Comparing Liquefaction Evaluation Methods Using Penetration-V S Relationships

Transcription

1 Comparing Liquefaction Evaluation Methods Using Penetration-V S Relationships Ronald D. Andrus,* Paramananthan Piratheepan, 1 Brian S. Ellis, 2 Jianfeng Zhang, and C. Hsein Juang Department of Civil Engineering, Clemson University Clemson, SC , USA Ph: (864) ; Fax: (864) ; randrus@clemson.edu *Corresponding author ABSTRACT Three methods that follow the general format of the Seed-Idriss simplified procedure for evaluating liquefaction resistance of soils are compared in this paper. They are compared by constructing relationships between penetration resistance and small-strain shear-wave velocity (V S ) implied from cyclic resistance ratio (CRR) curves for the three methods, and by plotting penetration-v S data pairs. The penetration-v S data pairs are from 45 Holocene-age sand layers in California, South Carolina, Canada, and Japan. It is shown that the V S -based CRR curve is more conservative than CRR curves based on the Standard Penetration Test (SPT) and Cone Penetration Test (CPT), for the compiled Holocene data. This result agrees with the findings of a recent probability study where the SPT-, CPT-, and V S -based CRR curves were characterized as curves with average probability of liquefaction of 31 %, 50 %, and 26 %, respectively. New SPT- and CPT-based CRR equations are proposed that provide more consistent assessments of liquefaction potential for the Holocene sand layers considered. Key words: Cone Penetration Test, earthquake, liquefaction; in situ tests, probability, shear- wave velocity, Standard Penetration Test. 1 Moved to Leighton and Associates, Inc., Chino, CA , USA 2 Moved to United States Air Force, Minot AFB, ND , USA 1

2 INTRODUCTION The occurrence of liquefaction in soils is often evaluated using the simplified procedure originally proposed by Seed and Idriss [1] based on the Standard Penetration Test (SPT). This procedure has undergone several revisions and updates since it was first proposed in 1971, including the development of methods based on the Cone Penetration Test (CPT), the Becker Penetration Test (BPT), and small-strain shear-wave velocity (V S ) measurements. Youd et al. [2] provide a recent review of the Seed-Idriss simplified procedure and the in situ test methods commonly used to evaluate liquefaction resistance of soils. In situ V S measurements provide a promising alternative to the penetration tests, which may be unreliable in some soils, such as gravelly soils, or may not be feasible at some sites, such as capped landfills. In addition, V S is an engineering property, directly related to small-strain shear modulus, and required for dynamic soil response analyses. On the other hand, some factors that affect V S may not equally affect resistance to liquefaction, which is a medium- to large-strain event. Also, V S testing usually does not produce samples for classification or may not be conducted with sufficient detail to detect thin liquefiable strata. Youd et al. [2] and Andrus et al. [3] provide further discussion on the advantages and disadvantages of the V S - and penetration-based liquefaction evaluation methods. The purpose of this paper is to compare the V S liquefaction evaluation method, or curves, proposed by Andrus and Stokoe [4] and updated in Andrus et al. [3, 5] with the SPT and CPT curves summarized in Youd et al. [2] using relationships between penetration resistance and V S. The approach of using penetration-v S relationships to compare curves was applied earlier by Andrus et al. [6] with data from 25 Holocene-age (< 10,000 years) sands with < 10 % fines (particles < mm). In this paper, the SPT-V S and CPT-V S databases are expanded to include 2

3 20 additional sand data pairs. Regression analyses are performed on the expanded databases and the resulting penetration-v S relationships are used to develop new, more consistent liquefaction evaluation curves. REVIEW OF LIQUEFACTION EVALUATION METHODS The Seed-Idriss simplified procedure for evaluating liquefaction resistance basically involves the calculation of two parameters: 1) the level of cyclic loading on the soil caused by the earthquake, expressed as a cyclic stress ratio; and 2) the resistance of the soil to liquefaction, expressed as a cyclic resistance ratio. The cyclic stress ratio, CSR, at a particular depth in a level soil deposit is calculated from (Seed and Idriss [1]): CSR = 0.65( amax / g)( σ / σ ' ) r (1) where a max = peak horizontal ground surface acceleration, g = acceleration of gravity, v v d σ v = total vertical (overburden) stress at the depth in question, σ ' = effective overburden stress at the v same depth, and r d = a shear stress reduction coefficient. Three methods, or curves, for determining the cyclic resistance ratio, CRR, are shown in Figures 1a, 1b, and 1c. In Figure 1a, the curve for determining CRR from energy- and overburden stress-corrected SPT blow count, (N 1 ) 60, by Seed et al. [9] and modified by Youd et al. [2] is shown. This curve is for earthquakes with moment magnitude, M w, of 7.5 and sands with fines content, FC, < 5 %. To apply the curve to soils with FC > 5 %, I. M. Idriss with the assistance of R. B. Seed developed the following correction of (N 1 ) 60 to an equivalent clean sand value [2]: ( N cs = β N (2) 1 ) 60 α + ( 1) 60 where (N 1 ) 60cs = equivalent clean sand value of (N 1 ) 60, and α and β = coefficients determined using the following relationships: 3

4 α = 0.0 for FC < 5 % (3a) 2 α = exp[ / FC ] for 5 % < FC < 35 % (3b) α = 5.0 for FC > 35 % (3c) β =1.0 for FC < 5 % (4a) 1.5 β = [ FC /1000] for 5 % < FC < 35 % (4b) β =1.2 for FC > 35 % (4c) Equations 3 and 4 are suggested for routine liquefaction resistance calculations [2]. In Figure 1b, the curve for determining CRR from overburden stress-corrected CPT tip resistance, q c1n, by Robertson and Wride [10] is shown. This curve is for earthquakes with M w of 7.5, and sands with FC < 5 % and median grain size, D 50, of mm. To apply the curve to soils with FC > 5 %, Robertson and Wride [10] developed the following correction of q c1n to an equivalent clean sand value: ( = (5) qc1 N ) cs Kcqc1N where (q c1n ) cs = equivalent clean sand value of q c1n, and K c = a correction factor for grain characteristics determined using the following relationships: K =1.0 for I c < 1.64 (6a) c c = c c c c K 0.403I I 21.63I I for I c > 1.64 (6b) where I c = soil behavior type index, defined by: where and 2 I c = [( 3.47 logq) + ( log F) ] (7) c v a a v 2 n Q = [( q σ ) / P ][ P / σ ' ] (8) 0.5 4

5 F = [ f s /( q c σ v )]100% (9) where q c = measured cone tip resistance, f s = measured cone sleeve resistance, P a = a reference stress of 100 kpa (or 1 atm), and n = an exponent that depends on soil type. The values of q c, f s, P a, σ v, and σ ' are all in the same units. The value of n ranges from 0.5 for clean sands to 1.0 v for clays [11], and can be approximated through an iterative approach [10]. In Figure 1c, the curve for determining CRR from overburden stress-corrected shearwave velocity, V S1, by Andrus and Stokoe [4] is shown. This curve is for earthquakes with M w of 7.5 and young, uncemented sands and gravels with FC < 5 %. To apply the curve to soils with FC > 5 % and/or older soils, V S1 can be corrected to an equivalent young, clean soil value by: ( V S1 ) csa1 K a1( VS1) cs = Ka1K csvs1 = (10) where (V S1 ) csa1 = equivalent young clean soil value of V S1, (V S1 ) cs = equivalent clean soil value not corrected for age, K cs = a fines content correction factor, and K a1 = an age factor to correct for high V S1 values caused by aging. Juang et al. [12] suggested the following relationships for estimating K cs : where K =1.0 for FC < 5 % (11a) cs K cs = 1+ ( FC 5) T for 5 % < FC < 35 % (11b) K cs = T for FC > 35 % (11c) ( V S 1 /100) ( VS1 /100) T = + (12) Andrus and Stokoe [4] assumed K a1 = 1.0 for all Holocene-age soils. Because the three CRR curves shown in Figure 1 are all for M w = 7.5 earthquakes and sands with FC < 5 %, they imply relationships between SPT, CPT and V S. One can obtain these 5

6 relationships by plotting values of (N 1 ) 60cs, (q c1n ) cs and (V S1 ) csa1 with the same CRR values. The implied (N 1 ) 60cs -(V S1 ) csa1, (q c1n ) cs -(V S1 ) csa1 and (q c1n ) cs -(N 1 ) 60cs relationships are presented in Figures 2, 3 and 4, respectively. One advantage of studying penetration-v S relationships is they provide comparisons of the liquefaction evaluation methods without needing to calculate CSR. Thus, data from sites not shaken by earthquakes can also be used to validate the consistency between liquefaction evaluation methods. HOLOCENE SAND DATA Data from 45 Holocene-age sand layers with FC < 20 % or I c < 2.25 are also plotted in Figures 2, 3 and 4. The data are summarized in Table 1. They are from California, South Carolina, Canada, and Japan, and are based on measurements performed by various investigators [13-24]. The data were originally compiled by Andrus et al. [6], Piratheepan [25], and Ellis [26]. Three of their compiled Holocene sand data (Coyote Creek with depth of m; Bay Bridge Toll Plaza, SFOBB1 with depth of m; and WPC , SC1 with depth of m) are not considered in this paper, because penetration or V S measurements are not consistent with the data plotted in Figures 2, 3 and 4. The reason for selecting sands with FC < 20 % or I c < 2.25 is so that a significant number of data points are available for regression analysis, while limiting the FC or I c corrections. According to a relationship proposed by Robertson and Wride [10], sands with I c < 2.25 typically have values of FC < 20 %. Average values of D 50 for the sand layers listed in Table 1 range from 0.08 mm to 1.68 mm. These sands classify as SP, SP-SM, SP-SC, and SM by the Unified Soil Classification System. The general criteria used for selecting the penetration and V S measurements are as follows: 1) Measurements are from below the ground-water table where reasonable estimates of 6

7 effective stress can be easily made. 2) Measurements are from thick, uniform soil layers identified primarily using CPT measurements. When no CPT measurements are available, exceptions to Criterion 2 are allowed if there are several SPT and V S measurements within the layer that follow a consistent trend. 3) Penetration test locations are within 6 m of the V S test locations. 4) At least two V S measurements, and the corresponding test intervals, are within the uniform layer. 5) Time history records used for VS determination exhibit easy-to-pick shear wave arrivals. Thus, values of V S determined from difficult-to-pick shear-wave arrivals are not used. When the time history records are not available, exceptions to Criterion 5 are allowed if there are at least 3 V S measurements within the selected layer. The 45 Holocene-age sand layers range in depth from 1.7 m to 13.0 m. Of the 45 selected sand layers, 27 were tested by seismic cone, 7 by crosshole, 3 by both seismic cone and crosshole, 6 by suspension logger, and 2 by downhole techniques. Values of (V S1 ) cs are calculated using average FC values. Where no FC information is available, an apparent FC value is calculated using the I c value and the relationship suggested by Robertson and Wride [10], where FC 1.75Ic for 1.26 < I c < 3.5. Calculated (V S1 ) cs values are 0 % to 7 % higher than values of V S1. SPT blow counts are available for 38 of the 45 selected sand layers. Values of (N 1 ) 60 are determined from measured SPT blow counts using reported test equipment and procedure information. Where no energy measurements are available, average corrections recommended by Youd et al. [2] are assumed based on the type of hammer used. Calculated (N 1 ) 60cs values are 0 % to 76 % higher than values of (N 1 ) 60. 7

8 CPT resistances are available for 41 of the 45 selected layers. All of the CPT measurements are from 10-cm 2 cones. Values of q c1n and I c are averaged over the interval of the selected V S measurements. They are calculated using the electronic CPT data files, when available. When the electronic files are not available, average values are determined from the reported graphical profiles. Because values of I c are not available for the six sand layers in Canada, they are approximated using Robertson and Wride s [10] I c -FC relationship. Calculated (q c1n ) cs values are 0 % to 77 % higher than values of q c1n. REGRESSION ANALYSIS Regression equations are determined for the Holocene sand data from nonlinear regression analysis by power curve fitting. The decision to use power curve fitting is based primarily on results of earlier studies. The regression equation developed for 38 (N 1 ) 60cs -(V S1 ) cs data pairs is expressed as: ( S1 cs cs B2 V ) = B [( N ) ] (13) where B 1 = 87.7 ± 14.4 (95 % confidence interval) and B 2 = ± 0.053, with (V S1 ) cs in m/s and (N 1 ) 60cs in blows/0.3 m. These values of B 1 and B 2 are most similar to values obtained in earlier SPT-V S regression studies by Yoshida et al. [27] for fine sand and Fear and Robertson [28] for Ottawa sand. The coefficient of multiple regression, R 2, and standard deviation of the residuals (or errors), s, associated with this regression are and 19 m/s, respectively. The equation developed for 41 (q c1n ) cs -(V S1 ) cs data pairs is expressed as: ( S1 cs 1 c1n B2 V ) = B [( q ) ] (14) where B 1 = 67.6 ± 20.4 and B 2 = ± 0.063, with (V S1 ) cs in m/s and (q c1n ) cs is dimensionless. These values of B 1 and B 2 are most similar to values obtained in earlier CPT-V S regression cs 8

9 studies by Robertson et al. [29] for mainly quartz sands and Hegazy and Mayne [30] for various sands. Values of R 2 and s associated with this regression are and 22 m/s, respectively. The equation developed for 34 (q c1n ) cs -(N 1 ) 60cs data pairs is expressed as: ( 1 60cs 1 q c 1N B2 N ) = B [( ) ] (15) where B 1 = ± and B 2 = ± with (N 1 ) 60cs in blows/0.3 m and (q c1n ) cs is dimensionless. It should be noted that similar B 1 and B 2 values (0.357 and 0.842, respectively) are obtained when Equations 13 and 14 are set equal to each other and solved for (N 1 ) 60cs, indicating that the three equations are in general agreement. For this regression, R 2 = and s = 7 blows/0.3 m. This high s value of 7 blows/0.3 m associated with Equation 15 is not likely the result of grain size characteristics. Robertson and Campanella [31] and Seed and de Alba [32] developed relationships between median grain size, D 50, and the ratio of CPT tip resistance to energycorrected SPT blow count. Their relationships exhibit penetration ratios increasing from about 2.5 at D 50 = 0.01 mm to about at D 50 = 1 mm. This increasing trend is not seen in the energy-, overbuden-, and fines content-corrected penetration resistances listed in Table 1. Presented in Figure 5 are the ratios of corrected penetration resistances compiled for this study versus corresponding values of D 50. Because there is little or no increasing trend in the plotted (q c1n ) cs /(N 1 ) 60cs values, it appears that the fines content correction accounted for most, if not all, of the effects of grain size characteristics. Equations 13, 14 and 15 are also plotted in Figures 2, 3 and 4, respectively. Although somewhat better fits of the plotted data can be obtained using more complex regression models, these equations appear to be adequate for the comparison of liquefaction evaluation methods. cs 9

10 COMPARISON OF EVALUATION METHODS As explained by Andrus and Stokoe [4], both the SPT and V S evaluation methods provide similar predictions of liquefaction resistance when the data point lies on the implied curve in Figure 2. When the data point plots below the implied curve, the V S method provides the more conservative prediction. When the data point plots above the implied curve, the SPT method provides the more conservative prediction. Because most of the data points plot below the implied curve, the V S method provides an overall more conservative prediction of liquefaction resistance than does the SPT method below (N 1 ) 60cs of 26 for the plotted Holocene sand data. Above (N 1 ) 60cs of 26, both methods appear to provide similar predictions on average. This finding agrees with the probability assessment of Juang et al. [12], where the SPT-based CRR curve (see Figure 1a) and the V S -based CRR curve (see Figure 1c) are characterized with average probability of liquefaction, P L, of 31 % and 26 %, respectively. Both the CPT and V S evaluation methods provide similar predictions of liquefaction resistance when the data point lies on the implied curve in Figure 3. When the data point plots below the implied curve, the V S method provides the more conservative prediction. When the data point plots above the implied curve, the CPT method provides the more conservative prediction. Because the majority of the data points lie below the implied curve, the V S method provides an overall more conservative prediction of liquefaction resistance than does the CPT method for the plotted data. This finding also agrees with the assessment of Juang et al. [12], where the CPT-based CRR curve (see Figure 1b) is characterized with average P L of 50 %. The flatter slope exhibited by the implied curves below (N 1 ) 60cs of 6 (see Figure 2) and (q c1n ) cs of 30 (see Figure 3) can be explained by different assumed minimal values of CRR. A minimum CRR value of 0.05 is assumed for the SPT and CPT curves, whereas is assumed 10

11 for the V S curve for the lowest V S1 value (100 m/s) of most soils with FC < 5 %. More liquefaction/no liquefaction case histories are needed at these lower values of CSR, (N 1 ) 60cs, (q c1n ) cs, and (V S1 ) cs to fully assess these assumptions. Both the CPT and SPT methods provide the same predictions of liquefaction resistance, when the data point lies on the implied curve in Figure 4. When the data point plots below the implied curve, the SPT method provides the more conservative prediction. When the data point plots above the implied curve, the CPT method provides the more conservative prediction. Because more of the data points between (q c1n ) cs of 40 and 120 plot above the implied curve, the CPT method provides more conservative predictions of liquefaction resistance than does the SPT method in this range. Above (q c1n ) cs of 120, the mean curve for the data points plots below the implied curve, indicating the SPT method is more conservative in that range. Liquefaction resistance curves that are consistent, on average, may be obtained using Equations 13 and 14 and the V S -based CRR curve defined by [4]: 2 ( V = S1) csa1 1 1 CRR 7.5cs (16) ( VS1) csa1 215 Substituting Equations 13 and 14 into Equation 16 leads to the following relationships: [ ] [ ] CRR 7.5cs = ( N1) 60cs (17) ( N ) cs CRR 7.5cs = [( qc1n ) cs ] (18) ( [ q ) ] 215 c1n cs Equations 17, 18 and 16 are compared with the original curves in Figures 6a, 6b and 6c, respectively. The ranges where the V S -based CRR curve is more conservative than the SPT- and CPT-based CRR curves can be clearly seen in these figures. 11

12 Because Equation 16 is characterized with P L = 26 % [12], Equations 17 and 18 should also define curves of similar P L. To verify this assumption, results of various probability studies are plotted in Figures 7a, 7b and 7c. In Figure 7a, Equation 17 is compared with six P L = 26 % curves determined from SPT-based liquefaction case histories. The curves by Liao et al. [33], Youd and Noble [34], Toprak et al. [35], and Juang et al. [12] Model 1 are derived from logistic regression analysis. The curves by Cetin et al. [36] and Juang et al. [12] Model 2 are derived from Bayesian analysis. Five of the P L = 26 % curves suggest upper bounds for liquefaction occurrence greater than (N 1 ) 60cs of 30, the value traditionally assumed as the limiting upper bound [9]. These larger upper bound values could be real, or they could be the result of the model assumed. Nevertheless, the agreement is remarkable given the fact that Equation 17 is derived from V S -based liquefaction case histories and the SPT-V S regression equation. In Figure 7b, Equation 18 is compared with three P L = 26 % curves determined from CPT-based liquefaction case histories. The curves by Toprak et al. [35] and Juang et al. [12] Model 1 are derived from logistic regression analysis. The Model 2 curve by Juang et al. [12] is derived from Bayesian analysis. It can be seen that Equation 18 generally agrees with all three curves below (q c1n ) cs of 100. Above (q c1n ) cs of 100, each curve suggests a different limiting upper bound value of (q c1n ) cs for liquefaction occurrence. Equation 18 and the Juang et al. [12] Model 1 curve both suggest upper bounds for liquefaction occurrence greater than (q c1n ) cs of 160, the value traditionally assumed as the limiting upper bound [10]. These results support I. M. Idriss suggestion [2, page 821] that the limiting upper value of 160 be increase by %. Nevertheless, the agreement between Equation 18 and the three P L = 26 % curves is remarkable. In Figure 7c, Equation 16 is compared with three P L = 26 % curves determined by Juang et al. [12]. Model 1 is derived from logistic regression analysis using a model similar in form to 12

13 the logistic model equation assumed in the SPT and CPT probability studies [33-35]. Model 2 in Figure 7c is also derived from logistic regression analysis, but is different from the Model 1 equation by an additional term. Model 3 is the Andrus and Stokoe [4] curve and is characterized as a P L = 26 % curve from Bayesian analysis. It can be seen that all three curves are in general agreement below (V S1 ) csa1 of 210 m/s. The high limiting upper (V S1 ) csa1 value of 235 m/s suggested by Model 1 is believed to be the result of the form of the assumed logistic model equation. RECOMMENDATIONS FOR DESIGN EVALUATIONS The Building Seismic Safety Council (BSSC) [37] suggests a factor of safety of 1.2 to 1.5 is appropriate when applying the SPT-based CRR curve by Seed et al. [9] in engineering design evaluations, where factor of safety, F S, is defined as CRR/CSR Traditionally, liquefaction is predicted to occur when F S < 1; and not occur with F S > 1. Juang et al. [12] characterize the Seed et al. [9] curve as a P L = 31 % curve, and interpret F S values of 1.2 to 1.5 as corresponding to P L of 20 % to 10 %. The SPT-, CPT-, and V S -based CRR curves defined by Equations 16, 17 and 18, respectively, are shown earlier in this paper to be approximately P L = 26 % curves. When applying these equations in engineering practice, the appropriate range of F S values that correspond to the BSSC s [37] suggested range is 1.1 to 1.4 [12]. Greater care should be exercised when applying the V S -based CRR curves to soils older than Holocene age. Preliminary values of K a1 for Pleistocene-age (10,000 to 1.8 million years) sands are given in Andrus and Stokoe [4] and Andrus et al. [3, 5]. These values of K a1 should be used when applying the V S -based CRR curves to Pleistocene sands. Work is under way to develop a continuous relationship between age and K a1, and will be presented in another paper. 13

14 CONCLUSIONS Regression analyses were performed on penetration and V S data pairs from Holocene sands, and the resulting equations were compared with relationships implied by CRR curves for three liquefaction evaluation methods. Based on the comparisons, the following conclusions can be made: 1. For the compiled Holocene sand data, the SPT-based CRR curve [9] between (N 1 ) 60cs values of 8 to 20 was shown to be less conservative, on average, than the V S - and CPT-based CRR curves [4, 10]. The CPT-based CRR curve above a (q c1n ) cs value of about 120 was shown to be less conservative than the SPT- and V S -based CRR curves. These results are in general agreement with a recent probability study [12]. 2. New equations were developed for estimating CRR from (N 1 ) 60cs and (q c1n ) cs by substituting the developed regression equations into the equation defining the V S - based CRR curve. These new equations compared well with P L = 26 % curves developed by various investigators using SPT and CPT liquefaction case histories. 3. More high-quality penetration-v S data are needed from other deposit and soil types to further compare the liquefaction evaluation methods. One advantage of studying penetration-v S relationships is that they provide comparisons of the evaluation methods without needing to calculate CSR. Thus, data from sites not shaken by strong earthquakes, which have been largely ignored in the past, can be used in the comparisons. ACKNOWLEDGMENTS This work was funded in part by the U.S. Geological Survey, Department of the Interior under USGS award number 01HQGR0007; and by the South Carolina Department of 14

15 Transportation (SCDOT) and the Federal Highway Administration under SCDOT Research Project No The views and conclusions contained in this document are those of the authors and should not be interpreted as necessarily representing the official policies, either expressed or implied, of the U.S. Government or the State of South Carolina. The authors acknowledge the insights shared by K. H. Stokoe, II of The University of Texas at Austin during earlier collaborative studies and by T. L. Holzer of USGS during parts of this work. The authors also express their sincere thanks to the many individuals who generously assisted with data compilation. In particular, T. L. Holzer, M. J. Bennett, J. C. Tinsley, III, and T. E. Noce of USGS, S. Iai of the Port and Harbour Research Institute in Japan, R. Boulanger of the University of California at Davis, and T. J. Casey and W. B. Wright of Wright Padgett Christopher. REFERENCES [1] Seed, H.B., and Idriss, I.M. Simplified procedure for evaluating soil liquefaction potential. Journal of the Soil Mechanics and Foundation Division, ASCE, 1971; 97(9): [2] Youd, T.L., Idriss, I.M., Andrus, R.D., Arango, I., Castro, G., Christian, J.T., Dobry, R., Finn, W.D.L., Harder, L.F., Jr., Hynes, M.E., Ishihara, K., Koester, J.P., Liao, S.S.C., Marcuson, W.F., III, Martin, G.R., Mitchell, J.K., Moriwaki, Y., Power, M.S., Robertson, P.K., Seed, R.B., and Stokoe, K.H., II. Liquefaction resistance of soils: summary report from the 1996 NCEER and 1998 NCEER/NSF workshops on evaluation of liquefaction resistance of soils. Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 2001; 127(10): [3] Andrus, R.D., Stokoe, K.H., II, and Juang, C.H. Guide for shear wave-based liquefaction potential evaluation. Earthquake Spectra, EERI, 2004; 20(2): in press. 15

16 [4] Andrus, R.D., and Stokoe, K.H., II. Liquefaction resistance of soils from shear-wave velocity, Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 2000; 126(11): [5] Andrus, R.D., Stokoe, K.H., II, Chung, R.M., and Juang, C.H. Guidelines for evaluating liquefaction resistance using shear wave velocity measurements and simplified procedures. NIST GCR , National Institute of Standards and Technology, Gaithersburg, MD, [6] Andrus, R.D., Stokoe, K.H., II, and Chung, R.M. Draft guidelines for evaluating liquefaction resistance using shear wave velocity measurements and simplified procedures. NISTIR 6277, National Institute of Standards and Technology, Gaithersburg, MD, [7] Chrisley, J.C. Consistency between liquefaction prediction based on SPT, CPT, and V S measurements at the same site. M.S. Report, University of Texas at Austin, [8] Toprak, S., and Holzer, T.L. Liquefaction potential index: field assessment. Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 2003; 129(4): [9] Seed, H.B., Tokimatsu, K., Harder, L.F., and Chung, R.M. Influence of SPT procedures in soil liquefaction resistance evaluations, Journal of Geotechnical Engineering, ASCE, 1985; 111(12): [10] Robertson, P.K., and Wride, C.E.. Evaluating cyclic liquefaction potential using the Cone Penetration Test. Canadian Geotechnical Journal, 1998; 35(3): [11] Olsen, R.S. Cyclic liquefaction based on the cone penetration test. Proceedings, NCEER Workshop on Evaluation of Liquefaction Resistance of Soils, National Center for Earthquake Engineering Research, State University of New York at Buffalo, 1997;

17 [12] Juang, C.H., Jiang, T., and Andrus, R.D. Assessing probability-based methods for liquefaction potential evaluation. Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 2002; 128(7): [13] Mitchell, J.K., Lodge, A.L., Coutinho, R.Q., Kayen, R.E., Seed, R.B., Nishio, S., and Stokoe, K.H., II. Insitu test results from four Loma Prieta earthquake liquefaction sites: SPT, CPT, DMT, and Shear Wave Velocity. Report No. UCB/EERC-09/04, Earthquake Engineering Research Center, University of California at Berkeley, [14] Youd, T.L., and Bennett, M.J. Liquefaction sites, Imperial Valley, California. Journal of Geotechnical Engineering, ASCE, 1983; 109(3): [15] Bierschwale, J.G., and Stokoe, K.H., II. Analytical evaluation of liquefaction potential of sands subjected to the 1981 Westmorland earthquake. Geotechnical Engineering Report GR-84-15, University of Texas at Austin, [16] Boulanger, R.W., Mejia, L.H., and Idriss, I.M. Liquefaction at Moss Landing during Loma Prieta earthquake. Journal of Geotechnical and Geoenvironmental Engineering, ASCE, 2002; 123(5): [17] Fuhriman, M.D. Crosshole seismic tests at two northern California sites affected by the 1989 Loma Prieta earthquake. M.S. Thesis, University of Texas at Austin, [18] Hryciw, R.D. Post Loma Prieta earthquake CPT, DMT and shear wave velocity investigations of liquefaction sites in Santa Cruz and on Treasure Island. Final Report to the U.S. Geological Survey, Award No G1865, University of Michigan at Ann Arbor, [19] Holzer, T.L., Bennett, M.J., Noce, T.E., Padovani, A.C., Tinsley, J.C., III. Liquefaction hazard and shaking amplification maps of Alameda, Berkeley, Emeryville, Oakland, and 17

18 Piedmont: A digital database. U.S. Geological Survey Open-file Report , 2002; [20] WPC. Various unpublished project reports, Wright Padgett Christopher, Inc., Mount Pleasant, SC, [21] Wride (Fear), C.E., Robertson, P.K., Biggar, K.W., Campanella, R.G., Hofman, B.A., Hughes, J.M.O., KÜpper, A., and Woeller, D.J. Interpretation of in situ test results from the CANLEX sites. Canadian Geotechnical Journal, 2000; 37: [22] Iai, S. Personal communication on sites in Hakodate Port, Japan, [23] Iai, S., Morita, T., Kameoka, T., Matsunaga, Y., and Abiko, K. Response of a dense sand deposit during 1993 Kushiro-Oki earthquake. Soils and Foundations, Japanese Society of Soil Mechanics and Foundation Engineering, 1995; 35(1): [24] Ishihara, K., Kokusho, T., Yasuda, S., Goto, Y., Yoshida, N., Hatanaka, M., and Ito, K. Dynamics properties of Masado fill in Kobe Port Island improved through soil compaction method. Summary of Final Report by Geotechnical Research Collaboration Committee on the Hanshin-Awaji Earthquake, Obayashi Corporation, Tokyo, Japan. [25] Piratheepan, P. Estimating shear-wave velocity from SPT and CPT data. M.S. Thesis, Clemson University, Clemson, SC, [26] Ellis, B.S. Regression equations for estimating shear-wave velocity in South Carolina sediments using penetration test data. M.S. Thesis, Clemson University, Clemson, SC, [27] Yoshida, Y., Ikemi, M., and Kokusho, T. Empirical formulas of SPT blow counts for gravelly soils. Penetration Testing 1988, ISOPT-1, Orlando, FL, 1988; 2:

19 [28] Fear, C.E., and Robertson, P.K. Estimating the undrained strength of sand: a theoretical framework. Canadian Geotechnical Journal, 1995; 32: [29] Robertson, P.K., Woeller, D.J., and Finn, W.D.L. Seismic CPT for evaluating liquefaction potential. Canadian Geotechnical Journal, 1992; 29: [30] Hegazy, Y.A., and Mayne, P.W. Statistical correlations between V S and cone penetration data for different soil types. Proceedings, International Symposium on Cone Penetration Testing, CPT 95, Linkoping, Sweden, Swedish Geotechnical Society, 1995; 2: [31] Robertson, P.K., and Campanella, R.G. Liquefaction potential of sands using the CPT. Journal of the Geotechnical Engineering Division, ASCE, 1988; 111(3): [32] Seed, H.B., and de Alba, P. Use of SPT and CPT tests for evaluating the liquefaction resistance of sands. Use of In Situ Tests in Geotechnical Engineering, ASCE, 1986; [33] Liao, S.S.C., Veneziano, D., and Whitman, R.V. Regression model for evaluating liquefaction probability. Journal of Geotechnical Engineering, ASCE, 1988; 114(4): [34] Youd, T.L., and Noble, S.K. Liquefaction criteria based on statistical and probabilisitic analysis. Proceedings of the NCEER Workshop on Evaluation of Liquefaction Resistance of Soils, Technical Report NCEER , National Center for Earthquake Engineering Research, State University of New York at Buffalo, 1997; [35] Toprak, S., Holzer, T.L., Bennett, M.J., and Tinsley, J.C., III. CPT- and SPT-based probabilitistic assessment of liquefaction. Proceedings of the Seventh US-Japan Workshop on Earthquake Resistant Design of Lifeline Facilities and Counter-measures Against 19

20 Liquefaction, Technical Report MCEER , Multidisciplinary Center for Earthquake Engineering Research, Buffalo, NY, 1999; [36] Cetin, K.O., Seed, R.B., and Der Kiureghian, A. Probabilistic assessment of liquefaction initiation hazard. Proceedings of the Twelth World Conference on Earthquake Engineering, Auckland, New Zealand, [37] Building Seismic Safety Council (BSSC). NEHRP Recommended Provisions for Seismic Regulation for New Buildings and Other Structures, FEMA 368, Federal Emergency Management Agency, Washington, DC, 2000; Part 2: page

21 Table 1. Data from Holocene soil deposits with FC < 20 % or I c < Andrus et al Site Name Depth (m) USCS Soil Type D 50 (mm) FC a (%) V S Test Type b V S1cs (m/s) (N 1 ) 60cs I c q c1ncs Source California, USA Bay Bridge, SFOBB SP-SM CH [13] Bay Bridge, SFOBB SP-SM CH [13] Bay Farm Island-Dike SP-SM CH [13] Bay Farm Island-Dike SP-SM CH [13] Heber Road, Point Bar SM CH [14,15] Port of Oakland, P SP-SM CH/SCPT [13] Port of Oakland, P SP-SM CH/SCPT [13] Port of Oakland, P SP-SM CH/SCPT [13] Sandholt Road, UC SP SCPT [16] Sandholt Road, UC SP SCPT [16] State Beach, UC SP SCPT [16] State Beach, UC SP SCPT [16] State Beach, UC SP SCPT [16] State Beach, UC SP SCPT [16] State Beach, UC SP SCPT [16] State Beach, UC SP SCPT [16] Treasure Island, B1-B SP-SM CH [17] Treasure Island, B1-B SM CH [17] Treasure Island, UM SP SCPT [18] Treasure Island, UM SP-SC SCPT [18] Treasure Island, UM SP na c 3 SCPT [18] Treasure Island, UM SP SCPT [18] Treasure Island, UM SP-SC SCPT [18] USGS Alameda, ALC na na 7 d SCPT 233 na [19] South Carolina, USA WPC , SC na na 6 d SCPT 193 na [20] WPC , SC na na 6 d SCPT 160 na [20] WPC , SC5A SM SCPT [20] WPC , SC5B SM na 7 d SCPT 210 na [20] WPC , SC na na 20 d SCPT 247 na [20] WPC , SC na na 6 d SCPT 198 na [20] WPC , SCPT na na 9 d SCPT 253 na [20] Canada Fraser River Delta, Kidd SP 0.20 <5 SCPT <1.64 d 68 [21] Fraser River Delta, Massey SP 0.20 <5 SCPT <1.64 d 53 [21] HVC Mine, LL Dam SP-SM SCPT d 43 [21] HVC Mine, Highmont Dam SP-SM SCPT d 52 [21] Syncrude, J-Pit SM SCPT d 28 [21] Syncrude, Mildred Lake SP-SM SCPT d 87 [21] Japan Hakodate Port No SM SL [22] Hakodate Port No SP-SM SL [22] Hakodate Port No SP-SM SL [22] Hakodate Port No SM SL [22] Kushiro Port, No. 2 (PB-1) SP-SM SL na na [23] Kushiro Port, No. 2 (PB-1) SP-SM SL na na [23] Port Island, Common Factory SP-SM na 6 DH na na [24] Port Island, Common Factory SP-SM na 6 DH na na [24] a FC = fines content (silt and clay) b CH = crosshole; SCPT = seismic CPT; SL = suspension logger; DH = downhole c na = not available d Estimated fines content or I c from: FC = 1.75I c for 1.26 < I c < 3.5 (Robertson and Wride [10]) 21

22 22 Cyclic Resistance Ratio, CRR M w = M w = D 50 = mm Liquefaction Modified Seed et al. (1985) 0.1 No Liquefaction Corrected SPT Blow Count, (N 1 ) 60cs, blows/0.3 m (a) (b) M w = 7.5 (c) Cyclic Resistance Ratio, CRR Liquefaction Robertson & Wride (1998) 0.1 No Liquefaction Cyclic Resistance Ratio, CRR 0.1 No Liquefaction Corrected CPT Tip Resistance, Corrected Shear-Wave Velocity, (q c1n ) cs (V S1 ) csa1, m/s Liquefaction Andrus & Stokoe (2000) Figure 1. Liquefaction resistance curves based on SPT by Seed et al. (1985), CPT by Robertson and Wride (1998), and V S by Andrus and Stokoe (2000) Andrus et al. 2003

23 Corrected Shear-Wave Velocity, (V S1 ) cs Curve implied from CRR relationships Mean curve: (V S1 ) cs = 87.7 [(N 1 ) 60cs ] Location 150 California Canada Japan So. Carolina Corrected SPT Blow Count, (N 1 ) 60cs Figure 2. Relationships between (V S1 ) cs and (N 1 ) 60cs for uncemented, Holocene sands Corrected Shear-Wave Velocity, (V S1 ) cs Location California Canada Japan So. Carolina Curve implied from CRR relationships Mean curve: (V S1 ) cs = 67.6 [(q c1n ) cs ] (q c1n ) cs = 319 = 332 = Corrected CPT Tip Resistance, (q c1n ) cs Figure 3. Relationships between (V S1 ) cs and (q c1n ) cs for uncemented, Holocene sands 23

24 Corrected SPT Blow Count, (N 1 ) 60cs Location California Canada Japan So. Carolina Curve implied from CRR relationships (q c1n ) cs = 321 = 332 = 319 Mean curve: (N 1 ) 60cs = [(q c1n ) cs ] Corrected CPT Tip Resistance, (q c1n ) cs Figure 4. Relationships between (N 1 ) 60cs and (q c1n ) cs for uncemented, Holocene sands 24

25 Corrected Penetration Ratio, (q c1n ) cs /(N 1 ) 60cs Location California Canada Japan So. Carolina (q c1n ) cs /(N 1 ) 60cs = Median Grain Size, D 50, mm Figure 5. Relationship between corrected penetration ratio and median grain size for uncemented, Holocene sands 25

26 26 Cyclic Resistance Ratio, CRR M w = M w = D 50 = mm Liquefaction Liquefaction 0.4 Modified Seed et al. (1985) Eq. 17 No Liquefaction Corrected SPT Blow Count, (N 1 ) 60cs, blows/0.3 m (a) (b) M w = 7.5 (c) Cyclic Resistance Ratio, CRR Robertson & Wride (1998) Eq. 18 No Liquefaction Cyclic Resistance Ratio, CRR Liquefaction Eq. 16 Andrus & Stokoe (2000) No Liquefaction Corrected CPT Tip Resistance, Corrected Shear-Wave Velocity, (q c1n ) cs (V S1 ) csa1, m/s Figure 6. Comparison of liquefaction resistance curves by Seed et al. (1985), Robertson and Wride (1998), and Andrus and Stokoe (2000) with curves derived from penetration-v S equations Andrus et al. 2003

27 27 Cyclic Resistance Ratio, CRR M w = 7.5 P L = 0.26 Liao et al. (1988) Cetin et al. (2000) Youd & Noble (1997) Juang et al. (2002) Model 1 Toprak et al. (1999) Juang et al. 0.1 (2002) Eq. 17 Model Corrected SPT Blow Count, (N 1 ) 60cs, blows/0.3 m (a) M w = 7.5 (b) M w = 7.5 (c) P L = 0.26 Juang P L = Juang et al. et al. (2002) (2002) Model Model 2 Cyclic Resistance Ratio, CRR Toprak et al. (1999) Eq. 18 Juang et al. (2002) Model Cyclic Resistance Ratio, CRR Eq. 16, Andrus & Stokoe (2000); Juang et al. (2002) Model 3 Juang et al. (2002) Model Corrected CPT Tip Resistance, Corrected Shear-Wave Velocity, (q c1n ) cs (V S1 ) csa1, m/s Figure 7. Comparison of liquefaction resistance curves derived from the CRR curve by Andrus and Stokoe (2000) and penetration-v S equations with P L = 26 % curves developed by various invesitgators Andrus et al. 2003