Design of Foundations in Seismic Areas: Principles and Applications, Bhattacharya (ed)

NICEE, ISBN: 81-904190-1-3

LIQUEFACTION SUSCEPTIBILITY, ASSESSMENT AND REMEDIATION

A.J. Brennan a), L. Govindaraju b) and Subhamoy Bhattacharyac)

ABSTRACT

The occurrence of liquefaction has been long recognised as a major hazard during earthquakes, and recent earthquakes including that in Gujarat in 2001 continue to exhibit liquefaction behaviour. This paper explains why liquefaction occurs and the possible effects. Based on this, the identification of liquefiable soils based on laboratory and site investigations is introduced. A review on the Indian code of practice and the International practice on assessment of liquefaction hazard is presented. Strategies for combating or controlling liquefaction are then discussed. These include traditional methods such as densification and drainage, as well as more novel techniques such as limiting shear strain. Guidance on the applicability and appropriateness of each method is given, along with simple design considerations, case studies and design examples.

Key words: earthquakes, densification, grouting, liquefaction, liquefaction remediation, site amplification, vertical drains

INTRODUCTION

The occurrence of liquefaction has been long recognised as a major hazard during earthquakes, and recent earthquakes including that in Gujarat in 2001 continue to exhibit liquefaction behaviour. The phenomenon of soil liquefaction during earthquakes has been recognised since the 1960s, when observed during the earthquakes in Niigata and Alaska. Simple laboratory tests followed (Seed and Lee, 1966) to elucidate the effects of cyclic loads on saturated sands. Liquefaction is now acknowledged as a crucial problem in seismically-active regions and much research has been done on the exact nature of cyclic soil behaviour, the composition of susceptible soils, and techniques to combat the associated problems. The frequent observation of liquefaction during earthquakes, and its associated damage, has made liquefaction a major concern in geotechnical earthquake engineering.

OCCURRENCE OF LIQUEFACTION

Loose Sand

A loose dry sand subjected to shaking (i.e. cyclic shearing) would be expected to reduce in volume, as particles are given the energy to drop into inter-particle voids (Fig. 1a). If the soil is instead saturated and these voids are filled with water, then the timescales are insufficient to allow the sand particles to drop into the gaps during shaking. In other words, an undrained response with no volume change occurs (Fig. 1b). Soil particles lose contact and effective stress reduces. Excess pore pressure rises as the floating grains pressurise a) University of Dundee (U.K) b) University of Oxford (U.K) c) University of Oxford (U.K)

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the fluid. At zero effective stress, looser soils are entirely in a state of sedimentation, as postulated by Florin and Ivanov (1961) and observed in glass beads by Sasaki et al. (2001). Shear strength is reduced greatly and large settlements occur. This phenomenon is called complete liquefaction. Only after a period of time long enough to allow the fluid to drain away will the excess pore pressures dissipate and the grains regain contact.

The generation of pore pressures has been investigated by Seed et al. (1976) by cyclic laboratory tests. N/Nl is the ratio between current cycle number to the number of cycles at the current shear stress level required to cause full liquefaction. ru is the excess pore pressure ratio, defined in Equation 1 as the ratio between excess pore pressure and initial vertical effective stress v0', and is a frequently used term in liquefaction problems because it equals 0 before shaking and 1 at complete liquefaction. The relationship between the two is given by Equation 2, which models a rapid initial rise and a rapid final rise in excess pore pressure. Parameter should be approximately 0.7. A similar curve is used for design purposes in Japan, by the PHRI (1997).

=0v

uur

Equation 1

+= 12arcsin1

21

1

lu N

Nr Equation 2

Dense Sand

Dense sands under shearing experience an initial volume drop followed by an increase in volume as the particles ride over each other. The soil dilates. As demonstrated by Seed and Lee (1966) and others, dense saturated sands under shaking can also generate large excess pore pressures. A difference in behaviour is observed when shear is applied to the soil; attempting to shear dense sands in this condition causes the soil to dilate. This means that the soil attempts to increase in volume (rather than decrease), dropping the excess pore pressure and increasing the effective stress (and hence strength). Dense sands differ because, even when they generate these large excess pore pressures, they do not exhibit catastrophic failures as they regain strength rapidly when sheared. This state is therefore not called liquefaction but variously cyclic mobility (Castro, 1975) or liquefaction with limited strain potential (e.g. Seed, 1976).

Figure 1: Schematic behaviour of loose sand particles under rapid shaking.

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Figure 2: Generation of excess pore pressure as a function of number of cycles (adapted from Seed et al.,

1976).

EFFECTS OF LIQUEFACTION

Liquefaction causes loss of strength, settlement or tipping of buildings, failure of earth structures; lateral spreading of gently sloping ground, densification causing vertical ground settlements and slope instability.

Loss of Strength and Settlement

As the particles lose contact, so the shear strength drops. Slopes and dams can slip (e.g. Lower San Fernando Dam, described by Seed et al., 1975, or various dams described by Muhunthan and Schofield, 2000), and structures can settle and tilt. Figure 3 shows some images of such failures from past earthquakes; there are many similar occurrences on record. This is the primary problem that the remediation measures described below attempt to address.

Additional settlement also occurs as the large excess pore pressures dissipate during the hours that follow liquefaction. The amount of settlement occurring here has been approximated by Ishihara (1985) in relation to the shear strain in the ground (as plotted in Figure 4) or Shamoto and Zhang (1996) who proposed Equation 3. In Equation 3, vp is permanent volumetric strain, max is the maximum shear strain caused by the earthquake, e0 is in-situ void ratio and emax/emin are the reference state void ratios. Note that Equation 3 and Figure 4 can give rather different results.

+

+=

0

minmax076.0max 1

3.13.06.1

eeee

vp Equation 3

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Figure 3: Example failures associated with liquefaction strength loss. a) Taiwan, 1999, photo Uzarski and Arnold (2001); b) Turkey, 1999, photo Youd et al. (2000); c) Niigata, 1964, photo National Information Service for Earthquake Engineering, University of California, Berkeley; d) Taiwan, 1999, photo S.P.G.

Madabhushi.

Figure 4: Relationship between shear strain and volumetric strain in liquefaction events, from Ishihara

(1985).

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Loss of Stiffness

As pore pressures rise and effective stress reduces, so the stiffness of the soil decreases. Soil stiffness is usually considered to be a function of the square root of effective stress (see, e.g., Hardin and Drnevich, 1972). This has the associated effect of reducing the natural frequency of the soil layer, and input motions may amplify or attenuate in a different manner to that expected (e.g. Madabhushi and Schofield, 1993). It is therefore important that the site response analysis for liquefiable conditions allows for the generation of pore pressures. Downhole investigations by Youd and Carter (2005) go further, suggesting that accelerations at high frequency will be significantly attenuated and low frequencies amplified if, and only if, the large excess pore pressures are generated at the start of shaking. The same authors suggest the cutoff between high and low frequency is 1 Hz (or 1 s period) which seems somewhat misleadingly exact. As these figures are based on case histories, it seems intuitive to concentrate instead on the actual natural frequencies of the given site in question and its likely frequency shift in the event of excess pore pressure generation. A simple dynamics calculation can guide.

Example of Altered Site Response

In a simplified site, perhaps some h = 12 m of liquefiable soil might be encountered. Site investigation reveals a saturated bulk density of 1800 kg/m3 and an average voids ratio e of around 0.8. According to the equation of Hardin and Drnevich (1972) (Equation 4, for units of kN and m), the small strain shear modulus will be approximately 48 MPa at mid-depth. This implies a shear wave velocity Vs (Equation 5) of 163 m/s and, taking this average as representative of the whole layer, hence a natural frequency f0 (Equation 6) of 3.4 Hz (or 0.29 s natural period). If excess pore pressure ratios of 0.9 are generated throughout the layer, mid-depth vertical effective stress drops to 4.8 kPa, stiffness to 15 MPa and hence natural frequency to 1.9 Hz (0.52 s period). If an expected earthquake would contain significant frequency content at around 2 Hz (0.5 s period) then these would amplify differently through the unliquefied and the liquefying sands. Figure 5 gives an approximate guide to dynamic amplification factors. For unliquefied sand, normalised frequency is 2 Hz input / 3.4 Hz natural = 0.59. Guessing a viscous damping coefficient of about 15%, this corresponds to an amplitude amplification of about 1.5 times according to Figure 5. For the liquefying soil, normalised frequency is 2 Hz / 1.9 Hz = 1.1. The same curve in Figure 5 now suggests an amplification of about 3 times, illustrating how the site response has been somewhat altered by the pore pressure generation.

Ignored from the above analysis are the effects of shear strain. In the liquefying soil then somewhat larger shear strains would be expected, which in themselves will reduce shear modulus yet further, and also cause an increase in the degree of damping (selected somewhat arbitrarily above).

( )

++

= vK

eeG

321

133230 0

2

max Equation 4

GVs = Equation 5

hV

f s40

= Equation 6

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Figure 5: Dynamic amplification of forced vibrations (amplitude only). (see Plate 5 for coloured graph)

Denser soils, as mentioned above, dilate on shearing. Model tests on such materials show not only the low stiffness attenuation of accelerations but a sudden stiffening twice per loading cycle (described in many places, e.g. Kutter, 2006) transmitting sudden large accelerations. This also changes the frequency content of the motion experienced by surface structures.

Lateral Forces

As excess pore pressure increases, so the lateral pressure from the soil increases. Retaining structures such as walls and quays come under increased load for a long duration (hours or even days) that may be more damaging than the transient lateral accelerations of the earthquake itself. For example, back analysis and model testing of failed quay walls in Kobe showed that this had indeed happened (Inagaki et al., 1996).

Piping/Boiling

Excess pore pressures are limited to the value of the initial vertical effective stress. Liquefaction conditions therefore represent an excess pore pressure profile as shown in Figure 6, that is, with an upwards hydraulic gradient of /, the critical hydraulic gradient. Water rushes upwards rapidly and piping can occur. More common is the phenomenon of sand boiling. This fast moving fluid carries sand particles upwards (possible, as the hydraulic gradient is critical) and out across the ground surface. If the surface is a different material to the liquefiable sand, clear volcano-like structures are seen as shown schematically and photographically in Figure 6. Sand boils are commonly used as indicators of damage, e.g. Juang et al. (2005).

Sand boils are a relatively unresearched field. Ishihara (1985) suggests a relationship between layer thicknesses and earthquake accelerations in order to create boils. Simple photographic studies have been performed by, e.g., Fiegel and Kutter (1994) and Butterfield and Bolton (2003), showing that thinner surface layers and imperfections attract sand boils, and Ritchie (2006), who observed that the phenomenon seems to occur only after the end of shaking.

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Figure 6: Schematic illustration of sand boiling. Photo: National Information Service for Earthquake

Engineering, University of California, Berkeley.

Lateral Spreading

The final liquefaction-related issue described in this paper is that of lateral spreading. Shallow slopes (3 - 5) such as riverbanks can move downslope as shown in Figure 7. As a consequence, rivers can get smaller and objects such as piles in the ground can be subjected to lateral forces (Abdoun et al., 2003, Dobry et al., 2003, Haigh and Madabhushi, 2002a). There is still some debate over the magnitude and influence of such forces. Existing code provision is given by the Japanese, although Haigh and Madabhushi suggests that transient forces due to soil dilation could be somewhat larger. Lateral spreading is diagnosed by the characteristic lateral cracks shown in Figure 7, which occur when more brittle materials are present at the surface.

Figure 7: Schematic illustration of lateral spreading. Photo: lateral spreading at Navlaki port, January 2001,

S.P.G. Madabhushi.

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IDENTIFYING LIQUEFIABLE SOILS

It is obviously important to be able to identify whether a soil at site is susceptible to liquefaction or not. In order to liquefy, soils must be:-

saturated,

able to contract under shear (rules out clays), and

not so permeable that drainage can occur during shaking (rules out coarse gravels).

The first criterion is fairly straightforward to determine. The second and third show that bands of particle grading are capable of liquefaction. Ishihara et al. (1989) reported Japanese tests producing limits on particle sizes, as shown in Figure 8.

1E-3 0.01 0.1 1 100

10

20

30

40

50

60

70

80

90

100

Bhuj Sand

Boundariesfor mostliquefiable soil

Boundariesfor potentially liquefiable soil

Perc

ent F

iner

(%)

Particle Size (mm) Figure 8: Limits on particle sizes, following Ishihara et al. (1989).

Material recovered from sand boils (that has liquefied) in many earthquakes was used to define these limits. For example, Figure 9 shows the site where extensive liquefaction occurred during Bhuj earthquake of 2001. The particle sizes of liquefied soil recovered from the site is also plotted in Fig. 8 which fits well in to these limits.

Liquefaction

Sampling

Figure 9: Liquefied Ground in Bhuj. Photo: January 2001, T.G.Sitharam.

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However, the curves are not to be regarded as infallible, as an artificial fill, Masado gravel, used to construct artificial islands in Kobe was considered too coarse to liquefy and yet exhibited liquefaction behaviour during the 1995 Hyogoken-Nambu earthquake.

Other factors that may affect liquefiability are intensity of earthquake and its duration, location of ground water table, soil type, relative density, particle size gradation, particle shape, depositional environment of soil, soil drainage conditions, confining pressures, aging and cementation of the soil deposits, historical environment of the soil deposit and building/ additional loads on these deposits, see for example Kramer (1996)

Evaluation of Liquefaction Potential by Field Testing

A number of investigations have been reported correlating the liquefaction potential of a site with in situ tests. The most used methods for determining liquefiability are the in-situ Standard Penetration Test (SPT) (Seed and Idriss, 1971) and Cone Penetration Test (CPT) (Robertson and Campanella, 1985). Of these, the SPT presents a dynamic load to the soil and the CPT a continuous push, so therefore the SPT should be more likely to pick out a liquefiable soil.

Standard Penetration Test (SPT)

The original work for SPTs was by Seed and Idriss (1971), and perhaps the most up-to-date version was presented by Idriss and Boulanger (2004). The method works by estimating the expected resistance of the soil, the cyclic resistance ratio or CRR, and the expected load on the soil, the cyclic stress ratio or CSR. The ratio between these is the factor of safety against liquefaction.

CSR is determined by Equation 10, amax is the maximum acceleration expected during shaking and g is the acceleration due to gravity. rd is a depth reduction factor to account for soil not being rigid, and may be given as 1 0.015z where z is depth in metres. MSF is the magnitude scaling factor to account for earthquakes of magnitude other than the 7.5 for which the method was originally derived, and given by Equation 11 for earthquake magnitude M.

MSFr

ga

CSR d

v

v

v

eq max

0

0

0

65.0 ==

Equation 10

058.04

exp9.6

=

MMSF Equation 11

CRR is determined graphically. The chart in Figure 10 presents a curve which, for a given corrected SPT blowcount, gives the cyclic resistance ratio for clean sands (less than 5% fines content). Marked on the figure are some case histories where liquefaction has (filled markers) or has not been (open markers) observed, from the extensive data set used to derive the curve. The curve may be seen as a border between liquefiable (above the line) and non-liquefiable (below the line) soils, and constantly evolves with the data set. In truth, this is a very rudimentary method and highly unsatisfactory for a number of reasons, but it remains the current best liquefaction assessment tool available to the engineer at site. Engineers should aim to support the results from this simplified procedure by personal inspection of soil samples and use of judgement.

Note that different curves are sometimes presented for different fines contents. Expectation at time of production was that fines, being non-liquefiable, would reduce the likelihood of liquefaction. More recent work (e.g. Thevanayagam and Mohan, 2000, Thevanayagam and Martin, 2002) has suggested that this is not

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the case and that a small fines content can reduce the stability of a soil and hence increase its liquefaction potential. Therefore such curves should be used with caution when significant fines content is present. In particular, the so-called Chinese criteria should be avoided (Polito, 2001).

CRR may be further multiplied by factors K and K to account for initial static shear and normal effective stresses respectively. These are most recently considered by Boulanger (2003a, 2003b). The effect of such parameters is small compared to the accuracy of the method as a whole, and they are therefore omitted for brevity. The SPT forms the basis for the liquefaction hazard evaluation in several codes as outlined below.

Figure 10: Chart for determining CRR for clean sands (Idriss and Boulanger, 2004).

Cone Penetration Test (CPT)

In this test, the resistance from the cone penetration is used as a measure of liquefaction resistance. Stark and Olson (1995) have proposed a chart as shown in figure 11 based on CPT and field case histories to evaluate CRR for clean sands, silty sands and sandy silts.

Figure 11: Chart for determining CRR for clean sand, silty sand and sandy silt (Stark and Olson, 1995).

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Shear Wave Velocity Test

The shear wave velocity can be measured in situ employing different geophysical techniques, such as the uphole, down-hole, or cross-hole methods. Andrus and Stokoe (2000) have proposed a method to evaluate liquefaction potential of a site containing clean sands, silty sands and sandy silts based shear wave velocities as shown in figure 12. Here, vs1 represents the corrected shear wave velocity.

Figure 12: Chart for determining CRR for clean sand, silty sand and sandy silt (Andrus and Stokoe, 2000).

Laboratory Testing

Experiments have been employed for liquefaction assessment using mainly

Cyclic Triaxial Test

Cyclic Direct Simple Shear Test

Cyclic Torsional Shear Test

Cyclic Triaxial Test

The cyclic triaxial test has been used since the middle 1960s by Seed and Lee (1966) and Lee and Seed (1967). It is the most widely used laboratory method for evaluating liquefaction potential of soils. During an earthquake, an element of soil in the ground is subjected to a complex system of deformations resulting from the sequence of ground motions induced by the earthquake. During such earthquakes, the soil deformations may be attributed to the upward propagation of shear waves from underlying soil layers in the deposit. As shown in Figure 13, an element of soil may be considered to be subjected to a series of cyclic shear strains that reverse directions many times during the earthquake. During the earthquake, the normal stress on the horizontal plane remains constant while cyclic shear stresses are induced for the duration of ground shaking (Seed and Lee, 1966).

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(a) (b)

Figure 13: Simulation of geostatic and cyclic stresses in the sample of cyclic triaxial load application (a) Shear wave induced in soil by horizontal earthquake vibrations (b) Shear deformation resulting from

propagated wave, Seed and Lee (1966)

In cyclic triaxial testing, cylindrical soil specimens (consolidated either isotropically or anisotropically) are subjected to a cyclic axial stress, which is intend to simulate the cyclic shear stresses experienced by an element of soil in the field during an earthquake. The resulting stress conditions on soil specimens in the triaxial tests are illustrated in figure 14. When the axial stress d is applied on the specimen in undrained condition, the shear stress induced on the 45o plane is d / 2. The normal stress d / 2 is also induced on this plane but this is purely compressive component of d / 2 which is transmitted to pore water without inducing any change in the existing effective confining stress o. On the other hand, during application of the extension portion of the loading cycle, the shear stress on the plane reverses and the maximum shear stress is equal to d / 2. Since, d / 2 is equal in magnitude; a symmetrical reversing shear stress condition is obtained.

A

B

C

Figure 14: Cyclic triaxial stress conditions in an element of soil, Prakash (1981)

The cyclic stress-controlled technique recommended by ASTM: D 5311 92 (Reapproved 1996) covers the determination of the cyclic strength (also called liquefaction potential of saturated soils) in either undisturbed or reconstituted states. In this method, a saturated and consolidated cylindrical soil specimen is subjected to cyclic load in undrained condition by means of a load rod connected to the specimen top platen.

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This cyclic loading causes an increase in the pore water pressure in the specimen, resulting in a decrease in the effective stress and an increase in the axial deformation of the specimen.

Cyclic Direct Simple Shear Test

The cyclic direct simple shear test is also capable of reproducing earthquake stress conditions similar to cyclic triaxial shear test. In this test, a short cylindrical specimen is restrained against lateral expansion and by applying cyclic horizontal shear stresses to the top or bottom of the soil specimen, the soil can be deformed in the same way as an element of soil subjected to vertically propagating shear waves as shown in figure 13.

Cyclic Torsional Shear Test

Cyclic torsional shear tests allow isotropic or anisotropic initial stress conditions on the soil specimens and can impose cyclic shear stresses on horizontal planes with continuous rotation of principal stress axes. Dobry et al. (1985) used strain-controlled cyclic torsional loading with stress-controlled axial loading of specimens to develop CyT-CAU test for measurement of liquefaction behaviour.

Example: Assessment of Liquefaction Potential of a Site by Cyclic Triaxial Shear Test

Here the method involves two parts (a) Evaluation of Cyclic Stress Ratio (CSR) and (b) Use of CSR, earthquake magnitude and soil characteristics in a deposit

(a) Evaluation of Cyclic Stress Ratio (CSR)

Cyclic strength tests are destructive and hence failure may be defined on the basis of the number of stress cycles required to reach a limiting strain or 100 % pore pressure ratio (sometimes called initial liquefaction). It is customary to consider the state of 100% pore water pressure build up or alternatively the development of 5% double amplitude axial strain as a yardstick to recognize a state of cyclic instability of a soil. Also, it is customary to consider 10 or 20 cycles in view of the typical number of significant cycles being presented in many of actual time histories of accelerations recorded during the past earthquakes. Thus the onset condition of liquefaction or cyclic strength is specified in terms of the magnitude of cyclic stress ratio (CSR) required to produce 5% double amplitude axial strain in 20 cycles of uniform load application. This cyclic stress ratio is often referred to as cyclic strength (Ishihara, 1993).

For example, the liquefaction potential of a soil that has liquefied during the Bhuj 2001 earthquake (see Figure 9) was subjected to cyclic triaxial testing as per ASTM: D 5311 92 (Reapproved 1996) at an effective confining pressure of 100 kPa, frequency of 1 Hz with sinusoidal loading and for three different relative densities. The following figures 15 (a) & (b) illustrate the double amplitude axial strain, excess pore water pressure generation with number of loading cycles and figures 15 (c) & (d) indicate stress path and cyclic strength plots respectively of the soil.

The cyclic strength of soil is specified in terms of the magnitude of cyclic stress ratio

pressure)onfiningeffectivec and stressdeviator cyclicin which 2

CSR( /3cdc/3c

dc === required to

produce 5% double amplitude axial strain in 20 cycles of uniform load application (as described by Ishihara, 1993). The cyclic strengths of Bhuj sand obtained are 0.075, 0.09 and 0.182 for soil samples at relative densities 51%, 60% and 69.7% respectively. It is evident that the cyclic strength (resistance to liquefaction) of soil increases as the relative density increases.