What is Static Liquefaction Failure of Loose Fill Slopes?

7/24/2019 What is Static Liquefaction Failure of Loose Fill Slopes?

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1 INTRODUCTION

Slope failures occur in many parts of the world. Aslope will become unstable when its shear resistanceis smaller than any external driving shear stress,which may be induced by mechanical and hydraulic

means such as rainfall, earthquake, vibration andseepage. Alternatively, a slope will also become un-stable if its shear resistance is deteriorated and re-duced due to weathering and any other mechanismssuch as static liquefaction. Very often the terminolo-gy static liquefaction is used to describe soil slopefailures and reported in literature. However, it is evi-dent that different researchers and engineers may re-fer to different failure mechanisms. Some use debrismobility (travel angle or run out distance) to judgewhether a slope failure is caused by liquefaction ornot. Clearly there is no direct relationship betweenliquefaction and mobility. For instance, a levelground can liquefy (at zero/small effective stress un-der seismic loading) with zero run out distance. Onthe contrary, a steel ball can run down a bare slope toreach a very long travel distance and this is nothingto do with liquefaction or not (Ng 2007).

What is static liquefaction? How is it triggered?What is the effective stress at failure, if the slope isfully saturated initially such as undersea slopes?How can we identify and define static liquefactionfailures? Does a strain-softening material necessarily

mean static liquefaction? Is there any difference be-tween slide failure and flow failure? What is the roleof hydrofracture? How the angle of a slope affectsthe so-called static liquefaction? Is there any differ-ence between fluidization and liquefaction? Will

static liquefaction occur in unsaturated soil slopes?How does the angle of a slope affect the potential ofstatic liquefaction? Is there any relationship betweenthe so-called static liquefaction failure and run outdistance? Can soil nails be used to stabilize anyloose fill slopes? Some of these questions have not

been well understood and addressed and some ofthem may be even controversial. In this paper, someselected issues from above are investigated via la-

boratory triaxial element tests and centrifuge modeltests on loose fill slopes using gap-graded LeightonBuzzard (LB) sand and completely decomposed gra-nite (CDG), which is a well-graded silty sand. Ob-served key failure mechanisms of static liquefactionin the LB sand and non-liquefied slides of CDG fillslopes are identified and discussed, mainly followingon the papers by Ng (2005, 2007 & 2008).

2 CLARIFICATION OF SOMETERMINOLOGIES RELATING TO STATICLIQUEFACTION

Figure 1 shows some typical results from monotonictriaxial tests on saturated, anisotropically consoli-dated sand specimens (Ng 2008). As shown in Fig.1a, a very loose sand specimen, A, exhibits a peakundrained shear strength at a relatively small shear

to much smaller shear

strength at large strains. This behaviour is oftencausally referred to c flow liquefac- by many researchers and engineers. No matter

c- e-

What is Static Liquefaction Failure of Loose Fill Slopes?

Charles W. W. NgThe Hong Kong University of Science and Technology, Hong Kong SAR

ABSTRACT: Static liquefaction failure of soil slopes has often been reported in literature. It appears thatsome researchers and engineers use different criteria to define and describe static liquefaction and they refer todifferent failure mechanisms. What is static liquefaction? How is it triggered? How can we identify and definestatic liquefaction failures? Does a strain-softening material necessarily mean static liquefaction? These arenot all easy questions to answer and some of them may be even controversial. Based on some centrifuge mod-el and triaxial element tests, suggested answers to some of these questions are explored, discussed and veri-fied in this paper.


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haviour observed in the laboratory is rather confus-ing and, strictly speaking, incorrect. Would it beclearer and more precise to describe the material be-haviour of the loose specimen, A, and a dense

- - ctively, in the deviator stress-axial

strain space (see Fig. 1a)? In the mean effectivestress-deviator stress space (see Fig. 1b), would it be

more precise to use the terms -called collapse (Sladen et al. escribe

the strength changes of specimen A and specimen B,respectively? Of course, it is well-recognised that areduction and an increase in undrained shear strengthare caused by the respective tendency of sample con-traction and dilation, leading to a respective increaseand a reduction in pore water pressure ( u) forspecimens A and B during undrained shearing (seerelationship between u and axial strain in Figure1c). It must be pointed out that these are just materialelement behaviour that does not necessarily captureand represent the global behaviour of an entire fillslope or an earth structure.

3 INVESTIGATION OF THE FAILUREMECHANISM OF LIQUEFIED FLOW IN SANDFILL SLOPES BY CENTRIFUGE TESTS

3.1 Model materialCentrifuge model tests were carried out at the GCF

of HKUST (Ng et al. 2002a, Ng et al. 2006a) to in-vestigate the failure mechanisms of static liquefac-tion of loose fill slopes subjected to rainfall, a risingground water table and dynamic earthquake loadings(Zhang 2006, Zhang et al. 2006, Ng 2007). LeightonBuzzard (LB) Fraction E fine sand was selected asthe fill material for the model tests. Fig. 2 shows thegap-graded particle size distribution of LB sand. D10 and D50 of the sand were 125 m and 150 m, re-spectively. Following BSI (1990), the maximum andminimum void ratios of the LB sand were found to

be 1.008 and 0.667, respectively (Cai 2001). The es-timated saturated coefficient of permeability was 1.6

10 -4 m/s. LB sand was chosen because of its pro-nounced strain-softening characteristics with its highliquefaction potential, LP, i.e., a substantial strengthreduction in shear strength when it is subjected toundrained shearing (see Fig. 3a). The results fromfour loose specimens with different initial void ratios(eo) shown in the figure are obtained from isotropi-cally consolidated undrained compression triaxialtests. The loose sand clearly shows pronouncedstrain-softening behaviour and substantial strength

reduction in the deviator stress and shear strain ( q- q) space and contractive responses in the mean ef-fective stress ( p ) and deviator stress ( q) space, i.e. p decreases continuously as q increases until the peakstate is attained (see Fig. 3b), where p and q are

equal to )2( 31 /3 and )( 31 , respectively.After the peak state, q drops (the soil collapses) witha large deformation develops until the quasi-steadystate (a shear strain of about 15%) or the criticalstate (shear strain = 30%) is reached. The criticalstate friction angle ( c) of the sand is 30 (Cai2001). Following the approach proposed by Lade(1992), the angle of instability ( ins) determined for

the sand is 18.6. It is well-known that ins is dependent on void ratio and stress level (Chu & Leong2002). For engineering assessment and design of re-medial work for loose fill slopes, it may be reason-able to assume this angle is a constant as the firstapproximation.

Figure 1. Liquefaction, limited liquefaction, and dilation inmonotonic loading tests (modified from Castro 1969, Kramer1996).

3.2 Model package and test proceduresFigure 4 shows an instrumented 29.4 o loose sand fillslope model together with the locations of the porewater pressure transducers (PPTs) (Zhang & Ng2003, Ng 2008, Ng et al. 2009). The model slopewas prepared by moist tamping. The initial relative

compaction was 68%.The body of the sand slope was instrumented withseven PPTs and arrays of surface markers were in-stalled for image analysis of soil movements. Linearvariable differential transformers (LVDTs) and a la-

Limitedliquefaction

Dilation

D e v

i a t o r s t r e s s

Axial strain

ALiquefaction

A

B

C

B Dilation

Limited li-quefaction

Liquefaction

C

A

B Dilation

Limitedliquefaction

Liquefaction

CPhase trans-formation

point

Strain hardening

Strain softening Strain hardening

Strain softening

Undrained strength reductiondue to contractive tendency

contractive tendency

Contractive tendency

contractive tendency

Dilative tendency

Mean effective stress

Axial strain

D e v

i a t o r s t r e s s

E x c e s s p o r e p r e s s u r e

Undrainedstrength increasedue to dilativetendency

(b)

(c)

(a)


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ser sensor were mounted at the crest of the slope tomonitor its settlement.

0

20

40

60

80

100

0.001 0.01 0.1 1 10Particle size (mm)

P e r c e n t a g e

f i n e r

( % )

LB-SandSKW-CDGCKL-CDGBH-CDGWTS-CDG

Figure 2. Particle size distributions of LB sand and CDG.

Figure 3. Contractive behaviour of loose LB sand under con-solidated undrained tests (a) in the q - q and (b) in p - q planes(modified from Zhang 2006, data from Cai 2001).

3.3 Observed static liquefaction mechanismAlthough the initial angle of the loose slope was

prepared at 29.4o

at 1 g, the slope was densified to80% of the maximum relative compaction due toself-weight compaction at 60 g. The slope angle wastherefore flattened to 24 o (see Fig. 5a), which issteeper than the angle of instability of 18.6. This

implies that the slope was vulnerable to instability,which could lead to liquefaction (see Fig. 3). At 60g, the 18 m-height (prototype) slope was de-stabilised by rising ground water from the bottom ofthe model (Zhang 2006, Ng et al. 2009). The loosesand slope liquefied statically and flowed rapidly(see Fig. 5b), i.e., it followed a process in which theloose slope was sheared under undrained conditions,

lost its undrained shear strength as a result of the in-duced high pore water pressure (see Fig. 6) and thenflew like a liquid, called

Sand

Modelcontainer

Drainage board

LVDT & Laser sensor

PPT7

PPT 5 PPT 6

PPT4

PPT2PPT1

PPT3

LVDT

2 9

. 4

3 0 5

1130,7

Inlet hole

Reflector O utlet hole

Temporaryreservoir

x

y

Figure 4. Centrifuge model of a loose sand fill slope subjectedto rising ground water table at 60 g (Zhang & Ng 2003).

Figure 6 shows the measured rapid increases inthe excessive pore water pressure ratio ( u/ v)within about 25 seconds (prototype) at failure at anumber of locations in the slope during the test. Themaximum measured u/ v was about 0.6, whichwould be much higher if a properly scaled viscous

pore fluid were used to reduce the rate of dissipationof excess pore pressure in the centrifuge. This meansthat the slope would liquefy much more easily. Asshown in Figure 5b, the completely liquefied slopeinclines at about 4 o to 7 o to the horizontal after thetest. The observed fluidization from in-flight videocameras and the significant rise in excessive porewater pressures during the test clearly demonstratedthe static liquefaction of the loose sand fill slope. Itshould be noted that measurements of sudden and

significant rise of excessive pore water pressures areessen liquefaction of loose fill slopes if no video recordingis available. The liquefaction of the loose sand slopewas believed to be initially triggered by seepageforces in the test (Ng et al. 2009). It is obvious thatsoil nails cannot be used to stabilize a loose sand fillslope which has a high liquefaction potential (seeFig. 3a).

Figure 7 shows five postulated zones, Z 1-Z5, rep-resenting the sequence of the failure and liquefaction

process of the slope (Ng et al. 2009). Z 1 is a failureregion de-stabilised by the loss of its toe due to theseepage force in the gully (drained failure). The soilmass at the toe of Z 1 slid with the soil at the gullyhead to trigger the failure of Z 2. The soil mass in Z 2

(a)

(b)

0 10 20 30 400

100

200

300

400

500

600

700

800

e0=0.970

e0=0.992

e0=0.983

e0=0.973

q ( k P a

)

q (%)

Quasi-steady state

700

500

300

100

Liquefaction potential (LP)

Quasi-steady state

1130.7

Model scale


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collapsed rapidly (undrained) which was then fol-lowed by the collapse of Z 3 (undrained) without in-ducing obvious deformation in the lower part. Thecollapses of Z 2 and Z 3 were due to the strain-softening associated with the significant strength re-duction (i.e. high liquefaction potential) of the looseLB sand as illustrated in Figure 3. The rapidundrained collapses of Z 2 and Z 3 were evident from

the measured large excess positive pore pressures atPPT7 (see Fig. 6).

Figure 5. Slope profile in a loose sand fill test (a) before risingground water table; (b) after static liquefaction (Zhang & Ng2003).

37.8 38.2 38.6 39.0 39.4 39.8 40.2 40.6 41.0 41.4 41.8 42.2 42.6 43.0-1.0

-0.8

-0.6

-0.4

-0.2

0.0

0.2

0.4

0.6

0.8

1.0

E x c e s s p o r e p r e s s u r e r a

t i o

( u w

/ v '

)

Duration (min)

Slope failure

PPT6

PPT7

PPT5PPT4

PPT3

PPT2

PPT1

Figure 6. Measured sudden and substantial increases in porewater pressure at seven locations inside the slope (Zhang & Ng2003).

Subsequently, Z 4 also collapsed as a result of thestrain-softening associated with the significantstrength reduction (high liquefaction potential) of theloose LB sand (see Fig. 3). The dotted line drawn be-tween Z 4 and Z 5 in Fig. 7 represents the upper

boundary of the stable region (Z 5), monitored bymarkers and the small excess pore pressures at PPT1

and PPT2 (see Fig. 6) during the liquefaction proc-ess.Based on the observed mechanism, it is fair to

suggest that soil nails cannot be used to stop any liq-uefied flow of loose sand fill slopes. However, theuse of soil nails can reduce the magnitude of any ex-cessive positive pore water pressure generated in aloose sand slope, minimize the chance of liquefac-tion and reduce damages after liquefaction (Zhang etal. 2006).

Figure 7. Postulated failure zones during the liquefaction ofslope SG30 (Ng et al. 2009).

4 OBSERVED EXCESSIVE SETTLEMENTS OFTHICK LOOSE CDG FILL SLOPES INCENTRIFUGE

4.1 Monotonic and cyclic behaviour of CDG from Beacon Hill (BH)

Prior to centrifuge model tests, a series of undrainedmonotonic and cyclic triaxial tests on normally con-solidated CDG specimens (70 mm in diameter and140 mm in height) were performed to assist in theinterpretation of centrifuge tests on loose CDG fillslopes. Figure 2 shows the particle size distributionof the well-graded CDG samples obtained from ChaKwo Ling (CKL), Kowloon. In the figure, the well-graded CDG taken from Beacon Hill (BH) is also in-cluded for comparison. The mean particle size, D50,of the CDG from CKL is 1.18 mm and the samplecontains about 15% fines content. According to theBritish Standard, BS1377 (1990), CDG can be clas-sified as well-graded silty sand.

The triaxial specimens tested (Fig. 8) were prepared by moist tamping at the optimum moisture

content (Ng et al. 2004a). The initial relative compaction of the specimens was 70% before saturation.Enlarged lubricated end platens were used in thetests to reduce the end constraints on the soil speci-mens. In the undrained monotonic triaxial compres-

Gully head

Z5

Z1Z2

Z3

Z4

Gully erosion

Slope profile before failure

Water

Sand movement

Water flow

PPT1PPT2

PPT4PPT5

PPT7

PPT6

PPT3

Final slope profile

0 5 10 15 20 25 30 35 40 45 50 55 60 650

5

10

15

20

(m)

B

B

A

A


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Figure 10 illustrates the changes of the initiallymoist-tamped structure of the model fill at the crestduring the test. At 1 g, the very loose soil had an ini-tially very open structure (see Fig. 10a), which con-sisted of large voids supported by capillary suction.One such void is circled in the figure. At 60 g, manyof these macro-voids were observed to collapse (Fig.10b). However, not all the voids collapsed. In par-

ticular, the voids at low stress levels (i.e. shallowdepths) such as the highlighted void in Figure 10asimply settled along with the fill. The observationsof the collapse and the mechanisms shown in thesetwo figures cannot be easily obtained from the fieldor numerical analyses even with large-strain formu-lations.

After the initial self-weight consolidation, the fillslope was subjected to the equivalent of six weekly

periods of rainfall infiltration in centrifuge. A sig-nificant portion of the soil suction was destroyedvery rapidly at the shallow location after the arrivalof rainfall on the slope surface (Take et al. 2004).The loose model fill responded immediately to thisloss of surface tension by collapsing the macro-voids

that had survived self-weight consolidation (Fig.10c). Although the slope was suffered from exces-sive settlement, no flow slide and no liquefactionwere observed in the test. This finding is consistentwith the test in CDG reported by Ng et al. (2002b).

Figure 9. Model geometry of CDG fill slope (Take et al. 2004).

Figure 10. The observed changes of soil structure of the crest region due to rainfall infiltration (Take et al. 2004).

4.3 Response of loose fill slope subjected to rising ground water in centrifuge (Ng et al. 2002b)

To complement the rainfall infiltration tests carriedout at Cambridge, a series of centrifuge model testson loose CDG fill slopes with and without soil nailswas subjected to rising ground water at HKUST (Ng et al. 2002b, Zhang 2006). The CDG fill materialused for the tests in Hong Kong was also from BH.A model slope was initially prepared to incline at 45 o to the horizontal and the initial relative compactionof the fill was less than 80%. At 60 g, a 300 mmhigh model slope was equivalent to an 18 m high

prototype slope. Figure 11 shows the measured displacement vectors of a 45 o unreinforced loose CDGfill slope destabilised by the rise of the ground water.Excessive settlement was measured but no sign ofliquefied flow or slide of the slope was observed

during and after the test. This was probably becauseof the small liquefaction potential of the CDG (Ng etal. 2004a).

Figure 11. Displacement vectors in unreinforced slope (CG45)(Ng 2007).

2 4

. 2 4 0

6 .

2 4 00.600

18.900

Prototype Scale

Unit in metre Scale


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4.4 Response of loose CDG fill slopes toearthquake loading in centrifuge

4.4.1 Centrifuge model and test procedures (Ng etal. 2004b, Ng 2007)

To further investigate the possibility of flow lique-faction of loose CDG fill slopes, uni-axial and bi-axial dynamic centrifuge tests were carried out usingsoil samples taken from BH (Ng et al. 2004b). Themodel CDG fill slopes were subjected to shakingranging from 0.08 g to 0.28 g (prototype) in the cen-trifuge at HKUST. All the models were essentiallythe same in geometrical layout and made of looseCDG with the same initial dry density. Figure 12shows a typical model slope (6 m in prototype) ini-tially inclined at 30 o to the horizontal with its in-strumentation. A rigid rectangular model box wasused to contain the CDG samples compacted to aninitial dry density of about 1.4 g/cm 3 (or 77% ofrelative compaction). Five pairs of miniature accel-

erometers were installed in the slope. Each pair wasarranged to measure soil accelerations in two hori-zontal directions (i.e., X- and Y-directions). Fourminiature pore pressure transducers were installed inthe soil near the accelerometers to record pore water

pressures during shaking. On top of the slope, threeLVDTs were mounted to measure the crest settle-ment, and one LVDT and one laser sensor (LS) wereused to measure horizontal movement of the crest.

To simulate the correct dissipation rate of exces-sive pore pressures in the centrifuge tests, sodiumcarboxy methylcellulose (CMC) powder was mixedwith distilled deionized water to form the properlyscaled viscous pore fluid and to saturate the looseCDG model slopes.

After model preparation, the speed of the centri-fuge was increased to 38 g. Once steady state pore

pressure condition was reached at all transducers, awindowed 50 Hz (1.3 Hz prototype), 0.5 s (19 s pro-totype) duration sinusoidal waveform was then ap-

plied (Ng et al. 2004b). After triggering each earth-quake, the centrifuge acceleration was maintainedlong enough to allow the dissipation of any excess

pore pressure. Due to page limits, only some resultsfrom one biaxial shaking test are discussed here.Other details of all the tests are presented in Ng et al.(2004b).

140

660

712

3 0

1 5 0

LS-h1

LVDT-v3

LVDT-h1

LVDT-v1

Z

X

ACC-T-X,Y,Z

LVDT-v2

ACC3-X

ACC2-X

ACC1-YPPT1

PPT2

ACC3-YACC4-X

ACC1-XACC5-X

PPT3 ACC5-Y

PPT4 ACC4-Y

ACC2-Y

Figure 12. Configuration of the model slope and instrumenta-tion (Ng et al. 2004b).

4.4.2 Measured responses of the loose CDG fill slope subjected to bi-axial shaking (M2D-0.3)(Ng et al. 2004b)

Figure 13 shows some measured horizontal accelera-tion time histories in the X- and Y-directions to-gether with their normalized amplitudes in the Fou-rier domain. In the biaxial shaking test, the baseinput accelerations (recorded by ACC-T-X & ACC-

T-Y as shown in the figure) were 11.26 g (0.28 g prototype) and 7.77 g (0.19 g prototype) in the X direction and Y-direction, respectively. The win-dowed sinusoid waveform applied in the Y-directionlagged the X-direction input signal by 90. Recorded

by the accelerometer near the crest, the peak accel-eration in the X-direction increased by 45% atACC4-X, higher than that measured in a correspond-ing uni-axial shaking test (Ng et al. 2004b). A simi-lar trend of variations in the acceleration was alsofound in the Y-direction. The normalized spectralamplitudes of acceleration at the predominant fre-quency of 50 Hz decreased by about 9% in the X-direction but increased by about 4% in the Y-direction in the upper portion of the embankment.

Figure 13. Seismic acceleration history and Fourier amplitudespectrum (M2D-0.3) (extracted from Ng et al. 2004b).

Figure 14 shows the time history of the excess pore pressure ratios along the height of the modelembankment during shaking. Peak acceleration oc-curred at about 0.25 s after the start of shaking. Themaximum pore pressure ratio occurred at about 0.33s at each of the three transducers (PPT1, PPT2 &PPT4). PPT1 and PPT2 recorded about the samemaximum pore pressure ratio of 0.87, whereas PPT4registered the smallest of 0.75. These measured val-ues were less than the theoretical value of 1.0 forliquefaction, even though the pore fluid was cor-rectly scaled in the test. The excess pore pressuresdissipated to zero at about 12 s (6.8 minutes in pro-totype) after the start of shaking.

Figure 15 is a photograph of the model taken afterthe completion of a shaking test. The deformation

profile for the slope was similar in both the uni-axial


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and bi-axial shaking tests. The observed profile ofthe deformed slope clearly illustrates that no lique-fied flow and non-liquefied slide took place duringthe shaking. The significant difference between theobserved physical test results from the loose LB sandand CDG fill slopes may be attributed to the differ-ence in fine contents, gradation and liquefaction po-tential of the two materials (see Fig. 3).

-0.5

0.0

0.5

1.0

0 5 10 15Time (s)

PPT4 (Z=10mm)

PPT1 (Z=145mm)

PPT2 (Z=100mm)

u

/ v

-0.5

0.0

0.5

1.0

0.0 0.2 0.4 0.6 0.8 1.0Time (s)

PPT2PPT1

PPT4

Figure 14. Measured excess pore-water pressure ratios in bi-axial shaking test M2D-0.3 (Ng et al. 2004b).

LVDT

Laser sensor

Original liquid surface

Laser sensor

LVDT Laser sensor

Laser sensor

Ground water level

Figure 15. A typical profile of a loose fill slope after shaking(Ng et al. 2004b, Ng 2007).

5 OBSERVED EXCESS SETTLEMENT OF CDGFILL SLOPE IN THE FIELD

Tang & Lee (2003) reported a large-scale field trialon a partly reinforced 33 o loose CDG fill slope (seeFig. 16). The bulk fill material was taken from BH.The height and width were 4.75 m and 9 m, respec-tively. It was constructed by the end-tipping method

and resulted in a loose state with an initial dry den-sity ranging from 70% to 75% of the maximum drydensity. It was considered that the stress state of thisslope would represent reasonably well to that of

most of the existing fill slopes formed before 1977in Hong Kong.

Figure 16. General view of the slope (from Tang & Lee 2003).

Two rows of grouted nails were installed at a gridof 1.5 m x 1.5 m at an inclination of 20 o from the ho-rizontal. Holes of 100 mm diameter with two differ-ent lengths (8 m and 6 m) were drilled. A 25 mm di-ameter steel ribbed bar was inserted into each holeand the hole was filled with grout from the bottomup using a plastic hose.

In order to destabilize the slope, a water re-chargesystem was used. This re-charge system comprisedcrest recharge trench, buried piping system and sur-face sprinkler and they were installed separately sothat a rise in groundwater table in combination witha rainfall event could be simulated in the field.

To increase destabilising forces, 1 m x 1 m x 0.6m concrete blocks were stacked up to 3m high at thecentral area of the slope crest. They imposed a sur-charge loading of 72 kPa.

The slope was heavily instrumented. Details ofthe instrumentation are described by Tang & Lee(2003). After water was being recharged into theslope through the piping system, it was observed thatthe deformation of the slope increased rapidly. The

total deformations at the crest and mid-slope were139 mm and 33 mm, respectively. Due to the largedeformation, the surcharged blocks tilted and col-lapsed (Fig. 17). The settlement-induced topplingfailure of the blocks was restricted at the crest zoneand the slope remained intact. No sign of static li-quefaction and flowslide was observed in this large-scale field test. The observed excessive settlementsand large measured nail forces in the field are similarto those measured in the centrifuge model tests, asshown in Figure 18 (Ng et al. 2002b, Ng 2008).


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Figure 17. General view of slope after failure (from Tang &Lee 2003)

Figure 18 compares the displacement vectors ofthe loose CDG fill slopes obtained from two centri-fuge tests, one without and one with soil nails. Thesoil nails were installed in-flight at 60 g and it can beseen that the soil nails substantially reduced soilmovements by at least a factor of 5. No sign of staticliquefaction of the slopes was observed during andafter the tests. Similar findings are also reported byTake et al. (2004) from independent centrifugemodel tests using the same loose CDG fill at Cam-

bridge University and by Tang & Lee (2003) fromlarge-scale field tests conducted at Hong Kong Uni-versity.

Figure 18. Comparisons of measured soil displacements with-out (CG45) and with soil nails (CGN45) in two centrifuge testsusing CDG loose fill at 60 g (dimensions in metres at prototypescale) (Ng et al. 2002b).

6 OBSERVED NON-LIQUEFIED SLIDEMECHANISMS OF SHALLOW CDG FILLSLOPES IN CENTRIFUGE

6.1 Destabilisation of loose shallow CDG fill slopes

near the crest (Ng et al. 2007)The Housing Department of HKSAR has been ac-tively looking for innovative methods to preserve theenvironment by minimizing the need for felling treeswhen improving the stability of existing shallow

loose CDG fill slopes. Centrifuge model tests werecommissioned to investigate possible failure mecha-nisms of loose fill slopes. Figure 19 shows an in-strumented centrifuge model created to study the po-tential static liquefaction of a loose shallow CDG fillslope subjected to a rising ground water table. The

particle size distribution of the CDG used is denotedas WTS in Figure 2. The initial fill density was 66%.

This model was used to simulate a 1.5 m thick, 24 mhigh layered fill slope when tested at 60 g. In addi-tion to laser sensors (LSs) installed for monitoringsoil surface movements, PPTs were installed tomeasure excess pore water pressures during the tests.Effects of layering were considered by tilting themodel container during model preparation. The slopewas destabilised by downward seepage created by ahydraulic gradient, which was controlled by the wa-ter level inside the upstream temporary reservoir andthe conditions of the outlet hole located downstream(see Fig. 19). Two failures were induced in the test.

Figure 19. Model package of an instrumented shallow fill slope(Ng et al. 2007).

Figures 20 and 21 show the occurrence of a non-liquefied slide and the measured excessive pore wa-ter pressure during two failures, respectively. Theslide was initiated near the crest. Based on the ob-served failure mechanisms and the small excessive

pore water pressures measured, it was concluded thatnon-liquefied slide of loose shallow CDG fills slopescould occur but static liquefaction was very unlikelyto happen in the slopes.

6.2 Destabilisation of loose shallow CDG fill slopeat the toe in centrifuge (Take et al. 2004)

Take et al. (2004) also carried out centrifuge modeltests to investigate the possible slide-flow failuremechanism of a loose thin CDG fill layer. The CDGused was taken from Beacon Hill. Figure 22 showsthe geometry adopted. The slope angle was 33 o. At

30 g, the model corresponded to a fill slope of 9 m inheight, with a vertical depth of fill of 3 m. The cho-sen soil profile for the model fill also represents anidealized case of layering in which the CDG fill ma-terial has been sieved and separated into its coarse

PPT1

PPT8

PPT3

PPT4

PPT5

PPT6

PPT7

PPT

Unit: mm

Model box

Downstream drainage board

Upstream drainage board

Outlet hole

Upstream temporaryreservior

Downstream temporaryreservior

Inlet hole

PPT B

Loose CDG (WTS)

Wood block

Coarse soil

Coarse soil

PPT2

PPT9 PPT C

LS3

LS1

LS2

reservoirreservoir


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and fine fractions and placed one on top of the otherto form a layered backfill. The layer ends blindly atthe toe of the slope to generate elevated transient

pore pressures (Take et al. 2004). This ensures thatthe rate of arrival of the seepage water at the toegreatly exceeds that of the leakage, thereby ensuringa more rapid local transient build up of pore water

pressures in this region than would have existed in

the absence of layering. In this experiment, the im- permeable bedrock layer was modelled by a solidwooden block, the top surface of which was coatedwith varnished coarse decomposed granite to ensurea high interface friction angle.

The density of the fill material in the first layeredslope model was very loose, with an approximaterelative compaction of 77%. After preparation, themodel fill slope was installed on the centrifuge andslowly brought to the testing acceleration of 30 g.

Figure 20. Top view of the model showing a non-liquefied slide(Ng et al. 2007).

Figure 21. Variations in the measured pore water pressure atthe crest (PPT2) and at the toe (PPT7) of the slope with time(Ng et al. 2007).

Figure 23a shows the arrival of the transient porewater at the toe of the slope. Once the line source ofseepage water was activated, the high transmissivityof the coarse layer quickly delivered water to the toe

of the fill slope. As intended, the rate of water trans-fer into the toe region exceeded the seepage velocitythrough the model fill, causing a transient increase inthe pore water pressure at the toe. The local pore wa-ter pressure was observed to increase at a nearly con-stant rate reaching a maximum value of 16 kPa at

point B in Figure 23a. As this seepage front pro-gressed towards the toe, the slope was slowly creep-

ing (Fig. 23b).After time B, the slope mass is observed to accel-erate (points B-C on Figure 23b). By analysing im-ages captured by PIV (White et al 2003) at the onsetof more rapid failure, it is found that the toe acceler-ated horizontally with an average velocity of ap-

proximately 6 mm/s (Fig. 24). The observed displacement field over this time interval indicates thatthe surface of the model fill moved down-slope at aslower velocity. When the fill material finally cameto rest, it formed a low-angle run-out. This failuremechanism differs from that of the slope destabilised

by downward seepage in the test for the HousingDepartment in which the slope was not blinded hy-draulically at the toe (see Fig. 19). The initiation ofthe non-liquefied slides differed in these two slopes.

Figure 22. A slide-to-flow landslide triggering mechanismmodel (Take et al. 2004).

Figure 23. Observed behaviour of slide-to-flow models (Takeet al. 2004).


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Figure 24. Displacement field prior to final acceleration ofloose fill model (Take et al. 2004).

6.3 Destabilisation of a dense shallow CDG fill slope at the toe in centrifuge (Take et al. 2004)

Unlike the static liquefaction mechanism of loosesand fill slopes, the non-liquefied slide triggeringmechanism is argued to be independent of soil den-sity (Take et al. 2004, Ng 2007). In order to verifythis hypothesis, the experiment was therefore re-

peated with a fill compacted to 95% maximum Proc-tor density while all other factors remained constant(Take et al. 2004).

As before, seepage water was introduced to thecrest of the model fill slope and it was quicklytransmitted to the toe of the slope, building up local-ized transient pore water pressures at an identicalrate as in the loose fill model (Fig. 23a). Since theslope material was dry, the position of the wettingfront could be observed. This dense slope exhibiteda much stiffer response to the build up of pore water

pressures, with less than one half of the pre-failuredisplacements signalling the onset of failure (seeFig. 23b). Just before reaching the failure pore water

pressure, the brittle fill material cracked and waterrapidly entered the fill. As high-pressure water en-

tered the crack, the acceleration of the slide in-creased. The extent to which this crack injected wa-ter into the fill material at time B is shown in Fig.23a. After time B, the slope mass accelerated, al-though at a slower slide velocity than observed in theloose fill slope (points B-C in Fig. 23b). The subse-quent behaviour of the model fill slope is laid out

pictorially in the remainder of Fig. 25. As the toecontinued to accelerate horizontally, the surface ofthe model fill accelerated towards the toe (Fig. 25b),with the velocity increasing to such a point that itexceeded the shutter speed of the camera (Fig. 25c).Eventually, the slope came to rest (Fig. 25d). Simi-larly to in the shallow loose fill slope, the landslideevent triggered from localized transient pore water

pressures formed a low-angle run-out. The densifica-tion of the fill slope slightly increased the pore water

pressure required to initiate failure (see Fig. 23a), but it made the failure more brittle (Take et al.2004).

Based on the non-liquefied slides observed in both loose and dense CDG shallow fill slopes, it isevident that soil nails these non-liquefied slides since the CDG still pos-sess sufficiently large shear strength after the peak

(see Fig. 8b).

Figure 25. Failure mechanism in the dense fill model (modifiedfrom Take et al. 2004).

7 CONCLUSIONS

Both static and dynamic model tests on LB and CDGwere carried out. In-flight rainfall infiltration, risingground water and dynamic loadings were simulated.Based on the tests, it can be concluded that staticliquefaction/fluidization of the loose LB sand fillslope due to a rising ground water table was success-fully created in the centrifuge. The occurrence ofliquefaction in sand was observed by in-flight videocameras and verified by the significant and sudden

build-up of excessive positive pore water pressuresmeasured at various locations in the slope. It isfound that strain softening of the material is a neces-sary but not a sufficient condition to cause flow liq-uefaction. A trigger such as seepage force or addi-tional loading is needed.

No liquefied flow and slide was observed in thickloose CDG fill slopes when they were subjected torising ground water tables, heavy rainfall infiltrationand very strong bi-axial shaking. Only excessive soilsettlements were induced. Consistency was found

between centrifuge model tests and full-scale fieldtrial of a loose CDG fill slope. The significant dif-

ference between the observed physical test results onthe LB sand and CDG models may be attributed tothe difference in the fine contents, gradation and liq-uefaction potential of the two materials.

(a)

(c)

(b)

(d)


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Although static and dynamic liquefaction did notoccur in loose CDG fill slopes because ofsmall liquefaction potential, non-liquefied shallowslides were observed in both loose and dense shal-low fill slopes. Depending on the boundary condi-tions, different initiations of non-liquefied shallowslides were captured in the centrifuge. The landslideevent triggered by highly localized transient pore

water pressures at the toe results in a low-angle run-out in both shallow loose and dense CDG fill slopes.For improving the stability of loose fill slopes, it isvital to differentiate the potential differences be-tween a liquefied flow and a non-liquefied slide. It isevident that a potentially non-liquefied slide can bestabilized by soil nails.

8 ACKNOWLEDGEMENTS

The work presented here was supported by researchgrants DAG00/001.EG36 and HKUST3/CRF-SF/08

provided by HKUST. The author is grateful for re-search contracts provided by the Geotechnical Engi-neering Office of the Civil Engineering and Devel-opment Department and the Housing Department ofthe HKSAR. Moreover, the author thanks Dr RobinZhou for checking and formatting the paper.

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