ML19209C859
| ML19209C859 | |
| Person / Time | |
|---|---|
| Site: | La Crosse File:Dairyland Power Cooperative icon.png |
| Issue date: | 09/28/1979 |
| From: | DAMES & MOORE |
| To: | |
| References | |
| TASK-02-04, TASK-2-4, TASK-RR 11166-003-27, 11166-3-27, NUDOCS 7910180382 | |
| Download: ML19209C859 (150) | |
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I I
I LIQUEFACTION POTENTIAL AT LA CROSSE BOILING WATER REACTOR (LACBWR) SITE NEAR GEN 0A, VERNON COUNTY, WISCONSIN I
I Prepared by DAMES & MOORE 7101 Wisconsin Avenue Washington, D.C.
20014 I
I Prepared for I
Dairyland Power Cooperative La Crosse, Wisconsin 54601 Contract Number 11166-003-27 September 28, 1979 I
l'73 010 I
^
~
September 28,1979 Lacrosse Boiling Water Reactor Dairyland Power Cooperative Post Office Box 135 Genoa, Wisconsin 54632 Attention:
Mr. R.E. Shimshak Plant Superintendent Gentlemen:
We submit herewith six copies of our r3 port, " Liquefaction Potential at La Crosse Boiling Water Reactor (LACBWR) Site, Near Genoa, Vernon County, Wisconsin" for your use.
This report includes:
a) Brief summaries of all previous liquefaction analyses performed at LACBWR Site and relatad background studies; b) Details of field, laboratory and analytical investiga-tions that were performed to verify the earlier find-ings regarding liquefaction potential at LACBWR Site; and 4
c) Our conclusions based on analyses and testing performed on undisturbed samples obtained by utilizing state-of-the-art techr.' ques.
We have concluded in our study that a threshold liquefaction resistance level for the LACBWR site corresponds to an SSE producing an acceleratio between 0.18 g and 0.20 g at the ground surface.
The scope of services for this report was prepared by us after discussions with Mr. Richard Shimshak of Dairyland Power Co-operative. Three draft copies of this report were submitted to you on August 10, 1979, for your review and comment.
n;
I ranraes e raconc I
Dairyland Power Cooperative I
September 26, 1979 Page Two E
It has been a pleasure to work on this very interesting I
and challenging project. We look forward to a continued association with Dairv'=nd Power Cooperative in any of their ventures involving geotechnia and environmental studies.
I I
Very truly yours, DAMES & MOORE
(-fer)had Harcharan Singh, Ph.D.
I Partner t
/
ysore Nataraja, Pi.D., P.E.
Project Engineer HS/MN:amc Enclosures I
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l'173 012 I
CONTENTS Page List of Figures.........................
iii List of Tables iii I
1.0 INTRODUCTION
1 1.1 General 1
1.2 Purpose and Scope
2 2.0
SUMMARY
OF DAMES & MOORE GE0 TECHNICAL INVESTIGATION I
0F 1973 3
2.1 Geology 3
2.2 Seismology.........
4 2.3 Liquefaction Patential...................
4 3.0
SUMMARY
OF WES uEPORT OF 1978 6
3.1 Background.........................
6 3.2 Scope and Purpose of Repu t 6
3.3 Conclusions by WES.....................
6 3.4 Summary 7
4.0 BRIEF REVIEW AND DISCUSSION OF WES REPORT 8
5.0 REEVALUATION OF DAMES & MOORE REPORT OF 1973........
9 6.0 DAMES & MOORE RECOMMENDATIONS OF MARCH 1979 10 7.0 NRC/WES COMMENTS ON DAMES & MOORE RECOMMENDATIONS 11 8.0 TEST BORING PROGRAM 12 8.1 General 12 8.2 Drilling and Sampling Procedures..............
12 8.2.1 Standard Penetration Tests................
12 8.2.2 Undisturbed Sampling...................
14 8.3 Handling of Undisturbed Samples 16 9.0 LABORATORY TESTING PROGRAM.................
17 9.1 General 17 9.2 Specific Gravity......................
17 9.3 Particle Size Analyses...................
17 9.4 Minimum and Maximum Densities 17 9.5 Dry Density of Undisturbed Samples.............
21 I
9.6 Cyclic Triaxial Tests 21 9.6.1 Sample Preparation....................
21 9.6.2 Testing 22 1173 013 I
i
CONTENTS (cont'd)
Page 10.0 LIQUEFACTION ANALYSES...................
26 10.1 General..........................
26 I
10.2 Liquefaction Potential 26 10.3 Evaluation of Liquefaction Pote:1tial, Approach 1 27 10.3.1 Simplified Procedure 27 10.3.2 Japanese Procedure 35 10.4 Evaluation of Liquefaction Potential, Approach 2 46 10.4.1 Soil Model Used in the Responde Analysis 46 10.4.2 Soil Properties Used in the Response Analysis.,....
46 10.4.3 Decip Earthquake Used in the Response Analysis.....
48 10.4.4 One-Dimensional Wave Propagation Analysis........
52 I
10.4.5 Cyclic Shear Strength..................
52 10.4.6 Conversion of Irregular Stress History Into Equivalent Uniform Cyclic Stress Series 52 lt.4.7 Correction Factor, C 59 r..................
10.4.8 Factor of Safety Computation 60 10.4.9 Discussion and Conclusions 60 11.0
SUMMARY
OF LIQUEFACTION ANALYSES AT THE LACBWR SITE....
65 REFERENCES 66 APPEi4 DIX: BORING LOGS I
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I 1173 014 ii I
I TABLES Number Page 1
Particle Size Characteristics..............
20 2
Summary of Cyclic Triaxial Test Results....
24 E
3 Sumary of Liquefaction Analysis, Approach 1, Procedure 1 36 4
Summary of Liquefaction Analysis, Approach 1, Procedure 2 44 5
Generalized Soil Profile and Model for One-D,mensional Wave I
Propagation Analysis 47 6
Sumary of Liquefaction Analysis, Approach 2 62 FIGURES 1
Plot Plan.
13 2
Variation of SPT N-Values with Depth 15 3
Particle Size Analysis--Range for Sands Encountered at Site 18 4
Particle Size Analysis--Range for Gravels Encountered at Site 19 5
Variation of Dry Density with Depth.
25 I
6 Correla'. ion Between Field Liquefaction Behavior of Sands for Level Ground Conditions and Penetration Resistance g
( max = 0.10 g)....................
29 7
Correlation Between Field Liquefaction ';ehavior of Sands for Level Ground Conditions and Penetratior. Resistance (a
= 0.12 g) 30 max 8
Correlation Between Field Liquefaction Behavior of Sands for Level Ground Conditions and Penetration Resistance (a
= 0.14 g)....................
31 mu I
9 Correlati. 3etween Field Liquefaction Behavior of Sands for Level Ground Conditions and Penetration Resistance (a* * =
0.16 g)....................
32 10 Correlation Between Field Liquef actie Behavior of Sands for Level Ground Conditions and Penetra.1on desistance 33 c a,x = 0.1s 9).............. q g. 01 5 m
iii
FIGURES (Cont'd)
Number Page 11 Correlation Between Field Liquefaction Behavior of Sands for Level Ground Conditions and Penetration Resistance I
(a
= 0.20 g)....................
34 max 12 Comparison of Cyclic Shear Stress and Strength-Approach 1, I
Procedure 1 (a
= 0.10 g)............
37 max I
13 Comparison of Cyclic Shear Stress and Strength-Approach 1, Procedure 1 (a
= 0.12 g)..............
38 max I
14 Comparison of Cyclic Shear Stress and Strength-Approach 1, Procedure 1 (a
= 0.14 c)..............
39 max 15 Comparison of Cyclic Shear Stress and Strength-Approact 1, Procedure 1 (a
= 0.16 g)..............
40 max l
16 Comparison of Cyclic Shear St.2ss and Strength-Approach 1, Procedure 1 (a
= 0.1 g)..............
41 max 17 Comparison of Cyclic Shear Stress and Strength-Approach 1, Procedure 1 (a
= 0.20 g)..............
42 mu 18 Comparison of Cyclic Shear Stress and Strength-Approach 1, Procedure 2 (a
= 0.10 g to 0.20 g).........
45 mu 19 Typical Reduction of Shear Modulus with Shear Strain 49 20 Damping Ratio for Saturated Sands............
50 21 SSE Horizontal Component (digitized time history )....
51 I
22 Comparison of Cyclic Shear Stress and Strength-Ap? roach 2 (a
= 0.10 g to 0.20 g)...............
53 I
max 23 Cyclic Shear Stress Ratio vs Number of Cycles for Initial Liquefaction, Test No. 1, Depth 16 to 1) Feet.
54 24 Cyclic Shear Stress Ratio vs Number of Cycles for Initial Liquefaction, Test No. 2, Depth 31 to 37 Feet.
55 25 Cyclic Shear 3 tress Ratio vs Number of Cycles for Initial Liquefaction, Test No. 3, Depth 41 to 52 Feet.
56 I
} } } ]'} b iv I
I FIGURES (Cont'd)
Number Page I
26 Cyclic Shear Stress Ratio vs Number of Cycles for Initial Liquefaction, Test No. 4, Depth 87 to 92 Feet......
57 I
27 Sunmary Curve Showing Effects of Density and Soil Fabric on Number of Cycles to Initial Liquefaction.......
58 I
28 Relation Between Cyclic Shear Strength mi Confining Pressure Based On Cyc l i c Tr i ax i a l Te s t s..................
61 I
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1173 017 I
I 1.0 IN1RODUCTION 1.1 General In 1973, Dames & Moore (D&M) performed a Geotechnical Investigation of Geology, Seismology, and Liquefaction Potential at the Lacrosse Boiling Water Reactor (LACBWR) si 3 (Ref. 1).
This study a s conducted for Gulf United Nuclear Fuels Corporation.
D&M's report was submitted I
to the U.S. Nuclear Regulatory Conmission (NRC) in 1974, as part of the application for an operating license for the LACBWR plant (Ref. 2).
In the study, D&M concluded that the LACBWR plant had adequate factors of safety against potential for liquefaction under the design Safe Shutdown Earthquake (SSE) of.12 g.
NRC initiated a review of the LACBWR site and plant under its Systematic Evaluation Program (SEP) in 1978.
As a part of SEP, the U.S. Arry Engineer Waterways Experiment Station (WES) was regt.ested by NRC to review the 1973 D&M soils investigation.
After reviewing the data and analyses presented by D&M, WES performed its own analyses based on interpretations of the same data. The WES report submitted to NRC, and made public in 1978 (Ref. 3), concluded that the factors of safety against liquefaction potential were considerably lower than I
those calculated by D&M.
Upon request of the Dairyland Power Cooperative (DPC), D&M reviewed I
the WES report and reevaluated its 1973 report in view of the WES analyses.
Based on this effort, D&M presented to NRC a position which was essentially consistent with its 1973 study.
It was decided during the meeting with NRC on February 9, 1979, that a written report should be prepared summarizing the meeting, the reviews made, and the various analyses on liquefaction potential for the LACBWR site.
Accordingly, a report (Ref. 4) was submitted to NRC in which D&M rei;erated its earlier stand that the LACBWR site had adequate factors of safety against potential for liquefaction under the design SSE.
However, certain questions I
raiad by NRC regarding the lack of test data on undisturbed samples and the lack of continuous standard penetration test results could not be satisfactorily answered with the existing data.
Therefore, DPC agreed to perform modest field and laboratory investigations and limited analyses to verify the earlier findings on liquefaction potential.
un M8 3
In its March 1979 report (Ref. 4), D&M recomended a modest program consisting of a minimum of four test borings, undisturbed sampling, and cyclic triaxial testing and analyses.
After review of the D&H report, NRC approved the proposed gectechnical program and suggested minor modifica-tions.
I
1.2 Purpose and Scope
The Jurpose of this report is to sumarize all of the liquefaction I
analyses performed at the LACBWR site--(a) D&M (1973), (Ref.1); (b) WES (1978), (Ref. 3); (c) D&M (1979), (Ref. 4); and (d) NRC/WES (April 1979),
(Ref. 5).
Additionally, new analytical investigations have been conducted to verify prior findings on liquefaction potential.
The report is organ-ized as detailed below:
a.
Brief sumary of the Dames & Moore soils investigation of 1973 (Section 2.0).
b.
Brief sumary of the WES analysis of 1978 (Section 3.0).
g c.
Brief review of the WES report and discussions on the approach g
taken by NRC (represented by WES), (Section 4.0).
d.
Summary of the reevaluation of the report of 1973 (Sectior. 5.0).
I e.
Conclusions and recommendations for further work presented by D&M in March 1979 (Section 6.0).
f.
Review coments by NRC/WES or. D&M conclusions and recomendations I
(Secticn 7.0).
g.
Details of the current fie", laboratcry, and analytic'al investi-gations performed by D&M to verify earlier findings on liquefac-I tion analyses at the LACBWR site (Sections 8.0, 9.0, and 10.0).
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1173 c19 -
2.0
SUMMARY
OF DAMES & MOORE GEOTECHNICAL INVESTIGATION OF 1973 Two st 2ies were performed by Dames & Moore i7 1973--a study of ge-ology and engineering seismology, and an investigation of static and dy-namic soil properties and evaluation of liquefaction potential. A report containing the results of these studies was prepared in 1973 (Ref. 1) and was presented to NRC as a part of the application for an operating license for tM LACBWR plant.
The conclusions of this report are discussed in Sectic,is 2.1, 2.2, and 2.3.
9 Geology LACBWR is situated within the Central Stable Region of the North B
American continent. This region includes the dense igneous and meta-morphic rocks of the Canadian Shield and adjacent early Paleozoic sedi-mentary strata.
The geologic structure of the Central Stable Region is relatively simple.
Other than uplift and subsiderce, very little struc-tural activity hn occurred in tl.is quiescent area since Proterozoic time.
The region is characterized by a system of broad, circular-to-ellip-tical erosional uplifts--the Wisconsin and Ozark Domes, and three sedi-mentary basins--the Forest City, Michigan, and Illinois Basins.
The site is located on the western flank of the Wisconsin Arch, a southern extension of the Wisconsin Dome.
Minor structures, consisting primarily of synclines and anticlines of low relief, show no preferred orientation.
They are superimposed on the broader features in the region.
Faults in the region are believed to have I
been dormant since late Paleozoic time, i.e., for at least 200 million years.
The Paleozoic strata and overlying unconsolidated sediments are essentially undeformed within about 50 miles of the site.
LACBWR is located within the Wisconsin Driftless section of the Cent-I ral Lowland physiographic province.
This section is characterized by flat-lying, maturely dissected sedimentary rocks of early Paleozoic age.
Moderate-to-strong relief has been produced on the unglaciated landscape which has been modified only slightly by a mantle of loess and glacial outwash in the larger stream valleys of the area.
I 1173 020 I
3
The LACBWR facilities are situated on about 20 feet of hydraulic fill overlying 100 to 130 Teet of glacial outwash and fluvial deposits on the east flood plain of the Mississippi River Valley.
The surface configura-tion of the underlying bedrock is unknown because of the relative paucity of borehole data. The bedrock below the site consists of nearly flat-lying sandstone and shales of the Dresbach Group (Upper Cambrian).
Dense Pre-cambrian crystalline rock underlying these sedimentary rocks is estimated to be at a depth of 650 feet.
2.2 Seismology Based on the seismic history and the tectonics of the region, D&M concluded that the site will not experience any significant earthquake-induced ground motion during the remaining economic life of the nuclear facility.
Historically, there is no basis for expecting ground motion of more than a few percent of gravity at the LAC 8WR plant site.
- However, three possible sources of earthquake motion at the site were considered:
The nearest zone of repeated earthquake activity, which is in a.
northern Illinois-southern Wisconsin.
~
b.
The effect of a series of events such as those which occurred.n 1811-1812 near New Madrid, Missouri, c
The effect of several shocks in the region which have not been related to any currently identifiable geologic structure or tec-tonic feature.
Y Af ter a careful evaluation of these possible sources of earthquake motion and their possible effect on the LACBWR site, it was concluded that the SSE should be considered as the occurrence at an MM Intensity VI shock with its epicenter close to the site.
It was estimated that the maximum horizontal ground acceleration induced by such an event would be 12 per-cent of gravity at the ground surface.
2.3 Liquefaction Potential The liquefaction potential of the granular soils underlying the ex-isting plant was analyzed by comparing the anticipated shear stresses due
~
to the SSE with the shear stresses required to produce liquefaction at various depths.
The analysis was confined to the upper strata (from the ground surface to a depth of 100 feet) of the zone of potential liquefac-tion.
To provide pertinent subsurface data for this analysis, a field exploration and laboratory program of index properties tests and dynamic tests was conducted.
4 1173 021
--m
I The factors of safety against liquefaction were calculated for vari-ous depths. The calculations were based on 10 significant stress cycles, l
following the engineering practice of 1973.
The results of the analysis indicated that the calculated minimum factor of safety against liquefac-tion under the SSE w1s 1.47.
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I173 022' I
s
I 3.0
SUMMARY
OF WES REPORT OF 1978
3.1 Background
The NRC requested that WES review the foundation conditions at the LACBWR site and prepare a report (Ref. 3) specifically examining the earthquake safety of the pile foundation which supports the contain-ment vessel.
3.2 Scope and Purp6se of Report The scope of WES's report included the following:
a.
Review of Chapter 3, Soil Engineering Properties, in the DPC's Application for Operating License for the Lacrosse Boiling Water I
Reactor (Ref. 2), including portions of Appendix A, entitled
" Field Exploration and Laboratory Tests," and associated design drawings.
b.
Performance of a liquefaction analysis using the Seed-Idriss Sim-plified Procedure (Ref. 6), assuming peak ground surface accelera-ti "S
'*9 "d
9' N
3 c.
Performance of a liquefaction analysis using Seed's empirical method (Ref. 7), assuming both the 0.12-g and 0.20-g earthquakes, and comparison with a " rule of thumb" based on the Japanese ex-I perience at Niigata in 1964 (Ref. 8).
3.3 Conclusions by WES The liquefaction potential was evaluated for two earthquakes--an SSE with a peak ground acceleration of 0.12 g and an SSE with a peak ground acceleration of 0.20 g.
Two methods were employed in the analysis--
the Seed-Idriss Simplified Procedure and an empirical procedure. Also, a Japanese " rule of thumb" based on blowcounts from standard penetration tests was used to predict liquefaction. The fo: lowing were the conclu-
. ions:
a.
Liquef action was predicted between depths of 32 and 48 feet by Seed-Idriss calculations for 0.12 g ground acceleration.
b.
Liquefaction was predicted between depths of 24 and 35 feet by the empirical procedure for 0.12 g ground acceleration.
c.
Liquefaction was predicted below a depth of 25 feet by Seed-I Idriss calculations for 0.20 g ground acceleration.
d.
Liquefaction was predicted between depths of 25 and 60 feet 3
and 85 and 105 feet by the empirical procedure for 0.20-g 5
ground acceleration.
e.
Japanese experience, based on the Niigata earhquake of 1964, also indicated liquefaction potential below a depth of 15 feet (for both cases of 0.12 g and 0.20 g).
j j7} {}
6
I f.
If lateral support was lost at the depths indicated abt
, piles would be in danger of ful lure due to buckling.
3.4 Summary Based on judgements concerning the density and strength data and on analyses presented in the WES report, the soils below the reactor at the LACBWR site were predicted to strain " badly" under an SSE which produces 0.12 g acceleration ct the ground surface. The soils beneath the reactor vessel at the site were predicteJ to experience excessive strains and liquefaction under an SSE with a peak acceleration at the ground surface I
of 0.20 g.
According to the WES report, because of limitations and the limited data available, it was conclrded that the reactor vessel foundation was unsafe under the 0.20 g SSE, iut no conclusion was reached on whether the reactor vessel foundation was safe under the 0.12 g SSE.
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~1173 024 7
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I 4.0 BRIEF REVIEW AND JISCUSSION OF WES REPORT The WES report, Liquefaction Analysis for Lacrosse Nuclear Power Sta-tion (Ref. 3), now a oublic document, was discussed for the first time at a meeting with NRC cn January 9, 1979.
DPC took exception to the contents of I
the WES report and requested that NRC arrange another meeting for discus-sion of the recort after D&M had an opportunity to review it. As a result, a second meeting was held at NRC on February 9, 1979, during which D&M presented its review cf the report to NRC.
Dr. W. F. Marcuson, the principal author of the WES report, was present at the meeting.
The details of the February 9, 1979, meeting are precented in Ref. f.
The following review conments were expressed by D&M and DPC:
a.
In generai, WES adopted a very conservative approach in inter-preting the available data.
I b.
WES postulated an earthquake of MM Intensity IX as a design SSE for the LACBWR site; this was considered unrealistic.
I c.
The report consistently reflected conservatism in selection of soil parameters, selection of cyclic shear stress ratio, and selection of stress reduction factor, which resulted in a cumulative underestimation of safety factors.
d.
WES performed empirical analysis based on standard penetration results and compared the LACBWR site with sites which have experienced much higher seismic activity.
e.
A Japanese " rule of thumb" developed afi.er the experiences of the 1964 Niigata earthqueke and based on standard penetration I
test results was applied to the LACBWR site; such a direct application was considered inappropriate.
In summary, D&M felt that the conservative approach taken by WES in each individual step of the analysis resulted in low factors of safety against liquefaction.
I I
I I
1173 02s I
I 5.0 REEVALUATION OF DAMES & MOORE REPORT OF 1973 I
After the January 9, 1979, meeting with NRC, D&M reevaluated its 1973 report (Ref.1) i' light of the conments and concerns raise / in the 1978 WES report (Rcf. 3).
In general, the D&M Tpproach was found to be consis-I tent with the state-of-the-art in 1973.
Ine obvious limitation of the 1973 study was the lack of liquefaction test data on " undisturbed" samples.
This limitation was indeed realized in the D&M analysis of 1973 and, therefore, a conservative approach was followed.
Two possible mod'fications to the D&M analysis of 1973 were consid-ered:
a.
Redrawing of the strength curves based on densities, rather than relative de.1sities, in a manner similar to the procedure used in the WES report of 1978.
I b.
Select'ng the design shear stress ratio corresponding to five equivalent cycles to represent more realistically the postulated design SSE.
Based on these modifications, factors of safety against liquefaction were recomputed and found to be essentially similar to those cited by D&M in I
1973.
I I
I I
I I
I 1173 026 I
I 9
6.0 DAMES & MOORE RECOMMENDATIONS OF MARCH 1979 The 1973 data and the analyses indicated that the factors of safety against liquefaction under the design SSE were adequate at the LACBWR site.
However, two basic issues needed to be addressed to further I
verify the results obtained--better definition of in situ densities at the LACBWR site, and development of continuous standard penetration test data at the site.
Also, it was necessary to have better estimates of cyclic shear stresses that would result from the design SSE, and better estimates of shear strength data from test results on relatively undisturbed samples.
To address the above concerns, D&M recommended a testing and analysis program for the LACBWR site, consisting of a test boring program with a minimum of four borings, a modest laboratory testing program, and modest analyses.
Details on the procedures and the actual program are discussed in Ref. 4.
I I
I I
I I
I I
lo 1173 027 I
I
I 7.0 NRC/WES COMMENTS ON DAMES & MOORE RECOMMENDATIONS The March 1979 D&M recommendations (Ref. 4) were reviewed b2 ;.he authors of the WES report and NRC staff and approved by NRC subjs t to WES comments.
The following are the main points of the review as outlined in Ref. 5:
I a.
The, rogram outlined by DSM is acceptable for determining the potential for liquefaction in the immediate vicinity of the containmer.t building of the LACPWR plant.
I b.
Additional borings may be required near the turbine building and the cribhouse.
c.
State-of-the-art techniques should be employed to obtain undis-turbed samples in cohesionless soils.
d.
Conmercial transportation of samples should be avoided.
I e.
A sufficient number of cyclic triaxial tests should be performed to cover all the deptha and confining pressures of interest and to obtain a good definitior. of strength at different con-fining pressures and at different significant stress cycles.
I f.
Analyses should be performed to cover a range of assumed peak grocnj acceleration levels between 0.12 g and 0.20 g, so that a threshold liquefaction resistance level can be estimated I
for the LACBWR plant site.
The remaining sections of this report describe the actual test boring program, the laboratory testing program, and the analyses performed to estimate a threshold liquefaction resistance level for the LACBWR plant site.
I I
I I
1173 023 I
11
I 8.0 TEST BORING PROGRAM 8.1 General Five additional borings were drilled at the LACBWR site to obtain more complete data for verification of the earlier liquefaction analyses.
Two borings were drilled on eac.n side of the reactor building near 1973 borings DM-1 and DM-3, and a fifth hole was drilled near the cribhouse I
naar boring DM-5.
Approximate locations of the new borings (DM-7 through DM-11) are shown along with the earlier borings on Figure 1.
Detailed descriptions of the soils that were encountered are presented on boring logs in the Appendix.
Three of the borings (DM-8, DM-10, and DM-11) provided standard penetration test (SPT) blow counts at 5-foot intervals throughout the depth of the holes; continuous blow counts had been precluded in previous borings by the use of several types of samplers in each hole. Borings DM-7* and DM-9 yielded relatively undisturbed samples suitable for density determinations and laboratory cyclic triaxial strength testing.
The split-spoon samples from the SPT holes were used for field classifica-I tion and laboratory confirmation of index prcperties.
8.2 Drilling and Sampling Procedures Drilling operations were performed by Raymond International of Chicago, using a Mobile B-61 truck-mounted rotary wash drill rig.
I The rig was leveled before beginning each hole to ensure vertical drilling.
Drilling to the specified sampling depths was done with a 4 1/8-inch tri-cone roller bit attached to A-size drill rods, with side discharge of drilling fluid to minimize disturbance to soil belcw the bit.
Casing was advanced at intervals to keep the hole open as orilling progressed, and a thick drilling mud was mixed and maintained above the groundwater level in the hole at all times. When completed, each hole was grouted without pressure with a thick cement slurry to prevent caving.
8.2.1 Standard Penetration Tests. The standard penetration tests (SPT) were performed at 5-foot intervals in borings DM-8, DM-10, and DM-11 to provide blow-count values for all depths to be considered I
- In DM-7 samples were taken alternately every 5 feet by the Osterberg piston sampler and the SPT split-spoon.
1173 029 12 I
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4 TEST BORING 8Y RAYMONO Y INT'l fN JULY 1962
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____y g TEST BORfNG FOR D&M INVESTIGATION IN 1979 1173 030 I
13 FIGURE 1
I in the analysis, and to provide samples for field classification and laboratory verification of index properties.
Sampling was done in accordance with ASTM D-1586-67 (Standard Penetration Test) specifications, using a calibrated 140-pound pin-held hammer dropping 30 inches. The pull rope used was old and flexible, wrapped two turns around the cathead, and was oiled frequently to minimize friction and approach as free a fall of the hamer as possible.
The 2-inch split-spoon was driven 18 inches into the soil, and blow counts were recorded in 6-inch incre-ments.
The split-spoon was then slowly withdrawn and the disturbed I
sample was preserved for classification and testing after field iden-tification.
Figure 2 shows N-values plotted with depth for all borings; these values represent blow counts for the last 12 inches of each sample.
8.2.2 Undisturbed Sampling. Relatively undisturbed samples were obtained at 5-foot intervals in boring DM-9 and at 10-foot intervals in DM-7.
Samples were taken in thinwall tubes by means of an Osterberg piston sampler.
The tubes were coated with polyurethane to minimize frictional disturbance.
Before each sampling operation, the piston sampler was cleaned and oiled and extended by hydraulic pressure applied by the rig to ascertain that grit would not hinder its even extension into the soil once it was lowered to sampling depth.
When clean, the sampler I
was lowered to rest on the bottom of the hole, and the tube was extended 30 inches into the soil by even F.iraulic pressure.
The rig was chained down during the sampling to pres it uplift of the rig and uneven pressure application on the sampler.
The sampler was then slowly withdrawn from the nole, maintaining the mud level near the top of the hole.
When the sampler cleared the top of the casing, a small amount of soil was removed from the bottom of the tube (and dimensions recorded) to permit insertion of a solid cap in the end of the tube.
The purpose of the end cap was to prevent loss of sample material and moisture during removal from the sampler.
The sample tube was then severed at the tcp of the sample with a pipe cutter to release any vacuum within the tube and minimize disturbance.
while disengaging the tube from the sampler.
The tube was capped on I
top, while maintaining its vertical orientation, and carried by.a D&M field engineer to the onsite laboratory for measurement and storage.
I 1173 031 14 I
I N(biows/ foot)
I 0
20 40 60 80 100 I
i 0
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A l
A{OO 20-e\\
I dtOh KEY
$^
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-- DESIGN N VALUES N
+ N NOTE: ONLY D&M 1979 RESULTS ARE PRESENTED HERE g
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I VARIATION OF SPT N-VALUES WITH DEPTH I
1173 032 g
FIGURE 2 15
I I
8.3 Handling of Undisturbed Samples Field density measurements were made in an onsite temperature-controlled laboratory accessible only to the site security chief and D&M personnel. Upon arrival in the laboratory, a sample tube was imme-diately measured and weighed, using appropriate tare weights, to deter-mine a field density. A small amount of soil was then removed from the top and bottom of the sample to determine moisture content. A drainage cap, consisting of two perforated metal disks separated by a rubber grommet, was installed in the bottom of each sample tube to prevent displacement of the soil as drainage occurred.
The rubber giommet could be tightened or loosened by means of a wing nut. The I
sample was covered with a non-airtight cap and allowed to drain at least 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> in a vertical tube rack.
It was anticipated that, after drainage of the free water, freezing of the remaining capillary moisture would create minimal, if any, disturbance of the structure of the sand samples.
This freezing techniq'Je currently is considered the best means of preserving the structure of clean, loose sands below the water table for transport and testing (Ref. 9).
Af ter draining, the samples were placed in vertical racks in 55-gal-lon drums and packed with dry ice surrounded by shredded insulation.
The samples were allowed several hours to freeze, checked by a short length of tube filled with water ('4 ich frc a completely within a half I
1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br />).
The drums were then transported to a commercial cold-storage plant in Lacrosse, where they were stored at -20 F for the duration of the field operations.
Upon completion of the drilling program, the samples were repacked in dry ice and insulation and driven to Chicago for laboratory testing, where they were unloaded and stored in a freezer 0
maintained at about -10 F.
The sample transport was performed by a D&M field engineer to ensure careful handling.
I
~
I I
ii73 033 16
I 9.0 LABORATORY TESTING PROGRAM 9.1 General The purpose of the testing program was to provide additional strength data from undisturbed samples for the liquefaction analysis, and to I
make a limited number of confirmations of index properties.
In addition to 15 stress-controlled cyclic triaxial tests, testing inclujed specific gravity determinations, particle size analyses, minimum and maximum density determinations, and measurements of dry density of the undis-turbed samples.
9.2 Specific Gravity Determinations of specific gravity of sands at the site were made in the D&M laboratory in accordance wth ASTM D-854-58 (Specific Gravity of Soils).
Tests of four samples from depths of 31 to 47 feet yielded specific gravity values ranging from 2.60 to 2.65.
These results cor-relate with expected values for such soils and with 1973 results.
9.3 Particle Size Analyses Particle size analyses were performed on 25 samples from boring DM-10 by a D&M laboratory according to ASTM D-422-63 (Particle Size Analysis of Soils).
Results are shown as ranges for the sandy and I
gravelly soils respectively, on Figures 3 and 4.
The coefficient of uniformity, Cu (Table 1), or ratio of D60 10, provides a useful
- to D comparison of grain-size distribution at various depths and can be used in relative density calculations.
9.4 Minimum and Maximum Densities Minimum and maximum densities for a composite of samples between 31 and 47 feet were determined in the laboratory by using equipment and methous similar to those specified by ASTM D-2049-69 (Relative Density of Cohesionless Soils). These tests produced an average minimum dry density of 97.2 pounds per cubic foot (pcf) and an average maximum of 114.3 pcf. Relative density, D, could then be calculated by compar-I r
ing these densities with measured in situ densities.
- D refers to the grain size which is coarser than 60 percent of the sklebyweight.
D is defined similarly as coarser than 10 percent 10 of the sample 17 I
1173 334
M M
M M
M M
M W
W W
W W
W W
M M
M M
U.S. e rANDARD SIEVE SIZE 3 IN. l.5 IN. 3/4 IN. 3/3 IN.4 IO 20 40 60 100 200 l
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DESCRIPTION: Brown or gray-Creen fine I
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to medium sand with
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and fine gravel l
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3 N
kg <31%..g,
D PARTICLE SIZE ANALYSIS e ~ee,ea S ~oS e~coe~1eeeo A1S,1e 25 0
ty, c
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d DESCRIPTION: Fine to coarse gravel and 1
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u o
O PARTICLE SIZE ANALYSIS Jg,f ]plg l
1 O
u RANGE FOR GRAVELS ENCOUNTERED AT SITE 3
I g
O C:o 3
m 6
M TABLE 1 PARTICLE SIZE CHARACTERISTICS Depth D
- 0 D
u 10 50 60
=D
/0 7
(ft)
(mm)
(mm)
(mm) 60 10 10 0.12 0.40 0.42 3.5 20 0.20 0.42 0.43 2.2
=
30 0.15 0.44 0.51 3.4 40 0.15 0.41 0.49 3.3 50 0.16 0.39 0.42 2.6 60 0.16 0.54 0.67 4.2 70 0.18 0.39 0.43 2.4 80 0.18 0.41 0.43 2.4 90 0.20 0.66 0.72 3.6 100 0.19 0.62 0.71 3.7 n
- D refers to the grain size which is coarser thar 10 percent of the 10 sample by weight.
D nd D are defined similarly, i.e., they are 50 60 coarser than 50 percent and 60 percent, respectively.
1173 037
L Relative densities were also calculated by means of an expression developed by Meyerhof in 1957 (Ref. 10) which relates relative density to blow coints and overbur - stress at a particular depth.
These values were used in the
,p, n:
analysis (Ref. 8) and in the shear modulus calculat.ons for the one-dimensional analysis.
Anothr* che('
relative demit'es was made by means of the Marcu-son /Bie p nousky (nef. 11) expreosion involving uniformity coefficients, blow counts, and overburden stresses.
These values compared satisfactor-ily with those used in the analysis.
9.5 Dry Density of Undist rbed Samples As described previously, field densities were calcula ed by measuring and weighing the undisturbed samples immediately on extraction from tha boring.
We believe that these values are fairly accurate because the measurements were made for the entire sample and as soon as possible after sampling.
Dry density values were derived after field moisture contents were taken from each end of a sample.
Densities were also measured on the frozen samples in the laboratory.
Sample tubes were cut into smaller sections and accurately measured, and the weight of soil solids was determined by drying the sample.
Dry densities were calculated by dividing this weight of solids by the frozen volume.
9.6 Cyclic Triaxial Tests Fifteen stress-controlled cyclic triaxial tests were performed on undisturbed samples in the laboratory of the University of Illinois at Chicago Circle by Proftisor Marshall Silver.
Samples f0. testing were chosen in the depth ranges of 10 to 20 feet (hydraulic fill),
30 to 40 feet, 40 to 50 feet, and 80 to 90 feet.
9.6.1 Sample Preparation.
In preearation for testing, sample tubes were removed from the freezer and cut into sections with a cube cutter to produce test specimens. An inch of possibly dis b rbed material was wasted from the bottom of each tube. A vertical band-saw was then used to split one side of the tube, which allowed the frozen sample to be extruded vertically into a split brass cyclinder for trimming and transporting.
The specimen was placed in the triaxial cell in a membrane with filter paper at top and bottom, and a small vacuum 1173 038 21
I I
of minus 5 inches of mercury was applied while thawing the specimen.
The sample was then consolidated under a pressure slightly greater than the in situ effective confining pressure.
Specimen dimensions were recorded before and after thawing and after consolidation.
9.6.2 Testing. Cyclic triaxial testing of the specimens was performed according te procedures outlined by Silver (Ref. 12).
Samples were placed in a triaxial cell capable of being loaded with a periodic cyclic stress of constant amplitude.
Cyclic loading was begun and continued until double amplitude strains exceeded 10 percent, axial compressive or extensive strains exceeded 20 percent, or the predetermined number of load cycles was achieved.
These test results were evaluated with respect to the magnitude of cyclic axial stress and the number of cycles required to produce double amplitude, compressive or extensive strains of 5 percent and 10 percent. Also recorded was the first cycle at which the induced excess pore pressure became equal to the cell pressure, which is referred to as initial liquefaction. Ranges of stress ratios at failure were selected to obtain relationships between stress ratios and number of cycles required to cause liquefaction.
Using the weight of the solid particles, determined by drying and weighing the sand particles after completion of the triaxial test, three density calculations were made for the tested specimens. The density calculated using the frozen dimensions was called the frozen density.
After the sample was allowed to thaw for 2 or 3 hours3.472222e-5 days <br />8.333333e-4 hours <br />4.960317e-6 weeks <br />1.1415e-6 months <br /> in the laboratory under vacuum confinement, new diameter values were measured with the Pi tape at three locations on the specimen. The change in I
height in the vertical dial gage was noted. A new volume calculation for the specimen was made with the new height and diameter.
By dividing the new volume into the weight of the solid particles of the specimen, a secor.d density was determined.
This density was called the thawed density.
The cell was then assembled around the specimen and the specimen was saturated and consolidated.
Vertical dial readings and volume change readings were made and used to calculate the consolidated volume of the specimen. By dividing the dry weight of the solids by the consoli-dated volume, a third density, the consolidated density, was determined.
l 1173 039 '
22
i i
A summary of densities and the triaxial test results is given in Table 2, and the variation of dry density with depth is shown in Figure 5.
z.
-m 3
11~73 040'
m m
M M
M M
M m
M M
M M
M M
m m
m m
M TABLE 2
SUMMARY
OF CYCLIC TRIAXIAL TEST RESULTS*
Nanter of Cycles Dry DensityJpf)
Effective Saturation to Liquefaction Sample /
Thael anxl Confining Stress Details N
FJ g
5 "10 Test S m imen Depth Frozen Thard Consolidated Pressure Datio
'n (5% m (10t m Nunber Nunber (ft) Condition Corvli tion Cork'I t ion I!"c ar w ter c
(Ir. i t ial) Strain) Strain) Pemarks 5/2 21.5 102.2 104.6 105.4 2,000 0.22 0.97 5
5 5
1 4/3 16.0 101.8 103.5 104.1 2,000 0.18 0.98 11 10 11 4/2 16.5 103.9 105.4 106,4 2,000 0.32 0.97 3
2 3
7/2 31.5 99.7 101.0 101.4 2,500 0.20 0.86 10 10 14 tow "B" value 2
7/3 31.0 102.0 103.7 104.2 2,500 0.28 0.99 10 12 16 ru
^
8/1 37.0 104.1 105.4 105.7 2,500 0.39 0.96 6
11 25 8/3 36.5 102.9 104.3 104.8 2,500 0.12 0.99
> 1,000 > 1,000 > 1,000 Did not l inon ry 10/1 47.0 101.6 103.1 103.7 4,000 0.32 0.95 4
3 5
10/2 46.5 103.0 104.7 105.4 4,000 0.23 0.97 5
4 7
3 9/2 41.5 103.0 104.8 106.1 4,000 0.43 0.96 5
3 6
10/3 46.0 3.5 104.8 105.5 4,000 0.18 0.99 8
8 11 11/2 52.0 103.0 104.5 105.1 4,000 0.13 0.43 84 86 92 18/1 87.0 104.7 105.8 107.8 8,000 0.35 1.00 7
8 16 4
10/2 86.5 103.6 104.6 105.6 8,000 0.26 0.96 14 16 18 19/l 92.0 103.2 104.8 106.0 8,000 0.45 0.99 3
2 3
Nu
- b est were performed on soil type SP from boring no. DM-9.
C)
-c-
7 (PCF) d 89 90 100 110 120 130 140 0
0 10-es e og O
O O
20-e se e
o e e O
30-e ggg O
KEY:
O e 1973 OSTEHBERG SAMPLES e
O 1979 OSTERBERG SAMPLES (froien)
O O
40-es es O
=
O O
gw-O o
e O
m-ee O
e e
O 70-ee 9 O
e OO 80 e
ege oO O
90-e O
100-
=_
VARIAT!ON OF DRY DENSITY WITH DEPTH '
1173 042-names a moons 25 FIGURE 5
I 10.0 LIQUEFACTION ANALYSES 10.1 General The liquefaction analyses performed for the LACBWR site in the past were based on test borings and laboratory data from the D&M investigations of 1973 (Ref. 1). As mentioned in earlier sections of this report, the findings of the 1973 studies were evaluated by NRC in 1978 under its Safety I
Evaluation Program, and several questions were raised regarding the fac-tors of safety under the design SSE.
D&M also reevaluated its earl;;er findings as a result of the NRC review using current state-of-the-art methods. Although D&M concludeu that the factors of safety against lique-faction under the design SSE remained unchanged, there was agreement among NRC, WES, D&M, and DPC in the following:
I a.
" Undisturbed" sampling in the cohesionless soils was necessary.
b.
In situ dry densities must be estimated with greater accuracy I
using " undisturbed" samples.
c.
There was a need for development of continuous standard penetra-tion "N" values under carefully controlled conditions.
I d.
Cyclic shear strength parameters of the liquefiable soils had to be obtained by performing cyclic triaxial tests on " undisturbed" samples.
e.
Estimates of the cyclic shear stresses resulting from the design SSE must be made by performing a one-dimensional wave propagation analysis.
f.
The seismicity of the LACBWR plant site and the potential for liquefaction under an acceleration level which realistically rep-resents the seismicity of the site must be analyzed.
With consideration of these requirements, a limited but carefully con-trolled field and laboratory investigations program was undertaken.
Using the data developed in these investigations, detailed liquefaction analyses d
were performe.
10.2 Liquefaction Potential There are two basic approaches for evaluating the liquefaction poten-tial of a deposit of saturated sand when it is subjected to earthquake loading.
The first approach uses the information available on the perform-I ance of various sand deposits during past earthquakes.
This approach is essentially empirical, and the response of soil to earthquake loading is un M3
I not evaluated by any direct means.
Simplified methods of analysis, with known limitations, have been proposed by various investigators.
l Also, a large number of factors that significantly affect the liquefac-tion characteristics of a given sand have been recognized and may be studied in detail to confirm the conclusions of such an analysis.
In the second approach, stress conditions in the field are evaluated by using an analytical technique, such as the one-dimensional wave propagation analysis.
Laboratory investigations are conducted to deter-mine the cyclic shear stresses required to cause liquefaction at various I
depths.
At a given depth, a factor of safety against liquefaction can be evaluated by dividing the cyclic shear stress required to cause I
liquefaction by the cyclic shear stress induced during the design earth-quak e.
Methods based on these two approaches were used to assess the liquefaction potential of the granular soils at the LACBWR site.
10.3 Evaluation of Liquef action Potential, Aporoach 1 10.3.1 Simolified Procedure (Proccdure 1).
In the first approach, I
the procedure recommended by Seed (Ref. 7) was used to estimate the cyclic shear stress required to cause liquefaction.
The cyclic shear stress induced during shaking was computed by the Seed and Idriss Simpli-fied Procedure (Ref. 6).
The following steps are used in Procedure 1:
Convert the "N"* values from the Standard Penetration Tests to N) a.
values (N) is the penetration resistance, corrected to an effec-tive overburden pressure of 1 ton /ft ) using the relationship:
I where:
C = 1 - 1.25 log 5 /5 y
c 1 2
5 = effective overburden pressure (tons /f t )
c 2
dy = a constant equal to 1 ton /ft.
I
- N = number of blows required to advance a standard split-spoon 12 inches into the ground, when driven by a hammer weighing 140 pounds dropping I
a T stance of 30 inches.
27 117.3 044
I b.
Based on a collection of data from actual field performance and a few additional site studies, the lower bounds for the cyclic shear stress ratios that cause liquefaction in the I
field and which correspond to different N1 values and magnitudes of earthquakes have been established (Ref. 7).
Using this values can then be converted to the cyclic relationship, the Ng shear stress ratios, T/5, requ, mi to cause liquefaction c
for the design earthquake.
I c.
Compute the cyclic shear stress ratio at any depth fn the ground that is induced by the design earthquake (Ref. 6) using the relationship:
T II = 0.65 (a I9) ( o/#c) #d av c mn where:
Oc
= effective overburden pressure 00 sand layer I
a
= maximum acceleration at the max ground surface (ft/sec )
I o
= total initial overburden pressure on 2
U sand layer under consideration (tons /ft )
r
= a stress reduc'. ion factor varying from d
a value of 1.0 at the ground surface to a value of 0.9 at a depth of about I
30 feet.
I av
= average cyclic shear stress in the sand 2
layer under consideration (tons /ft )
2 g
= acceleration due to gravity (ft/sec ).
d.
The lower bound cyclic shear strength values that are obtained from Step "b" can then be compared with the average cyclic I
shear stresses obtained in Step "c," and the liquefacthon potential at various depths can be evaluated.
SPT(N) values from the recent D&M investigations were plotted as a function of depth (Figure 2).
Average design N values were chosen for different depths and were converted to corrected blow counts (N,).
Relative boundaries between " liquefaction" and "no liquefaction" condi-tions for various magnitudes of earthquakes corresponding to ground surface acceleration levels between 0.10 g and 0.20 g were drawn using the data and principles oresented by Seed (Ref. 7).
Figures 6 through 11 I
28 1173 045
0.6 2
S u.
UN hg, 0.5
^
C O
'^
3C (o
gS f
C $
0.4 5E 8GO k2 0.3
/
eq C
b%
ts 9 $
0.2 25
/
amax.10g
=
g
^
$5
- 0. i as og o
os j
U 0
to 20 30 40 50 MODIFIED PENETRATION RESISTANCE Nj - Wm/ foot E
CORRELATION BETWEEN FIELD LIQUEFACTION BEHAVIOR OF SANDS FOR LEVEL GROUND CONDITIONS AND PENETRATION RESISTANCE j
'1173 046 29 FIGUPE 6
9 i
0.6 9 u.
UC 0.5
(
8d 38 J
4 E $
0.4 zs bb
?5 m m EE u4 Q'- m 9 h 0.2 5
/
max.12g a
=
yg
$5 C3 n
gs ggr o o
Jb S5 00 10 20 30 40 50 MODIFIED PENETRATION RESISTANCE N3 - blows / foot CORRELATION BETWEEN FIELD LIQUEFACTION BEHAVIOR OF SANDS FOR LEVEL GROUND CONDITIONS AND PENETRATION RESISTANCE 1173 047-DAMES B MOORM 30 FIGURE 7
I I
I I
I 0.6 a
E "_
4 l
lJ o5
~8 3
I 0.4 Ei kz 0.3 oE
/
e 2 h 0.2 I
E5 Es@
./.
max 14*9 a
+
3 e,
==
n ds I
b 0
10 20 30 40 50 MODIFIED PENETRATION RESISTANCE Ng - blows / foot I
I CORRELATION BETWEEN FIELD LlOUEFACTION BEHAVIOR OF SANDS FOR LEVEL GROUND CONDITIONS AND PENETRATION RESISTANCE I
I i 173 048 I
31 FIGURE 8
0.6 2
Og bN 0.5 oD
]C Jk N
0.4 EE 8a2 k2 0.3 UE
/
t*E 4
9 6 0.2
=
4m C O hN Ah A
amax.16g
=
c:: 5
/
g3 3.# f6 0.1
~
u:
Ub S8 00 10 20 30 40 50 MODIFIED PENETRATION RESISTANCE Nj - blows / foot CORRELATION BETWEEN FIELD LIQUEFACTION BEHAVIOR OF SANDS FOR LEVEL GROUND CONDITIONS AND PENETRATION RESISTANCE 1173 049-DAMES 8 MOORE 32 FIGUFIE 9
0.6 z
o u.
EU
@i U
k7 M
g, 0.5 O D 38 J u.
N 0.4 EE Es
@2 a
z 0.3 UE
/
e*$
=c yw 0.2
<G E
m x
max.18g
]
a
=
0.1 3
u=
Jb SE o
00 10 20 30 40 50 MODIFIED PENETRATION RESISTANCE N1 - blows / foot CORRELATION BETWEEN FIELD LIQUEFACTION BEHAVIOR OF SANDS FOR LEVEL GROUND CONDITIONS AND PENETRATION RESISTANCE 1173 050 DAMEES O MOORE 33 FIGURE 10
- s a
.i 0.6 z
9 u.
UU Y
Em 1
g, 0.5 a o 3E 8O N
o.4 Ei 80 ao n.
3<z 0.3 oq
/
8*
u
> 6 0.2 9
$E 0 0 mO
.0[]
amax.20g
=
c h
.1 U
- 0 u=
Jb SE OO 10 20 30 40 50 MODIFIED PENETRATION RESISTANCE N1 - blows / foot M
m_
CORRELATION BETWEEN FIELD LIQUEFACTION BEHAVIOR OF SANDS FOR LEVEL GROUND CONDITIONS AND PENETRATION RESISTANCE 1173 051 DAMES 8 MOORE 34 FIGURE 11
I show such boundaries on which are plotted LAC 8WR plant site data of N values corresponding to certain cyclic shear stress ratios induced y
by the respective earthquakes of different magnitudes.
The SPT(N) values, the corresponding Ny values used to compute the strength, the cyclic shear strengths, and the cyclic shear stresses computed using I
the Simplified Procedure, and the resulting factors of safety at different depths for various acceleration levels are presented in Table 3.
The stresses and strengths are presented as functions of depth for the six acceleraion levels in Figures 12 through 17.
Data in Figures 12 through 17 and Table 3 show that the site does not liquefy at acceleration levels less than or equal to 0.12 g at any depth.
As the acceleration level increases from 0.14 g to 0.20 g, the depths susceptible to liquefaction increase from 20 to 30 feet to lu to 40 feet.
10.3.2 Jaoanese Procedure (Procedure 2).
A procedure being used in Japan (Ref. 8) also falls under the general category of Approach 1.
The procedure for computing the cyclic shear stresses is the same as described in Step "c" above using Seed and Idriss (Ref. 6) simplifications.
The estimation of the cyclic shear strength is as follows:
a.
Estimate the relative density of the liquefiable soil using the relation developed by Meyerhof (Ref. 10) based on laboratory tests performed by Gibbs and Holtz (Ref. 13):
D * = 21 N/(T + 0.7) r y
D*
= estimated relative density N
= blow count from SPT 6
= effective overburden pressure at the y
2 depth of interest (kg/cm ).
b.
Estimate the cyclic shear strength using the appropriate equation:
/0.35)
R
= 0.0042 D * - 0.225 log 10 (D50 r
for 0.04 m 1 050 1 0.6 m R
= 0.0042 D * - 0.05 L
r for 0.6 m i D50 5 1. 5 m 1173 052 35
T N M M W!E E
m M
uum TABLE 3
SUMMARY
OF LIQUEFACTION ANALYSIS APPROACP 1, PROCEDURE 1 and Factors of Safety for Various AcceleraM ons*
Average _ Cyclic Shear Stresses J clic Shcar Strengthsg 0.12 g a
= 0.14 9 a
- 0.16 y a
= 0.18 g a
= 0.20 g a
= 0.10 g a
=
- 8I9" "/"I
'ov i
FS av i
FS
'ay FS av i
FS av t
FS
'av t
FS t
I pepth (ft}
10 6/8 73 138 1.89 88 138 1.57 10.?
127 1.23 118 127 1.08 132 127 0.96 147 127 0.86 20 t,/ 6 145 167 1.15 174 174 1.00 203 148 0.73 232 148 0.64 261 148 0.57 290 148 0.51 10 10/9 213 317 1.49 255 117 1.24 298 292 0.98 341 292 0.86 3h3 292 0.76 426 292 0.69 40 14/ll 270 525 1.94 324 494 1.52 378 49s 1.31 432 463 1.07 486 463 0.95 540 463 0.86 50 25/17 301 1,005 3.34 161 968 2.68 421 894 2.12 482 894 1.85 542 856 1.58 602 856 1.42 60 l #,/ 20 328 1,371 4.18 194 1,327 3.38 459 1,283 2.80 525 1,239
- 2. 16 590 1,239 2.10 6"5 1,239 1.89 70 12/16 343 1,256 3.66 411 1,206 2.93 480 1,155 2.41 548 1,105 2.02 617 1,105 1.79 685 1,105 1,61 or 80 15/15 16 5 1,317 3.63 435 1,317 3.03 508 1,260 2.48 581 1,203 2.07 653 1,201 1.84 726 1,203 1.66 90 37/14 lH9 1,414 3.63 467 1,349 2.89 545 1,285 2.36 622 1,285 2.07 700 1,285 1.04 778 1,285 1.65 100 40/12 409 1,349 3.30 491 1,278 2.60 573 1,207 2.11 655 1,207 1.84 737 1,207 1.64 818 1,207 1.48
- T
= average cyclic shear stress from Seed & Idriss Simplified Procedure (Ref. 6).
3y
= cyclic shear strength based on corrected blow counts and recorded cyclic behavior during earthquakes.
1 Nu Factor of safety (FS) = (cyclic shear strength) i (average cyclic shear stress).
O Cn u
CYCLIC SHEAR STRESS / STRENGTH (psf)
'#UO 0
's %
9\\\\
20-Q's NN s
\\ %
' 's
_ 4G-t
,'~~,,-o~
3 I
t
'~~s~s' ma 4
s,p so~
/
s'
~
5
\\
g k
80-
\\
\\
b l
I
/,
100-KEY:
0--e AVER AGE CYCLIC SHE AR STRESS (SIMPLIFIED PROCEDURE)
O---<> CYCLIC SHE AR STRENGTH B ASED ON PENETR ATION RESISTANCE COMPARISON OF CYCLIC SHEAR STRENGTH APPROACH 1 PROCEDURE 1 (a
= 0.10g) max 1173 054 DAMES B MOOsam 37 FIGURE 12
l I
I CYCLIC SHEAR STRESS / STRENGTH (psf)
'Y "W
U o
\\
\\
I
\\
20-
\\
I
'%'s N
_ 40-
%s I
j
~~~
j
s s
b
'~ss'y
- so.
I
/
NN I
T
"~
I
/
/
100-d I
KEY:
9----e AVERAGE CYCLIC SHEAR STRESS (SIMPLiflED PROCEDURE) 0- --O CYCLIC SHE AR STRENGTH B ASED ON PENETR ATION RESISTANCE I
COMPARISON OF CYCLIC SHEAR STRESS AND STRENGTH APPROACH 1 PROCEDURE 1 (amax 0.129)
=
I I
i 73 055 UAM ES B MOORE 38 FIGURE 13
I I
I CYCLIC SHEAR STRESS / STRENGTH (psf) o 200 400 soo soo 1000 t200 isoo g
N 20-k\\ ss
's\\s I
do-
'*%s~~s'N t
=
~s
,7 60-
/
k L
NN Bo-
\\
h
/
d 100-KEY:
M AVER AGE CYCLIC SHEAR STRESS (SIMPLIFIED PROCEDURE)
I C>--O CYCLIC SHE AR STRENGTH B ASED ON PENETR ATION RESISTANCE COMPARISON OF CYCLIC SHEAR STRESS AND STRENGTH APPROACH 1 PROCEDURE 1 (a
= 0.14g) max I
1173 056 DAMES B MOORW 39 FIGURE 14
CYCLIC SHEAR STRESS / STRENGTH (psf) 4K 6@
800 1000 1200 14po 0
s t
1 20-h s'N N
% 's N
_ 40-
,'~~s
~~.'%'~
=
E
~ s ~. s w
Q N
60-s,w
/
N
\\
so-R
\\\\
N
/
/
=
100-d KEY:
H AVER AGE CYCLIC SHE AR STRESS (SIMPLIFIED PROCEDURE)
C>--O CYLIC SHE AR STRENGTH BASED ON PENETR ATION RESISTANCE COMPARISON OF CYCLIC SHEAR STRESS AND STRENGTH APPROACH 1 PROCEDURE 1 (a
= 0.16g) max i
1173 057 UAMES 8 Moostm 40 FIGURE 15
I I
I CYCLIC SHEAR STRESS / STRENGTH (psf) 0 I
\\
\\
20-
's I
N N s 's I
_ 40-
~
g
'~s
=
s~~s E
N'~s
=
I w
s~s s,T so.
/
I
/
Cl's\\
I 80-
\\
\\
/
/
100-KEY:
O--e AVER AGE CYCLIC SHEAR STRESS (SIMPLIFIED PROCEDURE)
I O--O CYCLIC SHEAR STRENGTH B ASED ON PENETRATION RESISTANCE I
COMPARISON OF CYCLIC SHEAR STRESS AND STRENGTH APPROACH 1 PROCEDURE 1 (a
=0.18g) max 1173 058 g
41 FIGURE 16
I I
CYCLIC SHE AR STRESS STRENGTH (psf) 0 200 400 600 800 1MO 1200 14,00
's N
N I
(
20-
'N I
N h
' s.
N
_ 40-M~s I
'~~'~
e
~,
E
'w~ ~ s I
b
'~s s O
80 y
/
I
/
<'s\\
I PO-g
\\
I
/
/
100-KEY:
W AVERAGE CYCLIC SHEAR STRESS (SIMPLIFIED PROCEDURE)
O---O CYCLIC SHE AR STRENGTH BASED ON PENETRATION RESISTANCE I
I COMPARISON OF CYCLIC SHEAR STRESS AND STRENGTH APPROACH 1 PROCEDURE 1 (a
=0.20g) max I
1173 059 g
42 FIGURE 17
I where D*
= estimated relative density from Step "a" r
D
= particle size in nm corresponding to I
50 50 percent on the grain size curve and R
= cyclic shear stress ratio required to cause liquefaction (triaxial test conditions).
c c.
Apply suitable corrections to R to obtain R, the corrected cyclic shear stress ratio required to cause liquefaction.
c R
= (C ) (C ) (C ) R y
2 3
t where C
= 0.57 (to convert triaxial test conditions 1
to simple shear field condition)
C
= 1.3 to 1.5 (to account for N
= 5 rather 2
eq than N
= 20 used in Japanese study eq I
C
= 1 (to account for differences in failure 3
strain criteria **).
By comparing the estimated values of cyclic shear strengths from I
the Japanese procedure (Ref. 8) and the average cyclic shear stresses from the Seed and Idriss procedure (Ref. 6), factors of safety against potential for liquefaction can be estimated for different depths.
The cyclic shear stresses which were calculated under Procedure 1 were also used under Procedure 2.
The relative densities were estimated using SPT(N) values (Figure 2) and 0 values (Table 1:.
Table 4 presents 50 the estimated relative densities, the estimated cyclic shear strengths, the calculated cyclic shear stresses, and the resulting factors of safety at various depths for different accelerations. The stresses I
and strengths at differer.t depths for all the acceleration levels considered are plotted on Figure 18.
The data on Table 4 and Figure 18 suggest that there would be no liquefaction susceptibility under earthquakes producing ground surface accelerations of less than or equal to 0.16 g.
Under accelerations of 0.18 g and 0.20 g LACBWR site soils between depths of 20 to 40 feet may experience liquefaction.
I I
- The 5 to 6 percent double amplitude shear strain used in the Japanese studies was very close to the initial liquefaction criterion used in this D&M study.
Therefore C ~ 1 was used.
3 l
1173 060 43
M M
M M
M M
M M
M M
M M
M M
M TABLE 4
SUMMARY
OF LIQUEFACTION ANALYSIS APPROACH 1, PROCEDURE 2 Average Cyclic Shear Stresse s and Factors of Safety for Various Accelerations
- a,,,
= 0.10 g a,,,
= 0.12 g a,,, = 0.14 g a,,,
= 0.16 g a
= 0.18 g a,,, = 0.20 g max IleEt.lLi!tl L (l>s t )
[av_{ psf [
[E h
[E
[ tv
[E
,' a_v
{!
h
[E
'ay f!
10 46 161 73 2.21 88 1.83 103 1.57 118 1.37 132 1.22 147 1.10 20 41 222 143 1.53 174 1.28 203 1.09 232 0.96 261 0.85 290 0.77 10 48 365 213 1.72 255 1.43 298 1.23 341 1.07 383 0.95 426 0.86 40 53 494 270 1.83 324 1.52 378 1.31 432 1.14 486 1.02 540 0.91 50 tib 745 301 2.47 361 2.06 421 1.77 482 1.54 542 1.37 602 1.24 60 15 885 328 2.70 394 2.25 459 1.93 525 1.69 590 1.50 656 1.35 10 67 1,005 343 2.93 411 2.44 480 2.09 548 1.83 617 1.63 685 1.47 H0 66 1,145 36 3 3.16 435 2.63 508 2.25 581 1.97 653 1.75 726 1.58 90 65 1,092 389 2.01 461 2.34 545 2.00 622 1,76 700 1.56 778 1.40 100 6S 1,349 409 3.30 491 2.75 573 2.35 655 2.06 737 3.83 818 1.65 average cyclic shear stress computed using Seed & Idriss Simplified Procedure (Ref. 6).
- t
=
gy
[
cyclic shear strength estimated from relative density (Japanese procedure), (Ref. 8).
=
T u
Factor of safety (FS) = (cyclic shear strength) I (average cyclic shear stress).
0.5 N/f
+.7)
(Ref. 10).
relative density based on Dr " 21 v
o
- D
=
p
I I
I CYCLIC SHEAR STRESS / STRENGTH (psf)
'W 0
KEY:
AVERAGE CYCLIC SHEAR STRESS I
CYCLIC SHEAR STRENGTH BASED (SIMPLIFIED PROCEDURE)
N 0---O ON RELATIVE DENSITY A?sD GR AIN N
SIZE (JAPANESE PROCEDURE)
I N
h
_ 40-I s
s s S
's E
N
\\
I b
s so.
N N NN \\
s ss i
s e
~
1
~~
I s
2 2
5
,2 N
3%.
I I
I COMPARISON OF CYCLIC SHEAR STRESS AND STRENGTH APPROACH 1 PROCEDURE 2 (a
=0.10g to 0.20g) max I
I 1173 062 g
45 FIGURE 18
I I
10.4 Evaluation of Liquefaction Potential, Approach 2 Approach 2 uses more rigorous methods and site-specific data from I
sophisticated laboratory results.
A seismic response analysis was perfarmed to estimate the stresses, strains, and accelerations at different depths within the soil profile resulting from SSE loading at the LACBWR site. Also,._ aral 'iquefaction tests were performed on undisturbed l
samples to define their behavior under cyclic loading.
10.4.1 Soil Model Used in the Response Analysis.
Based on a review of the data from the most recent investigation, a representative, idealized soil profile was established for the one-dimensional wave propagation analysis.
This idealized soil profile corresponds to an average surface elevation of +639 feet.
Table 5 presents the idealized soil profile and the soil properties that were used in the response analysis.
The upper 135 feet of the soil deposit were divided into 12 sublay-ers.
Detailed descriptions of the soils encountered at the LACBWR site are presented in Table 5 and on the boring logs in the Appendix.
10.4.2 Soil Properties Used in the Response Analysis.
The soil properties required for the wave propagation analysis are unit weight, shear modulus, damping ratio, and coefficient of earth pressure at rest.
a.
Average values of unit weights based on field and laboratory test results were used in the analyses (Table 5).
b.
In general, the shear modulus of a soil is influenced by several variables, including effective confining pressure, void ratio, stress history, degree of saturation, soil structure, amplitude of strain, and frequency of vibration (Ref.14).
The in situ shear modulus can be estimated by reviewing data from the I
tolowstrainlevels(approximately10gmoduluscorresponds geophysical survey.
This value of shea percent).
Shear moduli corresponding to other strain levels can be determined from strain-controlled cyclic triaxial tests and resonant column tests. The extensive field and laboratory investigations I
of different soils conducted by indeper. dent researchers have generally established shear modults versus strain relationships of soils (Ref. 6).
The shear modulus versus strain relation-ships of sands at the LACBWR site were developed by considering the generalized tiends from the published data, the results of strain-controlled cyclic triaxial tests performed on recon-I stituted samples in 1973, and the relative density estimations based on SPT results.
46 g
1173 063
m M
M M
m M
M M
m M
M M
M M
M M
M TABLE 5 GENERALIZED S0Il PROFILE AND MODEL FOR ONE-DIMENSIONAL WAVE PROPAGATION ANALYSIS Shear Elevation Depth Design N Modulus (ft)
(ft) valve I Wet (pcf)*
D **
r 2***
G (ksf)
+639 0
Ground surface elevation +639
!!ydraulic fill: brown fine-115 376 to-a.ed i um sanJ wi t h occas ion-6 45 43 al fine gravel, trace of sitt 120 1,162 619 20 Ground water +627 to +631 Gray-green fine-to-medium 121 1,660 sand with occasional fine gravel, trace of silt, layers 6-14 48 45 of Clayey Silt 128 1,080 599 40 (Bottom of reactor vessel +610 to +618)
Brown fine-to-medium sand 126 66 55 2,550 with occasional fine gravel /
coarse sand, trace of silt 14-36 la, 2,795 579 60 (Bottom of piles ap m x.
+580)
Brown file-to-medium sand 122 3,L37 559 80 with occasional fine gravel /
133 3,125 coarse sand, trace of silt 30-40 132 3,537 539 100 130 69 57 3,718 Fine-to-coarse gravel and brown fine-to-medium sand 134 90 6,244 with little silt 519 115 Brown line-to-medium sand with occasional fine gravel /
134 90 79 5,273 coarse sand, trace of silt 499 135 Bedrock N
- Based on field measurements.
u
_ + 0.7) 0 5 (Ref. 10); laboratory D tests, and Ref. 11.
- Based on D = 21 N/(o r
y r
Om
- From expression G = 1,000 K2 ("m)0*5 based on data from Ref. 15.
p-
I Jased on the reported results in the literature (Ref. 15),
I the shear modulus, at low strain levels of coarse grained soils, can be expressed by the equation G
= (1000) (K ) ( m) 2 2
G
= shear modulus (lb/ft )
2 5
= effective mean confining pressure (lb/ft )
m K
= constant for a given compactness of I
9" soil.
The values of K used in the analysis are presented in Table 5.
7 The shear modulus varies nonlinearly with strain level.
This variation was assumed to follow the patt.; n of average data on the coarse-grained soils found in the literature (Ref. 15).
I The nonlinear strain dependence of shear modulus that was used in this anr. lysis is presented in Figure 19.
c.
Most of the parameters discussed in relation to shear modulus have an opposite effect on the damping value, which increases with increasing strain amplitude, decreases slightly with increasing ambient stress, and decreases with increasing void ratio.
Strain-controlled cyclic triaxial tests and resonant I
column tests on undisturbed samples are necessary to define experimentally the variation of the damping ratio with strain level. However, for the purposes of this study, it was con-I sidered satisfactory to assume that the damping ratio of the soils at the LACBWR site, both in magnitude and in variation with strain level, were similar to the average results found I
in the literature (Ref. 15).
These values are the average values obtained from the experimental investigation performed by several independent researchers on typical sands.
The strain-dependent values of the damping ratios for typical l
sands which were used in this response analysis are presented in Figure 20.
(The damping ratios measured experimentally in the laboratory during the 1973 D&M investigations were also reviewed before choosing the design values).
d.
A value of 0.45 was assigned for the coefficient of earth pressure for all the granular soils.
I 10.4.3 Design Earthquake Used in the Response Analysis.
The horizontal component of a digitized acceleration-time history was used as the I
input motion at the surface of the soil deposit.
The corresponding accelerogram is presented in Figure 21.
The duration of the design earthquake was assumed to be 15 seconds, and a range of maximum ground surface accelerations of 10 to 20 percent of gravity was used, based on the recommendations of NRC.
l 1173 06E I
I I
I 1.0 I
0.9 -
I 0.8-I 0.7-g I
5 z
2
- 0. 6 -
E !
5 0.3-5 N
O 0.4-I E
i 5
0.3 0.2-0.1-I 0
0.0001 0.001 0.01 0.10 1.00 10.00 I
SHEAR STRAIN, / -PERCENT I
TYPICAL REDUCTION OF SHEAR MODULUS WITH SHEAR STRAIN (coarse grained soils)
/J V
(SE ED & IDRISS,1970) 49 FIGURE 19
M
)
07 9
1 S,
S RI M
ID D
E E
M (S
0 M
1 M
M i i M
SD N
M A
S D
T E
I N
M E
T 0
C A
1 R
R E
P UT A
N S
M I
R AR O
T F
S O
R I
M i 0 S
R A
T 2
E A
H 1
G N
I F
MAD M
3 M
i 0 1
M M
b 0
1 5
0 5
0 5
1 5
0 0
3 3
2 2
1 M
5at i 2 Ge $E 2 c
M
~ NU OD~
M OIIa
@ igaE I
E 2@
g
'7 1
Bog - xeme 6 - NO11VB 31333, o
o o
o N
9 8,
m
+
+
8 w
oq 3
w_
o m
o N
w S
g 2
i d
~y i
.E.
o
~
O cn
~C
-q N
O
- 6 o
~
=
r E
Z
~
w w
2 Z
p O
o.
E
.8 e
a N
4 H
Z O
N me-O 8
I m
h w
8 a
v 8
a
'6 a
o o
m 9
q q
c
+
+
60L' = xeme 6 - NOl1VH3 I333V 1173 068 UAMES 8 MOORE 51 FIGURE 21
I 10.4.4 One-Dimensional Wave Propagation Analysis.
A mathematical model was used to evaluate the response of the soils at the LACBWR site when subjected to the SSE loading.
This model is based on one-dimensional strain-compatible shear wave propagation through a layered system.
Each layer in the system is assumed to be isotropic, homogeneous, and of viscoelastic behavier (Ref. 16).
A computer program developed by Schnabel e_t a_1. (Ref. 16) was modified by D&M to include additional irput and output.
The nonlinearity of the shear modulus and damping ratio is accounted for in this program by the use of equivalent linear properties.
This computer program (Ref. 17) was verified for various practical problems, certified in accordance with quality assurance requirements, and used to analyze the soils at the LACBWR site.
The average shear stress levels in the stress histories obtained by peferming one-dimensional analysis were computed assuming the ground I
water to be at 10 feet below ground surface.
This represents the average condition; the groundwater level actually fl<ctuates slightly.
The I
average cyclic shear stresses computed by performing one-dimensional analyses are plotted as functions of depth for various acceleration levels on Figure 22.
10.4.5 Cyclic Shear Strength.
The next step in Approach 2 is to deter-mine the cyclic shear strength of undisturbed samples obtained from various potentially liquefiable layers.
Fifteen samples, representing four depths of the soil profile, were chosen for stress-controlled cyclic triaxial testing using standard procedures described in Section 9.6.
Figure 3 shows the envelope of particle size curves for the samples uscd. The results of these tests are surmiarized in Table 2.
The test results were plotted on a semilogarithmic plot to define the relationship I
between stress ratio and number of cycles required to cause initial liquefaction (Figures 23 through 27).
10.4.6 Conversion of Irregular Stress History Into Equivalent Uniform Cyclic Stress Series.
In Approach 2, the calculated cyclic shear stresses are compared directly with those required to cause liquefaction of representative soil samples in the laboratory.
It is usually more convenient to perform laboratory tests us;ng uniform cyclic stress 1173 069 52
I I
I CYCLIC SHE AR STRESS / STRENGTH (psf)
O I
KEY:
s 3
9.---e AVER AGE CYCLIC SHE AR STRESS N
(ONC-DIMENSIONAL ANALYSIS)
\\
O---O STRENGTHS BASED ON CYCLIC I
TRI AXt AL TESTS ON UNDISTURBED
'\\
SAMPLES 40-I 3
'N N %
E
's S
's D
60-s
's I
N
' N.
,\\
N I
80-N s 's N
I 100-5.,
5.,
k.. k..
k,.
u g
j g
g g
120-140-I COMPARISON OF CYCLIC SHEAR STRESS AND STRENGTH APPROACH 2 (a
= 0.10g to 0.20g) max I
1173 070 g
53 FIGURE 22
M 00 0,
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0 i
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i i i i i ii i
.,,ii.
i e i i ii BORING: DM-9 CONFINING PRESSURE: 2500 psf DEPTH: 31-37 ft DRY DENSITY: 101-106 psf NOTE:
0.8 UPPER CURVE IS INDICATED BY THE 3 VALID TEST RESULTS.
LOWER DESIGN CURVE WAS CHOSEN, HOWEVER, TO BE CONSERVATIVE AND TAKE INTO ACCOUNT THE TEST WITH OUESTION ABLE VALIDITY b
lb BECAUSE OF THE LOW d
"B" V ALUE.
0.6 -
a i
6
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d
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l TEST DISCARDED:
1 i
l l
~
C 0.0
)
u 1
5 10 100 1,000 10,000 I
O NUMBER OF CYCLES M
o 2'
I N
CYCLIC SHEAR STRESS RATIO VERSUS NUMBER OF CYCLES FOR INITIAL LlOUEFACT:ON, O
o j
o TEST NO,2, DEPTH 31 TO 37 FEET m
3
000 i
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/
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0.2 702
/
M
/
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/
w HYDRAULIC FILL
' ' ',000 U
0.0 1
10 100 1,000 10 g
l 5
NUMBER OF CYCLES o
I 3
SUMMARY
CURVE SHOWING EFFECTS OF DENSITY AND SOIL FABRIC ON NUMBER o
E O
OF CYCLES TO INITIAL LIQUEFACTION m
3 0
I I
applications than to attempt to reproduce a representative field stress history.
Therefore, it is necessary to convert the irregular stress I
history that is actually developed during earthquakes into an equivalent uniform cyclic stress series.
There are three basic methods by which this conversion can be accomplished.
However, it has been shown that the procedures used in this step of the analysis have little effect on the final analysis (Ref. 7). Based on the results of a statistical study of the representa-tive numbers of cycles developed during a number of different earthquake motions, a convenient basis for selecting an equivalent uniform cyclic stress series for earthquakes of different magnitudes has been presented by Seed (Ref. 7). Accordirq to Seed, for earthquake magnitudes between 5 and 6, the number of equivalent cycles, Neg, is approximately 5, corresponding to an average cyclic shear stress, Tav, of 65 percent of the maximum shear stresses. Therefore, the maximum cyclic shear stresses that were obtained by performing a one-dimensional wavc propaga-tion analysis were multiplied by 0.65 to obtain the average cyclic shear stress at any point.
10.4.7 Correction Factor. The cyclic triaxial test does not directly simulate the simple shear conditions actually induced during an earth-q uak e.
Also, the effect of multidirectional shaking is not included in this testing. As a result, the strass ratio obtained in the cyclic triaxial test is higher, and a correction factor, C, is applied to r
modify these values.
For normally consolidated sands (K ~ 0.4), a g
value of 0.57 is considered appropriate for Cr (Ref. 7), and the stress ratio causing liquefaction in the laboratory is multiplied by 0.57 I
to account for the field conditions.
Figure 22 represents the summary of cyclic shear stress computation I
using the one-dimensional analysis and the cyclic shear strengths from the laboratory tests.
The stresses plotted are 65 percent of the maximum shear stresses to correspond to an average condition. The cyclic shear strengths were obtained by following the procedures mentioned below.
Figure 27 is a summary of all liquefaction test results.
It can be seen that three distinct liquefaction curves can be drawn on the various data points--an upper bound and lower bound for natural materials below 1173 076 59
'"5 A
l the hydraulic fill material and an average for hydraulic fill material, a
Three stress ratios corresponding to these three curves can be chosen for a given number of cycles, N For N of 5, the three possible eq.
eq q
relations between confining pressure and the cyclic shear stress required 5
to cause liquefaction (cyclic shear strength) are plotted on Figure 28.
The shaded zone shows the scatter of data for natural soils below the
_j hydraulic fill.
However, the non? Mear effects of the relationship between confining pressure and the cyclic shear strength can best be
]
estimated by selecting the stress ratios from data on each individual test presented on Figures 23,24, 25, and 26.
These four tests represent
]
four different confining pressures ranging from 2,000 to 8,000 psf.
A design strength curve was drawn on Figure 28 by selecting the four different stress ratios from Figures 23 through 27.
It can be seen that the design curve chosen represents a lower bound of strength q
to almost 4,000 psf of confining pressure.
(This range of confining pressure represents the crucial depths up to 50 feet where liquefaction potential is of primary concern.) A field strength curve corresponding
_l to 57 percent of the laboratory triaxial strength curve has been drawn L
on Figure 28.
It is this curve that was used to determine the strengths j
at various depths.
I 10.4.8 Factor of Safety Computation. The cyclic shear stress required q
to cause liquefaction at a particular depth is found from the field strength curve of Figure 28 by reading off the ordinate corresponding
]
to the confining pressure at that depth.
The variation of the cyclic F~
shear strength, the induced average she r stress (obtained by the one-dimensional wave propogation analysis), and their ratio (that is, the H
factor of safety against liquefaction) with depth are summarized in a
1--
Table 6.
The cyclic shear stresses and strength are also presented as a function of depth on Figure 22.
Table 6 and Figure 22 show that even with a very conservative interpretation of strength data, no lique-faction is predicted up to an acceleration level of 0.20 g.
I 10.4.9 Discussion and Conclusions.
During the current investigation, the liquefaction potential at the LACBWR site was studied using a simpli-j fied approach and a rigorous approach.
In the simplified approach, stresses were computed using empirical equations and strengths were
=
N h
1173 077 60
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e
=
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Y I
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52 I
E g
5 E
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d o Od 8
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EE
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$5e bl'a y0 lb I
182 8?3
/
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8 egi
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2
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h
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s I
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5
~
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m I
a a
i 5
m
~
~
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W (15d) H1DN381S HV3HS Ol'IDA3 g
l 1175 078 I
61 FIGURE 28
M M
M M
M M
M M
M M
M M
M M-M M
M M
M TABLE 6
SUMMARY
OF LIQUEFACTION ANALYSIS APPROACil 2 Avera_ge_ Cyclic Shear Stresses and Factors of Safety for Various Accelerations
- a
= 0.10 g a,,,
- 0.12 9 a
= 0.14 g a,,,
= 0.16 g a 9 " max 9
max max max FS
'av FS
'av FS
'av FS
'av FS
'av FS Deptl_Qt[
.v _ gin fl
'av "U
l 10 150 72 2.08 86 I.74 101 1.49 114 1.32 128 1.17 141 1.06 20 250 134 1.H7 163 1.53 189 1.32 214 1.17 238 1.05 262 0.95 30 350 194 1.R0 233 1.50 26R 1.31 302 1.16 340 1.03 367 0.95 40 460 244 1.R9 290 1.59 332 1.39 371 1.24 411 1.12 450 1.02 50 590 284 2.08 339 1.74 387 1.52 433 1.36 479 1.23 525 1.12 60 730 322 2.27 384 1.90 439 I.66 490 1.49 539 1.35 SR7 1.24 Q
70 860 156 2.42 420 2.05 478 1.80 532 1.62 585 1.47 639 1.35 80 1,050 381 2.76 449 2.34 512 2.05 573 1.83 633 I.66 695 1.51 90 1,240 407 3.05 482 2.57 550 2.25 62,0 2.00 684 1.81 750 1.65
- r
= average cyclic shear stress = (maximum cyclic shear stress from one-dimensional analysis) x (0.65).
av
[
= cyclic shear strength = (triaxial cyclic shear strength) x (0.57).
T Factor of safety (FS) = (cyclic shear strength) > (average cyclic shear stress).
N u
O N
E I
estimated using past experience during earthquakes at various sites that liquefied and also using relative densities.
In the rigorous approach, stresses were computed using a one-dimensional model and wave propagation response analysis.
The strengths were measured by performing cyclic triaxial tests on undisturbed samples.
The following conclusions were based on these analyses:
a.
Using Approach 1, Procedure 1 (Seed and Idriss stress and strength based on SPT, N values), no liquefaction is predicted up to a maximum acceleration level of 0.12 g.
The potential I
for liquefaction is suggested for various maximum surface acceleration levels as stated below:
a depths prone to liquefaction max 0.10 g none 0.12 gnone 0.14 g 20 to 30 feet I
0.16 g 20 to 30 feet 0.18 g 10 to 40 feet 0.20 g 10 to 40 feet.
I b.
Using Approach 1, Procedure 2 (Seed and Idriss stress and strength based on relative densities), no liquefaction is predicted up to a maximum acceleration level of 0.16 g.
The potential for liquefaction is predicted for various maximum I
acceleration levels as stated below:
a depths prone to liquefaction I
max 0.10 g none 0.12 g none 0.14 g none 0.16 g 20 feet 0.18 g 20 to 30 feet 0.20 g 20 to 40 feet.
I c.
Using Approach 2 (stresses from one-dimerisional wave propagation analysis and strength from stress-controlled cyclic triaxial I
tests peformed in the laboratory on undisturbed samples),
no liquefaction is predicted up to a maximum acceleration level of 0.20 g as stated below:
I 63 1!173 080' I
I I
a depths prone to liquefaction max 0.10 g none
. I 0.12 g none 0.14 g none 0.16 g none 0.18 g none l
0.20 g 20 to 30 feet.
The rigorous analysis made during the current investigation is relatively more accurate than all other analyses made at the LACBWR plant site.
We believe that a high degree of confidence can be assigned to the rigorous analysis made during the current investigations for I
the following reasons:
a.
The test boring and sampling program was performe_d under care-I fully controlled conditions using state-of-the-art techniques.
b.
The undisturbed samples were drained and frozen at the site I
before transporting for storage and were kept frozen until just before testing (practically eliminating sample disturbance at the site).
I c.
The frozen samples were carefully packaged and transported by D&M field engineers to minimize any possible sample disturb-ance during transport.
I d.
State-of-the-art testing techniques were used to determine the in situ densities and the cyclic shear strengths of samples.
e.
All the field and laboratory investigations were subject to I
stringent quality assurance and quality control requirements of D&M, DPC, and NRC.
In summary, given our present knowledge and understanding of the seismicity of the region and the behavior of soils under dynamic loading, it is our opinion that there is little threat of liquefaction at the LACBWR site under a maximum acceleration level corresponding to a realis-tic design SSE that can be assigned to the site.
I I
g e4 I
1175 08I I
I I
11.0
SUMMARY
OF LIQUEFACTION ANALYSES AT THE LACBWR SITE I
The D&M analysis of 1973 (Ref. 1) was conservative and concluded that the factors of safety against potential for liquefaction under the design SSE at various depths were adequate.
WES performed a very conservative analysis (Ref. 3) and concluded that the minimum factor l
of safety was close to unity under a 0.12 9 ground acceleration.
As a result of D&M's review of its past work and the WES report, and reevalua-tion of the various analyses, it was concluded that the factors of safety were indeed adequate.
However, there were certain questions that were raised by NRC regarding the lack of test data on undisturbed I
samples and the lack of continuous standard penetration test results.
Since the existing data did not satisfy these new concerns, DPC decided to perform modest field and laboratory investigations and modest analyses to verify the earlier findings on liquefaction potential.
As a result of the new state-of-the-art investigations performed at
' LACBWR plant site, it is now concluded that the LACBWR plant site has an adequate factor of safety against potential for liquefaction under any realistic design SSE that can be assigned to the plant site.
However, in the absence of an NRC decieion regarding a design SSE and a corresponding design acceleration level, a range of acceleratico between 0.10 g and 0.20 g was assumed and factors of safety were estimated.
The minimum factors of safety are listed below for the different acceleration amax Depth (ft)
Factor of Safety 0.10 g 30 1.80 I
0.12 g 30 1.50 0.14 g 30 1.31 0.16 g 30 1.16 0.18 g 30 1.03 0.20 g 20-30 0.95 Based on these results, it can be concluded that the threshold l
liquefaction resistance at the LACBWR site occurs for a design SSE which yields a maximum ground surface acceleration greater than 0.18 g and less than 0.20 g.
1173 082 65
I REFERENCES 1.
Dames & Moore, Geotechnical Investigation of Geology, Seismology, and Liquefaction Potential, Lacrosse Boiling Water Reactor LACBWR)
Near Genoa, Vernon County, Wisconsin, October 1973 (prepared for Gulf United Nuclear Fuels Corporation).
2.
Dairyland Power Cooperative, Application for Operating License I
for the Lacrosse Boiling Water Reactor, 1974 (submitted to the U.S. Nuclear Regulatory Commission).
3.
Marcuson, W. F. and W. A.
Bieganousky, Liquefaction Analysis I
for Lacrosse Nuclear Power Station, U.S. Army Engineer Waterways Experiment Station, December 1978 (submitted to the U.S. Nuclear Regulatory Comission).
4.
Dames & Moore, Review of Liquefaction Potential, Lacrosse Boiling Water Reactor (LACBWR Near Genoa, Vernon County, Wisconsin, March 1979 (submitted to the U.S. Nuclear Regulatory Commission).
5.
U.S. Nuclear Regulatory Comission, letter of April 30, 1979 (Docket No. 50-409), to Dairyland Power Cooperative's General Manager.
I 6.
Seed, H. B. and I. b' Idriss, " Simplified Procedure for Evaluating So.' Liquefaction Potential," Journal of the Soil Mechanics and I
Foundations Division, ASCE, Vol. 97, No. SM9, Proceedings Paper 8371 (September 1971), pp. 1249-1273.
7.
Seed, H. B., " Soil Liquefaction and Cyclic Mobility Evaluation I
For Level Ground During Earthquakes," Journal of the Geotechnical Engineering Division, ASCE, Vol. 105, No. GT2, Proc. Paper 14380 (February 1979), pp. 201-255.
8.
Ohashi, M., T. Iwasaki, F. Tatsuoka, and K. Tokida, "A Practical Procedure for Assesing Earthquake-Induced Liquefaction of Sandy I
Deposits," Proceedings--Tenth Joint Meeting U.S.-Japan Panel on Wind and Seismic Effects (Public Works Research Institute Ministry of fonstruction, 1978).
9.
Marcuson, W. F. and A. G. F'anklin, " State of the Art of Undisturbed Sampling of Cohesionless Soils," Proceedings--International Symposium on Soil Sampling, Preprint, Singapore (July 1979).
- 10. Meyerhof, G. G., " Discussion of Gibbs and Holtz Paper," Proceedin 5 of 4th International Conference of Soil Mechanics and Foundation Engineering, Vol. III, London (1957).
- 11. Marcuson, W. F. III and W. A. Bieganousky, "SPT and Relative Density in Coarse Sands," Journal of the Geotechnical Engineering Division, I
ASCE, Vol.103, No. GT11, Proc. Paper 13350 (November 1977), pp.1295-1309.
56 l
1173 083
I 12.
Silver, M.
L., Laboratory Triaxial Testing Procedures to Determine the Cyclic Strength of Soils, NUREG-0031 (U.S. Nuclear Regulatory Commission, June 1977).
I
- 13. Gibbs, H. S. and W. G. Holtz, "Research on Determining the Density of Sands by Spoon Penetration Tests," Proceedings of 4th International I
Conference of Soil Mechanics and Foundation Engineering, Vol. I, London 1957).
14.
I Hardin, Bobby 0. and Vincent P. Drnevich, " Shear Modulus and Damping in Soils:
Design Equations and Curves," Journal of the Soil Mechanics and Foun.dations Division (American Society of Civil Engineers, July 1972).
15.
SW-AJA (Shannon-Wilson and Agbabian-Jacobsen Associates), Soil Behavior Under Earthquake Loading Conditions (Union Carbide Corpora-tion 1972), (submitted to U.S. Atomic Energy Commission).
16.
Schnabel, B., J. Lysmer, and H. B. Seed, SHAKE, A Computer Program for Earthquake Response Analysis of Horizontal Lavered Sites, I
Report No. EERC 72-12 (Earthq mke Center, University of California, 1972).
I 17.
Dames & Moore, SHAKE--One-Dimensional Wave Propogation for Multi-Layered Soil System, Computer Program EP55 (1975).
I I
I I
I I
I I
67 I
l'75 U84
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I I
I I
I BORING LOGS I
I I
I I
I I
I 1173 085 I
I I
KEY TO LOG OF BORINGS I
LEGEND:
12 3
INDICATES DEPTH OF STANDARD SPLIT SPOON SAf1PLE.
INDICATES fiUMBER OF ELOWS REQUIRED TO DRIVE STANDARD SPLIT SPOON ONE FOOT IN STAtlDARD PENETRATION TEST.
I If1DICATES DEPTH OF SPT SAMPLING ATTEMPT WITH NO RE-O C0VERY.
INDICATES DEPTH OF RELATIVELY UNDISTURBED SAMPLE 03-I p
TAINED WITH OSTERBERG PISTON SAf1PLER.
INDICATES THAT SAMPLE TUBE WAS PUSHED INTO S0IL BY HYDRAULIC PRESSURE.
I INDICATES DEPTH OF DISTURBED SAMPLE OBTAINED WITH OSTERBERG PISTON SAMPLER.
I I
ELEVATIONS REFER 10 THE USGS MEAN SEA LEVEL DATUM.
APPROXIf1 ATE LOCATIONS OF BORINGS ARE SHOWN ON PLOT PLAN.
CLASSIFICATION SYMBOLS REFER TO UNIFIED CLASSIFICATION SYSTEM, PLATE A-2.
DISCUSSION IN TEXT IS NECESSARY FOR C0f1PLETE UNDERSTANDING OF TPE SUBSURFACE f1ATERIALS.
I I
I H D U" g
PLATE A 1
SOIL CLASSIFICA GRAPH MAJOR O/ VISIONS Syygg S..s e e
o, e
GRAVEL CLEAN GR Av E L S *****
AND su't6s Pj!
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.RAVEtL, a.
p r
8,8,.,..,
- 301, j
g 88 e1 CO A R$1 at b
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S OIL S I
[
,,9,8
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,,g,,,,
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g
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SAND CL E AN SA%D AND
<6.v et s oo =o SANDY SOILS some v.a. so g or oa'teia6 is
','f,
',9 LA!113 '*** **
7 aos s.t va s at i*
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esoes team so%
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b 6, M,L HIGHLY ORGANIC SolLS Nb
-r-t _
15
'_X NOTES:
1 OuAL SvuSOL; ARE USED TO esD6CATE 80#0ERL4E CLASSrFCA 2 WMEN SHown 08s TME SOR@G LOGS, TME FOLLOwWG TERMS ARE ComS81TEseCY OF CO**Elevf $0sLS AmD TME #EtativE Cou*acTNE COME SivE S0ILS COME
( APP 90miMATE SME ARW.
STR E NG_T'* DeJF) vEnv SOFT LEsS Tum.t$
vEev L00$f SOF T OfS T0 05 LOOSE ut0AAA ST#F 0 $ TO 10 ut0evu DEN STFF 10 TO 2 0 DENSE l
YERT STIFF t o TO 4 0 vERYDENSE t
MARC saE ATER TMau 4 0 1173 087
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