SdeComparing Liquefaction Evaluation Methods Using Penetration-VS Relationshipse Andrus Rd

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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 UniversityClemson, SC 29634-0911, USA

Ph: (864) 656-0488; Fax: (864) 656-2670; E-mail: [email protected]

*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


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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


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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]):

d vv r g aCSR )'/)(/(65.0 max σσ= (1)

where amax = peak horizontal ground surface acceleration, g = acceleration of gravity, vσ = total

vertical (overburden) stress at the depth in question, v'σ = effective overburden stress at the

same depth, and r d = a shear stress reduction coefficient.

Three methods, or curves, for determining the cyclic resistance ratio, CRR, are shown in


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0.0=α for FC < 5 % (3a)

]/19076.1exp[ 2 FC −=α for 5 % 35 % (3c)

0.1=β for FC < 5 % (4a)

]1000/99.0[ 5.1 FC +=β for 5 % 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, qc1 N , 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, D50, of 0.25-2.0 mm. To apply the curve

to soils with FC > 5 %, Robertson and Wride [10] developed the following correction of qc1 N to

an equivalent clean sand value:

N cccs N c q K q 11 )( = (5)

where (qc1N)cs = equivalent clean sand value of qc1N, and Kc = a correction factor for grain


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%100)]/([ vc s q f F σ−= (9)

where qc = 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 qc, f s,

P a, vσ , and v'σ are all in the same units. The value of n ranges from 0.5 for clean sands to 1.0

for clays [11], and can be approximated through an iterative approach [10].

In Figure 1c, the curve for determining CRR from overburden stress-corrected shear-

wave velocity, V S 1, 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 S 1 can be corrected to an equivalent young, clean soil value by:

111111 )()( S csacsS acsaS V K K V K V == (10)

where (V S 1)csa1 = equivalent young clean soil value of V S 1, (V S 1)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 S 1 values caused by aging. Juang et al. [12] suggested the following relationships for

estimating K cs:


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relationships by plotting values of ( N 1)60cs, (qc1 N )cs and (V S 1)csa1 with the same CRR values. The

implied ( N 1)60cs-(V S 1)csa1, (qc1 N )cs-(V S 1)csa1 and (qc1 N )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 3.6-6.0 m; Bay Bridge

Toll Plaza, SFOBB1 with depth of 10.0-12.8 m; and WPC 2000-344, SC1 with depth of 3.8-6.8

m) are not considered in this paper, because penetration or S V measurements are not consistent

i h h d l d i i 2 3 d 4


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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 S V measurements within the

layer that follow a consistent trend. 3) Penetration test locations are within 6 m of the S V test

locations. 4) At least two S V measurements, and the corresponding test intervals, are within the

uniform layer. 5) Time history records used for S V determination exhibit easy-to-pick shear

wave arrivals. Thus, values of S V 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 S V 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 S 1)cs are calculated using average FC values. Where no FC information is available, an

apparent FC value is calculated using the Ic value and the relationship suggested by Robertson


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CPT resistances are available for 41 of the 45 selected layers. All of the CPT

measurements are from 10-cm2

cones. Values of qc1 N and I c are averaged over the interval of the

selected S V 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

(qc1 N )cs values are 0 % to 77 % higher than values of qc1 N .

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 S 1)cs

data pairs is expressed as:

2])[()( 60111 B

cscsS N BV = (13)

where B1 = 87.7 ± 14.4 (95 % confidence interval) and B2 = 0.253 ± 0.053, with (V S 1)cs in m/s


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studies by Robertson et al. [29] for mainly quartz sands and Hegazy and Mayne [30] for various

sands. Values of R2

and s associated with this regression are 0.544 and 22 m/s, respectively.

The equation developed for 34 (qc1 N )cs-( N 1)60cs data pairs is expressed as:

2])[()( 11601 B

cs N ccs q B N = (15)

where B1 = 0.488 ± 0.468 and B2 = 0.779 ± 0.184 with ( N 1)60cs in blows/0.3 m and (qc1 N )cs is

dimensionless. It should be noted that similar B1 and B2 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, R2 = 0.709 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, D50, and the ratio of CPT tip resistance to energy-

corrected SPT blow count. Their relationships exhibit penetration ratios increasing from about

2.5 at D50 = 0.01 mm to about 5.5-8 at D50 = 1 mm. This increasing trend is not seen in the


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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 VS method provides the more conservative prediction When the


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for the V S curve for the lowest V S 1 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,

(qc1 N )cs, and (V S 1)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 (qc1 N )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 (qc1 N )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]:

−

−+

=

215

1

)(215

18.2

100

)(022.0

11

211

5.7csaS

csaS cs

V

V CRR (16)


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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


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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 S 1)csa1 of 210 m/s. The high limiting upper (V S 1)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 VS based CRR curves defined by Equations 16 17 and 18


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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 (qc1 N )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 (qc1 N )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 VS data are needed from other deposit and soil types to


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Transportation (SCDOT) and the Federal Highway Administration under SCDOT Research

Project No. 623. 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): 1249-1273.

[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


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[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): 1015-1025.

[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 03-854, National Institute of Standards and Technology, Gaithersburg, MD,

2003.

[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, 1999.

[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, 1999.

[8] Toprak, S., and Holzer, T.L. Liquefaction potential index: field assessment. Journal of

Geotechnical and Geoenvironmental Engineering , ASCE, 2003; 129(4): 315-322.

[9] Seed H B Tokimatsu K Harder L F and Chung R M Influence of SPT procedures in


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[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): 580-589.

[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, 1994.

[14] Youd, T.L., and Bennett, M.J. Liquefaction sites, Imperial Valley, California. Journal of

Geotechnical Engineering , ASCE, 1983; 109(3): 440-457.

[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, 1984.

[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): 453 467


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Piedmont: A digital database. U.S. Geological Survey Open-file Report 02-296 , 2002;

http://geopubs.wr.usgs.gov/open-file/of02-296.

[20] WPC. Various unpublished project reports, Wright Padgett Christopher, Inc., Mount

Pleasant, SC, 2000-2001.

[21] Wride (Fear), C.E., Robertson, P.K., Biggar, K.W., Campanella, R.G., Hofman, B.A.,

Hughes, J.M.O., K Ü pper, A., and Woeller, D.J. Interpretation of in situ test results from the

CANLEX sites. Canadian Geotechnical Journal , 2000; 37: 505-529.

[22] Iai, S. Personal communication on sites in Hakodate Port, Japan, 1997.

[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): 115-131.

[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


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[28] Fear, C.E., and Robertson, P.K. Estimating the undrained strength of sand: a theoretical

framework. Canadian Geotechnical Journal , 1995; 32: 859-870.

[29] Robertson, P.K., Woeller, D.J., and Finn, W.D.L. Seismic CPT for evaluating liquefaction

potential. Canadian Geotechnical J ournal, 1992; 29: 686-695.

[30] Hegazy, Y.A., and Mayne, P.W. Statistical correlations between VS and cone penetration

data for different soil types. Proceedings, International Symposium on Cone Penetration

Testing, CPT ’95, Linkoping, Sweden, Swedish Geotechnical Society, 1995; 2: 173-178.

[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): 384-403.

[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; 1249-

1273.

[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): 389-

410


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Liquefaction, Technical Report MCEER-00-0019, Multidisciplinary Center for Earthquake

Engineering Research, Buffalo, NY, 1999; 69-86.

[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, 2000.

[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 196.


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Table 1. Data from Holocene soil deposits with FC < 20 % or I c < 2.25.

Site Name Depth

(m)

USCS

Soil

Type

D50

(mm)

FC a

(%)

V S Test

Type b

V S1cs

(m/s)

(N 1 )60cs I c qc1Ncs Source

California, USA

Bay Bridge, SFOBB1 5.4 - 7.2 SP-SM 0.26 12 CH 152 7 2.15 67 [13]

Bay Bridge, SFOBB1 8.0 - 9.9 SP-SM 0.27 8 CH 151 20 1.90 77 [13]

Bay Farm Island-Dike 3.7 - 5.0 SP-SM 0.23 8 CH 211 53 1.35 321 [13]

Bay Farm Island-Dike 5.0 - 7.8 SP-SM 0.28 12 CH 250 48 2.09 185 [13]

Heber Road, Point Bar 1.8 - 4.2 SM 0.11 18 CH 233 34 2.00 319 [14,15]

Port of Oakland, P007-2 3.0 - 5.1 SP-SM 0.29 7 CH/SCPT 183 22 1.50 173 [13]



Sandholt Road, UC-4 2.1 - 3.5 SP 0.85 2 SCPT 161 15 1.42 188 [16]

Sandholt Road, UC-4 6.3 - 10.1 SP 1.11 3 SCPT 216 43 1.19 332 [16]

State Beach, UC-15 2.0 - 3.8 SP 0.28 2 SCPT 137 7 1.90 67 [16]






Treasure Island, B1-B3 2.2 - 4.0 SP-SM 0.21 7 CH 162 21 1.87 85 [17]

Treasure Island, B1-B3 9.0 - 11.5 SM 0.21 14 CH 183 17 2.11 64 [17]Treasure Island, UM-05 3.3 - 5.7 SP 0.33 4 SCPT 170 14 1.82 79 [18]

Treasure Island, UM-05 5.8 - 8.3 SP-SC 0.33 7 SCPT 188 18 1.88 72 [18]

Treasure Island, UM-06 2.2 - 5.0 SP nac 3 SCPT 175 12 2.10 44 [18]

Treasure Island, UM-06 5.0 - 10.4 SP 1.41 3 SCPT 193 21 1.82 73 [18]

Treasure Island, UM-09 2.7 - 6.3 SP-SC 0.15 11 SCPT 161 9 2.04 68 [18]

USGS Alameda, ALC026 4.0 - 10.0 na na 7d SCPT 233 na 1.73 237 [19]

South Carolina, USA

WPC 2000-344, SC2 6.4 - 10.4 na na 6d SCPT 193 na 1.67 108 [20]

WPC 2000-344, SC3 4.5 - 8.5 na na 6d

SCPT 160 na 1.72 118 [20]WPC 2000-344, SC5A 3.8 - 8.8 SM 0.13 29 SCPT 224 29 1.61 130 [20]

WPC 2000-344, SC5B 3.8 - 10.8 SM na 7d SCPT 210 na 1.77 105 [20]

WPC 2000-344 SC10 7 4 - 10 4 na na 20d SCPT 247 na 2 24 229 [20]


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23/27( V S 1 ) c s

300

Curve implied from

CRR relationships

Corrected SPT Blow Count, ( N 1)60cs

C o r r e c t e d S h e

a r - W a v e V e l o c i t y ,

( V S 1 ) c s

0 10 20 30 40 60100

150

200

250

300

50

(V S 1)cs = 87.7 [( N 1)60cs]0.253

Mean curve:

Curve implied from

CRR relationships

Location

California

Canada

Japan

So. Carolina

Figure 2. Relationships between (V S 1)cs and ( N 1)60cs for uncemented, Holocene sands

Andrus et al. 2003

Location

California

Canada


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Corrected CPT Tip Resistance, (qc1 N )cs

C o r r e c t e d S P T B l o w C o u n t ,

( N 1 ) 6 0 c s

0 50 100 200 3000

10

20

30

60

40

(qc1 N )cs

= 321

= 332

150 250

50

= 319Curve implied from

CRR relationships

( N 1)60cs = 0.488 [(qc1 N )cs]0.779

Mean curve:

Figure 4. Relationships between ( N 1)60cs and (qc1 N )cs for uncemented, Holocene sands

Andrus et al. 2003

Location

California

Canada

Japan

So. Carolina


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Median Grain Size, D50, mm

C o r r e c t e d P e n e t r a t i o n R a t i o , ( q c 1

N ) c s / ( N 1 ) 6 0 c s

0.01 0.1 1 100

2

4

6

12

8

10

Figure 5. Relationship between corrected penetration ratio and median grain size for uncemented, Holocene sands

Andrus et al. 2003

Location

California

Canada

Japan

So. Carolina

(qc1 N )cs/( N 1)60cs = 12.5


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SdeComparing Liquefaction Evaluation Methods Using Penetration-VS Relationshipse Andrus Rd

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