay et al., 2013) The numbering and 168 A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 Fig (a) Laboratory setup included lysimeter, tensiometer, pressure transducer, and datalogger; (b) Location of tensiometers installed on one side of the lysimeter with coordinates (x, y) referenced to the toe of the slope (0, 0) the coordinates of the tensiometers with respect to the toe of the slope are given in Fig 1b Bi-directional pressure transducers (ASDXRRX005PDAA5, Honeywell Sensing and Control, Golden Valley, MN, USA) were attached to the tensiometers to transmit pore-water pressure heads to datalogger (CR1000 w/multiplexer, Campbell Scientific, Logan, UT) in the range from 345 cm-H2O to ỵ345 cm-H2O (5 psi) at 10 s intervals In addition, cumulative discharge measurements of the seepage flow were recorded by a digital scale located at the outlet section of the lysimeter Specific time measurements were recorded during an experiment which included the “time of seepage” (elapsed time from the start of an experiment to seepage initiation at the slope face), “time of final failure” (the time the slope becomes stable because of the failed soil resistance at the toe), and “end of the experiment” The soil used in the construction of the slope was selected to be similar to that of sand used by Akay et al (2013) Geotechnical properties of soil used in this study are given in Table Particle size distribution (ASTM D 6913, 2009), specific gravity (ASTM D 854, 2010), and triaxial shear strength (ASTM D 4767, 2011) tests were conducted on randomly selected undisturbed core samples (20-cm long, 7.26-cm diameter) extracted from the compacted slope after completion of the experiments Similarly, the hydraulic properties of the soil were determined from nine undisturbed core samples on which constant head permeability (ASTM D 2434, 2006) tests were conducted that resulted in an average saturated hydraulic conductivity, Ksat, of 0.0295 cm/s (standard deviation, s ¼ 0.0035 cm/s) The average dry density of the core samples was equal to 13.93 kN/ m3 (s ¼ 0.11 kN/m3) Retention characteristics of the soil were determined by collecting pore-water pressure and water content pairs using undisturbed core samples (6.4-cm long, 7.26-cm diameter) following the procedure given by Akay et al (2013) In this study, three core samples were extracted for this purpose which resulted in a total of 147 pairs including the data from Akay et al (2013) The van Genuchten-Mualem model (van Genuchten, 1980) was used to represent the retention curve fitted by the RETC (RETention Curve) computer code (van Genuchten et al., 1991): qhị ẳ < qr ỵ : qs qr ẵ1ỵjahjn  m qs h < 0= h0 (1) ; Table Material properties of soil and geofoam blocks used in the laboratory lysimeter experiments Property Material: sand Classification Particle size distribution Effective size Uniformity coefficient Coefficient of curvature Specific gravity Dry unit weight Cohesion Angle of internal friction Poisson’s ratio Material: EPS-block geofoam Dry unit weight Cohesion Angle of internal friction Compressive strength Corrected initial Young modulus Poisson’s ratio Description & Unit Value Unified soil classification system Sand (%), Silt ỵ Clay (%) D10 (mm) Cu [e] Cc [e] Gs [e] gd (kN/m3) c’ (kPa) f0 (degrees) SP 98.0, 2.0 0.17 3.5 1.2 2.64 14 33.7 0.35 n gd (kN/m3) c (kPa) f (degrees) @ 5% strain (kPa) @ 10% strain (kPa) E (MPa) n 0.2 22.5 30 90e96 106e112 4.2e4.8 0.1 A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 169 Fig Embankment type configuration with four different heights of geofoam block assemblages: (a) 10 cm, (b) 15 cm, (c) 22.5 cm, and (d) 30 cm h   i2 1=m m Khị ẳ Ksat Sle   Se m ¼  ;n > n (2) (3) where h is the pore-water pressure head [L]; qs is the saturated water content; qr is the residual water content; a [L1] is the inverse of the air-entry pressure (bubbling pressure); n is the poresize distribution index; l is the pore-connectivity parameter (taken as 0.5); and Se ¼ (q  qr)/(qs  qr) is the effective saturation Based on the curve fit (R2 ¼ 0.92), the van Genuchten parameters of the soil were: qs ¼ 0.45, qr ¼ 0.0, a ¼ 0.0928 cm1, and n ¼ 2.3579 While the soil was compacted to construct the back-slope, geofoam blocks were placed in a certain configuration as a lightweight remediation material (dry unit weight ¼ 0.2 kN/m3) at the toe of the slope The compressive resistance of geofoam block was determined using a series of three standard 50-mm cube samples according to ASTM D1621 (2010) The geofoam block strength parameters used in this study are summarized in Table 170 A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 2.2 Geofoam block configuration As stated earlier, the configurations tested by Akay et al (2013) were ineffective to remediate deep-seated failures For this reason, the new configuration was expected to withstand earth and hydrostatic pressures exerted by the backslope in order to remediate the slope (FSGL > 1.0 and FSSL > 1.0) However, due to their lightweight nature, geofoam blocks are vulnerable against hydrostatic sliding Stark et al (2004) summarized remedial procedures to increase factor of safety against hydrostatic sliding and overturning under four main categories for geofoam block embankments: use separation material between EPS blocks and foundation soil which can provide higher interface friction angle then EPS/foundation soil, increase overburden stress on EPS blocks (i.e heavier pavements), implement a drainage system, or use anchorage system to keep the blocks intact Using a separation layer to increase the interface shear resistance along the geofoam blocksefoundation interface and implementing anchorage plates can bring additional cost to the projects In addition, the use of anchorage system is not a typical field application Therefore, the overburden concept was implemented to design a configuration that can resist hydrostatic forces that can occur if the subsurface drainage system malfunctions In order to account for the previously mentioned overburden concept to resist hydrostatic forces, embankment type configuration was implemented for remediation of marginally stable steep sandy slopes in which geofoam blocks were placed at the toe of the compacted sandy slope (Fig 2) As shown in Fig 2, compacted sandy slope applies overburden along the portion of the geofoam block assemblage which remains inside the slope Geofoam block assemblages were rested on the 2.5 cm thick bedding level (Fig 2) Geofoam blocks used in the lysimeter tests were: 2.5-cm high, 5-cm wide, and 15-cm long These dimensions were selected to ensure 1:20 scale of geofoam block which relates to a common manufactured size As shown in the previous study by Akay et al (2013), final failure surface entered at the face of the initial 1:1 slope (shallow-seated failure) for the 25 cm-H2O boundary condition, whereas final failure surface entered at the crest (deep-seated failure) for the 38 cm-H2O and 50 cm-H2O boundary conditions for the “Matrix” configuration They concluded that that replacement of the soil mass that is typically dislodged during a shallow-seated failure with the lightweight geofoam blocks, could remediate the slope However, it seemed to be ineffective to prevent “deepseated” failures For this reason, embankment type configuration with four different heights (10 cm, 15 cm, 22.5 cm, and 30 cm) of geofoam block assemblages were tested against three different constant water levels in the reservoir to quantify the effect of the overburden stress on both FSGL and FSSL All of the 24 lysimeter experiments (including duplicates) were conducted in 2012, and the complete list can be found in Table The title of an experiment included the geofoam block assemblage, constant water head level at the reservoir, and the date (ddmm), respectively 2.3 Numerical model simulations Numerical model simulations involved variably saturated flow modeling coupled with slope stability modeling for each geofoam block assemblage The first step in variable saturated flow modeling included a calibration process to match the cumulative seepage discharge computed by the variably saturated flow modeling to the cumulative seepage discharge recorded at the outlet section of the lysimeter during laboratory experiments This was achieved by the inverse estimation of saturated hydraulic conductivity, Ksat, a prevailing soil parameter affecting the time of seepage and magnitude of the seepage discharge (Akay et al., 2008) HYDRUS, a finiteelement computer model which solves the Richards equation for Table List of the laboratory lysimeter tests including “Matrixa” and “Embankment” type configuration Configuration a Matrix Embankment Height of geofoam block assemblage Constant water head level 25 cm 38 cm 50 cm e 10 cm 1004 1705 1707 0206 3107 0806 0507 1906 2106 2004 2305 0908 2905 0908 0806 0507 1306 2206 0404 1805 1207 2305 0308 0206 2906 1206 1506 15 cm 22.5 cm 30 cm a “Matrix” configuration refers to a non-remediated slope that no geofoam blocks placed at the toe (Akay et al., 2013)   saturated-unsaturated water flow (Sim unek et al., 2012), was used To access the predictive ability of the calibration model of HYDRUS, NasheSutcliffe model efficiency coefficient (NS) was calculated, where a value of indicates a perfect match between modeled and observed discharge: 2 Pn  obs sim i¼1 Yi  Yi NS ¼  2 Pn  obs mean Y i¼1 Yi (4) where Yiobs is the ith observed cumulative discharge, Yisim is the ith simulated cumulative discharge, and Ymean is the mean of observed cumulative discharge Once the observed hydrological response of the laboratory lysimeter experiments matched the simulated response of the numerical models after calibration of Ksat, the next step included a coupled analysis of variable saturated flow modeling and slope stability modeling by using SEEP/W, a finite-element model solving Richards equation for two-dimensional variably saturated flow (Geo-Slope International, 2012a) and SLOPE/W, a conventional limit equilibrium analysis to determine slope stability (Geo-Slope International, 2012b), respectively The initial pressure head conditions needed for the start of the numerical models were obtained from the tensiometer-pressure transducer-datalogger setup used during the laboratory lysimeter experiments Instead of prescribing 22 different pressure heads at tensiometer locations, a single averaged value was input as an initial condition for the entire computational domain since the variance between recordings were minimal at the start of the laboratory lysimeter experiments (initial water content of the soil was constant during compaction of the slope) The boundary condition of the computational domain was prescribed as “hydrostatic pressure head boundary” for the inlet boundary simulating the water reservoir of the laboratory lysimeter The prescribed pressure heads referenced to the base of the slope were 25 cm-H2O, 38 cm-H2O, and 50 cm-H2O, respectively A “seepage face boundary” condition was prescribed to the outlet boundary simulating the exfiltration from the 45 hill-side of the compacted slope A no-flux boundary was selected for the bottom of the computational domain to simulate the impervious base of the lysimeter Numerical simulations were terminated at times equal to the end of the laboratory lysimeter experiments The Spencer method (Spencer, 1967), a limit equilibrium model, was utilized during slope stability analysis by SLOPE/W to estimate FSGL Evaluation of geofoam block slope systems by using conventional limit equilibrium analysis required assigning appropriate strength parameters not only to soil but also to geofoam blocks A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 171 Following Akay et al (2013), both cohesion and internal friction angle was defined for the geofoam blocks, where a cohesion value equal to one-quarter of the compressive strength of the geofoam used in this study (Table 1) The dry unit weight of geofoam blocks was 0.2 kN/m3 However, during design of geofoam block slope systems, potential weight gain from long-term water absorption is taken into account While an increase in the unit weight due to water absorption increases the driving forces, it also increases the resisting forces since normal stresses along the failure surface increases In order to discern which unit weight value is more conservative, stability analysis should be performed using both the dry unit weight and water absorbed unit weight (Arellano et al., 2011) For this reason, this study used the water absorbed value of 1.0 kN/ m3 as suggested by Stark et al (2004), which generated lower FSGL during stability analysis Fig Free-body diagram of geofoam block embankment located at the toe of the slope indicating the prevailing forces taken into account during the calculation of the Factor of Safety against sliding (FSSL) 2.4 Hydrostatic sliding Lightweight geofoam blocks are mainly used in slope remediation works to reduce the driving forces that cause global stability failure However, seepage flow may trigger hydrostatic sliding failure of the geofoam block assemblage since the resisting force, a function of normal stress, is reduced by the lightweight geofoam blocks For this reason, stability analyses have not only been performed against global stability but also against hydrostatic sliding failures Hydrostatic sliding mechanisms were first considered for stand-alone geofoam block embankments by Stark et al (2004) along the interface of bedding level and bottom of the embankment as a part of external stability analysis In this study, analysis of horizontal sliding of the geofoam block slope system at the interface between the bottom of the geofoam block embankment and the underlying foundation soil was conducted For this purpose, coupled seepage and stress modeling were performed to calculate total horizontal driving forces and total resisting forces acting on the geofoam block embankment to calculate FSSL as follows: P Horizontal Resisting Forces FSSL ¼ P Horizontal Driving Forces (5) Horizontal resisting and driving forces along with vertical forces acting on the geofoam block configuration are shown in Fig Total resisting force against hydrostatic sliding is comprised of only the shear resistance along the EPS/foundation soil interface (FF): FF ẳ ca A ỵ X X Ntan d N ẳ WEPS ỵ FT;V (6) (7) where ca is the interface cohesion along the EPS/foundation soils, A is the area resisting horizontal sliding, SN is the total normal forces acting along the interface, WEPS is the weight of the geofoam block assemblage, FT,V is the total vertical earth force (comprised of vertical hydrostatic and vertical effective earth pressures), and d is the interface friction angle between the geofoam blocks and the foundation soil The total horizontal driving force is the summation of the effective horizontal earth force (FE,H), and horizontal component of the hydrostatic force (FH,H) As a result, by neglecting the weight of the geofoam blocks, FSSL can be dened as: FSSL ẳ ca A ỵ FT;V tan d FE;H ỵ FH;H (8) After evaluating the EPS/sand interface test results reported by Jutkofsky (1998), Bartlett et al (2000), and Jutkofsky et al (2000), Stark et al (2004) recommended an EPS/sand interface friction angle, d, of 30 to use in design Xenaki and Athanasopoulos (2001) performed a series of direct shear tests to determine d of the EPS/ sand interface and discussed the effects of geofoam block density, the mean grain size, void ratio and grain shapes of sand particles on the test results They concluded that the failure envelope at the EPS/ sand interface is a non-linear curve, and may be approximated by a piecewise linear curve Under low normal stress range (0e35 kPa) a d value of 32 for geofoam blocks with a density of 20 kg/m3 were recommended Within this low normal stress range, d is approximately equal to the friction angle of the sand and the apparent interface adhesion, ca, is zero Xenaki and Athanasopoulos (2001) indicated that by increasing normal stress acting at the interface, the interaction behavior becomes progressively adhesional; therefore, while the value of d decreases, the value of ca increases until it becomes equal to the shear resistance of geofoam block Based on the interface shear testing results reported in the literature, and considering the stress ranges acting on the geofoam blocks in the lysimeter, a d of 30 and ca of kPa were selected to use for the calculation of FSSL as follows: FSSL ¼ FT;V tan 30 FE;H þ FH;H (9) In order to calculate the forces in the above equation, coupled seepage and stress modeling were performed using SEEP/W and SIGMA/W SIGMA/W is a stress-deformation computer program which includes six different analysis types: in-situ, stress redistribution, load/deformation, coupled stress-pore pressure, volume change, and dynamic deformation analysis (Geo-Slope International, 2009) As shown in the free body diagram (Fig 3), both the resisting and the driving forces due to earth and hydrostatic pressure acting on geofoam blocks are related with the gravity (self weight) of the variably saturated sand In-situ analysis (gravity turn-on analysis) converts the weight of the soil into stresses (Geo-Slope International, 2005) In-situ analysis uses the unit weight which simulates the effect of gravity, and Poisson’s ratio ðyÞ which simulates the earth pressure at rest ðKo ¼ y=1  yÞ In-situ analysis ignores stiffness related parameters (Geo-Slope International, 2012a,b) since there is no need to calculate strains at this stage Poisson’s ratios used to calculate Ko in the models for both sand and geofoam block are presented in Table Once the stress distributions acting on the geofoam block assemblage were obtained for both horizontal and vertical directions, the related forces were calculated as the total areas of these distributions 172 A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 Fig Slope physical conditions and hypothesized failure surfaces at the end of the duplicate experiments of “10 cm Embankment” assemblage under (a) 25 cm-, (b) 38 cm-, and (c) 50 cm-H2O pressure head boundary conditions (d) Representative failure surfaces for each boundary condition Results and discussion 3.1 Model calibration Inverse estimation of Ksat required the use of cumulative seepage discharge measurements recorded during laboratory lysimeter experiments However, not all experiments were eligible for calibration due to change in their physical condition during the experiment While the physical condition of the geofoam block slope system may change due to global failure and/or hydrostatic sliding during laboratory lysimeter experiments, the computational domain of the HYDRUS calibration model remains unchanged (time-invariant mesh structure) during simulations For this reason, model calibration involved experiments with 25 cm-H2O constant pressure head boundary condition during which the physical condition of the constructed slope remained unchanged as shown in the upcoming sections Moreover, since the sediment entrainment by the seepage flow was negligible for these experiments, direct conversion of digital scale weight measurements to seepage cumulative discharge (volumes of water) was possible The calibration routine of HYDRUS iteratively improved the initial estimation of the Ksat to match the simulated cumulative seepage discharge to the laboratory measured cumulative seepage discharge For each calibration model, laboratory measured value of 0.0295 cm/s (average of nine samples) was used as the initial estimate of the Ksat The calibrated Ksat has a value of 0.02647 cm/s, 0.04510 cm/s, 0.04791 cm/s, and 0.04398 cm/s for geofoam block assemblages “10 cm Embankment”, “15 cm Embankment”, “22.5 cm Embankment”, and “30 cm Embankment”, respectively These calibrated values of Ksat were utilized by the coupled analysis of SEEP/W and SLOPE/W Using the calibrated values of Ksat, the numerical model revealed an acceptable level of performance for the prediction of the cumulative discharge with high values of NS (NS  0.94) for all selected experiments 3.2 Geofoam block assemblage: 10 cm embankment The physical conditions formed at the end of the duplicate experiments of the “10 cm Embankment” geofoam block assemblage are given in Fig As in the case of “Matrix” configuration A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 experiments by Akay et al (2013), the experiments were terminated when the steady-state conditions were well established in terms of pore-water pressures measured by the tensiometers (Fig 5a) Steady-state condition refers to the horizontal portion of the soil pore-water pressure versus time graph (Fig 5) For example, the experiment “10cmEmb25cmHead1705”, for which the physical condition at the end of the experiment was shown in Fig 4a, reached steady-state condition near the water reservoir section of the slope (tensiometer 5) at 1330 s, coinciding with the 173 time of seepage (Fig 5b) The experiment was terminated at 3600 s soon after all tensiometers in the slope reached steady-state condition at approximately 2400 s Pore-water pressures measured by tensiometers located along the bottom of the slope during experiment “10cmEmb25cmHead1705” were given in Fig 5b Considering the final slope surfaces obtained at the end of the experiments, hypothesized global failure surfaces (entering from the crest and exiting at the toe of the slope) were indicated for each experiment with 38 cm-H2O and 50 cm-H2O constant water level in Fig Pore-water pressures measured by tensiometers during selected lysimeter experiments 174 A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 the reservoir (Fig 4bed) One exception was the “10cmEmb38cmHead0908” experiment which did not show any indication of change in physical shape with 38 cm-H2O boundary condition, and hence slope failure was not evident (Fig 4b) In contrast to the non-remediated slope (“Matrix” configuration) which showed a shallow-seated type global stability failure with 25 cm-H2O boundary condition in the earlier study by Akay et al (2013), the geofoam block slope system experienced only a small crack formation due to sliding on the slope face for experiment “10cmEmb25cmHead1705” (Fig 4a) On the other hand, its duplicate experiment “10cmEmb25cmHead1707” showed no signs of either global stability failure or hydrostatic sliding (Fig 4a) By taking into account the hypothesized failure surfaces indicated at the end of experiments, a representative failure surface was drawn (Fig 4d) Experiments showed that global static failures took place along the failure surface immediately behind the geofoam block assemblage Consequently, the representative failure surface shown in Fig 4d was used to calculate FSGL by the coupled analysis of SEEP/W and SLOPE/W FSGL values remained above 1.0 at all times during the simulation with 25 cm-H2O boundary condition confirming the laboratory observations that no global static failure took place with this boundary condition (Figs 4a and 6a) On the other hand, it is obvious that considering the strength parameters of the sand (f0 ¼ 33.7, c’ ¼ kPa in Table 1), the 1:1 slope face above the geofoam block assemblage is inherently unstable This has also been quantified by the slope stability analysis for the hypothesized failure surface given in Fig 4a Even though, the FSGL value was below 1.0, no shallow-seated global stability failure above geofoam block assemblage was observed during the experiments (Fig 4a) This can be attributed to the apparent cohesion of the partially-saturated sand which clearly acted as an additional resisting force (Fox and Wilson, 2010) It can be noticed that FSGL values tend to decrease with time during the simulation as the seepage flow progress into the failure zone and phreatic surface rises above the toe of the geofoam block slope system (Fig 6a) The effect of seepage on FSGL was even more severe for the cases with 38 cm-H2O and 50 cm-H2O boundary conditions (Fig 6a) The FSGL value decreased from 1.3 at the start of the simulation to 0.9 and 0.7 at the end of the simulation with 38 cm-H2O and 50 cm-H2O boundary conditions, respectively (Fig 6a) The FSGL value of 0.9 at steady-state condition, being slightly under the critical value of 1.0, indicated a marginally stable slope with 38 cm-H2O boundary condition This situation could be observed during laboratory testing that its duplicate, experiment “10cmEmb38cmHead0908”, did not show any indication of change in physical shape (Fig 4b) Failure due to hydrostatic sliding was observed with 25 cm-H2O boundary condition (experiment “10cmEmb25cmHead1705” in Fig 4a) Hydrostatic sliding was induced by seepage progressing into geofoam block assemblage only 590 s after the time of seepage (Fig 5b) At this time of crack formation, almost all tensiometers reached steady-state condition The in-situ stresses (total vertical stress, effective horizontal stress, and pore-water pressure) developed at three time intervals (beginning of the test, time of seepage, end of experiment) is presented in Fig The magnitudes of the resultant forces (kN/m) and computed FSSL values are given in Table The FSSL value of 1.0 at steady-state (end of experiment) with a 25 cm-H2O boundary condition reveals a critical condition in term of hydrostatic sliding type of failure This situation was experienced in the laboratory that only one of the duplicates formed a crack on the slope (experiment “10cmEmb25cmHead1705” in Fig 4a) while the other showed no signs of hydrostatic sliding Due to higher levels of the phreatic surface at steadystate, the FSSL value decreased to 0.9 and 0.8 with 38 cm-, and 50 cm-H2O boundary conditions, respectively Nevertheless, the geofoam block system experienced global static failures due to lower FSGL (Figs and 6a) In addition, the effect of seepage on FSSL was not severe as compared to its effect on FSGL The FSSL value Fig Change of FSGL values through the duration of the experiments A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 175 decreased from 1.0 at the start of the experiment to 0.8 at steadystate with 50 cm-H2O boundary condition whereas the decrease was from 1.3 to 0.7 for FSGL (Table 3) 3.3 Geofoam block assemblage: 15 cm embankment Fig Insitu stress [kPa] contours generated for “10 cm Embankment” assemblage under 25 cm-H2O pressure head boundary condition by SIGMA/W (a) Total vertical stress; (b) Effective horizontal stress; (c) Pore-water pressure The final slope surfaces of the duplicate experiments of the “15 cm Embankment” geofoam block assemblage are given in Fig Contrary to “10 cm Embankment” geofoam block assemblage, the slope showed no indication of a tension crack on the slope face at the end of the experiment with a 25 cm-H2O boundary condition; hence no hydrostatic sliding occurred In addition, calculated FSSL values throughout the test were greater than 1.0 which confirmed the laboratory slope model results (Table 3) Since the laboratory physical slope model tests showed no global stability failure at the end of the duplicate experiments (Fig 8a), a failure surface located immediately behind the geofoam blocks was selected (Fig 8d) to compute the magnitude of FSGL with a 25 cm-H2O boundary condition FSGL values were reported for this deep-seated global stability failure surface in Fig 6b Considering the change in FSGL with time during the experiment for this failure surface, it can be seen that the FSGL value remained above 1.0 throughout the experiment (Fig 6b), where steady-state conditions were well established (Fig 5c) The FSSL with 38 cm-H2O pressure head boundary condition decreased from 1.1 at the beginning of the experiment to 0.8 at the end of the experiment which demonstrated unstable condition against hydrostatic sliding (Table 3) This unstable condition for hydrostatic sliding was observed during only one of the duplicate experiments as a crack formed on the slope face of the experiment “15cmEmb38cmHead2905” (Fig 8b) As in the case of 25 cm-H2O pressure head, deep-seated global stability failure was not observed under 38 cm-H2O pressure head boundary condition (Fig 8b) Therefore, a failure surface located behind the geofoam blocks (Fig 8d) was selected to calculate the FSGL by SLOPE/W The FSGL value calculated by SLOPE/W for this representative failure surface at the end of the experiments was 0.9 which showed unstable condition against global failure (Fig 6b) However, the duplicate experiments showed no global stability failure (Fig 8b) The representative failure surface resulting from the 50 cm-H2O pressure head boundary condition was “deep-seated” type global stability failure (Fig 8c and d) The FSGL value decreased from 1.5 at the beginning of the experiment to 0.8 at the end of the experiment (Fig 6b), whereas FSSL value decreased from 1.1 to 0.8 (Table 3) Simultaneous reduction of both FSGL and FSSL below the critical value of 1.0 during the experiment (at the time of seepage; Table The magnitude of total vertical earth force (FT,V), effective horizontal earth force (FE,H), and horizontal component of hydrostatic force (FH,H), all in [kN/m], acting on the geofoam block embankment computed by using SIGMA/W stress contours at three time intervals under boundary conditions (BC) of 25 cm-, 38 cm-, and 50 cm-H2O Geofoam block assemblage 10 cm Embankment 15 cm Embankment 22.5 cm Embankment 30 cm Embankment BC 25 38 50 25 38 50 25 38 50 25 38 50 Beginning of the test Time of seepage End of experiment FT,V FE,H FH,H FSSL FT,V FE,H FH,H FSSL FT,V FE,H FH,H FSSL 0.233 0.233 0.233 0.406 0.406 0.406 0.710 0.710 0.710 1.060 1.060 1.060 0.133 0.133 0.133 0.220 0.220 0.220 0.327 0.327 0.327 0.450 0.450 0.450 0 0 0 0 0 0 1.0 1.0 1.0 1.1 1.1 1.1 1.3 1.3 1.3 1.4 1.4 1.4 0.233 0.249 0.266 0.414 0.431 0.462 0.741 0.782 0.860 1.108 1.200 1.280 0.131 0.123 0.129 0.213 0.203 0.210 0.305 0.286 0.306 0.391 0.420 0.410 0.003 0.024 0.050 0.013 0.055 0.101 0.048 0.100 0.209 0.058 0.145 0.270 1.0 1.0 0.9 1.1 1.0 0.9 1.2 1.2 1.0 1.4 1.2 1.1 0.253 0.277 0.313 0.427 0.461 0.517 0.766 0.827 0.900 1.150 1.220 1.270 0.115 0.110 0.121 0.180 0.192 0.183 0.292 0.267 0.262 0.400 0.376 0.354 0.033 0.075 0.098 0.054 0.137 0.194 0.087 0.230 0.352 0.113 0.270 0.405 1.0 0.9 0.8 1.1 0.8 0.8 1.2 1.0 0.8 1.3 1.1 1.0 176 A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 Fig Slope physical conditions and hypothesized failure surfaces at the end of the duplicate experiments of “15 cm Embankment” assemblage under (a) 25 cm-, (b) 38 cm-, and (c) 50 cm-H2O pressure head boundary conditions (d) Representative failure surfaces for each boundary condition FSGL ¼ 0.9 and FSSL ¼ 0.9) indicated that the “15 cm Embankment” geofoam block assemblage was ineffective to prevent global stability failure or hydrostatic sliding with 50 cm-H2O pressure head boundary condition (Fig 8c) 3.4 Geofoam block assemblage: 22.5 cm embankment The final physical conditions of the duplicate experiments of the “22.5 cm Embankment” geofoam block assemblage are given in Fig Since there is no global stability failure observed, the representative failure surface located immediately behind the geofoam blocks was selected with both 25 cm-H2O and 38 cm-H2O boundary conditions (indicated as 25 cm and 38 cm in Fig 9d, respectively) to report the FSGL values (Fig 6c) The FSGL values remained above 1.0 at the end of the tests (Fig 6c) with both 25 cmH2O and 38 cm-H2O boundary conditions in which steady-state conditions were well established (Fig 5d) The slopes showed no indication of crack formation on the slope faces at the end of the experiment with both 25 cm-H2O and 38 cm-H2O boundary conditions (Fig 9a and b) In addition, calculated FSSL values throughout the test were greater than 1.0 for both boundary conditions (Table 3) which is in agreement with the laboratory physical model results that indicated no signs of hydrostatic sliding (Fig 9a and b) Therefore “22.5 cm embankment” geofoam block assemblage was an effective remedial solution for the slopes tested with 25 cm-H2O and 38 cm-H2O boundary conditions against both hydrostatic sliding and deepseated global stability failure A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 177 Fig Slope physical conditions and hypothesized failure surfaces at the end of the duplicate experiments of “22.5 cm Embankment” assemblage under (a) 25 cm-, (b) 38 cm-, and (c) 50 cm-H2O pressure head boundary conditions (d) Representative failure surfaces for each boundary condition Crack formations were observed with 50 cm-H2O boundary conditions as a result of hydrostatic sliding (Fig 9c) Crack formation developed on the slope face at a location near the crest and extended diagonally into the slope in experiment “22.5cmEmb50cmHead0206”, whereas a crack formation developed on the crest in experiment “22.5cmEmb50cmHead2906” (Fig 9c) The calculated FSSL was 1.3 at the beginning of the experiment and reduced to 0.8 at the end of the experiment which confirmed the physical experiments On the other hand, no global stability failure was observed; however, the FSGL values were reported for a failure surface located immediately behind the geofoam blocks (indicated as 50 cm in Fig 9d) The FSGL value for this deep-seated global failure plane decreased from 2.1 at the beginning of the experiment to 1.1 at the end of the experiment (Fig 6c) Consequently, while the “22.5 cm embankment” geofoam block assemblage was ineffective at preventing hydrostatic sliding, it was an effective way of remediation to prevent deep-seated global failure observed by Akay et al (2013) for the “Matrix” configuration with the same boundary conditions 3.5 Geofoam block assemblage: 30 cm embankment The physical conditions at the end of the duplicate experiments of the “30 cm Embankment” geofoam block assemblage indicated that neither global stability nor hydrostatic sliding type failure occurred with the three different pressure head boundary 178 A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 conditions (Fig 10a, b, and c) Considering the previous representative failure surfaces that passed immediately behind the geofoam block assemblage, analysis of global stability failure by SLOPE/W was conducted for the failure surfaces shown in Fig 10d As expected and confirming the laboratory experiments, the FSGL values remained above the critical value of 1.0 throughout the experiments with 25 cm, 38 cm, and 50 cm-H2O pressure head boundary conditions (Fig 6d) The FSGL value decrease from 1.9 at the beginning of the experiment to 1.4 and 1.1 at the end of the experiment with 25 cm and 38 cm-H2O pressure head boundary conditions, respectively (Fig 6d) The FSGL values for the mentioned times with 50 cm-H2O pressure head boundary condition were 2.2 and 1.1 (Fig 6d) The “30 cm Embankment” geofoam block assemblage proved to be effective against hydrostatic sliding with values of FSSL higher than the critical value of 1.0 throughout the experiments with 25 cm, 38 cm, and 50 cm-H2O pressure head boundary conditions (Table 3) The FSSL values decreased from 1.4 to 1.3, 1.1, and 1.0 at the end of the experiments with 25 cm, 38 cm, and 50 cm-H2O pressure head boundary conditions, respectively (Table 3) It was also documented that, although a significant portion of the toe section of the slope was replaced by geofoam block assemblage (Fig 2d), the “30 cm Embankment” geofoam block assemblage did not alter the seepage flow dynamics at this location when compared to non-remediated “Matrix” configuration since the pore-water pressures recorded at the location of the Fig 10 Slope physical conditions at the end of the duplicate experiments of “30 cm Embankment” assemblage under (a) 25 cm-, (b) 38 cm-, and (c) 50 cm-H2O pressure head boundary conditions (d) Representative failure surfaces for each boundary condition A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 Tensiometer 12 with 25 cm-H2O pressure head boundary condition were almost the same for these two cases (Fig 5a and e) Nonappearance of excess pore-water pressure build-up behind the geofoam block assemblage indicated that flow could percolate through the joints of geofoam block assemblage Summary and conclusions Two types of failure mechanisms were observed during laboratory lysimeter experiments, namely, global stability failure of the slope and hydrostatic sliding failure of the geofoam block assemblage The physical condition of the slope at the end of the experiment was able to be explained by a rotational failure surface for tests that underwent global stability failure, whereas by a surficial crack formation for tests that underwent hydrostatic sliding of the geofoam block assemblage along the interface of bedding level and bottom of geofoam block assemblage Embankment type configuration could prevent deep-seated failures of a marginally stable sandy slope under seepage Deep-seated global stability failures were observed when the phreatic surface exits the slope above the top of the geofoam block assemblage This phenomenon was successfully predicted by the numerical model analysis in computing values of FSGL below the critical value (FSGL < 1.0) In addition, hydrostatic sliding analysis also demonstrated that crack formations could develop either on the slope face or on the crest if FSSL value decreases below the critical condition (FSGL < 1.0) Embankment type configuration could remediate the slope (FSGL > 1.0 and FSSL > 1.0) as long as the phreatic surface at steadystate did not rise above the top of the embankment on the slope face Even though current design precedence for geofoam usage in slope stabilization does not allow the phreatic surface to rise above the bottom of geofoam block assemblage with the use of permanent drainage system, the configuration tested in this study was able to withstand horizontal driving forces by using the overburden concept suggested for the design of stand-alone geofoam embankments over soft ground to increase resistance to sliding During the experiments, the inherently unstable 1:1 slope above the geofoam block assemblage did not experience shallow-seated global stability failure due to the apparent cohesion of the partially-saturated sand However, it should be noted that apparent cohesion cannot be considered as a permanent acting strength For this reason, additional preventive geotechnologies should be implemented for the part of the slope above the geofoam block assemblage Moreover, this study showed that conventional limit equilibrium analysis, in which appropriate strength parameters for both the soil medium and geofoam blocks were assigned, could be used for stability evaluation of geofoam block slope systems The geofoam block assemblages tested through laboratory experiments showed that geofoam block slope systems can be considered as a viable remediation technique where seepage induced global stability failure is a concern In addition, the quantification of the stability conditions by numerical stability analysis showed a similar trend with the laboratory models Thus, the results of the laboratory models provided a basis to understand the real behavior of the geofoam block configuration used to remediate a steep sandy slope under seepage However, scale effects, due to differences in the stress levels between the 1:20 scale 1-g laboratory model and field prototype, avoids the extrapolation of the model behavior to field-constructed slopes For future work, it is recommended to test embankment configuration in the field by using fully instrumented prototype models, in order to verify the performance of laboratory lysimeter experiments In addition, types of separation materials between the bottom of the geofoam block assemblage and foundation soil should be investigated in 179 terms of providing additional resistance against hydrostatic sliding of geofoam block assemblages Acknowledgments The lysimeter experiments were conducted at Okan University Civil Engineering Laboratory, Istanbul, Turkey The writers acknowledge the contributions of Halis S¸ahin, laboratory techniuz S¸ims¸ek, cian, and undergraduate student assistants: Gönüllü Og atay, and Mehmet Og uz Anıl Çag References Aabøe, R., 2011 40 years of experience with the use of EPS geofoam blocks in road construction In: Proceedings of 4th International Conference on Geofoam Blocks in Construction Applications, Lillestrøm, Norway Akay, O., Fox, G.A., 2007 Experimental investigation of direct connectivity between macropores and subsurface drains during infiltration Soil Sci Soc Am J 71 (5), 1600e1606   Akay, O., Fox, G.A., Sim unek, J., 2008 Numerical simulation of flow dynamics during macropore-subsurface drain interactions using HYDRUS Vadose Zone J (3), 909e918 Akay, O., Özer, A.T., Fox, G.A., 2012 Experimental investigation of failure mechanism of expanded polystyrene block geofoam slope system under seepage In: 5th European Geosynthetics Congress, EuroGeo5, 16th e 19th of September, Valencia, Spain Soil Improvement and Reinforcement, vol 4, pp 13e17 Akay, O., Özer, A.T., Fox, G.A., Bartlett, S.F., Arellano, D., 2013 Behavior of sandy slopes remediated by EPS-block geofoam under seepage flow Geotext Geomemb 37, 81e98 Arellano, D., Stark, T.D., Horvath, J.S., Leshchinsky, D., 2011 Guidelines for Geofoam Applications in Slope Stability Projects Transportation Research Board, Washington, D.C., USA Final report NCHRP Project No 24e11(02) http://onlinepubs trb.org/onlinepubs/nchrp/docs/NCHRP24-11(02)_FR.pdf ASTM Standard D 1621, 2010 Standard Test Method for Compressive Properties of Rigid Cellular Plastics American Society for Testing and Materials, Pennsylvania, USA ASTM Standard D 2434, 2006 Standard Test Method for Permeability of Granular Soils (Constant Head) American Society for Testing and Materials, Pennsylvania, USA ASTM Standard D 4767, 2011 Standard Test Method for Consolidated Undrained Triaxial Compression Test for Cohesive Soils American Society for Testing and Materials, Pennsylvania, USA ASTM Standard D 6913, 2009 Standard Test Methods for Particle-size Distribution (Gradation) of Soil Using Sieve Analysis American Society for Testing and Materials, Pennsylvania, USA ASTM Standard D 854, 2010 Standard Test Methods for Specific Gravity of Soil Solids by Water Pycnometer American Society for Testing and Materials, Pennsylvania, USA Bartlett, S., Negussey, D., Kimble, M., Sheeley, M., 2000 Use of geofoam as superlightweight fill for I-15 reconstruction In: Proceedings of Transportation Research Board 79th Annual Meeting, Washington, D.C., USA Benjamim, C.V.S., Bueno, B.S., Zornberg, J.G., 2007 Field monitoring evaluation of geotextile-reinforced soil-retaining walls Geosynth Int 14 (2), 100e118 Choudhary, A.K., Jha, J.N., Gill, K.S., 2010 Laboratory investigation of bearing capacity behavior of strip footing on reinforced flyash slope Geotext Geomemb 28, 393e402 Dash, S.K., Sireesh, S., Sitharam, T.G., 2003 Model studies on circular footing supported on geocell reinforced sand underlain by soft clay Geotext Geomemb 21, 197e219 Deb, K., Samadhiya, N.K., Namdeo, J.B., 2011 Laboratory model studies on unreinforced and geo-grid reinforced sand bed over Stone column-improved soft clay Geotext Geomemb 29, 190e196 Dronamraju, V.S., 2008 Studies on Field Stabilization Methods to Prevent Surficial Slope Failures of Earthfill Dams Ph.D Dissertation The University of Texas, Arlington Duncan, J.M., Wright, S.G., 2005 Soil Strength and Slope Stability John Wiley and Sons, Inc., New Jersey El Sawwaf, M.A., 2007a Behavior of strip footing on geogrid-reinforced sand over a soft clay slope Geotext Geomemb 25, 50e60 El Sawwaf, M.A., 2007b Uplift behavior of horizontal anchor plates buried in geosynthetic reinforced slopes Geotech Test J 30 (5), 1e9 Elragi, A.F., 2000 Selected Engineering Properties and Applications of EPS Geofoam Ph.D Dissertation State University of New York College of Environmental Science and Forestry, Syracuse, NY Fay, L., Akin, M., Shi, X., 2012 Cost-effective and Sustainable Road Slope Stabilization and Erosion Control In: A Synthesis of Highway Practice Transportation Research Board, Washington, D.C NCHRP Synthesis 430 http://onlinepubs.trb org/onlinepubs/nchrp/nchrp_syn_430.pdf Fox, G.A., Wilson, G.V., 2010 The role of subsurface flow in hillslope and streambank erosion: a review Soil Sci Soc Am J 74 (3), 717e733 180 A.T Özer et al / Geotextiles and Geomembranes 42 (2014) 166e180 Fox, G.A., Wilson, G.V., Periketi, R.K., Cullum, R.F., 2006 Sediment transport model for seepage erosion of streambank sediment J Hydrol Eng 11 (6), 603e611 Geo-Slope International Ltd, 2005 Geo-Slope Direct Contact In: Establishing the Insitu Stresses http://www.geo-slope.com/contact/directcontact_dec05.html Calgary, Alberta, Canada Geo-Slope International Ltd, 2009 Stability Modeling with SIGMA/W 2007, an Engineering Methodology, fourth ed Calgary, Alberta, Canada Geo-Slope International Ltd, 2012a Seepage Modeling with SEEP/W 2012, an Engineering Methodology, November 2012 ed Calgary, Alberta, Canada Geo-Slope International Ltd, 2012b Stability Modeling with SLOPE/W 2012, an Engineering Methodology, November 2012 ed Calgary, Alberta, Canada Jutkofsky, W.S., 1998 Geofoam Stabilization of an Embankment Slope, a Case Study of Route 23A in the Town of Jewett, Greene County Geotechnical Engineering Bureau, New York State Department of Transportation, Albany, NY, USA Jutkofsky, W.S., Sung, J.T., Negussey, D., 2000 Stabilization of an embankment slope with geofoam J Transp Res Board 1736, 94e102 Latha, M.L., Somwanshi, A., 2009 Bearing capacity of square footings on geosynthetic reinforced sand Geotext Geomemb 27, 281e294 Lee, K.M., Manjunath, V.R., 2000 Experimental and numerical studies of geosynthetic-reinforced sand slopes loaded with a footing Can Geotech J 37 (4), 828e842 Mann, G., Stark, T.D., 2007 Slope stabilization using geofoam In: Proceedings of Geo-Denver 2007: New Peaks in Geotechnics GSP 161 Embankments, Dams, and Slopes ASCE Portelinha, F.H.M., Bueno, B.S., Zornberg, J.G., 2013 Performance of nonwoven geotextile-reinforced walls under wetting conditions: laboratory and field investigations Geosynth Int 20 (2), 90e104 Reuter, G.R., 2001 Use of geofoam for landslide stabilization-CTH “A”, Bayfield County, Wisconsin In: Proceedings of the 3rd International Geofoam Conference, Salt Lake City, Utah Reuter, G., Rutz, J., 2000 A lightweight solution for landslide stabilization Geotech Fabr 18 (7), 42e43 Schmertmann, J.H., 2006 Estimating slope stability reduction due to rain infiltration mounding J Geotech Geoenviron Eng 132 (9), 1219e1227 Shah, B.H., 2008 Field Manuel on Slope Stabilization United Nations Development Program, Pakistan http://www.preventionweb.net/english/professional/ publications/v.php?id¼13232 Sidle, R.C., Ochiai, H., 2006 Landslides: Processes, Prediction, and Land Use In: Water Resources Monograph No 18 American Geophysical Union, Washington, D.C., p 312    Sim unek, J., van Genuchten, M.Th., Sejna, M., 2012 The HYDRUS Software Package for Simulating Two- and Three Dimensional Movement of Water, Heat, and Multiple Solutes in Variably-saturated Media, Technical Manual, Version 2.0, PC Progress, Prague, Czech Republic Spencer, E., 1967 A method of analysis of embankments assuming parallel inter slice forces Geotechnique 17 (1), 11e26 Stark, T.D., Arellano, D., Horvath, J.S., Leshchinsky, D., 2004 Guideline and Recommended Standard for Geofoam Applications in Highway Embankments Transportation Research Board, Washington, D.C NCHRP Report 529 http://trb org/publications/nchrp/nchrp_rpt_529.pdf Tsukamoto, H., 1996 Slope stabilization by the EPS method and its applications In: Proceedings of the International Symposium on EPS Construction Method (EPS Tokyo ’96) EPS Construction Method Development Organization, Tokyo, Japan, pp 362e380 van Genuchten, M.Th, 1980 A closed-form equation for predicting the hydraulic conductivity of unsaturated soils Soil Sci Soc Am J 44 (5), 892e898 van Genuchten, M.Th., Leij, F.J., Yates, S.R., 1991 The RETC Code for Quantifying the Hydraulic Functions of Unsaturated Soils, Version 1.0 EPA Report 600/2e91/ 065 U.S Salinity Laboratory, USDA, ARS, Riverside, California Xenaki, V.C., Athanasopoulos, G.A., 2001 Experimental investigation of the interaction mechanism at the EPS geofoam-sand interface by direct shear testing Geosynth Int (6), 471e499 Wilson, G.V., Periketi, R.K., Fox, G.A., Dabner, S.M., Shields, F.D., Cullum, R.F., 2007 Soil properties controlling seepage erosion contributions to streambank failure Earth Surf Proc Land 32, 447e459 Yoo, C., 2001 Laboratory investigation of bearing capacity behavior of strip footing on geogrid-reinforced sand slope Geotext Geomemb 19, 279e298  ... evaluate the use of geofoam block for remediation of sandy slopes that experience deep-seated global stability failures under seepage The overall objective of this study was to evaluate a geofoam. .. results of this research are relevant to revealing insights of using the proposed geofoam block configuration for remediation of sandy slopes susceptible to seepage forces at 1:20 scale This study also... groundwater table may rise above the bottom of the geofoam blocks which may cause global stability failure of the slope and/or hydrostatic sliding failure of geofoam block assemblage The behavior of