4 boulanger soa 5icege

CYCLIC FAILURE AND LIQUEFACTION: CURRENT ISSUES

Ross W. BOULANGER 1, I. M. IDRISS2

ABSTRACT Three current issues regarding the procedures for evaluating cyclic failure and liquefaction of soils during earthquakes are discussed. The three issues pertain to: (1) liquefaction susceptibility criteria for fine-grained soils, (2) SPT-based liquefaction triggering correlations, and (3) residual strength of liquefied cohesionless soils. Background information regarding each issue is presented, followed by discussion of the key underlying technical aspects. Keywords: Liquefaction, cyclic failure, residual strength, susceptibility criteria

INTRODUCTION Engineering procedures for evaluating cyclic failure and liquefaction of soils during earthquakes progressively advance through the accumulation and synthesis of findings from experimental, theoretical, and case history studies. The theoretical and experimental components provide the framework for understanding and assimilating the lessons provided by the case histories, and the means for extrapolating outside the range of these case histories. Since practice often results in the need for such extrapolation, it is important that the experimental and theoretical bases that guide the extrapolation produced by different engineering procedures be well understood and recognized. Not surprisingly, the differences in engineering procedures produced by different investigators tend to be largest when the data are sparse and the theoretical understanding still developing, and to then decrease as data are accumulated and the theoretical understanding becomes more mature. This maturation of engineering procedures benefits from the ongoing technical discussions regarding the sources of differences or concerns that arise between different recommended procedures. Three current issues regarding the procedures for evaluating cyclic failure and liquefaction of soils are discussed in this paper. The three issues addressed pertain to the topics of: (1) liquefaction susceptibility criteria for fine-grained soils, (2) SPT-based liquefaction triggering correlations, and (3) residual strength of liquefied cohesionless soils. The purpose of this paper is to provide background information regarding each issue, followed by a discussion of the key underlying technical aspects.

LIQUEFACTION SUSCEPTIBLITY CRITERIA Predicting seismic deformations requires estimating strength loss and strains for soils ranging from cohesionless (e.g., saturated gravels, sands, and nonplastic silts) to cohesive (e.g., clays and plastic silts). 1 Professor, Department of Civil & Environmental Engineering, University of California at Davis, USA, e-mail: [email protected] 2 Professor Emeritus, Department of Civil & Environmental Engineering, University of California at Davis, USA.

5th International Conference on Earthquake Geotechnical Engineering January 2011, 10-13

Santiago, Chile The choice of engineering procedures for estimating strength loss and strains depends on the soil's characteristics with possible choices including liquefaction correlations for cohesionless soils and cyclic softening procedures for clays. For saturated cohesionless soils, such as clean sands or nonplastic silts, the concern is for loss of shear strength and development of large strains associated with liquefaction. The potential for liquefaction triggering is most commonly evaluated using correlations based on in-situ tests (SPT, CPT, Vs etc.) rather than performing laboratory tests on field samples because conventional tube sampling techniques cause excessive disturbance to cohesionless soils and frozen sampling techniques are usually uneconomical. For saturated cohesive soils, such as clays and plastic silts, the loss of strength and development of strains under earthquake loading is referred to as cyclic softening. Procedures for evaluating cyclic softening focus on estimating monotonic and cyclic undrained strengths using information from laboratory testing, in situ testing, and empirical correlations. In contrast to cohesionless soils, conventional tube sampling techniques can usually be used to obtain reasonably high-quality samples for laboratory strength testing. When soft low-plasticity silts and silty clays with plasticity indices, PI, less than about 18 are encountered, guidance on how best to evaluate potential behavior has changed in recent years. The first widely-used guidance for identifying fine-grained soils as susceptible to significant strength loss was the Chinese criteria (Seed and Idriss 1982) based on the findings of Wang (1979), but the use of these criteria is now widely discouraged. Other liquefaction susceptibility criteria have been suggested over the years (e.g., Koester 1992, Youd 1998, Pollito 1999, Andrews and Martin 2000); these criteria have generally been interpreted in practice as meaning that soils judged to be susceptible to liquefaction should be evaluated using SPT- or CPT-based liquefaction triggering correlations. Bray and Sancio (2006) used cyclic test results for a wide range of soils from Adapazari and the observed field performances of those soils during the 1999 Kocaeli earthquake to conclude that silts and clays with PI ≤ 12 and water contents (wc) greater than 85% of the Liquid Limit (LL) are liquefiable, while soils with 12 < PI < 18 and wc > 0.8LL are more resistant to liquefaction but still susceptible to cyclic mobility (Figure 1a). Boulanger and Idriss (2006) suggested that the emphasis should be put on determining which engineering procedures are most appropriate for evaluating cyclic strengths, and recommended that clays and silts with PI ≥ 7 should be evaluated using cyclic softening procedures, whereas silts and clays with lower PI should be considered as likely exhibiting sand-like behavior (and evaluated using liquefaction correlations) unless shown otherwise through detailed laboratory and in situ testing (Figure 1b). The differences between the guidance provided by Bray and Sancio (2006) and Boulanger and Idriss (2006) have sometimes been perceived in practice as being greater than they really are, in large part because of semantics. The commonality between these two sets of guidance is well illustrated by the following passage from Bray and Sancio (2006),

"Based on the results of the cyclic testing performed in this study, a soil may be susceptible to liquefaction if the ratio of the water content to liquid limit is greater than 0.85 (wc/LL>0.85) and the soil plasticity index in less than 12 (PI<12). Soils that do not meet these conditions but have plasticity index less than 18 (PI<18) and water content to liquid limit ratio greater than 0.8 (wc/LL>0.8) may be moderately susceptible to liquefaction. These soils, especially those satisfying the first set of requirements, should be tested in the laboratory to assess their liquefaction susceptibility and strain potential under the loading conditions existing in the field. Soils with PI>18 did not liquefy at low effective stresses. However, structures founded on these soils, and for that matter, any soil, may


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undergo significant deformations if the cyclic loads approach or exceed the dynamic strength of the soil." [Emphases added]

Bray and Sancio (2008) further clarify their recommendations in their closure to discussions on that paper, by stating,

"The authors contend that field sampling and laboratory testing currently offer the most reliable way to evaluate the liquefaction susceptibility, resistance, and response of fine-grained soils."

(a)

(b)

Figure 1. Liquefaction susceptibility criteria by: (a) Bray and Sancio (2006) and (b) Boulanger and Idriss (2006)


Santiago, Chile The above recommendations are in good agreement with the recommendations of Boulanger and Idriss (2006), as illustrated by the following passage from their closure on discussions (Idriss and Boulanger 2008),

"Effective communication regarding issues of liquefaction requires a clear understanding of the technical definitions used by different individuals. For the paper, definitions for terms were chosen such that the names for soil type, soil behavior, and analysis methodology were reasonably consistent: (1) "liquefaction" was reserved for describing the behavior of sand-like or cohesionless soils that would be appropriately evaluated using semi-empirical SPT- or CPT-based "liquefaction" correlations; (2) "cyclic softening" was used to describe the behavior of clay-like or cohesive soils that would be appropriately evaluated using procedures developed for, or modified from those for clays; and (3) the recommended criteria were called "liquefaction susceptibility criteria," because they distinguished between these two cases.

Thus, the two sets of guidance differ in the terminology used to describe the cyclic behavior of low-plasticity soils (e.g., cyclic softening versus liquefaction), but they both agree in recommending laboratory testing as the preferred basis for evaluating the cyclic strengths and potential strains for low-plasticity, fine-grained soils, regardless of what the behavior may be called. In this regard, there is no significant consequential difference between the two sets of guidance. In the authors' experience, the greatest source of confusion in practice is the tendency to equate the classification of a soil as "liquefiable" by some liquefaction susceptibility criteria with the requirement that it be evaluated using a "liquefaction correlation." This inevitably leads to the application of SPT- or CPT-based liquefaction triggering correlations to fine-grained soils that classify as "liquefiable" or "moderately susceptible to liquefaction" by the Bray and Sancio (2006) criteria, even though those authors clearly expressed that these soils should be tested in the laboratory. Use of SPT- or CPT-based liquefaction triggering correlations will generally result in the under-estimation of cyclic strength for fine-grained soils having some plasticity; while this may be conservative, it is important that practice not overlook the potential benefits that a program of high-quality sampling and laboratory testing may offer. The above observation is not an issue with the criteria, but rather with the interpretation of the criteria in practice. This source of confusion was the primary reason that Boulanger and Idriss (2006) suggested limiting the use of the term "liquefaction" for describing those soils that should be evaluated using the liquefaction triggering correlations that were developed for cohesionless soils. The key issues for practice are selecting the appropriate engineering procedures and then evaluating the potential deformations using those procedures. As recommended by both Bray and Sancio (2006) and Boulanger and Idriss (2006), site specific laboratory testing of low plasticity silts and clays can provide important insight into the evaluation of potential strength loss and deformations (e.g., Boulanger et al. 1998, Yilmaz et al. 2004; Bray and Sancio 2006, Sanin and Wijewickreme 2006; Dahl et al. 2010a), rather than relying solely on index test-based liquefaction susceptibility criteria and empirical liquefaction or cyclic softening correlations. Detailed observations of soil response in the laboratory, including behavior during sample preparation (i.e., trimming and extruding), index testing, consolidation testing, and cyclic and monotonic loading, give insight into the fundamental behavior of the soil and improved guidance for the evaluation of potential field behavior. Sample quality and the effects of sample disturbance on monotonic and cyclic strengths for low-plasticity silts and silty clays are a concern that should be carefully considered in relation to the laboratory testing procedures [e.g., Recompression (Bjerrum and Landva 1966) versus SHANSEP (Ladd and Foott 1974) methodologies) and the anticipated loading conditions in the field. For example, the effects of sample disturbance for a low-plasticity silt may not be a concern if the field loading is expected to greatly increase the consolidation stresses (e.g., constructing an embankment on top of it), such that the laboratory specimens are consolidated to effective stresses that are significantly greater than the effective stresses that existed at the time of sampling. In


Santiago, Chile other situations, the effects of sample disturbance can be found to render the results of cyclic testing unreliable (Dahl et al. 2010b). For these and other reasons, the evaluation of cyclic strengths by laboratory testing of field samples requires recourse to fundamental procedures, with each site potentially representing unique and challenging considerations. Further research and discussions will hopefully lead to clearer guidance and implementation in practice.

SPT-BASED LIQUEFACTION TRIGGERING PROCEDURES Differences between liquefaction triggering correlations A current issue regarding SPT-based liquefaction triggering correlations is determining the source of the differences between published liquefaction triggering relationships, and particularly between those published by the late Professor H. Bolton Seed and colleagues (Seed et al. 1984, 1985), which were adopted with slight modifications in the NCEER/NSF workshops (NCEER 1997, Youd et al. 2001), and those published more recently by Cetin et al. (2004) and Idriss and Boulanger (2004, 2008). These three liquefaction triggering correlations are compared in Figure 2, along with the relationships by Tokimatsu and Yoshimi (1983) and Yoshimi et al. (1994). The Cetin et al. (2004) correlation is significantly lower than the other correlations for SPT (N1)60cs values less than 20. The question of concern is why the correlation of CRR adjusted for M = 7.5 and σ'v = 1 atm versus (N1)60cs proposed by Cetin et al. (2004) is significantly lower than those by Seed et al. (1984) and Idriss and Boulanger (2004) although they are based on largely the same case histories. This issue became a concern in 2010 when Professor Raymond B. Seed at the University of California at

Figure 2. Comparison of some SPT-based liquefaction triggering correlations


Santiago, Chile Berkeley stated that the use of the Idriss-Boulanger (2004, 2008) correlation, and presumably the similar Seed et al. (1984)/Youd et al. (2001) correlation, was "dangerously unconservative" in a series of visits to major consulting firms and regulatory agencies, a series of e-mail messages to the EERI Board of Directors with copies to over 100 prominent individuals in the USA and abroad, and a UCB Geotechnical Report titled, "Technical Review and Comments: 2008 EERI Monograph (by I.M. Idriss and R.W. Boulanger)." The statements took us by surprise and to some degree left the profession in a state of confusion and asking, "Why the controversy" and "Why in this way?" We cannot answer these questions, but can openly and carefully examine the technical reasons for the differences in these liquefaction triggering correlations and see if there is any justification for the statements made by Professor Raymond B. Seed. A re-examination of SPT-based liquefaction triggering procedures was performed and documented in Idriss and Boulanger (2010), with the purposes of re-examining and updating the case history database, re-examining how the case-history based correlations compare to the database of cyclic test results for frozen sand samples, development of a probabilistic version of the Idriss-Boulanger (2004, 2008) liquefaction triggering correlation, presenting new findings regarding components of the analysis framework used to interpret and extend the case history experiences, and examining the reasons for the differences between these liquefaction triggering correlations. An important aspect of the latter question is whether these differences represent scientific (epistemic) uncertainty, or whether they are due to aspects of the analysis frameworks or case history interpretations that could be debated, refined, or resolved. The results of that last task are briefly summarized herein for the purpose of facilitating discussion within the profession. The key to examining the difference among these liquefaction triggering correlations is assessing how well the case histories with σ'v close to 1 atm support any given correlation. Accordingly, the case histories that control the position of the different correlations (i.e., the data points that plot close to the triggering curve) with σ'v close to 1 atm were identified and examined in detail. This examination showed that the case history interpretations by Seed et al. (1984), which are also the basis for Youd et al. (2001), and Idriss and Boulanger (2004, 2008) provide rational support for their respective correlations. The case history interpretations by Cetin et al. (2004) were also found to support their correlation, but that support was found to hinge on a number of significant discrepancies and problems in their interpretations of the database. The primary cause for this conclusion was found to be the interpretations by Cetin et al. (2000, 2004) for 11 key case histories for which the effective vertical stresses range from σ'v = 0.65 to 1.5 atm in their database. The differences in the rd, Kσ, and CN relationships used in developing these three liquefaction triggering correlations were found to be a lesser contributor to the differences in the triggering curves. In particular, the differences in the Kσ and CN relationships are generally small near σ'v = 1 atm, such that they have even smaller effects on the interpretations of the 11 key case histories discussed subsequently. The case histories interpreted by, and listed in, Cetin et al. (2004) with σ'v ranging from about 0.65 to 1.5 atm are shown in Figure 3. The values of CSR adjusted for M = 7.5 and σ'v = 1 atm are plotted versus the corresponding (N1)60cs, which shows that the case histories interpreted using the parameters listed by Cetin et al. (2004) support their liquefaction triggering curve for M = 7.5 and σ'v = 1 atm.


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The eleven case history points that appear to control the position of the Cetin et al. (2004) liquefaction triggering curve (for σ'v = 1 atm and M = 7.5) were found to require the following significant adjustments. For ease of reference, these points are numbered 1 through 11 in Figure 3.

• Four cases (points 1, 2, 3, & 10) were identified and interpreted by the original investigators as "no liquefaction" sites, but were listed and used as "liquefaction" sites by Cetin et al. (2004). These sites are:

o Point 1: Miller Farm CMF-10, 1989 Loma Prieta earthquake, M = 6.9 (Holzer et al. 1994, Holzer and Bennett 2007)

o Point 2: Malden Street, Unit D, 1994 Northridge earthquake, M = 6.7 (Holzer et al. 1999, O'Rourke 1998)

o Point 3: Kobe #6, 1995 Kobe earthquake, M = 6.9 (Tokimatsu 2010, pers. comm.) o Point 10: Shuang Tai Zi River, 1975 Haicheng earthquake, M = 7.0 (Shengcong and

Tatsuoka 1984). The observations and interpretations provided by the original investigators regarding each site are well supported, and thus these four points should be re-classified as "no liquefaction" cases.

• For cases where Cetin et al. (2004) report using their regression equation for computing rd, the values of rd listed and used in their database do not agree with the values computed using their rd regression equation. The differences between the computed and listed rd values for the Cetin et al. database increase with depth, as shown in Figure 4. The key sites affected by this discrepancy are:

o Point 3: Kobe #6, 1995 Kobe earthquake, M = 6.9, critical depth = 5.8 m

(N1)60cs

0 10 20 30 40

CS

R (

adju

sted

to M

= 7

.5 &

σσ σσ'

v =

1 at

m)

0.0

0.1

0.2

0.3

0.4

12

3

45

67

8

9

1011

Triangles : 1984 cases; Circles : 2000 cases;Squares : Kobe proprietary cases.Filled-in symbols : liquefaction;Open symbols : no liquefaction;Cyan symbol : marginal.

Cases for σσσσ'v = 0.65 to 1.5 atm

Data and parameters fromCetin et al (2004); Points 1 -- 11identified for further examinationas described in text.

Cetin et al (2004)M = 7.5; σσσσ'v = 1 atm

Figure 3. Case histories for σσσσ'v = 0.65 to 1.5 atm published by Cetin et al. (2004) with eleven

points (cases) identified for further examination


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o Point 4: Kobe #7, 1995 Kobe earthquake, M = 6.9, critical depth = 6.3 m o Point 6: Rail Road #2, 1964 Niigata earthquake, M = 7.6, critical depth = 10.0 m o Point 9: Panjin Chemical Fertilizer Plant, 1975 Haicheng earthquake, M = 7.0, critical

depth = 8 m o Point 10: Shuang Tai Zi River, 1975 Haicheng earthquake, M = 7.0, critical depth = 8.5 m o Point 11: San Juan B-3, 1974 Argentina earthquake, M = 7.4, critical depth = 11.7 m

The values of rd were recalculated for these sites using the Cetin et al. (2004; also Cetin et al. 2002 and Cetin and Seed 2004) rd equation, as the authors indicated that the equation was correct (Cetin 2010, personal communication). This increased the CSR by 5% to 25%, which resulted in these case histories plotting farther above the liquefaction triggering curve.

• The boring data obtained at Kobe site No. 7 (Point 4) included a N=8 value in the sand below the water table, but this N value was not utilized by Cetin et al. (2004) in assigning a representative (N1)60=26.5 for this site. Including the N = 8 blow count in the computation gives an average (N1)60= 21.9, which results in locating the point considerably above the Cetin et al. (2004) liquefaction triggering curve for M= 7.5 and σ'v = 1 atm. If the shallower N = 8 zone had been considered the critical zone by itself, then the resulting (N1)60 = 10.4 would be even farther above the Cetin et al. triggering curve.

• The Rail Road #2 site in the 1964 Niigata earthquake (Point 6), which is a marginal case (near the boundary between liquefaction and no liquefaction), is the only point that remained close to the Cetin et al. (2004) triggering curve. Cetin et al. plotted this point considerably lower than did other investigators, with one contributing reason being that Cetin et al. used considerably smaller unit weights than did other investigators (Figure 5). Using a more realistic total unit weights for this

Figure 4. Discrepancy between rd values computed using the Cetin et al. (2004) equation and the values listed and used in their database


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site, the point moves farther away from the Cetin et al. liquefaction triggering curve for M= 7.5 and σ'v = 1 atm.

The case histories interpreted by, and listed in, Cetin et al. (2004) with σ'v ranging from about 0.65 to 1.5 atm are shown in Figure 6 with the above-described adjustments to 8 of the 11 key points. The Cetin et al. database, with the above adjustments, no longer supports the Cetin et al. (2004) liquefaction correlation for M= 7.5 and σ'v = 1 atm. Instead, the adjusted data points, which still utilize the Cetin et al. analysis framework, are now in better agreement with the liquefaction triggering curves proposed by the late Professor H. Bolton Seed and colleagues (Seed et al. 1984, Youd et al. 2001) and Idriss and Boulanger (2004, 2008). The Cetin et al. (2004) triggering correlation, if it were updated after correcting the above problems, would be expected to move close to the Idriss-Boulanger and Seed et al. (1984)/Youd et al. (2001) correlations at effective vertical stresses ranging from about 0.65 to 1.5 atm. This would also cause the Cetin et al. Kσ relationship to become flatter because it is regressed as part of their analyses, and higher CRR values at higher confining stresses would dictate a flatter Kσ relationship. The combination of these changes would be expected to reduce the degree to which the Cetin et al. procedure predicts significantly smaller CRR values than the other liquefaction triggering correlations as depth increases. The above findings indicate that the lower position of the Cetin et al. (2004) triggering correlation, particularly at (N1)60cs values less than 20, does not represent epistemic uncertainty, but rather is the result of case history interpretations and numerical errors that can be resolved and corrected.

Depthbelowgroundsurface(m)

Figure 5. Comparison of average total unit weights used by different researchers in developing their liquefaction correlations


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Frozen sampling test data SPT-based relationships for the cyclic resistance ratio of clean sands have also been developed based on the results of cyclic laboratory tests on specimens obtained using frozen sampling techniques (e.g., Tokimatsu and Yoshimi 1983, Yoshimi et al. 1994). Frozen sampling techniques have been shown to produce high-quality samples of loose to dense sands when properly implemented (e.g., Singh et al. 1982; Goto and Nishio 1988; Yoshimi et al. 1989, 1994). The volumetric expansion of saturated sand during one-dimensional freezing has been shown to depend on the confining (or surcharge) pressure on the sand (Yoshimi et al. 1978). At low confining stresses, the freezing process can produce significant volumetric expansion, which loosens the sand and reduces sample quality. As confining stress increases, the volumetric expansion during one-dimensional freezing progressively decreases. Volumetric expansion was found to become negligible for confining stresses greater than about 10 to 50 kPa, with the critical value of confining stress being different for different sands. Specimens that have been one-dimensionally frozen under sufficient confinement, and hence have developed very small net volumetric strains, have been shown to retain the memory of prior stress and strain loading histories throughout the sampling, handling, and mounting processes (e.g., Singh et al. 1982; Goto and Nishio 1988; Yoshimi et al. 1989, 1994). Results of field and laboratory test programs using frozen sand sampling techniques are summarized in Figure 7 in terms of the CRR values for 15 uniform loading cycles to cause ≈ 3.8% shear strain in undrained cyclic loading versus the corrected SPT (N1)60cs values. The cyclic resistance ratios from the

(N1)60cs

0 10 20 30 40

CS

R (

adju

sted

to M

= 7

.5 &

σσ σσ'

v =

1 at

m)

0.0

0.1

0.2

0.3

0.4

12

34 5

67

8910

11

Triangles : 1984 cases; Circles : 2000 cases;Squares : Kobe proprietary cases.Filled-in symbols : liquefaction;Open symbols : no liquefaction;Cyan symbol : marginal.

Cases for σσσσ'v = 0.65 to 1.5 atm

NCEER/Youd (2001)M = 7.5; σσσσ'v = 1 atm

Idriss & Boulanger (2004)M = 7.5; σσσσ'v = 1 atm

Data and parameters fromCetin et al (2004); Changes toPoints 1 -- 11 described in text. Cetin et al (2004)

M = 7.5; σσσσ'v = 1 atm

Figure 6. Case histories for σσσσ' v = 0.65 to 1.5 atm published by Cetin et al. (2004) with the

corrections to the eleven points (cases) as described in the text


Santiago, Chile isotropically-consolidated undrained (ICU) triaxial tests are converted to equivalent DSS strengths using the relationship,

1 2

3o

DSS ICU TX

KCRR CRR −

+ =

(1)

where Ko is the coefficient of lateral earth pressure at rest. The value of Ko was taken as possibly ranging from 0.5 to 1.0 in-situ, as was suggested by Yoshimi et al. (1989, 1994). The data by Pillai and Stewart (1994) for Duncan Dam involved cyclic DSS and TX tests, from which the correction of TX to DSS conditions was established directly. The DSS and equivalent DSS strengths from all studies are multiplied by a factor of 0.9 to account for the effects of two-directional shaking in-situ. The cyclic strengths are further corrected to an equivalent effective confining stress of 1 atm by dividing each by the Kσ factor determined using the relationships by Boulanger and Idriss (2004); this final correction was very small for most of the available data because they were performed at effective consolidation stresses close to 1 atm, with the exception of the Duncan Dam tests which had been conducted at stresses of 2 to 12 atm. The liquefaction triggering relationships by Tokimatsu and Yoshimi (1983), Yoshimi et al. (1994), and Idriss and Boulanger (2004, 2008) are also shown in Figure 7. These three correlations are in good agreement with the frozen sand test data for (N1)60cs values less than about 15. At greater (N1)60cs values, the Tokimatsu and Yoshimi (1983) and Yoshimi et al. (1994) curves follow the results of the frozen sand test data in curving sharply upward near (N1)60cs of about 40, whereas the Idriss and Boulanger (2004, 2008) relationship and the similar Seed et al. (1984)/Youd et al. 2001 relationship curve sharply upward near (N1)60cs of about 30. The liquefaction triggering relationships by Tokimatsu and Yoshimi (1983) and Idriss and Boulanger (2004, 2008) are compared in Figure 8 to the updated case history data from Idriss and Boulanger (2010). For values of (N1)60cs less than about 15, these two correlations are consistent with the case history data. For (N1)60cs values greater than about 15, the Idriss and Boulanger correlation follows the case history data, whereas the Tokimatsu and Yoshimi (1983) correlation increasingly falls below the liquefaction case history points with an increasing number of no-liquefaction cases above it. The case history database contains no cases of liquefaction for (N1)60cs values greater than 26 despite the very strong shaking levels and numerous sites with representative (N1)60cs in the range of 25 to 40. There were also 10 no liquefaction cases with values of (N1)60cs between 21 and 28 that plotted above the Idriss-Boulanger triggering curve. There are a number of influencing factors to consider when examining the differences in liquefaction triggering relationships at high (N1)60cs values shown in Figures 7 and 8. One consideration is illustrated in the upper part of Figure 7, which shows the "limiting shear strain" as described in Seed et al. (1984). The limiting shear strain shown in the figure represents the shear strain that can be expected to develop in 15 uniform loading cycles under very high CSR values, and is related to the observed cyclic loading responses of dense sand specimens, in which the cyclic stress-strain response shows a very slow, progressive accumulation of shear strains with each cycle of loading, despite the extremely large cyclic stress ratios (e.g., Yoshimi et al. 1984). For very dense sands, the peak cyclic strains, and hence any permanent shear or reconsolidation strains, may be sufficiently small that the resulting ground deformations cause no observable settlement, movements, boils, or damage. If the permanent shear strains in the field are only a fraction of the peak shear strains for dense sands, then the triggering curve based on field observations of permanent ground deformations would be expected to plot to the left of a triggering curve based on peak shear strains in the laboratory.


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A second consideration is whether the effects of volumetric expansion during in-situ radial freezing are truly negligible. Volumetric expansion is expected to have been small for the majority of the studies because the effective overburden stresses were greater than 50 kPa for all but one site. Nonetheless, there is always a concern that some expansion could develop due to adverse freezing conditions in-situ. If volumetric expansion does occur during freezing, the specimen can be expected to also develop volumetric contraction strains during thawing under confinement in the laboratory. The anticipated net effect of any disturbance induced by volumetric expansion during in-situ freezing would be a small increase in cyclic strengths for loose sands and a more significant decrease in cyclic strengths for dense sands. Thus, the effects of even minor disturbance for dense sands would be expected to cause the laboratory-based triggering curve to plot lower than the in-situ triggering curve.

Figure 7. Comparison of the Idriss-Boulanger (2004, 2008) liquefaction triggering correlation with frozen sand sample test data and the guidance by Seed et al. (1984), Tokimatsu and

Yoshimi (1983), and Yoshimi et al. (1994).


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A third consideration discussed by Castro (1975), Peck (1979), Seed (1979), and Casagrande (1976, 1980) is whether measurements of cyclic strengths and cyclic strain accumulation for dense sands in the laboratory are reasonable approximations of in-situ cyclic strengths. Castro (1975) and Casagrande (1980) expressed the belief that the stress and strain concentrations that develop at the top and bottom caps in a cyclic triaxial test enable the accumulation of strains faster than would develop in-situ. Evidence for this effect was obtained when their spatial measurements of relative densities within laboratory specimens showed a systematic loosening of the soil near the top platen. The effects of stress and strain concentrations in a cyclic triaxial or DSS test can be expected to reduce cyclic strengths by an amount that varies with the relative density of the sand. For loose sands with low SPT blow counts, the triggering of high excess pore pressures is quickly followed by the development of large shear strains; it seems unlikely that the boundary conditions can significantly reduce the strengths of these looser specimens, and thus the cyclic strengths are likely reasonable. For dense sands, the accumulation of shear strains is much more gradual; for this reason, it seems reasonable to suspect that the boundary conditions would have the greatest effect on strain accumulation for dense sands under large cyclic loads. Consequently, these effects would cause the laboratory-based triggering curve to plot lower than the applicable in-situ triggering curve.

Figure 8. Comparison of the Idriss-Boulanger (2004, 2008) and Tokimatsu-Yoshimi (1983) liquefaction triggering correlations with the updated case history database by Idriss and

Boulanger (2010).


Santiago, Chile RESIDUAL SHEAR STRENGTH OF LIQUEFIED SOILS

Procedures for estimating the residual shear strength, Sr, of liquefied cohesionless or nonplastic soils have evolved considerably over the past 25 years and continue to represent an area of significant technical uncertainty. These procedures include those that require laboratory testing of field samples obtained by frozen sampling techniques (e.g., Robertson et al. 2000) or by high-quality tube sampling techniques coupled with approaches for "correcting" the shear strength for the estimated volume changes that occur during sampling and testing (e.g., Castro 1975, Castro and Poulos 1977, Poulos et al. 1985; Castro et al. 1992). Empirical correlations have also been developed based on back-analyses of liquefaction flow slide case histories (e.g., Seed 1987, Davis et al. 1988, Seed and Harder 1990, Ishihara 1993, Wride et al. 1999, Olson and Stark 2002). The case history-based correlations are preferred for practice, in part because they implicitly account for void redistribution (i.e., where the diffusion of excess pore pressures leads to localized loosening of a soil near an impeded drainage boundary) and other in-situ mechanisms of strength loss that laboratory element testing cannot recreate (Seed 1987). Practice often requires estimates of Sr for liquefiable soils spanning a greater range of in-situ relative densities than are represented in the flow slide case history database. For example, the schematic cross-section of an earth embankment dam in Figure 9 shows a potential slope failure passing through the dam shell with a representative (N1)60cs = 20 and the underlying alluvium with a representative (N1)60cs = 10. The case history database of flow slides only contains cases with (N1)60 values less than about 14, such that the shell material in this example is outside the range of the case history data. However, if a liquefaction triggering analysis shows that both zones are expected to liquefy during strong earthquake shaking, then the analysis of post-earthquake stability requires that estimates be made for the shear resistance in both zones. The range of the case history data, as interpreted by Olson and Stark (2002) and incorporating the fines adjustment recommended by Seed (1987), and some recent correlations for Sr/σ'v- are illustrated in Figure 10. This figure is illustrative of the range of conditions covered by the various databases in the literature, although it should be recognized that the individual databases do vary in their estimates of the representative SPT blow counts and residual strengths for each case history. The information in Figure 10 indicates: (1) the (N1)60cs values are limited to less than about 14, and (2) the residual shear

Figure 9. Schematic of an earth dam with two zones in which liquefaction may be triggered by strong earthquake shaking; what residual shear strengths should be assigned to each zone for

the purpose of evaluating post-earthquake stability?


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strength ratios are mostly limited to less than about 0.12, which is also generally less than about 20% of the pre-earthquake drained strength for these types of soils. The four correlations shown in Figure 10 also illustrate the range of approaches used by different investigators. The Olson and Stark (2002) correlation for Sr/σ'v is based on representative (N1)60 values without a fines content correction and is expressly limited to (N1)60 less than or equal to 12 (i.e., close to the limit of the available data). The Idriss and Boulanger (2007) correlation is based on representative, equivalent clean sand (N1)60cs values computed using the fines content correction by Seed (1987), and closely follows the Olson and Stark (2002) correlation in this range of (N1)60cs values. The Kramer and Wang (2011) correlation uses representative (N1)60 values without a fines correction, includes a dependency on confining stress, and spans across the other two correlations. The Yoshimine et al. (1999) correlation (as presented in Wride et al. 1999) is based on the minimum or 20th percentile (N1)60cs value (which is partly why it plots more to the left than the other correlations), uses the Seed (1987) fines correction, and tends to bound the data with a sharply upward-curving relationship. Despite their differences, the four correlations are consistent in indicating that the post-liquefaction shear resistance of a cohesionless soil with a representative (N1)60 value less than about 14 can be expected to be between 5% and 20% of its pre-earthquake drained strength, assuming a drained friction angle ranging from about 28 to 34 degrees. An important challenge faced in practice is deciding how to extrapolate these procedures to (N1)60 or (N1)60cs values greater than 14. Some of the technical issues involved in guiding this extrapolation are discussed below.

Figure 10. Comparison of some SPT-based residual shear strength ratio correlations over the range of available case history data


Santiago, Chile It is first useful to revisit the various definitions that have been used in the literature for the shear strength of liquefied soils. The ultimate shear resistance, or critical state strength, that is measured in undrained laboratory element tests may be denoted as SCS, whereas the quasi-steady state shear resistance, which corresponds to a local minimum in the stress-strain curve from an undrained laboratory element test, may be denoted as SQSS. The SQSS generally coincides with the soil's transition from an incrementally contractive to incrementally dilative response in undrained shearing, which has been termed phase transformation (Ishihara 1993). Residual shear strength, Sr, refers to the shear resistance that a liquefied soil mobilizes in the field, which can be complicated by void redistribution, particle mixing, and other field mechanisms that are not replicated in laboratory element tests. These three "strengths" are fundamentally different from a mechanics standpoint, and thus maintaining a distinction is essential. The undrained stress-strain response of saturated sand for a range of initial relative densities (DR) and consolidation stresses in direct simple shear loading is illustrated by the results in Figure 11 for Fraser river sand by Vaid and Sivathayalan (1996). The normalized shear resistance, τ/σ'vc, for this sand at σ'vc of 50 to 400 kPa are plotted versus DR showing that: (1) τ/σ'vc at phase transformation is relatively independent of σ'vc for σ'vc ranging from 50 to 400 kPa, and (2) τ/σ'vc increases with both increasing DR and increasing shear strain beyond phase transformation. These data are consistent with the trends in

0 20 40 60 80

Relative density, DR (%)

0

0.2

0.4

0.6

Tests with 'vc = 200 kPa

At phase transformation

At 10% shear strain

At 20% shear strain

Tests with 'vc = 50, 100, 200, & 400 kPa

At phase transformation

20% shearstrain

At phasetransformation

Vaid & Sivathayalan (1996), Vaid et al. (1996):Fraser river sand, water pluviated,undrained direct simple shear.

10% shearstrain

100 20 30

Equivalent (N1 )60 corresponding to Cd = 46

5 15 25

Figure 11. Undrained shear resistance ratios mobilized in direct simple shear tests on Fraser river sand at different relative densities, consolidation stresses, and shear strains

(after Vaid and Sivathayalan 1996, Vaid et al. 1996)


Santiago, Chile other laboratory testing programs (e.g., Ishihara 1993, 1996; Yoshimine et al. 1999) in showing that as the sand nears DR of 50-60%, its dilative tendencies at σ'vc less than 400 kPa result in the undrained shear resistance at large strains in simple shear exceeding the drained strength (e.g., τ/σ'v = tanφ' ≈ 0.6). Relationships between undrained shear resistance, relative density, and consolidation stress for sand in laboratory element tests have been converted to correlations with SPT (N1)60 or CPT qc1N values by a number of investigators (e.g., this was the approach used to guide the Yoshimine et al. (1999) correlation in Figure 10). The equivalent (N1)60 can be estimated as (Meyerhof 1957),

( ) ( )2

1 60 d RN C D= (2)

where the parameter Cd depends on the soil characteristics (e.g., Cubrinovski and Ishihara 1999). A typical value for clean sands is Cd = 46 (Idriss and Boulanger 2003), and this value was used to compute the equivalent (N1)60 values listed beneath the laboratory test data shown in Figure 11. In this manner, procedures for estimating undrained shear resistance of saturated sands at σ'v less than about 400 kPa shown the undrained shear resistance increasing sharply as (N1)60 exceeds values of about 12-16. Procedures developed in this manner are based on the assumption that the in-situ void ratio will not increase (or relative density decrease) in a critical zone or layer in response to the transient pore water seepage that occurs as excess pore water pressures dissipate during and after earthquake shaking. The concept of void redistribution was articulated by Whitman (1985), wherein he described situations where pore water seepage driven by earthquake-induced excess pore water pressure gradients could lead to the localized loosening of the liquefied soil, or "void redistribution", which could result in Sr being much lower in the field than would be obtained from laboratory tests of samples at the pre-earthquake void ratio. These situations require the presence of a soil layer of significantly lower permeability overlying the liquefied soil layer, thereby impeding the outward seepage, as illustrated for an infinite slope in Figure 12. The factors affecting void redistribution have been demonstrated through a range of physical modeling studies (e.g., Kokusho 1999, 2000, Kulasingam et al. 2004, Malvick et al. 2008) and analytical modeling studies (e.g., Yang and Elgamal 2002, Naesgaard et al. 2005, Seid-Karbasi and Byrne

Figure 12. Schematic of void redistribution in a confined sand layer due to upward seepage driven by earthquake-induced excess pore water pressure gradients (after Whitman 1985)


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2007). These studies show that as the DR of a sand deposit increases, the potential for void redistribution-induced strength loss decreases sharply because: (1) the consolidating portions of the deposit release less water because the post-liquefaction reconsolidation strains are smaller, and (2) the dilating portions of the deposit can accommodate the influx of more water without localizing because the sand has a greater dilatancy. Two possible scenarios were considered by Idriss and Boulanger (2007) in guiding the extrapolation of Sr/σ'v correlations to (N1)60 values greater than 14, as illustrated in Figure 13. The first scenario is that the field conditions and stratigraphy are such that post-liquefaction re-consolidation will not be impeded by lower-permeability layers, and thus the in-situ void ratios are expected to remain constant or decrease during the dissipation of excess pore water pressures. In this case, the in-situ post-liquefaction shear resistance can be expected to be reasonably represented by the relationships derived from laboratory element tests. This scenario was the basis for the upper curve which bends strongly upward at (N1)60cs values of 15-17 in Figure 13, which Idriss and Boulanger (2007) recommended only for those cases where void redistribution effects are expected to be negligible. The second scenario in Figure 13 is that the field conditions and stratigraphy are such that the post-liquefaction re-consolidation of a liquefied layer will, or could be, impeded by an overlying lower-permeability layer. In this scenario, the potential for void redistribution-induced loosening and strength loss may be judged to represent a serious possibility. In such a situation, the potential increases in void ratio near an impeded drainage boundary would mean that the laboratory element test-based relationships

Figure 13. SPT-based correlation for the residual shear strength ratio of liquefied cohesionless soils with σσσσ'v less than 400 kPa by Idriss & Boulanger (2007)


Santiago, Chile could significantly over-estimate the in-situ shearing resistance. For this scenario, Idriss and Boulanger (2007) recommended a lower relationship which curves more gently upward, eventually approaching drained strengths at an (N1)60cs of about 32; this latter point on the residual strength curve was selected to agree with the limits of the liquefaction triggering curve shown previously in Figure 8, with the expectation that a deposit of sand which was too dense for liquefaction to be triggered would have an undrained strength greater than its drained strength. The two scenarios presented by Idriss and Boulanger (2007) present an opportunity for discussing a few technical issues regarding the development of procedures for estimating Sr. • Can the absence of observations of flow slides in cohesionless soils with (N1)60 or (N1)60cs values

greater than about 14 be taken as evidence that in-situ shear strengths can always be expected to steeply approach drained strengths for (N1)60 or (N1)60cs values greater than 14?

• Are the mechanisms of in-situ void redistribution and particle intermixing sufficiently well understood that we know when it will, or will not, be significant? Can we omit these mechanisms in the development of procedures that will be used in practice?

• Are the case history data sufficiently well constrained or the mechanisms of strength loss sufficiently well understood to determine whether a strength (Sr) model, strength ratio (Sr/σ'v) model, or a hybrid model is the most appropriate?

• The epistemic uncertainties in Sr at high (N1)60 or (N1)60cs values can be accounted for in probabilistic analyses, but which models for Sr should be included and what weights should be used?

These and other questions regarding the estimation of residual shear strength warrant further discussions, and likely will continue to be topics for debate for years to come. In the meantime, further research on understanding and bounding the significance of void redistribution and other mechanisms of strength loss are needed to resolve concerns and advance this aspect of our practice.

CONCLUDING REMARKS Three issues regarding the procedures for evaluating cyclic failure and liquefaction of soils were described and discussed in this paper, with the hope of facilitating discussions during and after this conference. The three issues addressed pertained to the topics of:

(1) liquefaction susceptibility criteria for fine-grained soils; (2) SPT-based liquefaction triggering correlations; and (3) residual strength of liquefied cohesionless soils.

It is hoped that the continued discussion of these topics will facilitate better understanding of alternative view points, foster research that addresses the underlying technical issues, and contribute to improved guidance for engineering practice.

AKNOWLEDGEMENTS The authors appreciate the assistance of Dr. Dan Wilson with the examination of the different liquefaction triggering procedures and case history databases.


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