Mechanically Stabilized Earth Wall Failure at Two Soft and Sensitive Soil Sites

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Mechanically Stabilized Earth Wall Failure at Two Soft andSensitive Soil Sites

Debasis Roy1 and Raghvendra Singh2

Abstract: Two highway bridge approaches, about 10 and 12 m in height, near Kolkata �Calcutta�, India constructed with mechanicallystabilized earth �MSE� failed recently. These structures were founded on sensitive, soft and compressible, fine-grained soils of theintertidal flats and backswamps of the Ganges delta. One of these MSE walls, which failed in the final stages of its construction, wasconstructed after foundation soils were strengthened with prefabricated vertical drain installation and preloading. The second MSE wallthat failed within a month of its opening for traffic was constructed on unimproved ground. Fortunately, immediate collateral damage fromthese incidents was small. Using pre and postconsolidation shear strengths the MSE walls were redesigned. Reconstruction involvedprefabricated vertical drain installation at the second site and construction of stabilizing berms at both locations. The facilities are nowoperational and appear to be performing satisfactorily. Details of the failures, postfailure investigations, and monitoring, redesign, andreconstruction are presented in this paper.

DOI: 10.1061/�ASCE�0887-3828�2008�22:6�373�

CE Database subject headings: Geotechnical engineering; Embankment stability; Soil stabilization; Soft soils; Strain softening;Failures.

Introduction

Two highway embankments near Kolkata �Calcutta�, India failedrecently �Fig. 1�. The sites are located within the intertidal flatsand backswamps of the Hooghly River, a major distributary of theGanges �Fig. 2�, underlain mainly by fine-grained silts and claysof Holocene and Pleistocene age �Vaidyanadhan and Ghosh1993�. One of these sites, k18, remains waterlogged throughoutthe year, and the other, k26, also remains waterlogged over pro-longed periods. Both the embankments were constructed as me-chanically stabilized earth �MSE� walls composed of compactedriver sand. The sand backfill was reinforced with 8 m long gal-vanized steel strips, each 40 mm in width and 5 mm in thickness,placed horizontally at 750 mm centers measured vertically andhorizontally. The reinforcing strips were bolted to 180 mm thickinterlocking reinforced concrete facing panels with four stripsconnected to a single facing panel. The reinforcing strips had anultimate tensile strength of 490 MPa according to BS 8006:1995�British Standards Institution 1995� and the facing panels wereconstructed with M35 grade of concrete according to IS 456:2000�Bureau of Indian Standards 2000�. The facing panels werebacked by a filter cloth and a 600 mm thick zone of drain rock.For the backfill material, a friction angle of 35°, a cohesion inter-

1Assistant Professor, Dept. of Civil Engineering, Indian Institute ofTechnology, Kharagpur, West Bengal 721302, India �correspondingauthor�. E-mail: [email protected]

2Ph.D. Candidate �QIP Program�, Dept. of Civil Engineering, IndianInstitute of Technology, Kharagpur, West Bengal 721302, India. E-mail:[email protected]

Note. Discussion open until May 1, 2009. Separate discussions mustbe submitted for individual papers. The manuscript for this paper wassubmitted for review and possible publication on August 6, 2007; ap-proved on April 21, 2008. This paper is part of the Journal of Perfor-mance of Constructed Facilities, Vol. 22, No. 6, December 1, 2008.
©ASCE, ISSN 0887-3828/2008/6-373–380/$25.00.
JOURNAL OF PERFORMANCE OF CONST

J. Perform. Constr. Facil.

cept of 2 kPa, and a total unit weight of 16.5 kN /m3 weredeemed necessary for internal stability of the MSE wall. For thecompacted river sand backfill used in the project, these require-ments were met or exceeded.

Factual details on the incidents are presented first followed bythe outlines of postfailure investigation and monitoring programs,assessments of the writers on the possible causes of failure, andthe details of the redesign work and subsequent reconstructionwork.

Incidents

k26 Failure

The k26 event affected a highway interchange retained by a MSEwall along its outer shoulder and bounded by a 2 �horizontal� to 1�vertical� side slope along the inner shoulder �Fig. 3�. The con-struction of the highway interchange started in July 15, 2003. Theearth work was nearly complete at the time of the deep-seatedfailure that occurred in the early hours of January 12, 2005, whichdamaged the MSE wall. The incident followed 2 days of unsea-sonally heavy rains with 24 h precipitations of 18.4 mm on Janu-ary 10 and 0.4 mm on January 11. The height of the affected MSEwall was between 8.9 and 9.8 m at the time of the incident.Postincident inspection indicated that the MSE wall failed due toexternal instability without significant internal distress.

k18 Failure

The second incident involved a MSE wall that carried the north-bound lanes of a highway approach to a railway overpass �Fig. 4�.A 30 year old, 9 m high earth embankment with 3 �horizontal� to1 �vertical� side slopes along the eastern margins of the MSE wallcarried the southbound traffic. The incident occurred at around
11:00 p.m., February 9, 2006, about a month after the northbound
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lanes became operational. During the incident, a section of thenewly constructed MSE wall settled vertically by about 3 m andtranslated outward horizontally by about 4 m. Eyewitness ac-counts indicated that there were five or six heavy trucks on theaffected stretch of the highway at the time of the incident, ofwhich two overturned during the incident. The deformationslargely developed within the initial 15–20 min of triggering offailure; a rate slow enough to allow the crew of the overturnedtrucks to escape almost unhurt and the other trucks to be drivenoff the affected stretch. Postfailure monitoring indicated that thedeformations continued to increase for 9 days following the ini-tial incident �Fig. 4�a��. As at k26, the MSE wall at k18 appearedto have failed due to external instability without significant inter-nal distress.

The failed MSE wall was constructed for four laning of apreexisting two-lane, undivided highway that ran on the 30 year

Fig. 1. Failures at k26 and k18

Fig. 2. G

374 / JOURNAL OF PERFORMANCE OF CONSTRUCTED FACILITIES © AS


old embankment with 3 �horizontal� to 1 �vertical� side slopes.The construction of the old embankment also triggered severalslope failures. One such incident occurred in 1966, when the em-bankment reached its full height of 10.7 m. The instability wasassessed to be due to a deep-seated circular slip that day exitedthe ground surface just beyond the toe. The minimum height ofembankment affected during this failure was 6.7 m. As remedialmeasures, the failed embankment was removed, the highway el-evation was lowered, and the embankment was reconstructedalong with a 2.1 m high stabilizing berm along the edges of theembankment. Another slope failure occurred later immediately tothe south of the railway tracks. As poor subsurface conditions didnot allow construction of an embankment of required height at thelocation of this failure, the earth embankment there was replacedby extending the railway overpass southward by adding structuralspans supported on pile-supported piers. Details on the originalgeotechnical investigation and design of the highway embank-ments in this area can be found in Gangopadhyay et al. �1969�.

Incidentally, the north approach to the railway overpass at k18,constructed in the recent four laning project has remained stablesince becoming operational. This structure was of similar detail tothose of the failed MSE wall. The subsurface conditions at thenorth approach were also similar to those at the location of k18MSE wall failure except that unlike the failed MSE wall, thenorth approach was constructed on ground improved by prefabri-cated vertical drain �PVD� installation and preloading.

Subsurface Investigation and Monitoring

k26 Site

Before failure six boreholes, BH1–BH6, were drilled for founda-tion design for the overpass bridge structure �Fig. 3�b��. The otherboreholes were drilled during postfailure investigation. These in-vestigations included standard penetration testing �SPT�, extrac-tion of thin-tube samples, field vane shear testing �VST�, and

c setting
eologi
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laboratory tests such as unconsolidated undrained �UU� triaxialand one-dimensional incrementally loaded consolidation testingof selected thin-tube samples, and testing for grain size distribu-tion, natural moisture content, liquid limit, and plastic limit.

Subsurface investigations at k26 indicate that the site is under-lain by 15 m thick, grey, soft to firm, silty clay of Holocene ageover stiff, yellow-brown, silty clay of Pleistocene age containing

Fig. 3. Cross sect

Fig. 4. Cross sect



calcareous nodules and silt and sand interbeds. The top 10 m ofthe Holocene unit contained organics. Groundwater was within1.0–1.5 m of the original ground surface at the time of postfailureinvestigation. Both Holocene and Pleistocene deposits were clas-sified as CL according to ASTM D2487 �ASTM 2006�. The und-rained shear strength, su, of k26 soils from UU tests and VSTs are

d site layout: k26

d site layout: k18

ion an

ion an


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plotted in Fig. 5�a� against the effective vertical stress, �v�. All rawundrained shear strengths from VST measurements were cor-rected in this study following Bjerrum �1974�. The sensitivity ofthe native soil units was between 4 and 7.

The MSE wall at k26 was constructed after installing PVDs at1.5 m horizontal spacing to depths between 11.5 and 15.5 m in asquare grid pattern. The PVD-treated zone covered the entireMSE wall footprint and extended 4 m outside of the MSE wallface. To monitor the settlements and pore-water pressure devel-opment during construction, four settlement plates and four stand-pipe piezometers were installed near Station 800 �Fig. 3�a��. Themonitoring data from these installations indicate that the rapidconstruction rate after October 31, 2004 led to a rapid pore-waterpressure development and an accelerated settlement rate �Fig. 6�.

k18 Site

Prefailure geotechnical data basically originated from two bore-holes, BH8 and BH5 �Fig. 4�b��. The other boreholes were drilledafter the failure. The field work in these investigations includedconducting SPTs and VSTs, and extraction of thin-tube samples.The laboratory tests included UU triaxial and one-dimensionalincrementally loaded consolidation tests of selected thin-tubesamples, and tests for grain size distribution, natural moisturecontent, liquid limit, and plastic limit.

Data from these investigations indicate that the site is under-lain by a sequence of 10– to 12–m thick, Holocene silty clay, oversilty clay of Pleistocene age. The upper 5–8 m of the Holoceneunit was firm and overconsolidated with preconsolidation pres-sure, �c�, of up to 200 kPa, underlain by a 3 to 5 m thick soft,compressible, normally to lightly overconsolidated layer contain-ing organics and peat inclusions. The deepest part of the Ho-locene unit was firm and normally consolidated. The Pleistoceneunit contained sand or sandy silt partings and was stiff and over-

Fig. 5. Undrained shear strengths

consolidated with �c� of up to 300 kPa. Groundwater was at origi-



nal ground surface at the time of postfailure investigation. Thesoil samples were classified as CL according to ASTM D2487�ASTM 2006�. The values of su for k18 soils from UU tests andVSTs are plotted in Fig. 5�b�. The sensitivity of the soil layerswas between 4 and 7.

The site was not instrumented before MSE wall failure. Avail-able construction records indicate that the construction beganwith the placement of a 1 m high earth embankment in mid-January 2004 about 1 m to the west of the MSE wall alignmentfor dewatering the waterlogged site. Dewatering began in lateJanuary 2004, after which the site was stripped to a depth of about500 mm. A 500 mm thick compacted sand pad was placed on thestripped surface. Where the base of the MSE wall was to be at anelevation lower than the original ground level, excavation wascompleted for accommodating a 500 mm thick compacted sandpad underneath the base of the MSE wall. Sand filling com-menced in early February 2004 and MSE wall construction beganby the end of February 2004. The MSE wall was about 4 m highby mid-June 2004 and 5.253 m by early February 2005. Therewas virtually no earthwork between mid-June 2004 and mid-January 2005 and again between early February 2005 and end ofNovember 2005. The earthwork was completed in November andDecember of 2005. The paving work was completed by mid-January 2006 and the stretch of highway was opened for vehicu-lar traffic by mid-January 2006.

External Stability Reassessment of PrefailureConfiguration

The failure patterns at k26 and k18 sites were found to be ofdeep-seated nature caused by external instability of foundationsoils. Internally the MSE walls appeared to have been stable. Theembankments were to conform to IRC: 75—1979 �Indian RoadsCongress 1979�, which calls for a limit equilibrium factor ofsafety against external failure of 1.25 and allows a smaller factorof safety where the factor of safety is expected to increase with

Fig. 6. k26 monitoring data

time because of progressive strengthening of soils or where such


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smaller factors of safety represent shallow, inconsequential slips.IRC: 75—1979 also allows for a settlement of up to 600 mm.Brief accounts on limit equilibrium stability reassessments of k26and k18 embankments are as follows.

k26 Embankment

As indicated earlier, construction at k26 site began after theinstallation of PVDs �Colbonddrain CX1000 manufactured byColbond bv, Arnhem, The Netherlands, �www.colbond-geosynthetics.com�� to an average depth of 13 m in square grid at1.5 m centers. A mandrel of diamond-shaped cross section withdiagonals measuring 50 and 120 mm was used in PVD installa-tion.

The undrained shear strengths at the time of failure for exter-nal stability reassessment were estimated considering the state ofconsolidation at the time of failure in accordance with the solidline in Fig. 5�a�. For this, the average degree of consolidation wasobtained using cvh=0.02 m2 /day �cvh=coefficient of consolida-tion for flow in the horizontal direction� and kh=1�10−10 m /s�kh=horizontal hydraulic conductivity�. These inputs were basedon the results from one-dimensional consolidation tests and theassumptions that the ratio of vertical to horizontal coefficients ofconsolidation of 2 and smear zone diameter to be 2.5 times theequivalent mandrel diameter applicable for massive deposits�Hansbo 2004�. The results indicate that the average degree ofconsolidation for the foundation soils at the time of failure for thevertical pressure for embankment height of 3.5 m was 100% andthat for the stage above 3.5 m constructed relatively rapidly afterOctober 31, 2004 was 50%.

For the above-presented estimates of the input parameters,limit equilibrium stability analyses of the configuration of theMSE wall at failure indicate that the structure was marginallystable �Fig. 7�. In these analyses the Generalized Limit Equilib-rium method �Chugh 1986� and software package XSTABL Ver-sion 5.2 �Interactive Software Designs, Inc. 1994� were used.

In comparison, the minimum limit equilibrium factor of safetyagainst external failure in the original design under static loadswas 1.15. These computations were for an embankment height of9 m �against an actual of 9.8 m�. Further, a total unit weight of15 kN /m3 was used for the MSE wall backfill in the originaldesign instead of 18–19 kN /m3 representative of the compactedriver sand actually used to construct the MSE wall because of theinitial plan of constructing the MSE wall partly with fly ash. Mostimportant, the strength parameters used in the original design,shown in Fig. 5�a� for comparison, were based on the assumptionthat when the embankment construction reached a height of 9 m,the degree of consolidation for the native soil layers underneath

Fig. 7. Undrained stability of k26 MSE wall at failure

would be such that the minimum undrained shear strength would



be 50 kPa. Such a magnitude of undrained shear strength was notavailable at the time of failure. Thus, the bases of the originaldesign were, in general, unconservative.

k18 Embankment

Data from k18 site indicate that �1� the undrained shear strengthsat this location decrease westward as the distance increases fromthe MSE embankment and �2� the undrained shear strengths canbe expressed as functions of effective vertical stress, as shownwith the solid and a dashed lines of Fig. 5�b�.

External undrained stability of the configuration of the MSEwall was assessed using the simplified Bishop method �Bishop1955� and XSTABL Version 5.2. The input parameters used in theanalysis were based on the following assessments regarding thedegree of consolidation:• The native soil layers underneath the MSE wall were fully

consolidated under the vertical stress imposed by the old em-bankment;

• The degree of consolidation due to the MSE wall and thenewly constructed embankment was 50% over the top 2.5 mof native soils; and

• The deeper layers did not consolidate under the stress imposedby the MSE wall.

These assessments were based on one-dimensional incrementallyloaded consolidation test results obtained in the laboratory andborehole permeability measurements.

The strength parameters for the old embankment were backfigured from stability analysis to match the observed deep-seatedrotational instability during the construction of the embankmentin the 1960s described earlier. In this computation, the strengthparameters for native foundation soils were assumed in accor-dance with the dashed line of Fig. 5�b�. The results of externalstability assessment for the MSE wall at failure indicate that theMSE wall was indeed marginally stable at the time of failure �Fig.8�. The results of these stability analyses indicate that the loss ofshear strength due to remolding of sensitive foundation soils be-cause of continuous development of deformation behind and un-derneath the marginally stable MSE wall was the cause of failure.

In the original design the computed minimum factor of safetyagainst external failure involving a circular slip under static loadswas 1.42—a result that considerably exceeds the factor of safetyobtained by the writers. The difference appears to be mainly dueto the follwing facts. In the original design the embankmentheight including pavement structure was assumed to be 8.4 m�against the actual of 10.5 m� and for the embankment body the

3

Fig. 8. Undrained stability of k18 MSE wall at failure

total unit weight was assumed to be 15 kN /m �instead of


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18–19 kN /m3 considered representative of the compacted riversand actually used to construct the embankment� because of theinitial plan of constructing the embankment partly with fly ash.Thus, the bases of the original design were again unconservative.

Lessons Learned

Undrained Shear Strength

The main factor contributing to the instability at k26 appears to bein the estimation of postconsolidation undrained shear strength ofnative foundation soils. The estimates of undrained shear strengthratio, su /�v�, from k26 and k18 sites are plotted against the valuesof the overconsolidation ratio, OCR, obtained from incrementallyloaded one-dimensional consolidation tests conducted in the labo-ratory on thin-tube soil samples in Fig. 9. Also shown in Fig. 9 isthe range of correlations presented by Ladd et al. �1977� for fivesoft soil sites illustrating the SHANSEP framework. In spite ofthe scatter in the data, an approximate agreement between thedata from k26 and k18 and the correlations developed by Ladd etal. �1977� is apparent from the plot. The data also indicate that theundrained shear strength profiles of the floodplain deposits foundat k26 and k18 sites are approximated by su /�v�=0.25�OCR�0.76

�r2=0.80�.Considering that the SHANSEP approach is often employed

by designers in soil conditions similar to those at k26 and k18sites, it could be suggested that the overestimation of designstrength was partly because of inadequate geotechnical investiga-tion at the predesign stage that did not allow appropriate estima-tion of the degree of consolidation at various stages ofconstruction and consolidation-related strength increase for thenative foundation soils.

Trigger for External Instability

As is apparent from the results presented earlier, because of thefacts that original stability calculations were based on input pa-rameters not representative of the MSE wall configurations at thetime of failure or the backfill material used in construction, theembankments were assessed originally as stable against externalfailure although they were only marginally stable against externalfailure at the time of failure. Further, the undrained shear strength

Fig. 9. Undrained shear strength ratios as functions of OCR

for the foundation soils used as input in the original stability



assessments of the k26 MSE wall was based on an assumed de-gree of consolidation not achieved during construction.

It is evident from available instrumentation records that theembankment at k26 was undergoing rapid settlements for about2.5 months preceding failure. The construction rate was not con-trolled to allow settlement rates to decelerate primarily in order tomeet the completion deadline. As deformations increased, theshear strains within the foundation soils surpassed those at whichthe peak undrained shear strengths are mobilized leading first tostrain softening and eventually to failure.

The eyewitness accounts describing the failure presented ear-lier and post facto monitoring results shown in Fig. 3�b� suggestthat the failure at k18 was also progressive in nature. However, nodirect evidence of accelerated settlement rates over the periodbetween end of construction and failure is available because ofthe absence of monitoring data covering the construction andearly operational phases of the MSE wall. It is possible that be-cause progressive failure typically starts near the toe of an em-bankment �i.e., the toe of the berm that ran outside the MSE wallface� and there was a thick, flexible, reinforced earth structureunderneath the pavement at the time of failure, development ofdeformation within the foundation soils during the short periodover which the highway remained operational before failure didnot become noticeable. Thus, the k18 failure also appears to havebeen triggered by remolding of sensitive foundation soils becauseof development of deformations due to an inadequate factor ofsafety against external instability.

Project Management

Geotechnical design of the MSE walls was in the scope of thesystem supplier �internal stability� and the contractor �externalstability�. The designs were, in turn, approved by the engineer andthe owner. The construction scheduling and monitoring activitieswere also dealt with by the contractor duly, reviewed, and ap-proved by the engineer and the owner. The writers were involvedonly in the postfailure stages, during which they did not notice thepresence of a mechanism for an independent technical review andaudit of design, monitoring, and construction scheduling prior tothe failures. Of particular relevance to the failures was the over-sight in getting the geotechnical designs reviewed upon finaliza-tion of the MSE wall height, backfill material selection, andconstruction schedule.

Reconstruction

Design of Remedial Measures at k26 Site

Reconstruction work at the k26 failure site included constructionof a two-stepped, 5.2 m high and 13.4 m wide �measured outwardfrom the face of the reconstructed MSE wall to the toe of thelower bench� stabilizing berm along the outer face of the MSEwall, removal of the damaged MSE wall, and reconstruction ofthe MSE wall and the highway embankment. Schematic details ofthe final configuration of the MSE wall proposed for reconstruc-tion are presented in Fig. 10. The minimum factor of safety understatic loads against external instability for this configuration wasestimated to be 1.25 �based on modified Bishop approach andXSTABL Version 5.2� implying compliance with the IRC:75–1979 recommendations. As the berm was to be constructed beforeMSE wall reconstruction, the shear strengths for the stability as-
sessment were based on the assumption that the foundation soils

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within the PVD-treated zone that extends to a distance of 4 m
outward from the MSE wall face would consolidate under theweight of the berm by the time the MSE wall reconstruction wascomplete. With further consolidation of the native foundationsoils underneath the berm outside the PVD-treated zone, the fac-tor of safety is expected to reach 1.35 �modified Bishop, XSTABLVersion 5.2� after about 5 years of berm construction.

k26 Reconstruction

The berm construction was complete by mid-April 2005. TheMSE wall and highway embankment construction was completeby mid-June 2005. The pavement construction was complete byJuly 2005 and the reconstructed highway interchange was re-opened for vehicular traffic in November 2005. Since then thereconstructed MSE wall appears to be performing satisfactorily.

Design of Remedial Measures at k18 Site

The remedial measures at the k18 site involved installation ofPVDs to 13 m depth along a 17–19 m wide strip along the outerface of the failed MSE wall and construction of a two-steppedstabilizing berm over the PVD-treated area, removal of the dam-aged MSE wall, and reconstruction of the MSE wall and thehighway embankment. The stabilizing berm was to function as acounterweight against the driving forces as well as act as a pre-load over the PVD-treated foundation soils to ensure their con-solidation and strengthening. Schematic details of the finalconfiguration of the MSE wall proposed for reconstruction arepresented in Fig. 11. For this configuration the minimum factor ofsafety under static loads against external failure was estimated tobe 1.22 �modified Bishop, XSTABL Version 5.2�. This assessmentis based on the assumption that the consolidation of soils withinthe PVD-treated zone would be 75% complete at the end of re-

Fig. 10. Undrained stability of k26 reconstruction scheme

Fig. 11. Undrained stability of k18 reconstruction scheme



construction and the effect of consolidation outside the PVD-treated zone would be negligible. Although the computed factorof safety at the end of construction is marginally smaller than thatrecommended in IRC:75—1979, the factor of safety is expectedto reach 1.38 �modified Bishop, XSTABL Version 5.2� withinabout 5 years from completion of reconstruction.

k18 Reconstruction

PVDs �Colbonddrain CX1000� were installed in October 2006 at1.2 m horizontal spacing in a square grid pattern using a mandrelof diamond-shaped cross section with diagonals measuring 50and 120 mm. The width of the PVD-treated zone was 19 m to-ward measured west from the face of the failed MSE wall northof station 18.300. The width was 17 m south of station 18.300.The stabilizing berm was constructed between October 2006 andFebruary 2007. The height of the upper bench was 3 m lowerthan the finished road grade and its top width was 20 m. Themaximum overall berm width measured from the base of the re-constructed MSE wall to the toe of the lower bench was 37 m.Three standpipe piezometers and three settlement gauges wereinstalled through the stabilizing berm �Fig. 4�b��. In February2007 the failed embankment was removed to an elevation ofabout 5.75 m above the original ground surface and reconstruc-tion work for the MSE wall was taken up. Three standpipe pi-ezometers and three settlement gauges were installed within thefootprint of the MSE wall after removal of the failed embank-ment. These installations were maintained in serviceable condi-tions throughout the construction and early operational phases.Data from these instruments indicate that the settlements werecontinuing to develop and pore-water pressures were still dissi-pating after about 4.5 months PVD installation and berm con-struction. The instrument monitoring data covering thereconstruction period and the early operational phase are pre-sented in Fig. 12. The monitoring instruments at this site were inserviceable conditions at the time of preparing this paper, al-though their monitoring was discontinued after 2 months from theend of construction.

Drilling, sampling, and standard penetration testing were car-

Fig. 12. k18 monitoring data

ried out at BH1 and BH2 and VSTs were conducted at V4, V4A,


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and V5 at the end of April 2007 to check whether the gain inundrained shear strength due to accelerated consolidation of softsoils within the PVD-treated zone was as assumed in MSE wallredesign. Data from these tests, presented in Fig. 13, togetherwith the undrained shear strengths from nearby locations mea-sured immediately after failure, indicate that the increased un-drained shear strengths following accelerated consolidation of thesoils within the PVD-treated zone for the most part exceededthose assumed in the overall stability assessment for embankmentredesign shown in Fig. 11. Embankment reconstruction above8 m height was allowed from May 2007 after these data werereviewed. Subsequently the fill placement and paving work werecompleted by the beginning of June 2007. Since its opening forvehicular traffic in June 2007 the reconstructed MSE wall appearsto be performing satisfactorily.

Conclusions

Case histories pertaining to the failures of two MSE walls con-structed on the floodplain and backswamp deposits of the GangesDelta have been presented in this paper. The sites are underlainprimarily by soft and sensitive silty clay and clayey silt. Thefailed MSE wall at k26 was constructed on ground improved byPVD installation and preloading, whereas that at k18 was con-structed over unimproved ground. Both these failures appeared tobe due to external instability and progressive in nature.

The failure at k26 appeared to have been caused by:• A rapid construction rate that did not allow the foundation

soils to undergo consolidation needed for development of und-rained shear strengths assumed in the initial design; and

• Underestimation of driving force because of the use of anembankment height smaller than that of the actual configura-tion and an unrealistically low unit weight for the fill materialin the initial design.The failure at k18 site, on the other hand, appeared to be due

to:• The use of an embankment height smaller than that of the

actual configuration and an unrealistically low unit weight forthe fill material in the initial design; and

• The strength loss due to remolding of native soils found un-derneath the MSE walls because of inadequate factor of safety

Fig. 13. Undrained shear strength increase within PVD-treated zoneat k18 site



against external instability.Available information indicates that these incidents could beavoided if there was �1� a proper coordination between construc-tion and design offices and �2� a proper review and audit of geo-technical design, construction, and monitoring activities atdifferent stages of project execution.

The MSE walls were redesigned taking into account the actualembankment height and soil properties. The consolidation-relatedincreases in undrained shear strength for the foundation soils werealso considered employing the SHANSEP framework.

The remedial measure at k26 included construction of a stabi-lizing berm along the outer face of the MSE wall. After construct-ing the berm the damaged MSE wall was removed and theoverpass approach was reconstructed. At k18, on the other hand,PVDs were first installed along the outer margins of the failedMSE wall followed by construction of a stabilizing berm over thePVD-treated area. The stabilizing berm at k18 was to function asa counterweight against the driving forces as well as act as apreload over the PVD-treated foundation soils to ensure their con-solidation and strengthening. Reconstruction of the k18 MSE wallwas taken up only after adequate consolidation and strengtheningof soft soils within the PVD-treated zone could be demonstrated.From start to finish the reconstruction works at k26 and k18 sitesspanned about 6 months. Both these facilities are now operationaland appear to be performing satisfactorily.

References

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2008.22:373-380.

Documents

Mechanically Stabilized Earth Wall Failure at Two Soft and Sensitive Soil Sites