study of geotechnical failures through physical modeling

8/3/2019 Study of Geotechnical Failures Through Physical Modeling

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1 INTRODUCTIONGeotechnical failures often occur when sub-standard

construction procedures have been adopted. Highvariability and insufficient or inaccurate soilinvestigation results regarding subsurface soil layersand properties can also lead to erroneous designs.Many geotechnical failures have been reported. Forinstance, Ting et al. (1994) reported the failure ofpiles supporting an embankment due to landslip,Poulos (1994) reported the piled foundation failureof an office tower due to a nearby excavation, andthe failure of a pile-supported wharf structure due toriverbank movement has been described by Ting andTan (1997).

When failures occur, lives, money and confidenceare lost. Painful lessons can be learnt from eachfailure so that such failures can be avoided in thefuture. Nevertheless, it is also important tounderstand the behaviour of the structures duringand after failure so as to widen the knowledgebeyond the serviceability limits of structures. Real-life structures are not normally built and then testedto failure as tremendous amount of money and timeis involved. Therefore, full-scale testing of astructure to failure is generally deemed to be

uneconomical and undesirable.As such, physical modeling is considered an

attractive alternative to study geotechnical failures.Leung et al. (2000) reported the results of acentrifuge model study on the behaviour of piles

behind a collapsed retaining wall in sand. In thispaper, the responses of piles behind an unstableretaining wall in clay and adjacent to a tunnel that

subsequently collapses upon excavation are used asillustrative examples to demonstrate the benefits ofphysical modeling in the study of geotechnicalfailures. Centrifuge model tests can be carried outunder a controlled environment where the soilstrength profiles, soil deformation and elapsed timecan be measured with reasonable accuracy resultingin reliable test results. Besides that, the consistentrepeatability of centrifuge experiments also renderscentrifuge model study attractive and relativelyeconomical.

2 EXPERIMENT SET-UP2.1 Model pile and retaining wall

The first study involves the investigation of pileresponses behind a retaining wall that collapsesupon reaching certain excavation depth. Figure 1shows the centrifuge model setup. All theexperiments were conducted at 50g on the NationalUniversity of Singapore geotechnical centrifuge.

The model pile was fabricated from a hollowsquare aluminium tube with an outer diameter of9.53 mm and a wall thickness of 3.18 mm. Ten pairsof strain gauges were attached to the opposing facesof the model pile at vertical intervals of 25 mm. The

Study of Geotechnical Failures through Physical Modeling

C. F. Leung, D. E. L. Ong & Y. K. ChowCentre for Soft Ground Engineering, Department of Civil Engineering, National University of Singapore,Singapore 117576

ABSTRACT: The use and benefits of physical modeling in the study of geotechnical failures is demonstratedin this paper. Centrifuge model studies on the performance of piles behind an unstable retaining wall in clay as wellas the responses of piles due to tunnel collapse in sand are used as illustrative cases. An enhanced image processing

system is employed to visualize the soil movement patterns and trends during the failure of the wall or the tunnel. Theobservations together with the measured bending moment and deflection profiles along the pile and ground settlements

provide valuable information on the failure mechanism of the problems understudy without the occurrence of actual

failures in the field. Prediction of pile responses using a numerical method developed at the National University of

Singapore is also presented in this paper to evaluate the reliability of the centrifuge test observations and results.

Keywords: Soil flow, failure, tension cracks, soil movement, limiting soil pressure, image processing


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final width of the pile is 12.6 mm. At 50g, theequivalent pile width is 630 mm in prototype scale.The total length of the pile is 350 mm (17.5 m inprototype scale). The pile has an embedment depthof 250 mm (12.5 m) into the clay. A thin layer ofepoxy was applied to the entire pile length to ensurethat the strain gauges and all connections werewaterproof and well protected. The stiffness of the

epoxy coating is assumed to be negligible whencompared to that of the aluminium model pile. Theprototype bending rigidity, EI, of the model pile isapproximately 2.2 x 10

5kNm

2. This is somewhat

equivalent to a 12.7 mm thick, 610-mm diametersteel pipe pile in the field.

The model retaining wall is fabricated using a 3mm (150 mm in prototype dimension) thickaluminum plate. Its prototype bending rigidity, EI, is24 x 10

3kNm

2 /m, which is equivalent to that of a

FSP II A sheet pile.

Figure 1. Centrifuge model set-up (all dimensions in mm)(after Ong et al., 2003a)

2.2 Soil propertiesThe properties of the clay and sand used are

shown in Tables 1 and 2, respectively. Theundrained soil strength profiles measured by in-

flight bar penetrometer tests are given in Ong et al.(2003b).

Table 1. Physical properties of kaolin clay (after Ong et al.,2003a)

Specific gravity, Gs 2.6

Liquid limit, LL 80 %

Plastic limit, PL 40 %

Compression index, Cc 0.65

Swelling index, Cs 0.14

Void ratio at 115 kPa at NC line 1.67Permeability at 115 kPa at NC line 1.3x10-8 m/s

Table 2. Physical properties of sand (after Ong et al., 2003a)

Mean grain size 0.16 mm

Uniformity coefficient 1.3

Specific gravity, Gs 2.65

Friction angle (50-100 kPa) 43o

2.3 Test preparation and procedureKaolin powder was mixed with water at a watercontent of 120% in a de-airing mixer. The slurry wasthen placed in the model container under water.Subsequently, a 17-kg plate was placed on top of theslurry to stiffen it. Then, the sample was placed on aloading frame for 1-D consolidation under a load of20 kPa for 3 days. After that, self-weightconsolidation of clay was carried out under 50g forabout 6 to 7 hours. The ground surface settlementswere monitored by displacement transducers(LDVT). After about 90% consolidation had beenreached, the centrifuge was spun down. Porepressure transducers (PPT) were then embedded atpositions shown in Figure 1.

Subsequently, the model wall and pile were jacked vertically into place using a guide at 1g.Excavation was then carried out. The excavated claywas replaced by ZnCl2 solution contained in a latexbag. The density (16.5 kN/m3) and height of theZnCl2 solution were made to be identical to those ofthe clay that had been excavated.

After that, the front perspex face of the container

was removed so that soil movement markers couldbe placed on the clay at 20 mm square grids. LVDTswere installed to measure the ground settlementsbehind the excavation. The pile head deflection,during and after excavation was monitored by twonon-contact laser displacement transducers. A high-resolution image-processing camera was mounted infront of the perspex window of the model container.The camera is capable of recording a high-resolutionimage with a pixel-to-pixel spacing of less than 0.1mm.

Finally, the model package was put together andspun up to 50g for reconsolidation. When both porewater pressures and ground settlements behind thewall showed negligible changes, the ZnCl2 solutionwas then released to depict excavation at 50g. Inprototype scale, the simulated excavation rate wasabout 0.56 m per day, which is equivalent to normalreal-life average rate of excavation. The stresshistory of the clay sample was intended to simulatethat of a normally consolidated clay deposit with a2.8-m overconsolidated crust. The detail testprocedures are described in Ong et al. (2003a).

Two tests depicting the collapse of retaining wallwere conducted in the present study. Theconfigurations of the tests are shown in Table 3.


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Table 3. Configuration of Tests 1 and 2

Item Test 1 Test 2

Distance of pile from

wall

3 m 3 m

Excavation depth 1.8 m 2.8 m

Clay thickness 10.5 m 6.5 m

Sand thickness 2.0 m 6.0 m

Wall toe embedment insand

0 m; Floating inclay

1.5 m

3 TEST RESULTS AND DISCUSSIONSThe pile behaviour in Tests 1 and 2 are discussed

with particular reference to the measured soildeformation patterns and the pore water pressureresponses.

3.1 Pile responseThe bending moment profiles along the pile at

different excavation depths for Test 1 are shown inFigure 2. The maximum pile bending moment islocated at 8.75 m below the ground level. Amaximum value of 235.7 kNm is recorded at anexcavation depth of 1.2 m. As excavation continuesto its final depth of 1.8 m, the bending moment isnoted to reduce to 185.8 kNm.

Figure 2. Development of pile bending moment profile over

time (Test 1)

Figure 3 shows the development of bending

moment profiles along the pile for Test 2 in whichgreater excavation depth was involved. During theearly excavation stages, the measured pile bendingmoments show a similar trend as that in Test 1. Themaximum pile bending moment is about 186.7 kNm

also located at 8.75 m below the ground at anexcavation depth of 1.2 m. Subsequently, this valuedrops slightly with increasing excavation depth butsomehow increases again after the excavation depthexceeds 2.0 m, as shown in Figure 3. A maximumpile bending moment of 250.7 kNm is recorded atthe end of excavation of 2.8 m.

Figure 3. Development of pile bending moment profile over

time (Test 2)

3.2 Wall and soil deformation

Figure 4. Soil and wall deformation over time

0 50 100 150 200 250Bending moment (kNm)

12.5

10

7.5

5

2.5

0

Depth(m)

Excavation depth (m)

Symbol

1.0 1.2 1.4 1.6 1.8

0 100 200 300

Bending moment (kNm)

12.5

10

7.5

5

2.5

0

De

pth(m)


Symbol

0.4 0.8 1.2 1.6 2.0 2.4 2.8

0.1 1 10 100 1000

Time from start ofexcavation (days)

1.8

1.5

1.2

0.9

0.6

0.3

0

Ground

settlement(m)

0

0.5

1

1.5

2

2.5

3

Wallheadlateral

movement(m)

Wall headmovement

Test 1

Test 2

Settlement1.5 m

behind wall

Settlement4.5 m

behind wall


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Figure 5. Selected soil movement vectorial plots for Test 1

Figure 4 shows the wall head movements andground settlements over time for both tests. Asexpected, the wall and ground deformations in Test2 are larger than those in Test 1. Both the groundsettlement and wall deflection follow a similarpattern over time.

For both Tests 1 and 2, when excavation depthexceeds 0.6 m, the image processor shows that theclay surface starts to move past the pile head.When the excavation depth exceeds about 0.8 m, itis observed from the image processor that some

fissures develop at the ground surface in Tests 1 and2. Water can hence flow freely into the increasingnumber of fissures as excavation progresses. Thedevelopment of these fissures further reduces thestability of the wall, thus causing the wall to tilt andthe clay to be sheared progressively over time.

Therefore, the overall clay stiffness is reduced. Theobserved slip lines, which shows the extent of thedeformed zone around the pile, is similar to thecharacteristics meshes postulated by Randolph et al.(1984) for a plastically deforming cohesive soil.

-7

-5

-3

-1

Depth(m)

Soil movement(mm)

1000

5

15

25

35

45

55

65

75

50

100

150

200

250

300

-7

-5

-3

-1

Depth(m)

50

100

150

200

250

300

350

400

450

500

-7

-5

-3

-1

Depth(m)

1 3 5 7Distance (m)

-7

-5

-3

-1

Depth(m)

50

150

250

350

450

550

650

750

Excavation 0.6 m,

1.0 days

Excavation 1.0 m,1.7 days

Excavation 1.4 m,

2.4 days

Excavation 1.8 m,

3.0 days


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Large

tension

crack

Deformed

zone

Soil flowdirection

Deformed

zone

High-resolution pictures were taken duringvarious excavation stages of the tests, as shown inthe left-hand side photos of Figure 5. It is evidentthat tension cracks have developed when theexcavation depth exceeds 1.0 m. These cracks causethe loss of contact of clay in front of the pile. This inturn prevents the full transmission of soil pressureonto the pile. It is also probable that the soil would

flow past the pile and not to have exerted fullpressure on the pile. Figure 6 shows the top planview of the deformed soil around the pile afterexcavation. It is again evident that there is aconsiderable drop in soil-pile contact and this mighthelp to explain the reduction in pile bendingmoments after the development of tension cracks.Nevertheless, the pile head deflection remains fairlyconstant (see Figure 7) as the pile is continuouslybeing pushed by the separated soil mass behind thepile. However, such pile behaviour is only notedprior to the occurrence of a fully developed active

shear failure wedge as in Test 1.

Figure 6. Soil flow and tension cracks around the pile

Therefore, Test 2 can be used to assess the pilebehaviour during and after the occurrence of a fullydeveloped active shear failure since the excavationdepth is large enough for the shear failure wedge todevelop. When excavation exceeds 1.8 m, the photosshow that a shear band appears gradually. Asexcavation progresses further, an active shear failurewedge has initiated, starting from the location of thetension cracks and gradually moves down until itintersects the wall toe. As the wall is not braced, it isobserved to have rotated around the pivot, which is

located at the clay-sand interface. Theseobservations are consistent to the findings of Boltonand Powrie (1987) and Wei (1997). The soilmovement immediately after excavation and thedevelopment of the active shear failure wedge are

plotted in terms of soil displacement vectors andshown in Figure 8. It clearly shows that the suddenchanges in the length of vectors reflect thedevelopment and occurrence of the active shearfailure wedge.

Figure 7. Development of pile bending moment and

deflection

The development of the active shear failurewedge in front of the pile causes the pile bendingmoment to drop, similar to that observed in Test 1.However, the mass of soil behind the pile starts to

move forward in response to the increasingexcavation depth, thus causing the pile bendingmoment to increase again. This occurs when theexcavation depth exceeds 2.0 m as shown in Figure3.

3.3 Pore water pressureThe variations of excess pore water pressure with

time for Tests 1 and 2 are shown in Figures 9 and10, respectively. All PPTs register excess pore waterpressures immediately after excavation.Nevertheless, PPT 1 of Test 1 registers positiveexcess pore water pressures shortly after thecompletion of excavation. This is due to the waterlevel after excavation being higher than itshydrostatic level caused by ground surfacesettlements. Besides that, the flooded tension crackalso caused the free water and air to enter PPT 1.

Owing to relatively greater soil deformations nearthe wall, PPT 2 registers a greater dissipation ofexcess negative pore water pressures as opposed toPPT 3, which is located further away from the wall.

The difference in excess pore water pressures createsa hydraulic gradient that leads to pore water pressureredistribution and hence, the reconsolidation of theclay over time. This is somewhat analogous to thatobserved by Stewart (1992).

0 0.4 0.8 1.2 1.6 2


0

50

100

150

200

250

Bendingmoment(kNm

)

0

50

100

150

200

250

Deflection(mm)Bending moment

Deflection


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Figure 8 Development of active shear failure wedge during wall collapse (Test 2)

Soil movement(mm)

1800

Distance (m)

1 3 5 7-6

-4

-2

Depth(m)

-5 -3 -1-6

-4

Figure 9. Excess pore water pressure variation over time

(Test 1)

After excavation, the ground water level at the

excavated side will drop. This creates a waterpressure head difference between the retained andexcavated sides. PPT 4 in Test 2 shows a rebound ordissipation of excess negative pore water pressureover time. However, a rebound is not observed in

Test 1 because the wall in Test 1 is embedded

entirely in clay, whose permeability is much lowerand hence, the dissipation of excess negative porewater pressure is not as easy and obvious ascompared to Test 2, where the wall is embedded intothe sand layer so that seepage can occur easier.

Figure 10. Excess pore water pressure variation over time

(Test 2)

0 50 100 150 200 250 300 350

Time after start of excavation (days)

-25

-20

-15

-10

-5

0

5

Excessp

orewaterpressure(kPa)

PPT 1

PPT 2

PPT 3

PPT 4

0 50 100 150 200 250 300Time after start of excavation (days)

-40

-30

-20

-10

0

Excessporewaterpre

ssure(kPa)

PPT 1

PPT 2

PPT 4

PPT 3


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Figure 11. Measured and predicted bending moment and deflection of pile for Test 1 (after Ong et al. 200b)

Since the wall in Test 1 is floating in clay withoutembedding into the stiffer sand layer, thereconsolidation of soil caused by pore water pressureredistribution becomes more dominant than theseepage caused by the head difference afterexcavation, as revealed by PPT 4 readings in Test 2shown in Figure 10. The excess negative pore waterpressures in the retained side for Tests 1 and 2 arerelatively small as they may be partly cancelled outby the positive pore water pressures generated by theundrained shearing of the clay. Similar observationswere also reported by Kimura et al. (1994).

4 NUMERICAL ANALYSISThe numerical method developed by Chow and

Yong (1996) is used to back-analyze the responsesof a single pile due to excavation-induced soilmovement obtained from the centrifuge tests. Thisnumerical method has been used successfully byLeung et al. (2000) to back-analyze the pile

responses due to excavation-induced soil movementin sand.

The concept of analysis is based on finite elementmethod where the pile is represented by beamelements and the soil is idealized using the modulusof subgrade reaction. The non-linearity of the soilbehaviour can be incorporated to an extent bylimiting the soil pressure that could act on the pile.The numerical analysis requires the knowledge ofthe pile flexural rigidity (EI), the distribution oflateral soil stiffness (Kh) with depth, the limiting soilpressure (py) that acts on the pile and the lateral soilmovements. This approach is used in the presentstudy to predict the pile responses in Test 1.

4.1 Input soil propertiesThe distribution of lateral soil stiffness with depth

(Kh), Youngs modulus of the soil (Es) as well as thesoil resistance or limiting pressure (py) in kaolin clayare described in detail in Ong et al. (2003b).

12.5

10

7.5

5

2.5

0

Depth(m)

0 100 200 300 400 500


12.5

10

7.5

5

2.5

0

Depth(m)

-20 0 20 40 60 80 100 120

Deflection (mm)

Uncorrected

py = 9cu

Corrected

py = 3cu

Excavationdepth (m) Measured Predicted

0.6

0.8

1.0

1.2

1.4

(a) (b)

(c) (d)


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4.2 Prediction and discussionsWhen the excavation depth exceeds 1 m in Test

1, the predicted pile responses are grosslyoverestimated, as shown in Figures 11a and b. Thisis because when the clay starts to yield, the limitingsoil pressure may have been reached The yieldingand/or failure behaviour of the clay is evidentlydemonstrated by the bar penetrometer test results

(Ong et al., 2003b), where the undrained shearstrength (cu) of the clay reduces significantly afterthe soil has been excavated. By performing back-analysis of the centrifuge results, reasonablepredictions of the pile responses can be obtained byadopting py = 3cu (Ong et al., 2003b). Subsequently,the corrected predicted pile responses are shown inFigures 11c and d, which reveal considerably betteragreements with the measured pile responses.

5 TUNNEL-SOIL-PILE INTERACTIONThe second example involves the investigation ofresponses of piles close to a collapsed tunnel. Fenget al. (2002) presented results of centrifuge modelstudy to investigate pile responses prior, during andafter the collapse of a model tunnel. A polystyrenefoam core, which is placed inside a brass foil-mademodel tunnel lining, is dissolved by an organicsolvent during centrifuge flight to simulate theprocess of tunnel excavation. Strain gauges attachedon the model pile shaft are used to measure thebending moment and axial load profiles along thepile during tunnel excavation.

Figure 12. Centrifuge model set up for tunnel (after Feng et

al., 2002)

All the centrifuge tests were conducted at 100g.In one test, the brass foil is 0.05mm thick simulatinga relatively weak tunnel lining, and wrapped aroundthe tunnel-shaped polystyrene foam core. The liningwas soldered using tin solder and an electronicsoldering gun. The tunnel is 210mm long (21m atprototype scale) and 60mm (6m) in diameter. Figure

12 shows the centrifuge model setup and Table 4lists the test parameters in prototype scale.

Table 4. Test parameters (after Feng et al., 2002)

Tunnel diameter 6 m

Tunnel depth 16 m

Ground depth 15 m

Ground type Dry sand

Distance of pile from tunnel

vertical center-line

9 m

The weak model tunnel lining is not strongenough to support the soil and the tunnel completelycollapses during excavation. This test conditionrepresents the worst situation in practice.

The first pile, which is used to measure thebending moment profile, was positioned 90mm (9min prototype scale) from the tunnel verticalcenterline. The pile has free head and tip conditionsas well as a bending capacity of about 3000 kNm.Figure 13 shows the maximum induced bendingmoment profile of the pile due to the tunneling

process. The maximum bending moment is 1345kNm located approximately at the depth ofhorizontal tunnel center. As expected, the bendingmoment at pile head and pile tip is zero.

Figure 13. Bending moment along pile shaft (after Feng et

al., 2002)

Model container

LVDT

sandstraingauge

pile

tunnel


0

5

10

15

20

25

0 500 1000 1500

Depth(m)


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The second pile used to measure the axial loadprofile was positioned on the opposite side of thetunnel having the same distance from the tunnelvertical centerline as the first pile. However, this is afloating pile with its tip at 20 mm above thecontainer base, and its axial load capacity is about2500 kN. Figure 14 shows the induced axial loadincreased downwards from the pile head, and

reached a maximum value of 1232 kN atapproximately the depth of tunnel center, and thenreduced towards the pile tip. The soil movementswere tracked by the movement of the beads placedon the sand before the test. It is noted that the soilabove the tunnel had moved downwards, while thesoil below the tunnel had moved upwards during thetunneling process as shown in Figure 15.

Figure 14. Axial load of pile due to tunneling (after Feng et

al., 2002)

This helps to explain why the axial load transferof the pile below the tunnel elevation is positive.Compared with the soil movements above thetunnel, the soil movements below the tunnel weregenerally much smaller. The ground surfacesettlements due to tunneling were measured with anarray of LVDTs. The observed ground surfacesettlement trough is noted to resemble the classicalGaussian-shape settlement trough.

From this study, it can be observed that themaximum bending moment and axial load of the pileoccurred at approximately the depth of the tunnelhorizontal centre-line. However, the axial load

gradually reduced below the tunnel centre-line. Theinduced pile bending moment and axial load wereapproximately 55% and 50% of the bending andaxial capacities of the piles, respectively.

Figure 15. Soil movement due to tunneling

6 CONCLUSIONIn this paper, the benefits of physical modeling inthe study of geotechnical failures are demonstratedby means of examples using studies on pileresponses behind an unstable retaining wall andadjacent to a tunnel that collapses upon excavation.Excavation of soil causes lateral stress relief due tothe removal of overburden pressure. Therefore, the

strength and stiffness of the clay are reduced after anexcavation. Such reduction in soil strength may havecontributed to the soil flow phenomenon and thedevelopment of tension cracks, which subsequentlyaffect the behaviour of piles embedded in soilundergoing large deformation. The deformation ofclay can be correctly modeled by the numericalanalysis if appropriate limiting soil pressure valuesare used. This can be achieved by performing in-flight bar penetrometer tests to obtain the clayundrained shear strength profile before and afterexcavation.

Tunnel-soil-pile interaction has also been studied.Owing to the collapse of tunnel lining, it is observedthat about half of the ultimate bending and axial loadcapacities of the pile have been induced on anadjacent pile from the present centrifuge modelstudy. This would reduce the capability of theinstalled piles to sustain the serviceability loads ofthe structures. Ground subsidence due to thecollapse of tunnel lining also poses a great threat toexisting adjacent buildings, especially those foundedon pad footings, where differential settlement will

cause structural distress.The two studies demonstrate that centrifuge

modeling can be a versatile tool to studygeotechnical failures as reliable results can beproduced with reasonable accuracy. The

0

5

10

15

20

25

0 500 1000 1500

Axial load (kN)

D

epth(m)


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observations and results facilitate a much betterunderstanding of the behaviour of piles due tofailure of adjacent geotechnical structures.

ACKNOWLEDGEMENTS

The authors wish to acknowledge the contributions

of Mr S-H Feng on the study of tunnel-soil-pileinteraction study. The assistance of GeotechnicalCentrifuge Laboratory professional and technicalstaff is also gratefully appreciated.

REFERENCES

Bolton, M. D. and Powrie, W., The Collapse of DiaphragmWalls Retaining Clay, Geotechnique, Vol. 37, No. 3, pp.335-353, 1987.

Chow, Y. K. and Yong, K. Y., Analysis of Piles subject toLateral Soil Movements. Journal of The Institution ofEngineers, Singapore, Vol. 36, No. 2, pp. 43-49, 1996.

Feng, S. H., Leung, C. F., Chow, Y. K. and Dasari, R.,Centrifuge Modelling of Pile Responses due to Tunneling,15th KKCNN Symposium on Civil Engineering, Singapore,2002.

Kimura, T., Takemura, J., Hiro-oka, A., Okamura, M. and ParkJ., Excavation in Soft Clay Using an In-flight Excavator,Centrifuge 94, Leung, Lee and Tan (eds), pp. 649-654,1994.

Leung, C. F., Chow, Y. K. and Shen, R. F., Behaviour of Pilesubject to Excavation-induced Soil Movement, Journal ofGeotechnical and Geoenvironmental Engineering, Vol.126, No. 11, pp 947-954, 2000.

Ong, D. E. L., Leung, C. F. and Chow, Y. K., Time-dependentPile Behaviour due to Excavation-induced Soil Movementin Clay, Proc. 12th Pan-American Conference on SoilMechanics and Geotechnical Engineering, MIT, Boston,U.S.A, Vol. 2, pp. 2035-2040, 2003a.

Ong, D. E. L., Leung, C. F. and Chow, Y. K., Piles subject toExcavation-induced Soil Movement in Clay, In publication,XIIIth European Conference on Soil Mechanics andGeotechnical Engineering, Prague, Czech Republic, 2003b.

Poulos, H. G., Design of Piles subjected to Lateral SoilMovements. 5th Indonesian National GeotechnicalConference.

Randolph, M. F. and Houlsby, G. T., The Limiting Pressure ona Circular Pile Loaded Laterally in Cohesive Soil,Geotechnique Vol. 34, No. 4, pp. 613-623, 1984.

Ting, W. H., Chan, S. F. and Ooi, T.A., Design Methodologyand Experiences with Pile Supported Embankments,Symposium on Development in Geotechnical Engineering,AIT, Thailand, 1994.

Ting, W. H., Tan, Y. K., The Movement of a Wharf Structuresubject to Fluctuation of Water Level, Proc. of XIV th Int.Conf. on Soil Mechanics and Foundation Engineering,Hamburg, 1997.

Wei, J., Centrifuge Modelling of Deep Excavations., M.EngThesis, National University of Singapore, 1997.

study of geotechnical failures through physical modeling

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