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Geotextiles and Geomembranes 42 (2014) 302e311

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Geotextiles and Geomembranes

journal homepage: www.elsevier .com/locate/geotexmem

Instability of a geogrid reinforced soil wall on thick soft Shanghai claywith prefabricated vertical drains: A case study

Jian-Feng Xue a, b, Jian-Feng Chen a, *, Jun-Xiu Liu a, Zhen-Ming Shi a

a Department of Geotechnical Engineering, School of Civil Engineering, Tongji University, 1239 Siping Road, Shanghai 200092, Chinab School of Applied Sciences and Engineering, Monash University, Churchill 3842, Victoria, Australia

a r t i c l e i n f o

Article history:Received 21 January 2014Received in revised form25 April 2014Accepted 5 May 2014Available online 2 June 2014

Keywords:Reinforced soil wallSoft soilsPrefabricated vertical drainsPore water pressureGeneral sliding failure

* Corresponding author. Tel./fax: þ86 21 65983545E-mail address: jf_chen@tongji.edu.cn (J.-F. Chen).

http://dx.doi.org/10.1016/j.geotexmem.2014.05.0030266-1144/© 2014 Elsevier Ltd. All rights reserved.

a b s t r a c t

A 7.6 m high geogrid reinforced soil retaining wall (RSW) was constructed at the end of an embankmenton very thick, soft Shanghai clay with 12 m deep prefabricated vertical drains (PVDs). The settlement ofthe ground, the wall movement and pore water pressure were monitored during the construction. Fromday 118, halfway through the construction, unexpected pore water pressure increment was recordedfrom the pore water pressure meters installed in the PVD drained zone indicating a possible malfunctionof the PVDs due to large deformation in the ground. After the last loading stage, on day 190, a suddenhorizontal movement at the toe was observed, followed by an arc shaped crack on the embankmentsurface at the end of the reinforced backfill zones. The wall was analyzed with a coupled mechanical andhydraulic finite element (FE) model. The analysis considered two scenarios: one with PVDs fully func-tional, and the second one with PVD failure after day 118 by manually deactivating the PVDs in the FEmodel. The comparison between the measured and simulated ground settlement, toe movement, andpore water pressure supported the assumption on the malfunction of the PVDs. It is believed that thegeneral sliding failure in the wall was caused by the increase of pore water pressure in the foundation soiland soils in front of the toe. It is suggested that possible failure of PVDs should be considered in thedesign of such structures, and the discharge rate of the PVDs and the pore water pressure should beclosely monitored during the construction of high soil walls on soft soils to update the stability of thestructures, especially for grounds where large deformations are expected which may cause the failure ofthe PVDs.

© 2014 Elsevier Ltd. All rights reserved.

1. Introduction

Geosynthetic reinforced soil walls (RSWs) have been widelyused throughout the world in road embankment and retainingstructures. RSWs have many advantages including aesthetics, shortconstruction period, good wall stability, cost effectiveness, goodseismic response, strong adaptability on soft highly compressiblefoundation soils and the ability to tolerate large differential set-tlement (Tatsuoka et al., 1997; Bloomfield et al., 2001; Yoo and Jung,2004).

Research has been done on the behavior of RSWs on various soilfoundations, and various loading conditions through a large num-ber of in-situ and laboratory tests as well as theoretical analyses(Rowe and Skinner, 2001; Mandal and Joshi, 1996; Viswanadhamand Konig, 2009; Leshchinsky and Han, 2004; Yoo and Jung,

.

2008; Huang et al., 2011; Raisinghani and Viswanadham, 2011).Despite the fact that many geosynthetic-reinforced soil walls havebeen safely constructed to date, there is limited literature on theperformance and behavior of the reinforced soil wall on PVDdrained soft soils (Alfaro et al., 1997; Bloomfield et al., 2001;Skinner and Rowe, 2005; Tanchaisawat et al., 2008; Demir et al.,2013), and even fewer reports on failure case studies to a furtherreview of their original design and therefore the failure mechanism(Leonards et al., 1994; Tatsuoka et al., 1997; Collins, 2001; Ling et al.,2001; Borges and Cardoso, 2002; Yoo et al., 2004, 2006;Scarborough, 2005). There are still many areas that need in-depthstudies which will help in the safe construction of RSWs on softclay drained with PVDs (Koerner and Koerner, 2013).

This paper presents a failure case history of a 7.6m high geogrid-reinforced soil wall with vertical wrap-around facing constructedon a typical Shanghai multi-layer soft soil ground installed withPVDs. Unexpected increase of pore water pressure was observed inthe PVD drained soils during the construction, followed by a largedeformation in the toe, and an arc shape crack in the surface of the

Fig. 2. Detail of the RSW and the ground condition.

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embankment at the end of the reinforced zone, indicating thefailure of the wall after the last layer of backfill was placed. Nu-merical modeling was carried out to compare with the in-situmonitored data to identify the failure mechanism of the wall andpossible causes of the failure. Recommendations were made at theend for the design and construction of RSWs on soft ground.

2. Description of the RSW project

2.1. Site description

The project was located in Shanghai Botanic Garden, Shanghai,China. The RSWwall is a temporary retaining structure at the end ofa 530 m long embankment (see Fig. 1). The embankment is 37.2 mwide on top with a 1.5H: 1V side slope, as shown in Fig. 2 with theground condition: five layers of soft soils with a total thickness of31 m from the ground level, which are (from top to bottom): 2.6 mthick of silty clay, 4.4 m thick of mucky silty clay, 3.6 m thick of clay,7.4m thick of silty clay and 13.0m thick of silty clay. Underneath is alayer of 6.3 m thick stiff clayey silt with 31 of average SPT blowcount, underlain by stiff silty sand and clay. The ground water tabledepth is shallow and it fluctuates within 0.5 m below the groundsurface. Table 1 summarizes the properties of the subsoil obtainedfrom in situ and laboratory tests.

2.2. RSW construction

Thewall was constructed in layers via staged filling. The sitewasprefilled with 1.6 m thick soil excavated from layer 1 silty clay tocompensate the settlement. Above the fill, a 0.6 m thick mediumsand was compacted to a bulk density of 17.0 kN/m3 as a sandcushion for drainage as shown in Fig. 3. Prefabricated vertical drains(PVDs) were installed in a triangle patternwith a spacing of 1.5m tothe depth of 12 m below the surface of the sand cushion. Theembankmentwas constructed using the soil excavated from Layer 1silty clay compacted at the lift of 20 cm to the bulk density of19.0 kN/m3 using a 15 ton static double drum road roller (1.5 mbehind the wall face) with compaction width of 2.12 m, and alightweight walk-behind roller (within 1.5 m from the wall face).The loading stages are shown in Fig. 4.

Fourteen layers of 10 m long high density polyethylene (HDPE)uniaxial geogrid were placed at a vertical spacing of 0.6 m (top 6layers) and 0.5 m (bottom layers), with 3.5 m of wrap-aroundsegment at the wall facing. The axial stiffness of the geogrid is620 kN/m at 5% strain and the strength is 70 kN/m.

Fig. 1. The RSW on soft clay.

2.3. In situ measurement

Below instruments were installed along the centerline of theembankment as shown in Fig. 3:

1) : two settlement plates (C1 and C2) installed at the initialground surface 4m and 20m behind thewall toe tomeasure theground settlement;

2) : two sets of magnetic extensometers installed at 2.5 m (S1) and15m (S2) away from thewall toe. Five magnets were installed ineach extensometer at the depth of 4.4 m, 8.1 m, 11.8 m, 14.6 mand 20.6 m, namely S1-1 to S1-5 and S2-1 to S2-5;

3) : two rows of pore water pressure meters installed at 2 m (P1)and 18 m (P2) away from the wall toe at the depth of 2.6 m, 7 m,10.6 m,18 m and 24.5 m, namely P1-1 to P1-6, and P2-1 to P2-6;

4) : four displacement markers on the wall face at the height of 1.7,2.7, 3.7 and 4.7 m above the toe of the wall to monitor the walldisplacement from day 139;

5) : a settlement plate (D1) was installed at 0.5 m away from thewall toe to monitor the displacement of the toe.

3. Measurement results

3.1. Excessive pore water pressure

The measured excess pore water pressures at the monitoreddepth are shown in Fig. 5. Overall the observed excess pore waterpressure in the two rows of pore water pressure meters follows thesame trend. The figure shows that at slow loading rate, excess porewater pressure decreased steadily with time, e.g. from day 50e118.Pore water pressure increased rapidly during the days 36e40, 138to 140, and 187 to 189 when the loading rate was high, especially atshallower depth, e.g. above the depth of �10.6 m. The pore waterpressure increment at the deeper locations, e.g. below 18 m, is verylow, e.g. less than 10 kPa during the whole period.

There was an unexpected rise of pore water pressure on day 118in most of the pressure meters, while there were no constructionactivities. After day 118, the pore water pressure in the ground didnot seem to dissipate at all, especially at shallower depth(above �10.6 m). The increase of the pore water pressure observedin the ground on day 118 and slow dissipation of pore waterpressure after that is perhaps due to the failure of the PVDsresulting from the large deformation in the ground and wallmovement during construction. Chu et al. (2006) found that200mmof settlement in a vacuum pressure preloaded PVD drainedsoft ground can cause the buckling of PVDs and dramatic reduction,e.g. 84%, in the PVD discharge capacity. As shown in Fig. 6 and Fig. 7,till day 118, therewas a total settlement of more than 400mm in C1

Table 1Summary of geotechnical properties of the subsoil.

Layer USCS classification g wn Ip eo c' f' Eref50 kh kv

(kN/m3) (%) (%) (kPa) (�) (MPa) (10�7 cm/s) (10�7 cm/s)

1 CL 18.4 32.5 16.2 0.93 7.0 27.9 3.18 16.0 1.142 CL 17.9 38.0 15.1 1.06 7.0 28.0 2.12 16.4 1.083 CL 17.5 41.1 18.0 1.16 6.0 24.9 2.47 13.9 1.334 CL 19.5 24.0 15.1 0.70 17.0 29.8 8.37 1.6 0.125 CL 19.3 25.2 15.1 0.73 18.0 31.2 5.70 1.6 0.12

Note: USCS: unified soil classification system, g ¼ unit weight; wn ¼ natural water content; Ip ¼ plasticity index; e0 ¼ initial void ratio; c' ¼ effective cohesion; f' ¼ effectivefriction angle; Eref50 ¼ secant modulus in standard drained triaxial test; kh ¼ horizontal permeability; kv ¼ vertical permeability.

Fig. 3. Profile of the RSW and the ground surface, with the location of instruments.

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and C2, and more than 150 mm horizontal movement at the toe.With such a large deformation, there is a high possibility of mal-function or partial malfunction of the PVDs. This will be furtherdiscussed in the following sections.

3.2. Observed ground settlement and failure of the RSW

Fig. 6 shows the ground settlement measured from S1-1 to S1-4and S2-1 to S2-4, and C1 and C2. There is little settlement observedin S1-5 and S2-5, at the depth of 20.6 m, therefore the results fromthese two magnets are not shown. It can be seen that the settle-ment curves follow the loading stages. There were two sharp dropsin the settlement curve. The first one was during the days 31e40,with a settlement of 69 mm in C1 in 10 days, then followed by agentle curve during a three-week resting period. This agrees withthe observed dissipation of excess pore water pressure as shown inFig. 5. The second sharp settlement was after day 187 when the lastlayer of embankment was laid, then followed by the failure of thewall.

The figure also shows that, below the depth of 10.6 m, at the endof the PVDs, there is not much settlement occurred. This suggestedthat there was not much consolidation below that depth during the

construction period, which agrees with the observed low excesspore water pressure at the depth as shown in Fig. 5.

From day 187, sudden settlement was recorded in almost all themonitored points. On day 190, an arc-shaped crack was observed atthe end of the reinforced zone, which is 10 m away from the RSWfacing as plotted in Fig. 2. The shape and location of the crack mayindicate a deep-seated failure in thewall. To prevent further sliding,surcharge was placed in front of the wall on day 200, and 0.5 mdiameter and 28 m long concrete pipe piles were installed 5 maway from the wall facing in a square patternwith a spacing of 3 m.

3.3. Movement of the RSW

Fig. 7 shows the ground movement at the toe of the wall. Thefigure shows that, before failure, the horizontal movement at thetoe is less than the vertical movement, and the movement curvesreflect the loading stages well. From day 187, there is a suddenmovement away from the embankment at the toe on horizontaldirection, while the vertical settlement is relatively steady. Fromday 187 to day 189, the horizontal movement of the toe increasedabout 15 mm, followed by a 67 mm movement in 24 h to day 191,which indicates a possible block sliding of the wall.

Fig. 6. Ground settlement with time.

Fig. 7. Movement at the toe of the wall during construction.

Fig. 8. Settlement of the embankment at three monitored locations.

Fig. 5. Excess pore water pressure with time.

Fig. 4. Fill height vs. time curve.

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Fig. 8 compares the settlement of the toe and measurementsfrom the two plates C1 and C2. It can be seen that from day 136 today 187, the settlement curve of the toe concave upward instead ofdownward as shown in the other two curves. This may suggest that,during that period, the toe of the wall may have been subjected to aslightly heave due to rotation of the wall.

3.4. Deformation of the wall face

Fig. 9 shows the horizontal movement of the wall face from day139 at the level of 1.7 m, 2.7 m, 3.7 m and 4.7 m above the toe of thewall. It can be seen that the wall moved outward dramatically afterday 187, or start of the last loading stage. Fig. 10 (a) shows thebulging of the wall face, which indicates that the maximum bulgingis at the middle height (3.7 m) of the wall. Fig. 10 (b) shows the netincrement of displacement of the four monitored levels after day156, which indicates a block sliding outward with slightly tilting.More bulging was observed at the middle height of the wall duringthe days 190e195, which is a common phenomenon for flexiblefacings walls due to compaction stresses, self-weight of the backfill,lack of section modulus or toe fixity (FHWA, 2009; Ehrlich andMirmoradi, 2013).

4. Numerical modeling

PLAXIS has been used by researchers to solve large deformationof embankment on soft soils, e.g. more than 1 m of settlement(Brinkgreve et al., 2004; Bergadoa et al., 2002; Chaiyaput et al.,

Fig. 9. Movement at different locations of the wall face during construction.

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2014). The failure mechanism of the wall was studied using thetwo-dimensional (2D) coupled mechanical and hydraulic FE modelon PLAXIS Version 8.2 platform. The PVDs were modeled withdrains in PLAXIS. The failure of the PVDs was modeled by turningthe drains off in the analysis. To turn off all the PVDs may be aconservative choice in the simulation when there is limited infor-mation available. Another choice is to use equivalent permeabilityof soils obtained from in-situ measured discharge rate instead ofusing drains in the numerical simulation to consider the effect ofdrain failure.

The stability of the RSW was examined using the StrengthReduction Method (SRM) incorporated in the PLAXIS program. Inthe model, 50 extra days were simulated after the last loadingstage. It is worthwhile to note that after the last loading stage, thesimulation is only for comparison purpose, as the main purpose ofthe paper is to discuss the behavior of the wall before the failure.

Fig. 10. Deflection of the wall face during construction, (a): total displacement fromday one, (b) net displacement after day 156.

4.1. Conversion of axisymmetric drainage to plane drainage

Tang et al. (2013) proposed a closed-form solution for theconsolidation of vertically drained three layer soil system. Toanalyze the 3D effect of PVDs in a plane strain condition, Hird et al.(1995) proposed to use the following equation to obtain theequivalent permeability of the soil system:

khp ¼ 23$B2

D2e$1m$kh (1)

where kh ¼ the horizontal permeability of subsoil; khp ¼ theequivalent horizontal permeability of the soil system in plane straincondition; B ¼ the width of the plane-strain unit cell; De ¼ thediameter of the effective zone of drainage; The dimensionless factorm is to account for the smear effect, which can be defined as:

m ¼ lnðn=sÞ þ ðkh=ksÞln s� 0:75 (2)

where ks ¼ the horizontal permeability in the smeared zone; n¼De/dw. The factor of s is defined as the ratio of the diameter of thesmeared zone (ds) and the diameter of the vertical drain (dw), ds/dw.The ratio of kh/ks varies with soils and ground conditions. Hansbo(1987) found that the value can vary from 1 to 5 for clay samplesin the laboratory. According to the in-situ measurement on anearby site by Shen et al. (2005), the value of kh/ks is about 13.5 for

soft clays in the tested area. Therefore, a value of 13.5 was adoptedfor kh/ks in the following numerical analysis.

Rixner et al. (1986) recommended the following equation toaccount for the shape effect of the PVDs using the equivalentdiameter:

dw ¼ wþ t2

(3)

where w ¼ the width of a band-shaped PVD; t ¼ the thickness ofthe PVD. Thewidth and the thickness of the PVDs used in this studywere 100 mm and 4 mm respectively.

Jamiolkowski et al. (1983) found that the diameter of the smearzone, ds, is proportional to the diameter of the cross-sectional areaof mandrel (dm):

ds ¼ ð2e3Þdm (4)

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In this study, the diameter of the smear zone is defined as threetimes of the diameter of the mandrel, 110 mm in this study, asrecommended by Chai and Miura (1999).

For triangle pattern of PVDs, the diameter of unit cell, De, is

De ¼ 1:05d (5)

where d ¼ spacing of PVDs, and d ¼ 1.5m in this study.Table 2 summarizes the parameters used in Eqs. (1) and (2).

Using the parameters, we obtained m ¼ 25.8 and khp ¼ 0.023kh. Asthe discharge capacity of the PVDs is very high (up to 100 m3/year),and staged-loading is used in the construction to avoid excess porewater pressure buildup in the drains, the effect of well resistancewas not considered in the simulation.

4.2. Numerical model and parameters

As the embankment is very wide and long comparing to itsheight, the central section of the embankment was chosen tosimulate the plane strain condition. The lower boundary was set at31 m from the original ground surface to the bottom of layer 5.Ninetymeters longof longitudinal section is simulated in themodel,with 45 m behind the RSW and 45 m ahead of the RSW. Standardboundary condition was used in the model, with water table at theground surface. An unstructured finite-element fine mesh consist-ing of 15-node triangular elements was used as shown in Fig.11. Themeshes and pore water pressure were updated using update meshtechnique in the analysis to consider the large ground deformation.

The preliminary fill, the sand cushion, and the RSW backfill weremodeled with linear elastic-perfectly plastic Mohr-Coulombmodel,and the hardening-soil model was used for the foundation soils.The drain element inherent in PLAXIS was used to model PVDs, inwhich excess pore water pressures are set to zero in the drain el-ements. As discussed in Section 3.1, the equivalent horizontalpermeability was used for soils in layers 1 to 3, in which the PVDswere installed.

The geogrid reinforcements were modeled as elastoplastic ma-terial, using geogrid element in PLAXIS, with axial stiffness ofJ ¼ 620 kN/m and the tensile strength of T ¼ 70 kN/m. The soil andgeogrid interface behavior have been studied by many researchers(Zhou et al., 2012). For simplification and availability of the modelsin the software, a rigid model was adopted for the soil and geogridinterface, with an interface coefficient of 0.7 based on pullout tests.

Table 3 summarizes the values of the soil parameters for the FEanalysis. The following empirical relationships: Erefur ¼ 3Erefoed ¼ 3Eref50andm¼ 1 (Schanz et al., 1999) were used to develop the hardening-soil model, inwhich Erefoed ¼ tangent stiffness for primary oedometerloading and Erefur ¼ Young’s modulus for unloading and reloading.

4.3. Results of the modeling

The embankment was simulated under two scenarios: withdrains on during the whole construction period, and with drains offafter day 118 to simulate the failure of the PVDs. After deactivating

Table 2Parameters for Eqs. (1) and (2).

Item Symbol Value

Width of the plane-strain unit cell (m) B 1.5Equivalent drain diameter (mm) dw 52Ratio of kh over ks in field kh/ks 13.5Smear zone diameter (mm) ds 330ds/dw s 6.35diameter of unit cell (m) De 1.58De/dw n 30.3

the drains, the horizontal permeability kh was used for the soils inthe PVD drained zone instead of khp, whichmay not represent the insitu condition, as the soils have been consolidated which will resultin lower permeability.

4.3.1. Excess pore water pressureThe modeled excess pore water pressures at the monitored

points are shown in Fig. 12. The results show that with all the PVDson, the excess pore water pressure dissipates quickly after eachloading stage to values lower than 20 kPa. In the first few loadingstages (before day 118), the measured and modeled values agreewell with each other. After day 118, the measured values seemunreasonably higher than the modeled results.

As discussed in the previous section, from day 118, there mightbe a failure of PVDs due to large deformation in the ground. Tovalidate this assumption, the drains were turned off in PLAXIS fromday 118 and a second simulation was carried out with all the otherparameters unchanged except the horizontal permeability of thesoils in the PVD zone as explained earlier. The simulated excesspore water pressures were compared with measured values inFig. 12 (with PVDs off). The figure shows that the simulated valuecaptures the excess pore water pressure variation very well afterday 118, which supports the assumption that the PVDs have failedafter day 118.

4.3.2. Ground movementThe measured settlements at C1 are compared with the simu-

lated values in Fig. 13. It can be seen that both simulations agreewell with themeasured settlement curve of C1 before day 118.Withthe PVDs on, the predicted settlements of the locations agree wellwith the measured values till day 189, but underestimates theoverall settlement in the long term, e.g. till day 200. With the PVDsoff, the predicted total settlement matches the measured valuewell. This again supports the assumption that the PVDs may havefailed during the construction.

4.3.3. Movement of the toeThe movement of the toe is plotted in Fig. 14. The figures show

that, for vertical settlement, the numerical model provides rela-tively good estimation for both situations. For lateral movement,the model well overestimated the movement after turning off thedrains, and under estimated the total settlement after the lastloading with the drains on. The reasons for this bad performancecan be numerous (Tavenas et al., 1979; Ehrlich and Mirmoradi,2013), such as boundary condition, soil anisotropy, the soil modeladopted in the simulation and the oversimplification of the possiblepartial failure of PVDs by assuming a full failure. In this case, themovement of the toe area is a combination of rotation due to globalfailure, squeezing due to lateral pressure, ground settlement andheave, which makes it difficult to obtain comparable results. Theusing of kh for the soils in the PVD drained zone after the failure ofPVDs may also contribute to this difference.

5. Failure mechanism of the wall

The stability of the retaining structures can be analyzed usingeither limit equilibrium (LEM) or strength reduction method(Cheng et al., 2007). Since the soil strength increment, whichshould be considered in LEM during consolidation, is a complexproblem and is not the purpose of this study, the factor of safety(FS) of the structure was analyzed using the SRM built in PLAXIS inwhich the soil strength increment was considered in the hardeningsoil model (Brinkgreve et al., 2004):

Fig. 11. Numerical model of the RSW.

Table 3Material properties of the backfill and subsoil.

Material Backfill Sand cushion Layer 1 Layer 2 Layer 3 Layer 4 Layer 5

Material Model MC MC HS HS HS HS HSMaterial Type Drained Drained UU UU UU UU UUgunsat (kN/m3) 19.0 17 13.6 12.8 12.1 15.4 15.1gsat (kN/m3) 19.0 20 18.4 17.9 17.5 19.5 19.3khp (m/d) e 4.32 3.18E-05 3.26E-05 2.76E-05 1.38E-04 1.38E-04kvp (m/d) e 4.32 9.85E-05 9.33E-05 1.15E-04 1.04E-05 1.04E-05Eref (MPa) 5 20 e e e e e

Eref50 (MPa) e e 3.18 2.12 2.47 8.37 5.7Erefoed(MPa) e e 3.18 2.12 2.47 8.37 5.7Erefur (MPa) e e 9.54 6.36 7.41 25.1 17.1m e e 1 1 1 1 1n 0.33 0.2 0.25 0.25 0.25 0.2 0.2c0 (kPa) 16 0 7 7 6 17 18f0 (�) 30 30 27.9 28 24.9 29.8 31.2

Note: UU: unconsolidated undrained, MC: Mohr-Coulomb model, HS: Hardening soil model, gunsat ¼ unit weight above water table; gsat ¼ unit weight below water table;m ¼ power for stress-level dependency of stiffness; n ¼ Poisson's ratio; Erefoed ¼ tangent stiffness for primary oedometer loading and Erefur ¼ Young's modulus for unloading andreloading.

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FS ¼ ccr

¼ tan f

tan fr¼

XMsf (6)

where c, cr and f, fr are the input and reduced cohesion and frictionangle respectively;

PMsf is the total multiplier, which increases

until failure occurs in the numerical modeling.The stability of the wall at the end of the last loading stage was

analyzed using SRM. The computed FS value is 1.097 with the PVDson, and 1.01 with the PVDs off, which is at the limit state. It is

Fig. 12. Variation of excess pore water pressure at the monitored points.

worthwhile to note that using kh for the soil permeability afterdeactivating the PVDs is conservative without considering thedecrease of soil permeability due to consolidation. The predictedcritical slip surface is shown in Fig. 15. As shown in the figure, therewas no failure within the reinforced wall, and the slip surface fol-lows the general failure mode.

Fig. 13. Comparison of settlement at C1 obtained from numerical simulation with andwithout considering the failure of PVDs.

Fig. 14. Comparison of toe movement obtained from numerical simulation with andwithout considering the failure of PVDs, a): vertical displacement, b): horizontaldisplacement.

Fig. 16. Excess pore water profiles under the embankment at the end of last loadingstage with, a): PVDs on, b): PVDs off after day 118.

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The excess pore water pressure profiles after the last loadingstage under the two scenarios are plotted in Fig. 16. It shows that noexcess pore water pressure was generated in the PVD drained zonewith the PVDs working, but large excess porewater pressure can be

Fig. 15. Factor of safety and the failure slip surface predicted with strength reductionmethod with PVDs off after day 118.

created (up to 70 kPa) if the PVDs fail. In both scenarios, pore waterpressure in front of the toe area of the wall increased significantlyafter the construction. The figures show that in front of the wall, inthe second layer, where the slip surface passes through, the excesspore water pressure can increase up to the range of 20e40 kPa,even with the PVDs functional, and the value can be higher if thePVDs fail. The increase of pore water pressure in front the toecauses the dramatic decrease of effective stress in the zone,therefore the reduction of shear strength and the passive resis-tance. This may explain the failure of the slope. To avoid the buildupof excess pore water pressure in front of the toe area, additionalPVDs can be installed in this area for drainage purpose if spacepermits and the installation of PVDs does not cause instability ofthe soils.

6. Summary and conclusions

A fully monitored RSW built on very thick Shanghai soft claydrained with PVDs was studied. Unexpected excess pore waterpressure in the PVD drained zonewas observed in themiddle of theconstruction during resting period, when about 75% height of the7.6 m wall had been finished. Bulging and rotation of the wall faceoccurred in the last few loading stages, which is a common phe-nomenon in this type of wall. Large deformation in the wall andaccelerating horizontal movement at the toe of the wall wasobserved after the last loading stage, with relatively steady verticalsettlement, followed by an arc shaped crack on the embankmentsurface, which suggests a general shear failure mechanism.

A two-dimensional coupled mechanical and hydraulic finiteelement (FE) model was used to study the behavior and the failuremechanism of the structure. It was confirmed that a general slidingfailure occurred in the wall, with no shear failure in the geogridreinforced zone. The numerical simulation shows that excess porewater pressure can be dissipated quickly via the PVDs when thedrains are fully functional. The observed increase of excess pore

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water pressurewas captured bymanually deactivating the drains tosimulate its failure during the construction stages in the FE model.The FE models give relatively good estimation for the settlement ofthe wall. The horizontal movement of the toe compares well withthe FE results before day 118, when the PVDs are working. After day118, the FE model underestimates the lateral movement of toe withPVDs on, and overestimates the toe lateral movement with PVDsoff.

The stability of the RSWwas examined by the SRM incorporatedin the FE program. It was found that, the last loading stage is criticalfor the stability of the wall. With PVDs on, the FOS of the slope is1.097. With deactivating the PVDs halfway through the construc-tion, the embankment is at its limit state after the last loading stagewith FOS of 1.01. Based on the case studied, it has been found that:

1. the measured settlement at the instrumented location in theground showed that the settlement in the foundation soilsmainly occurred in the PVD drained zone, where the effectivestress increased the most;

2. the FE simulation showed that high excess pore water pressurecan be generated in the toe area due to lateral earth pressure,evenwith the PVDs functional when excess pore water pressurecan be quickly dissipated via the drains in the drained zone.From this point of view, if space allows, it is suggested to installone or two rows of PVDs in front of the toe, without causing toomuch ground disturbance and instability to the wall;

3. the observed high excess pore water pressure below the wall,where the PVDs were installed, was most likely due to thefailure of the drains. In the performed FE analyses, the failure ofthe drains could be well represented by manually deactivatingthe drains during construction. The SRM analysis showed thatthe factor of safety of the wall decreased almost to 1 by deac-tivating the drains. The simulation suggested that the failure ofthe wall was due to the increase of pore water pressure in thePVD zone, and soils in front of the wall, which reduced thepassive resistance against sliding;

4. in the design of embankments on PVD drained soft soils, it isimportant to check the stability of the wall considering possiblefailure of PVDs, especially when large deformation is anticipatedin the ground. The failure of the PVDs can be simulated bydeactivating the drains at certain stages, which is a conservativeyet passive choice.Where possible, it is recommended to use themeasured discharge rate to obtain an equivalent permeability ofthe soils for numerical modeling or to carry out a full 3Dsimulation.

Acknowledgments

The support from the National Natural Science Foundation ofChina under grant No. 41072200 is gratefully acknowledged. Thefirst author was a visiting scholar at Tongji University when thepaper was written, funding received from the Key Laboratory ofGeotechnical and Underground Engineeringwasmuch appreciated.

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