STRESS-STRAIN-STRENGTH ANISOTROPY OF VARVED CLAYS
by
Surachat Sambhandharaksa
BE, University of New South Wales
(1967)
ME, Asian Institute of Technology
(1970)
SUBMITTED IN PARTIAL FULFILLMENT
OF THE REQUIREMENTS FOR THE DEGREE OF
DOCTOR OF SCIENCE
at the
MASSACHUSETTS INSTITUTE OF TECHNOLOGY
August 1977
-1
j2
.UJ-.
Signature of Author ._
_
Department of Civil Engineering
Certified by
Thesis Supervisor
Accepted by
I
I
I~
I
IIII
C~an,'nt
II
I
Committee on Graduate Studies
Archives
JAN 5
1978
STRESS-STRAIN-STRENGTH ANISCMT PY OF VARVED CLAYS
by
Surachat Sambhandharaksa
Submitted to the Department of Civil Engineering
on August, 1977 in partial fulfillment of the
requirements for the Degree of Doctor of Science.
ABSTRACT
The objectives of this thesis are: (1) to contribute to a
fundamental understanding of the nature of undrained shear
strength anisotropy of varved clays; (2) to present best estimates of normalized stress-strain-strength parameters for undrained total stress analyses and for predictions of pore pressures and undrained deformations used in the design of structures on Connecticut Valley varved clays, with background information showing how these parameters were developed; and (3)
to recommend suitable testing procedures for developing normalized soil parameters for other varved clay deposits.
To achieve these objectives, several types of K o consolidated and isotropically consolidated strength tests were performed at various overconsolidation ratios (OCR). These include:
(1) SHANSEP CIUC tests on Amherst and East Windsor samples with
different inclinations; (2) SHANSEP CK UC and CKoUE tests on
vertical Amherst and East Windsor samples; (3) SHANSEP CKoUDSS
tests on vertical Northampton, Amherst, and East Windsor samples;
and (4) SHANSEP CIUE tests on vertical and horizontal East Windsor samples. The testing program also includes CKoU plane strain
tests on normally consolidated Northampton samples and CKoUDSS
tests on normally consolidated inclined East Windsor samples.
For the study of suitable testing procedures of varved clay, RECOMPRESSION CKoU triaxial compression and extension tests at
OCR's of 2 and 4, and RECOMPRESSION CIUC tests on East Windsor
samples with different inclinations at OCR=4 (the in situ OCR)
were performed along with UUC tests on East Windsor samples with
different inclinations.
The three factors that control the anisotropic strength behavior of varved clays are: (1) the layered nature of the soil
(material anisotropy); (2) the prior Ko consolidation (structural anisotropy); and (3) when Kofl, the stress system induced
anisotropy. An attempt was made to separate the influence on
undrained shear strength (qf) of material anisotropy from the
effects due to structural and stress system induced anisotropy.
The anisotropy in qf/avm is due to the anisotropy in pore
pressure response and in the normalized effective stress envelope
(NESE) at qf, the plot of qfium versus pvm.
A major finding
is the fact the anisotropy in NESE at (/
_ is onlydue to material
anisotropy. Thus, there are three possible NESE at (/C3)max,
ii
namely: NESE for failure across the varves, that for failure
within a clay layer, and that for failure within a silt layer.
The first two NESE are shown to be basic material properties.
Test data on a block sample from East Windsor show that
disturbance is mostly due to separation of the varves and is
therefore directional dependent. With block samples, the SHANSEP method of consolidation is thought to yield reasonable estimates of undrained shear strength but the stress-strain characteristics are less brittle and less sensitive compared to the
in situ soil.
Chapter 8 presents: recommended normalized soil parameters
for the design of structures on Connecticut Valley varved clays;
type of tests for checking these parameters; recommended testing
procedures for developing design strengths for other varved
clay deposits; and suitable methods of stability analysis wherein
anisotropic strengths are taken into consideration.
Thesis Supervisor
Title
Charles C. Ladd
Professor of Civil Engineering
iii
ACKNOWLEDGEMENTS
During the preparation of this thesis, I have received help
and encouragement from many people and organizations.
Their
mention here cannot begin to express my sincere thanks for
their support.
Foremost among these is Professor Charles C. Ladd who is
my faculty and thesis advisor.
His guidance in the preparation
of this thesis, his understanding of my writing problems, the
encouragement that he provided in the difficult periods are
sincerely appreciated.
The Massachusetts Department of Public Works and US Department of Transportation provided financial support during the
first two years of this research.
Professor R.P. Long of the University of Connecticut provided two block samples of varved clay from East Windsor.
I greatly appreciated the long hours Mr. Jacques N. Levadoux spent with me in the laboratory.
I also would like to
thank former MIT students who contributed valuable test data
presented in this thesis.
They are David, H. Connell, Suzanne
M. Lacasse and Francis D. Leathers.
To my thesis committee, Professor T.W. Lambe and Professor
Mohsen M. Baligh, I owe a special debt of thanks for their
suggestions to improve the quality of this thesis.
Special
thanks are also extended to Dr. W.A. Marr who kindly reviewed
iv
this thesis.
The thesis was typed by Michele P. Jarrabet and the figures
were drafted by Nancy Iffland.
V
TABLE OF CONTENTS
Page
TITLE
PAGE.
ABSTRACT
. .
.
i
. . . . . . .
ACKNOWLEDGEMENTS.
. . . . . . . ..
OF
TABLES.
2
2.1
2.2
2.2.3
2.3
..
.............
1
. . . . . .
.
.. . . . .
.
. . . . . . . . . . . . . . . . . ..
. . .. . . . .. . . . . . . . . . . . .
6
6
6
Anisotropy in Undrained Shear Strength
. .
Measurement of Undrained Shear Strength
6
Anisotropy
9
..
. . . . .
The Normalized Soil Properties (NSP)
.
.
.
.
. . . . . . . . . . .
. . . .
13
17
18
18
of Anisotropic Behavior . . .
19
2.3.1 Analytical Approach
.
2.3.1(a) Curve Fitting Method
2.3.1(b)
.
.
Theoretical Prediction
2.3.1(c)
Factors Affecting the
Anisotropic Shear Strength
Behavior of Varved Clay
Experimental
Approach
. . . . .
.
O
..
2.3.2(a) Evidence from UUC Tests . . .
2.3.2(b) Evidence from CIUC Tests
. .
2.3.2(c) Evidence from CK
Tests
. .
C~~re,,,,,
,~
r,-3
0,·
Z.J.J
ummarUy anaL ;olCUsln
. . . . . . . . .
2.3.4 Effect of Certain Stress System Variables
on Undrained Shear Strength Behavior. . .
2.3.4(a) Effect of Intermediate
Principal
2.3.4(b)
3
1
2
3
. . . . . . . . . . . . . . . . . . .
.
. . . . . . . . . . .
.
Concept
. .
Literature
Review
.
2.3.2
x
xiv
...................
Introduction
Background.
2.2.1
2.2.2
vi
...........
Background
Objectives
Scope
.
BACKGROUND
iv
. . . . . . . . . . . . . . . . . . . . . .
INTRODUCTION
1.1
1.2
1.3
ii
...............
LIST OF FIGURES ....
1
..
...................
TABLE OF CONTENTS.
LIST
.. .
Stress
. . . . . . .
Effect of K o Consolidation
EXPERIMENTALPROGRAM ....
. .
21
23
23
25
29
32
34
35
36
60
vi
Page
...
. . . . . . . ..
General
Scope and Objectives
. ...
. . . . . . . . . . .
Program
Experimental
Description
of Varved Clays
. . . . . . . . . . .
Varved Clay
. . . . . . . . .
3.3.1
East Windsor
Amherst
Varved Clay . . . . . . . . . . . .
3.3.2
Clay .
. . . . . . . .
Varved
3.3.3
Northampton
3.3.4
Hackensack
Valley Varved
Clay . . . . . . .
of Soil Properties
. . . . . . . . . . . .
Summary
. . . . . . . . . . . .
Procedures
General
Testing
60
3.5.1
3.5.2
Consolidated Undrained Triaxial Tests .
Testing Problems with Consolidated
73
3.5.3
3.5.4
.
. .
..
Tests.
Triaxial
Undrained
. . . . . . .
The Direct
Simple Shear Test
The Plane Strain
Tests
.
... . . . . . .
77
78
79
3.1
3.2
3.3
3.4
3.5
4
60
64
66
67
68
70
71
72
COMPARISON OF NORMALIZED BEHAVIOR OF DIFFERENT
VARVED
CLAYS .
. . . . . . . . . . . . . . . . . . . .
and
Background.
. . 100
. . . . . . . ..
4.1
Introduction
4.2
Effect of Sample Stress History (MPPR) .
4.3
Effect
of Soil
4.3.1
4.3.2
Results from CIU Tests.
.
.....
Results from CKoU Tests
4.3.2(a) Data from CKoUC Tests . . . . ..
Location
100
... .
102
. . . . . . . . . . . . . 105
.... ..
105
107
107
4.3.2(b) Data from CKUDSS Tests . . . . . 108
Data from CKoUE Tests . . . . ..
4.3.2(c)
4.4
4.5
5.
109
Discussion
. . . . . . . . . . . . . .
.. . . . . . . . . . .
Summary
and Conclusions
.
111
115
EFFECT OF STRESS SYSTEM ON UNDRAINED BEHAVIOR FROM
* - .
145
.
Introduction .
Effect of K o Versus Isotropic Consolidation. ....
145
148
SHANSEP
5.1
5.2
SAMPLES
of
.
. . . . . 148
Method
5.2.2
Volume Change Data from CIU and CKoU
149
...................
5.2.3
5.2.4
Results from CIUC and CKoUC Tests . . ...
Results from CIUE and CKoUE Tests . . . . .
5.2.5
Summary
and
Discussion.
. .
. .
..
..
..
.
.
.
Effect of Intermediate Principal Stress . . . . .
5.3.1
Method
5.3.2
5.3.3
.
.. .161
Results from CIU Tests
Results from K Consolidated Plane
of
Strain
5.3.4
Investigation
and
Triaxial
..
Tests
. . . .
. .
.
. .
157
160
. 160
. . . . . . 165
DSS
Tests.
.
...
5.3.5 Summary and Discussion..
Effect of Testing Method for Shearing Parallel
to Varves
5.4.1
I50
154
Interpretation and Results from
Special
5.4
. .
Investigation
.
5.2.1
Tests
5.3
.
.
.
.
..
.
.
.
.. . . . .
Introduction and Method of Investigation
vii
170
181
182
182
Page
5.4.2
5.4.3
Results from CDDS and CIUC Tests
Results from CKoUDSS Tests and
Their
5.4.4
Interpretation
.
..
. .
.
. .. .
184
.
*
186
Comparison of Results from CIUC
(8=450)and CKoUDSSTests. . . . . . ...
6
5.5
General Discussion ........
5.6
Summary
and
CONSIDERATION
OF
6.1
6.1.5
. . . . . . . . . ....
DISTURBANCE
. . . . .
191
. . . . . 251
Experimental
Program
. . . . . . . ...
256
Effect of Methods for Minimizing Disturbance
on Undrained Shear Strength Behavior . . . ...
6.2.1 Behavior from Vertical Loading. . . . . ..
6.2.2 Behavior from Horizontal Loading. .....
6.2.3 Behavior from Inclined Loading. ..
.. .
6.3
6.4
256
257
260
262
Discussion
. . . . . . . . . . . . . . . . . . . . 263
Summary
and Conclusions.
. . . . . . . . . . . . . 266
INTERPRETATION OF ANISOTROPIC BEHAVIOR OF CONNECTICUT
VALLEY
8
SAMPLE
188
.190
Introduction and Background
.251
6.1.1 Purpose .
.......
251
6.1.2 Types and Causes of Sample Disturbance
. 251
6.1.3 General Effect of Sample Disturbance . . . 252
6.1.4 Methods for Minimizing Sample Disturbance
6.2
7
Conclusions.
. ..
VARVED
CLAYS
. . . . . . . . . . . . . . . . . . 292
7.1
7.2
Introduction
. . . . . . . . . . . . .292
Method
of Investigation
. . . . . . . . . . ...
294
7.3
7.4
7.5
Interpretation of "Material" Anisotropic Behavior
Effect of Structural Aniostropy Component . . ..
Interpretation of Combined Anisotropy
......
296
299
301
7.6
Summary
305
and
Conclusions
.
. . . . . . . . . . . .
EVALUATION OF UNDRAINED STRESS-STRAIN-STRENGTH PROPERITES
OF CVVC
FOR
USE
IN DESIGN
8.1 Objectives and Scope ....
8.2
.
.
.
. .
.
.
.
.
.
.
....
.
322
322
Development of Undrained Strength Parameters of
CVVC for Total Stress Stability Analyses Using
SHANSEP.
.
.323
.
8.2.1
Criteria for Developing Design Undrained
8.2.2
Anisotropic Su Parameters Developed by
Strength
Ladd
8.3
. .
Parameter
(1975)
. . . .
.
. . . . . .
324
. . . . . .327
8.2.3 Design Parameters Developed by the Author
Comparison of the Use of SHANSEP Anisotropic
Su Versus Corrected FV Strength Data .
. .
.
8.3.1
Introduction
8.3.2
Recommended Correction Factor ()
for Connecticut Valley Varved Clay. ....
331
. . . . . . . . . . . . . . . 331
viii
332
Page
8.3.3
Compatibility of Using SHANSEP
Anisotropic Su with M-P Analyses . . . . . 334
Evaluation of Undrained Shear Strength for
8.4
Total Stress Analyses in Practice . . . . . ..
335
8.4.1
8.5
Initial Undrained Shear Strength for
Total Stress Analyses (Undrained Case) .
8.4.2 Undrained Shear Strength for Total Stress
Analysis in Partially Drained Case . . ..
Recommended Parameters for Undrained Analyses
of Pore Pressure and Deformation.
.339
8.5.1 Parameters for Predictions Using
Elastic
8.5.2
..
. .
. .
. . . . .
339
. . . . . . . . . . . . . . . . . . 343
Recommended Procedures for Checking CVVC NSP
and For Developing Normalized Shear Strength
Parameters for Other Varved Clays . . . . . ..
8.6.1
Parameters
. . . . . . . . . . . . . . . . 346
Recommended Laboratory Testing Procedures for Developing Normalized Soil
Parameters for Other Varved Clay
Deposits.
10
SUMMARY
AND
APPENDIX
. . . .
CONCLUSIONS
REFERENCES
BIBLIOGRAPHY
345
Recommended Method for Checking Soil
8.6.2
9
336
Parameters for Finite Element Program
FEECON
8.6
Theory
335
..
. .
..
....
346
. . . . . . . . . . . . . . . 369
. . . . . . . .
. . . . .
. 381
. . . . . . . . . . . . . . . . . . . . . . . 387
A - NOTATION
. . . . . . . . . . . . . . . . . . . 388
APPENDIX B - BEST ESTIMATE OF STRESS-STRAIN-STRENGTH
DATA OF CONNECTICUT VALLEY VARVED CLAYS. . . .
395
APPENDIX C - SUMMARY OF TEST RESULTS FROM CONNECTICUT,VALLEY
AND'HACHENSACK VALLEY VAVARVED
CLAYS
.406
APPENDIX D - TABULATION OF TEST RESULTS FROM AMHERST AND
EAST
WINDSOR
VARVED
ix
CLAYS
. . . . . . . . . .
431
LIST OF TABLES
Table No.
2.1
2.2
2.3
2.4
Page
General Problems with Testing Methods for
Investigating Direction Variation of Undrained
Shear Strength (Su )
38
Prediction of Anisotropy in S from Hansen and
Gibson's Equation for Plane Srain Loading
39
Influence of Inherent Anisotropy on Measured
Undrained Shear Strengths of Homogeneous Clay
from UUC Tests
40
Influence of Inherent Aniostropy on Measured
Undrained Shear Strength of Non-Homogeneous
Clay from UUC Tests
41
2.5
Anisotropy in Su Measured from CKoUPSC, CKoUDSS,
CPSE
Tests (Option 4) of Several NC Cohesive Soils
42
2.6
Anisotropy in Su Measured from CK-UC, CK UE,
and CKOUDSS Tests (Option 3) of Several
Soft Clays
43
Summary of Anisotropic Data from Several
NC Cohesive Soils
44
2.7
2.8
Effect of Stress History on Anisotropic Shear
Strength Behavior of Boston Blue Clay Measured
from Plane Strain Compression and Extension
Tests
45
Effect of 2 on Vertical and Horizontal
Undrained Shear Strength Behavior of Several
NC Cohesive Soils
46
3.1
Summary of the Experimental Program
83
3.2
Summary of Index Properties and Engineering
Properties of East Windsor Varved Clay
85
Laboratory Testing Program of East Windsor
Varved Clay
86
Laboratory of Testing Program of Amherst Varved
Clay
87
Laboratory Testing Program of Northampton
Varved Clay
88
2.9
3.3
3.4
3.5
x
Table No.
3.6
3.7
3.8
4.1
4.2
4.3
4.4
4.5
4.6
4.7
4.8
5.1
5.2
5.3
5.4
5.5
Page
Laboratory Testing Program of Hackensack Valley
Varved Clay
89
Summary of the Index Properties of Varved Clay
Samples from Amherst, East Windsor, Northampton
and Hackensack Valley
90
Comparison of the Undrained Shear Strengths
frnmthe CKoUE Tests from Different Testing
Procedures
91
Experimental Program for Investigating the
Effects of Soil Samples' Stress History and
Soil Location on NSP
118
Effect of Sample Stress History on the NSP
of Amherst Varved Clay
119
Effect of Stress History on the NSP of Northampton Varved Clay
120
Effect of Sample Stress History on the NSP of
East Windsor and Hackensack Valley Varved Clay
121
Effect of Soil Location on NSP from CIUC
Tests on Normally Consolidated Samples.
122
Effect of Soil Location on NSP from CKoUC
Tests
123
Effect of Soil Location on the NSP from
CKoUDSS Tests
124
Undrained Strength Data on Samples from Two
Depths of Matagami Varved Clay
125
Experimental Program for Investigating the
Effect of Stress System on SHANSEP NSP
196
Comparison of NSP from CIUC and CKoUC Tests
on Connecticut Valley Varved Clays
197
Comparison of NSP from CIUE and CKoUE Tests
on Connecticut Valley Varved Clays
198
Comparison of NSP from CIUE (8=900) and CIUC
(e=00) Tests on Connecticut Valley Varved
clays (Vertical Loading)
199
Comaprison of NSP from CIUE (8=0) and CIUC
(8=900) Tests on Connecticut Valley Varved
clays (Horizontal Loading)
xi
200
Page
Table No.
Comparison of NSP from CKoU Triaxial,
Plane Strain, and Special DSS Tests on
Normally Consolidated Samples
201
Load Increments in CKoUDSS(C) and CKoUDSS(E)
Tests on Normally Consolidated Inclined
Samples
202
Estimation of Af from DSS Test Sample (=450)
Using NESE Assumption
203
Summary of the Effect of Ac2 on Vertical
Loading
204
Summary of the Effect of Aa2 on Horizontal
Loading
205
Summary of NSP from CIUC Tests on Inclined
Samples (0=450)
206
Results from Drained Direct Shear Tests on
Northampton Varved Clay for Failure Within
a Clay Layer
207
5.13
Summary of the Test Results from CKoUDSS Tests
208
5.14
Comparison of Typical Values of qf/vc,
Tff/avc at (a /a 3 )max and Af from CIUC Tests on
Inclined Samples and CKoUDSS Tests
209
Experimental Program for Investigating the
Effect of Method for Minimizing Disturbance on
Undrained Shear Strength Behavior
271
Comparison of Test Results from CKoU Tests on
SHANSEP and RECOMPRESSION Samples from Connecticut Valley
272
Comparison of Test Results from CIUC Tests on
SHANSEP and RECOMPRESSION Samples
273
Summary of Test Data from CIUC Tests for the
Investigation of "Material" Anisotropy
308
Summary of Measured and Estimated Data from
CKoUPSC, CK UDSS, and CKoUPSE Tests for the
Study of Combined Anisotropy
309
5.6
5.7
5.8
5.9
5.10
5.11
5.12
6.1
6.2
6.3
7.1
7.2
8.1
Types of Stability Analyses and Recommendations for Embankments on Connecticut Valley
Varved Clays
xii
349
Page
Table No.
8.2
8.3
8.4
Input Parameters for Connecticut Valley Varved
Clays for Undrained Deformation and Pore Pressure Analyses Using FEECON
350
Soil Parameters for the Estimation of Initial
Settlement and Pore Pressure
351
Determination
of Su at OCR = 1.0 from Plane
Strain Loading Using Strain Compatibility
Method
352
Determination of S at OCR = 1.0 from Triaxial
Loading Using Strain Compatibility Method
353
Determination of S by Using Strain Compatibility Method for riaxial Loading at OCR=2
354
Determination of Su by Using Strain Compatibility Method for Triaxial Loading at OCR=4
355
8.7
Determination of Su by the Author's Approach
356
8.8
Comparison of Su values from Ladd (1975),
Author, and Corrected Field Vane Strengths
356
8.5
8.6(a)
8.6(b)
xiii
LIST OF FIGURES
Page
Fig. No.
1.1
2.1
Locations of Soil Samples of Varved Clays from
Connecticut Valley
Definition of Undrained Shear Strength Aniso-
47
tropy
2.2(a)
In Situ and Applied Stresses in Field Situation
of Elements
2.2(b)
2.3
2.4
2.5
2.6
48
A, B, and C.
Testing Methods for Investigating the Undrained
Shear Strength Anisotropy of Varved Clays in
the Laboratory
49
Davis-Christian (1971) Elliptical Anisotropic
Strength Relationship
50
Modified Mohr-Coulomb Failure Criteria for CrossAnisotropic Soil (After Merkle, 1971)
51
Summary of Results from UUC Tests oh Connecticut
Valley Varved Clays (Data from Healy, 1967;
Ladd and Wissa, 1970; and Long, 1975)
52
Anisotropic Behavior of London Clay at the Site
of the Ashford Common Shaft Measured from CIUC
Tests on Vertical and Horizontal Samples (Data
from Bishop
2.7
53
et al., 1965)
Test Data from CIUC Tests on OC Welland Clay
Samples with Different Inclinations (From Lo,
54
1965; and Lo, 1966)
2.8
Anisotropy in qf and Af of Fulford Soft Clayey
Silt Measured from CIUC tests on Samples with
Different Inclinations (Data from Perry and
Nadarajah,
2.9
2.10
55
1974)
Test Data from CIUC Tests on Laboratory Prepared
Vertical and Horizontal Samples of OC Kaolin
(Duncan and Seed, 1966)
56
Test Data from CIUC Tests on Horizontal and Vertical Samples of a Laboratory Prepared Lightly
Overconsolidated Kaolin (Data from Mitchell,
57
R.J., 1972)
2.11
5
Effect of
c on the Measurement
of Undrained
Shear Strength Anisotropy of Connecticut Valley
Varved Clay (After Long, 1975)
xiv
58
Fig. No.
2.12
2.13
3.1
3.2
Page
Anisotropy in S (S =
fma) from CK UDSS Tests
on Samples withuDi~feren inclinations (Option
5) of Three Soft Clays (Data from Soydemir, 1972)
58
Anisotropic Undrained Strength Ratio. Versus
Plasticity Index
59
Results of Consolidation Tests on East Windsor
Varved Clay
92
Soil Profile, Atterberg Limits, Recompression
Ratio, and Virgin Compression Ratio of Amherst
Varved Clay
93
3.3
Soil Profile, Stress History and Undrained Strength
Data at Amherst, Massachusetts
94
3.4
Summary of Atterberg Limits and Compressibility
of Northampton Varved Clay (From Ladd and Wissa,
1970)
3.5
95
Summary of Stress History, Undrained Shear Strength
of Northampton Varved Clay (From Ladd and Wissa,
1970)
3.6
96
Summary of Consolidation Data at Boring 65 and
65A, Secaucus, New Jersey (from Saxena, et al.,
1974)
97
3.7
Schematic Diagram of Essential Elements in Triaxial
Extension Test
98
3.8
Diagramatic Drawing of Plane Strain Sample and
Loading Systems
4.1
Normalized Stress-Strain Curves from CK UC Tests
on Overconsolidated Amherst Varved Clay at Two
vm(L)/vm Ratios
99
126
4.2
Normalized Stress-Strain Curves from CKUDSS Tests
on Normally Consolidated Northampton Varved Clay
127
4.3
Effect of MPPR on Values of 3 G/avc from CKoUDSS
Tests on Northampton, Amherst, East Windsor and
Hackensack Valley Varved Clays
128
Comparison of Stress-Strain Curves from CIUC and
CIUE Tests on Normally Consolidated Vertical
Samples from Amherst, East Windsor and Hackensack
Valley
129
4.4
xv
Fig. No.
4.5
4.6
4.7
4.8
4.9
4.10
4.11
4.12
4.13
4.14
4.15
4.16
4.17
Page
a/C -E Curves from CIUC and CUE Tests on
Nrm~lly Consolidated Vertical Samples from
Amherst, East Windsor and Hackensack Valley
130
Comparison of Stress-Strain Curves from CIUC
Tests on Inclined Samples (450) of Normally
Consolidated Amherst and East Windsor Varved
Clays
131
Comparison of Stress-Strain Curves from CIUC
Tests on Horizontal Samples of Normally Consolidated Amherst and East Windsor Varved
Clays
132
Normalized Stress-Strain Curves from CKoUC
Tests on Normally Consolidated Vertical Samples
from Connecticut Valley
133
Normalized Effective Stress Paths from CKUC
Tests on Normally Consolidated Vertical Samples
from Connecticut Valley
134
Pf Versus qf at Maximum Shear Stress from CK UC
Tests on Normally Consolidated Vertical Sampes
from Connecticut Valley
135
Comparison of Normalized Stress-Strain Curves
from CK UC Tests on Overconsolidated Amherst
and Ease Windsor Varved Clays
136
A and c Versus OCR from CK UC Tests on Amherst and East Windsor Varved Clays
137
Effects of Soil Location on Normalized Effective Stress Envelope at Maximum Shear Stress
138
Normalized Undrained Shear Strength of Connecticut Valley and Hackensack Valley Varved Clays
139
Normalized Stress-Strain Curves from CK UDSS
Tests on Normally Consolidated Northampton,
Amherst, East Windsor and Hackensack Valley
Varved Clays
140
Versus av/avm of ConnectiNormalized Thma /a
cut Valley an~ aclRnsack Valley Varved Clays
from CKoUDSS Tests
141
Normalized Stress-Obliquity-Strain from CK UE
Tests on Overconsolidated Amherst and East
Windsor Varved Clays
142
xvi
Fig. No.
4.18
Page
Af and Ef Versus OCR from CK UE Tests of East
Windsor and Amherst Varved Clays
143
Comparison of the Normalized Undrained Shear
Strengths of Connecticut Valley, Matagami and
New Liskeard Varved Clays
144
Volume Change Data from CIU and CK U Tests with
Horizontal Varved (=0 0 )
'Samples from East
Windsor
210
5.2
Ko Versus OCR for Connecticut Valley Varved Clay
211
5.3
Normalized Stress-Strain Data from CIUC and CKoUC
Tests on Normally Consolidated Amherst Varved
Clay
212
Normalized Stress-Strain Curves from CIUC and
CK0 UC Tests on Overconsolidated (OCR=4) East
Windsor Varved Clay
213
A Versus OCR from CUC, CKoUC, CIUE and CKoUE
Tests on Connecticut Valley Varved Clays
214
e Versus OCR from CIUC, CKoUC, CIUE and CKoUE
Tsts on Connecticut Valley Varved Clay.
215
Typical Normalized Effective Stress Paths for
CIU and CK U Triaxial Compression and Extension
Tests on N 8 rmally Consolidated and Overconsolidated Connecticut Valley Varved Clays
216
4.19
5.1
5.4
5.5
5.6
5.7
5.8
5.9
5.10
5.11
5.12
Normalized Effective Stress Envelopes from CIUC
and CK UC Tests on Connecticut Valley Varved
Clay:s
217
Eu/qf Versus Aq/Aqf from CIUC and CK UC Tests
on Samples from Connecticut Valley 0
218
Eu/avc at e = 0.2% Versus OCR from CIUC, CKoUC,
CIUE, and CKoUE Tests on Samples from East
Windsor
219
Normalized Stress-Strain Curves from CIUE and
CKoUE Tests on Normally Consolidated East Windsor
Varved Clay
220
Normalized Stress-Strain Curves from CIUE and
CK UE Tests on Overconsolidated (OCR=4)
Conne8ticut Valley Varved Clays
221
xvii
Fig. No.
5.13
5.14
Page
EU/Su Versus Aq/Aqf from CIUE and CKoUE Tests
on Samples from East Windsor
States of Stress
for CIUC
CIUE (8=900) and CIUC (=
5.15(a)
5.15(b)
5.16(a)
5.16(b)
(=0e),
CiUE
(=00),
900) Test Samples
Normalized Stress-Strain Curves for Normally
Consolidated CIUE (=900) and CIUC (=00)
Test Samples (Vertical Loading)
5.18
5.19
5.20
5.21
5.22
5.23
5.24
223
224
Normalized Stress-Strain Curves for Overconsolidated CIUE (=900) and CIUE (=0 e) Test Samples
(OCR=4)
225
Normalized Stress-Strain Curves for Normally
Consolidated CIUE (=0 °) and CIUC (=90)
Test
Samples (Horizontal Loading)
226
Normalized Stress-Strain Curves for Overconsolidated CIUE (=00) and CIUC (=00) Test Samples
(OCR = 4)
5.17
222
227
Normalized Effective Stress Envelopes from CIUC
and CIUE Tests on Vertical and Horizontal Samples
of Connecticut Valley Varved Clays
228
Effect of Intermediate Principal Stress on Normalized Octahedral Stress Paths
229
Effects of a2 : Normalized Effective Stress Paths
from CIUC and CIUE Tests on Vertical and Horizontal Samples of East Windsor Varved Clay
230
Normalized Stress-Strain Curves of Normally
Consolidated Samples from CKoUC and CKOUPSC
Tests
231
Normalized Stress-Strain Curves of Normally Consolidated Samples from CKoUE and CKoUPSE Tests
232
Normalized Effective Stress Paths for Plane
Strain and Triaxial Shear or Normally Consolidated Connecticut Valley Varved Clays
233
Comparison of Eu/F. from eK U Plane Strain,
Triaxial and Speciai DSS Tests
234
Estimation of qf/Uvc of Overconsolidated Plane
Strain Horizontal and Vertical Loading Samples
from Connecticut Valley Varved Clays
235
xviii
Fig. No.
5.25
Page
Assumed State of Stress (Pure Shear~for Measuring Su(V) and SU(H) of Inclined (45 ) Normally
Consolidated Samples from CKoUDSS(C) and
CKoUDSS(E) Tests
236
5.26
NESE at (/3)max
237
5.27
Normalized Stress-Strain Curves of'Normally
Consolidated Inclined Samples from CKoUDSS(C)
and CKoUDSS(E) Tests
238
Comparison of qf/Uvc from CKoU Triaxial Tests
and CKoUDSS Tes'ton Inclined Samples from East
Windsor (Using NESE Assumption)
239
Comparison of qf/v
Values from NESE Assumption
and the Pure ShearVAssumption
240
NESE at qf and (l/Gmaxfrom CIUE Tests on Inclined Samples and CDDS Tests
241
NESE at (al/a3)max for Failure Within the Clay
Layer
242
Typical Normalized Stress-Strain Curves from
CKoUDSS Tests on NC and OC Connecticut Valley
Varved Clays
243
Method of Computing qmax from CKOUDSS Test- on
Normally Consolidated Amherst Varved Clay
(Assumed NESE at qf from CIUC Tests on Inclined
Samples)
244
Method of Computing qf from CKoUDSS Test on OverConsolidated (OCR=2) Northampton Varved Clay
at Th
(Assumed NESE at qf from CIUC Tests on
InclileM Samples)
245
Method of Computing qf from CKoUDSS Testson OverConsolidated (OCR=4) East Windsor Varved Clay
at (Th)pax (Assumed NESE at qf from CIUC Tests
on Incllned Samples)
246
Determinations of
Tests
247
5.28
5.29
5.30
5.31
5.32
5.33
5.34
5.35
5.36
5.37
5.38
for Failure Across the Varves
ff at (Th/av)max from CKoUDSS
Failure Strains at Thma and (Th/ )max from
CKoUDSS Tests on Connecicut Varved Clay
Computation of Relative Magnitudes of qf/v,
Tof
and Thmax/aVC by Assuming Two Locations
ofNESE
v/249
xix
248
Fig. No.
5.39
6.1
6.2
6.3
6.4
6.5
6.6
6.7
6.8
6.9
6.10
6.11
Page
Comparison of the Effect of Stress System on
qf/vc Versus OCR for Three Different Directions of Loading
250
General Effects of Disturbance on UUC Test,
RECOMPRESSION and SHANSEP samples.
274
Effect of Method of Consolidation on Vertical
Undrained -c- Strength Behavior from Triaxial Compression Tests on East Windsor Varved
Clay OCR = 4
275
Effect of Method of Consolidation on Vertical
Undrained a-c- Strength Behavior from Triaxial
Compression Tests on East Windsor Varved Clay
at OCR=2
276
Normalized Effective Stress Envelopes from CU
Tests on SHANSEP and RECOMPRESSION Varved Clay
Samples from Connecticut Valley (Data from
Both Amherst and East Windsor for SHANSEP
Samples)
277
Normalized Stress Paths from CK U Tests on
SHANSEP and RECOMPRESSION Samples from Connecticut Valley
278
Compari-sonof qf(V)/vc from CKoUC Test SHANSEPSamples and UU Test Samples from Amherst and
Northampton
279
Comparison of qf(V)/Uvc Versus OCR from C UC,
CIUC and UU(V) Tests on SHANSEP and RECOPRESSION Samples of Amherst and East Windsor
Varved Clays
280
Effect of Method of Consolidation on Horizontal Undrained -eStrength Behavior from
Triaxial Tests of East Windsor Varved Clay
at OCR = 4
281
Effect of Method of Consolidation on Horizontal
Undrained ao-- Strength Behavior from Triaxial
Tests of Connecticut Varved ClayTat OCR = 2
282
Eu/avc at 1/2 Aq/Aqf from CKoUC and CKoUE Tests
on SHANSEP and RECOMPRESSION Samples from Connecticut Valley
283
Comparison of qf(H)/a c Versus OCR from CKoUE
CIUC (=90)
and UU ( 90°) Tests on SHANSEP
and RECOMPRESSION Samples of Amherst and East
Windsor Varved Clays
xx
-
284
Page
Fig. No.
6.12
6.13
6.14
Normalized Stress-Strain Curves from CIUC (8=900)
Tests on SHANSEP and RECOMPRESSION Samples (OCR=
4) from East Windsor
285
Normalized Stress Strain Curves from CIUC (=90 0 )
Tests on SHANSEP and RECOMPRESSION Samples (OCR=
2) from East Windsor
286
Normalized Stress-Strain Curves from CIUC (8=450)
Test on SHANSEP and RECOMPRESSION Samples (OCR=
4) from East Windsor
287
6.15
NESE from CIUC (8=450 ) Tests on SHANSEP and RECOM288
PRESSION Samples
6.16
Comparison of qf/avc Versus OCR from CIUC (=450°)
KOUD~SS, and UU (450 ) Test onSHANSEP
and RECOMPRESSION
Samples from Amherst and East Windsor
289
Comparison of G/a
aty = 1.0% from CKdUDSS Tests
on SHANSEP and RESMPRESSION Samples
290
Comparison of qf/avc from CIUC and CK UC Tests on
SHANSEP and RECOMPRESSION Samples
291
Effect of Sample Orientation on Stress Versus
Strain from CIUC Tests on Samples of Normally
Consolidated East Windsor Varved Clay
310
Effect of Sample Orientation on Stress Versus
Strain from CIUC Tests on Samples of Overconsolidated (OCR=4) East Windsor Varved Clay
311
Anisotropy in qf/ayc and Af of Connecticut Valley
Varved Clay Resulting from Material" Anisotropy
Component
312
Normalized Effective Stress Paths and Envelopes
from Connecticut Valley Varved Clay at Different Varved Inclinations
313
"Material" Anisotropy in Pore Pressure and Pore
Pressure Parameter A ate = 1.0%
314
7.6
Anisotropy in Undrained Modulus from CIUC Tests
315
7.7
Comparison of Anisotropy in qf Measured from
Plane Strain and CIUC Tests
316
6.17
6.18
7.1
7.2
7.3
7.4
7.5
7.8
Effect of Structural Anisotropy Component
Anisotropy
on
317
in q
xxi
Fig. No.
7.9
Combined Anisotropy in qf of Connecticut Valley
Varved Clay Measured from CKoU Plane Strain
Tests
7.10
Estimated Combined Anisotropy in NESE from Plane
Strain Loading
7.11
Anisotropy with Respect to G/3v of Normally
Consolidated East Windsor Varvea Clay at Various
Applied Shear Stress Levels
320
7.12
of Normally
Anisotropy with Respect to G/.
= 0.1% 321
Clay at
VarvXa
East
Windsor
Consolidated
8.1
Consideration of Anisotropy in Undrained Shear
Strength in Morgenstern-Price Analyses (After
8.2
Recommended Su/avc for Total Stress Analyses by
Ladd
8.3
357
1975)
Ladd,
-
358
(1975)
Comparison of Simplified Bishop Circular Arc
and Morgenstern-Price Analyses (from Ladd and
Foott,
359
1977)
Comparison of Normalized Undrained Shear
Strength Data from UUV, FV and Thesis's Recommended Undrained Shear Strength Parameters
for Design
360
8.5
SHANSEP EU/Su Versus OCR for Connecticut Valley
Varved Clays
361
361
8.6
Recommended Initial Stress Ratio Versus OCR
for Connecticut Valley Varved Clays for Plane
Strain Loading
362
Settlement Ratio Versus Applied Stress Ratio
for Load on Isotropic Homogeneous Foundation
(from D'Appolonia, Poulos and Ladd, 1971)
363
Recommended Pore Pressure Parameter A Versus
OCR for Predictions of Pore Pressure in Connecticut Valley Varved Clays for Plane Strain
Loading
364
8.4
8.7
8.8
xxii
Fig. No.
8.9
8.10
8.11
8.12
9.1
9.2
9.3
9.4
Page
Recommended Undrained Hyperbolic Stress-Strain
Parameters for Use in Finite Element Analyses
with Connecticut Valley Varved Clays
365
Recommended Vertical Undrained Strength for
Use in Finite Element Analyses with Connecticut Valley Varved Clays in Plane Strain Loading
366
Recommended Elliptical Anisotropic Undrained
Strength Parameters for Use in Finite Element
Analyses with Connecticut Valley Varved Clays
367
Recommended Pore Pressure Parameter a Versus
OCR for Use in Finite Element Analyses with
Connecticut Valley Varved Clays
368
Material and Combined Anisotropy in q
Connecticut Valley Varved Clays at OC
377
of
= 4
Combined Anisotropy in qf and in NESE at qf
at (l/53)max
378
Recommended Anisotropic Undrained Strength
Parameters for Use in Total Stress MorgensternPrice Stability Analyses with Connecticut Valley
Varved Clays
379
Effect of Method of Consolidation on Vertical
and Horizontal a-c- Strength Behavior from CKoUC
and CKoUE Tests on Connecticut Valley Varved
Clays at OCR = 2
380
xxiii
1.
1.1
INTRODUCTION
BACKGROUND
A varved clay is a rythmite or a rythmically layered sedi-
ment comprised of two layers and a couplet.
generally glacial lake deposits.
Varved clays are
Each varve represents the
deposition during a year, and is composed of a lower part of
coarser grained material deposited in the summer and an upper
finer grained part deposited in the winter.
The composition of
varved deposits can vary from alternating layers of highly
plastic clay and clayey silt (typical of varved clays in northern US and Canada) to alternating layers of silt and medium to
fine sand.
The thickness of the varves also varies consider-
ably, although 1/4 to 2 1/4 in,is most typical.
There'is gen-
erally a gradual transition of the soil composition within a
varve, but there is an abrupt change between varves resulting
from the rapid influx of sediments during the runoff season.
The strength, deformation and permeability characteristics of
varved clays are highly anisotropic as a result of this layering.
A three-year research program entitled, "The Behavior of
Varved Clays in Civil Engineering Structures," sponsored by the
Massachusetts' Department of Public Works in cooperation with
the US Department of Transportation Federal Highway Administration, was conducted by the Constructed Facilities Division of
the Department of Civil Engineering at MIT.
1
The final objective
of the program was development of field and laboratory procedures that can be used with confidencein conjunction with
realistic theoretical models for the design of civil engineering structures, and especially embankments, on varved clays
encountered in the Connecticut Valley.
In order to achieve this
objective, it was necessary to:
1.
Develop a better understanding of the fundamentals
controlling the anisotropic engineering properties
of varved clay;
2.
Develop an experimental approach and theoretical
model for predicting the engineering performance
of varved clays;
3.
Evaluate the reliability of these procedures
based on case studies of the actual performance
of structures on varved clays.
At the conclusion of the above research project, Ladd
(1975) presented detailed recommendations for the foundation design of embankments constructed on Connecticut Valley varved
clays.
This report had to be prepared before research on the
fundamentals controlling the undrained stress-strain-strength
properties of varved clay, the topic of this thesis, was completed.
1.2
OBJECTIVES
The thesis concentrates on a study of the fundamentals of
the undrained shear strength anisotropy of varved clays.
2
The
objectives are:
1.
To contribute to a fundamental understanding of the
nature of undrained shear strength anisotropy of
varved clays;
2.
To present best estimates of normalized stressstrain strength parameters for undrained total
stress analyses and for predictions of pore pressures and undrained deformations used in the design of structures on Connecticut Valley varved
clays, with background information showing how
these parameters were developed;
3.
To recommend suitable testing procedures for
developing normalized soil parameters for other
varved clay deposits.
1.3
SCOPE
In order to achieve these objectives, it is necessary to
study:
1.
The normalized behavior of several varved clays
using consolidated undrained shear tests;
2.
The effects of stress system* and sample disturbance on the normalized stress-strain-strength
parameters of varved clays;
3.
Suitable laboratory methods for measuring the
undrained shear behavior of varved clays;
*This denotes the applied stress conditions during consolidation and shear.
3
4.
The fundamental factors controlling the anisotropic behavior of varved clays.
A laboratory testing program was conducted at MIT on four
varved clays.
These were located in the Connecticut Valley at
Amherst and Northampton in Massachusetts and at East Windsor,
Connecticut and in the Hackensack Valley at Secaucus, New
Jersey.
The geological formation of the latter soil is similar
to that of the Connecticut Valley.
Fig. 1.1 shows the location
of the three Connecticut Valley varved clays.
The basic background and literature review of the anisotropic behavior of cohesive soils are presented in Chapter 2.
Chapter 3 presents the general testing program and description
of the varved clays, while Chapter 4
investigates their norma-
lized soil properties (NSP). Chapters 5 through 7 study the
fundamental factors influencing the anisotropic behavior of
varved clays.
The recommended NSP are presented in Chapter 8
with the necessary background showing how these parameters were
developed.
4
cn
J
0
w
-J
Uo '
a< p
CfZ
LL0
z0O)
J
5
FIGURE
1.1
2.
2.1
BACKGROUND
INTRODUCTION
The object of this chapter is to provide basic background
on the subject of undrained shear strength anisotropy.
Since
the study of this subject yielded rather conflicting results,
an attempt is made to include a fairly detailed summary of
available information.
The normalized soil properties (NSP)
concept of Ladd and Foott (1974) will also be presented because
it is used throughout the thesis.
Experimental data on undrained
shear strength anisotropy are also reviewed.
Appendix A lists
the notation used in this thesis.
2.2 BACKGROUND
2.2.1
Ansiotropy in Undrained Shear Strength
Definitions:
The term anisotropy means that the proper-
ties vary with direction of the measurement.
Anisotropy of un-
drained shear strength denotes variation in the undrained shear
strength with the direction of loading.
The convention that
will be used to define direction is shown in Fig. 2.1.
The
angle B is the angle between the major principal stress at
failure (alf) and the vertical direction in the ground.
Aniso-
tropy of the undrained shear strength, therefore, implies a
variation in undrained shear strength with
measured in the
laboratory using tests with different directions of alf
and
with controlled magnitudes of the applied intermediate principal
6
stress
(A62).
Causes and Types of Anisotropy:
The anisotropic undrained
strength behavior of cohesive soil is intimately connected with
"soil structure" (Lambe, 1958) which in turn depends on the environmental conditions during soil deposition, environmental
and stress changes subsequent to deposition, and the applied
stress system during shear.
Stress changes subsequent to the
deposition most frequently result from changes in the overburden
pressure and in the ground water table elevation, including desiccation.
It is convenient,
according
to Ladd et al.
(1977),
to divide undrained shear strength anisotropy, as measured in
the laboratory, into three components, as follows:
(1)
Inherent Anisotropy -- This component of strength
anisotropy is a direct result of the in situ soil structure.
Soil structure is affected by the arrangement of the soil particles, which in turn depends upon the environmental conditions
during deposition and the geological stress history.
The one-
dimensional consolidation which prevails during the geological
stress history of most sedimentary clays causes the platelyshaped soil particles to have a tendency to align themselves at
right angles to the direction of the major principal stress
(e.g., Martin and Ladd, 1975).
As a result, when a clay is
loaded in a direction such that the failure surface is horizontal or nearly horizontal, it may have the lowest shear strength
(Perry, 1971; Merkle, 1971).
This type of inherent anisotropy
is therefore termed the "structural" anisotropy.
7
Inherent
anisotropy also occurs if the material is non-homogeneous.
The prime example of this type of anisotropy is with varved
clays due to the presence of clay and silt layers, each of which
has different
stress-strain-strength
characteristics.
'Thiswill be
In addition, each layer may also
called "material" anisotropy.
exhibit structural anisotropy.
(2)
Stress System Induced Anisotropy -- This component of
undrained strength anisotropy results from the fact that different increments of shear stress are required to produce failure
as the major principal stress at failure,
lf' varies between
the vertical and the horizontal direction, when the coefficient
of earth pressure at rest (K ) is not equal to unity.
Hansen
and Gibson (1949) first discussed the existence of a stress system induced anisotropy and predicted large variations in undrained shear strength with direction of loading for isotropic
soils (i.e., no inherent anisotropy) with constant values of
the effective strength parameters (c and
lent of Skempton's
(1954) pore pressure
) and of the equivaparameter,
Af.
essence, stress system induced anisotropy results in
in effective stress as
In
changes
lf varies in direction during shear.
(3) Combined Anisotropy -- "Combined" anisotropy (used in
place of the term "in situ" anisotropy because anisotropic behavior is generally measured via laboratory tests) includes
the total effects of inherent and stress system induced anisotropy.
Ko consolidated undrained (CKoU) plane strain tests*
i.e., plane strain compression (PSC) and plane strain extension
(PSE).
8
run on vertical samples should measure the combined effect of
inherent and stress system induced aniostropy for plane strain
loading conditions in the vertical and horizontal direction because the anisotropic fabric is oriented in the same direction
as
in situ,and because shear starts from a K
consolidation.
However, with CK U triaxial tests, the directional variations
in strength shown by comparing data from triaxial compression
(TC) tests ( = 0) and triaxial extension (TE) tests (B = 900)
result both from the total effect of anisotropy (inherent and
stress system induced) and the differences in the magnitudes of
the intermediate
principal
tests while Ao>1
Ao2
=
A
3
stress
(Aa2 ).
[Aa1 = a2
> Aa 3 in TE
in TC tests.]
Section 2.2.2 will discuss the problems associated with
various types of laboratory strength tests for predicting in
situ shear strength anisotropy.
2.2.2
Measurement of Undrained Shear Strength Anisotropy
Figure 2.2(a) shows states of stress for three typical
soil elements along a potential sliding surface beneath an embankment.
For simplicity, straight lines are used to represent
the failure surface.
The stress components shown reflect both
the initial in situ stress on the failure surface (i.e., Tao'
Tao) and the incremental (decremental) effects (AT-) due to subsequent undrained shearing caused by construction of the embankment.
At Point A, the major principal stress at failure is in
the vertical direction (i.e., B = 0)
while it acts in the hori-
zontal direction at Point C ( = 90°).
9
At an intermediate
location such as Point B, the major principal stress at failure
is inclined.
Thus, unless K
is unity, the principal stresses
are rotated during shear at Points B and either A or C.
Types of laboratory consolidated-undrained shear tests
that might be used to measure the undrained shear strength of
Points A, B, and C are shown in Fig. 2.2(b).
For Point A, a
vertical sample can be sheared in triaxial or plane strain compression (TC or PSC), after Ko consolidation for Method Al.
An
inclined sample can also be tested in a direct simple shear
(DSS) test for Method A 2 (Bjerrum, 1973).
At Point B, Ko con-
solidated undrained direct simple shear (CKoUDSS) tests run on
vertical samples, or isotropically consolidated undrained triaxial compression (CIUC) tests performed on inclined samples
might be used, Method B1 and B 2, respectively.
K o consolidated
undrained (CKoU) triaxial and plane strain extension (TE and
PE) tests on vertical samples and isotropically consolidated
undrained (CIU) triaxial compression tests on horizontal* samples
could be used to simulate the state of stresses at Point C for
Methods C 1 and C 2.
A CKoUDSS test on an inclined sample, varved
inclination 8 = 450, which is sheared by reversing the direction
of the applied shear stress can be used to measure the strength
at Point C for Method C 3.
Figure 2.2(b) shows the states of
stress acting on the soil samples during consolidation and subsequent shear for each method.
It should be noted, however,
*Horizontal sample means that the axis of the sample is cut
from the in situ horizontal direction (i.e., °0 = 90 for varved
clay; and the specimen varves are vertical).
10
that only the stress components that can be measured are shown
for methods involving the direct simple shear test.
Both conventional and sophisticated methods of investigating undrained strength anisotropy are summarized in Table
2.1.
The problems associated with these approaches are also
outlined.
Many investigators (e.g., Bjerrum and Landva, 1966;
Beere, 1969,; Bjerrum, 1973; and Ladd et al.; 1977) are convinced that undrained shear strength and its anisotropy can
only be reliably measured in the laboratory provided:
the
samples are consolidated anisotropically and sheared with the
same stress system as in the field.
These researchers have de-
veloped their "standard" procedures for measuring "anisotropy."
They include K o consolidated triaxial compression and extension
tests (or less frequently plane strain compression and extension tests) for vertical and horizontal loadings respectively
(Points A and C) and a direct simple shear test for inclined
loadings (Point B).
Since the CK UDSS test can cause a rota-
tion of principal planes during shear, Bjerrum (1973) also suggested Methods A 2 and C 3 in Fig. 2.2, i.e., Option 5 in Table
2.1.
Table 2.1 also lists problems associated with each option.
Unconsolidated undrained triaxial compression (UUC) tests, Option 1, can at best yield only crude estimates of anisotropy
because of sample disturbance.
Furthermore, shear starts from
an isotropic state of stress, and therefore neglects the stress
system induced component.
With CIUC tests, wherein samples
11
are usually consolidated to the effective vertical overburden
stress (avo) or less, Option 2, the effect of sample disturbance
may be somewhat reduced, but isotropic consolidation will undoubtedly cause some change in the initial structure anisotropy.
Also, as with Option 1, there is no stress system induced anisotropy during the actual shearing process.
Option 3 uses Ko consolidated triaxial compression and extension tests and the direct simple shear test.
three different
pal stress
(2
=
values
of the intermediate princi-
3 in triaxial
rect simple shear, and
1 =
This involves
compression,
a1
>
2 >
3 in di-
2 in triaxial extension),
and
thus, can not be directly used to obtain the in situ directional
variation of undrained shear strength unless variations in the
intermediate principal stress have little effect on the undrained shear strength (or the data are corrected for estimated
effect of this factor).
The DSS test also has problems with
interpretation of the results, since the state of stress at
failure is unknown.
parameters, c and
Consequently, the effective shear strength
~, and Skempton's pore pressure parameter,
Af, can not be directly measured (Ladd and Edgers, 1972).
Fur-
thermore, it is not certain what the measured maximum horizontal shear stress (thmax) represents.
It is generally less than
the maximum shear stress (qf) and may be, even less than the
shear stress on the failure plane at failure, Tff
5.4).
(see Section
Because of the presence of the cylindrical wire rein-
forced rubber membrane,the non-uniform stresses and possibly
12
progressive action (especially at the edges of the sample) exist
in the DSS test.
In Option 4, use of plane strain equipment often presents
problems due to friction forces that are generated along the
loading piston and along the vertical boundaries of the samples.
Unless these friction forces are measured, one is never sure
of the reliability of the data.
In Option 5, the direct simple
shear test has the same problems associated with
tation of test results, as discussed above.
interpre-
In addition, there
are two problems with tests simulating Points A and C.
The
first is that K o must be known or estimated in order to apply
the correct value of a
and Tao (see Fig. 2.2(b)).
The second
is that the consolidation stresses perpendicular to this plane
can not be controlled.
Thus, based on the above discussion, all options involve
problems of a different nature and magnitude.
Option 4 is
theoretically the best if reliable plane strain equipment is
available; if not, Option 3 or 5 offer the most attractive
alternatives.
2.2.3
The Normalized Soil Properties (NSP) Concept
The normalized soil properties (NSP) concept is used to
study the undrained shear strength behavior of varved clays
throughout this thesis.
In this section, the basic ideas of
the NSP concept arepresented.
For further information on this
subject, the reader is referred to Ladd (1971), Ladd and Foot
13
(1974) and Ladd et al.
(1977).
The NSP concept was developed from the following empirical
observation that applies to many clays:
the results of consoli-
dated undrained (CU) laboratory shear tests performed on
samples with the same overconsolidation ratio (OCR), but different consolidation stresses and therefore, different maximum
past pressures ( vm
vm), exhibit very similar strength and stressstrain characteristics when normalized with respect to the
consolidation stress.
Thus, it is possible to run laboratory
tests at various OCR values and develop an unique normalized
plot of soil parameters versus OCR.
Examples of normalized
soil parameters that can be used in such plots are the normalized maximum.'shearstress at failure (qf/avc), the normalized
undrained Young's modulus (EU/ vc), Skempton's pore pressure
parameter (Af), and
(K).
the coefficient of earth pressure at rest
With a knowledge of the in situ stress history, perti-
nent soil properties can then be computed.
Measurement of NSP
Laboratory measurement of engineering properties always involves effects of sample disturbance.
In cohesive soils,
sample disturbance reduces the undrained shear strength (qf)
as measured in UUC tests (e.g., Ladd and Lambe, 1963; Davis
and Poulos, 1967).
But observations of compression curves from
consolidation tests (Schmertmann, 1955) indicate that "disturbed"
samples have approximately the same virgin compression characteristics as those of undisturbed samples provided that the
14
consolidation stress is beyond the maximum past pressure (and
the degree of disturbance is not excessive).
Thus, the effect
of sample disturbance can be reduced if the soil sample is consolidated beyond the maximum past pressure into the virgin compression range.
Based on this observation, Ladd and Foott (1974)
suggest that for soils exhibiting normalized behavior, undrained strength parameters should be obtained from samples
which are consolidated according to the SHANSEP approach.
This
involves consolidation of soil samples to stresses in excess of
the in situ maximum past pressure in order to minimize sample
disturbance effects, and subsequently reducing the effective
stress to obtain samples at the required OCR values.
Thus, this
method not only reduces the effect of sample disturbance, but
it also provides a sample of known OCR.
For obtaining the
appropriate SHANSEP NSP, the samples must be K
consolidated
to at least 1.5 to 2 times the in situ a
vm and sheared with
stress systems similar to the in situ conditions. Also, use
of SHANSEP NSP (obtained from the above testing procedure) for
estimating the in situ soil properties (qf, Eu etc.) requires
a knowledge of the in situ stress history (vo and avm ), since
NSP are related to OCR.
Limitation of SHANSEP Method of Consolidation
With highly structured soils (e.g., "quick" clays and
naturally cemented soils), consolidation above the in situ
stresses will cause radical changes in soil behavior.
The
SHANSEP method of consolidation, therefore, does not apply to
15
such soil.
With these materials, the in situ stresses should
be used for consolidation (Berre and Bjerrum, 1973) although
this requires samples of exceptionally high quality if meaningful data are to be obtained.
Recent work by Sangrey (1975)
illustrated that the NSP concept could be used in cemented soils
provided the soil parameters were normalized with respect to
yield stress, defined as the maximum stress to which the samples
can be consolidated without changing its basic behavior.
In theory, if a soil truely exhibits
normalized behavior,
NSP from the SHANSEP method of consolidation should represent
the in situ soil properties.
In reality, almost all natural
clays will undergo some change in structure when they are consolidated above the in situ
vm.
As a result, NSP from the
SHANSEP samples, especially those of overconsolidated (OC)
soils,
may be different from those of the in situ soil.
How-
ever, this does not imply that the SHANSEP method of consolidation necessarily yields unsatisfactory NSP, especially when compared to those obtained from other methods (e.g., UU tests or
CU with the RECOMPRESSION* Method).
Since sample disturbance
also causes changes in soil structure, data obtained by the
other methods also yield soil properties which are different
from those existing in situ.
SHANSEP samples (a
In fact, for most NC soils,
> (1.5 to 2) avm) should have a structure
*RECOMPRESSION: the method of consolidation whereby the
sample is consolidated at the desired OCR with a vertical consolidation stress is less than the in situ maximum past pressure.
16
similar to in situ normally consolidated clay and hence yield
reasonable
NSP
(Ladd et al. 1977).
Chapter 4 will present the experimental evidence showing
that for the stress range used in the study (which is rather
small), the varved clays under investigation do indeed exhibit
normalized behavior, at least with regard to certain properties.
Chapter 6 then compares properties obtained via the SHANSEP
method with those measured using other procedures.
2.3
LITERATURE REVIEW
Two areas of research for the study of the undrained shear
strength anisotropy are considered.
The first, the analytical
approach, concentrates on the analytical presentation of undrained shear strength anisotropy and includes (1) curve
fit-
ting methods to describe anisotropy using empirical equations
and semi-theoretical relations; and (2) theoretical predictions
wherein, after hypothesizing possible causes of shear strength
anisotropy, "theoretical" anisotropic shear strength relations
are derived.
The second approach is experimental, in which
possible causes and the factors influencing undrained shear
strength anisotropy are investigated.
Section 2.3.1 will review the analytical approach, including both curve fitting methods (Section 2.3.1(a)) and theoretical predictions (Section 2.3.1(b)).
Possible factors affecting
the aniostropic shear strength behavior of varved clay, using
isotropic Mohr-Coulomb failure theory, are presented in Section
17
2.3.1(c).
from:
CKoU
Section 2.3.2 summarizes the experimental results
UUC and CIUC tests on samples with different inclinations;
triaxial, direct simple shear (DSS), and plane strain
tests on vertical samples; and special CKoUDSS tests on samples
with different inclinations.
Since the effect of stress system
on NSP is studied in Chapter 5, a brief review of the influence
of the intermediate principal stress and the K
consolidation
on the strength behavior is presented in Section 2.3.3.
2.3.1
Analytical Approach
2.3.1(a)
Curve Fitting Method
The first curve fitting method,proposed by Casa-
grande and Carillo (1944),for presenting anisotropic shear
strength behavior is the equation:
S
)
=
Su(H) +(Su(V)
- Su(H))cos2
(Eq. 2.1)
Their relationship was suggested originally as a working hypothesis without experimental justification.
Davis and Christian
(1971) later showed that experimental data from UUC tests on
several relatively homogeneous soft clays (e.g., Welland clay,
San Francisco Bay mud, and Kaolinite) fitted this relationship
reasonably well.
Bishop (1966) proposed a modification of Casa-
grande-Carillo equation:
Su(B)
= Su(V)
(1 - m sin2 )
(1 - n sin2 2)
(Eq. 2.2)
to fit UUC test data on London clay in which the parameters
Su(V), m, and n are determined from laboratory tests (Bishop
suggested using Su(V), S(45)
and S(H)
lishing these parameters).
18
as a minimum in estab-
Davis and Christian (1971) proposed an elliptical anisotropic strength model to use in bearing capacity analyses.
They modified the anisotropic yield criteria for metals (Hill,
1950 and Scott, 1963) to apply to cohesive soil as follows:
(a
Y
-
a
S (V)-S (H)2
x
u
u
2
+
2
a
xy b
-
a
2
2
(Eq. 2.3)
Su(V) and SU(H) = the apparent undrained shear strength
in compression
in the vertical
and
horizontal directions.
a and b = constants
S U (V)+Su (H
)
to be found experimentally.
= a, and b/a =
2
(45)
Su(V)S
U
(H)
u
In the polar plot such as shown in Fig. 2.3, the variation in
undrained shear strength with angle 2
is elliptical.
This
equation was found to fit most existing UUC test data (Davis
and Christian, 1971).
2.3.1(b)
Theoretical Prediction of Anisotropic Behavior
Hanson and Gibson (1949) recognized that K
con-
solidation could cause anisotropy with respect to undrained
shear strength even in soil having isotropic Hvorslev (1960)
effective shear strength parameters and Skempton's (1948) X
parameter (the ratio of the slope of a swelling curve to the
slope of the virgin consolidation curve).
They first predicted
the existence of stress system induced anisotropy (Ladd et al.
19
1977).
Their expression for Su(=qf) for any orientation of the
failure plane was in terms of the initial shear stress
the Hvorslev shear strength parameters, the angle between the
failure plane and the horizontal plane (a), and the change in
pore pressure due to
changes in total stress based on
Skempton's (1948) "A theory."
Duncan and Seed (1966(a)) re-
viewed their work and modified the Hanson and Gibson equation
to use
isotropic Mohr-Coulomb effective stress
strength parameters 4 and c and Skempton's pore pressure parameter Af.
The equation in trms
of c, ~ and Af for plane
strain condition is
S
suu (a)
c-
Cc cos 4 + 1/2 (1+Ko)sin 4 - sinf(2Af-l)
(a)2
[(S)
Vc
+
/2 -)
zivc
1-K
+ (
S (a)
-u _
(1-Ko)cos2(45
2 1/2
(Eq. 2.4)
2 )
Although the theory is simplistic in that Af, c and
are
assumed constant, and therefore independent of the direction of
loading, Eq. 2.4 predicts trends (shown in Table 2.2) which are
amazingly realistic compared to test data on homogeneous clays
reported
in Tables
2.5 and 2.6.
Baker and Krizek (1970) and Markle (1971) considered directional variations in 4 and c with failure plane inclinations
(a). They independently developed equations for calculating
the shear strength (qf) and the failure plane inclination in a
20
cross-anisotropic soil, given the values of Band pf (Markle,
1971), or
and 03f (Baker and Krizek, 1970) provided the varia-
tions of
and c with a are known.
The governing equations (Eqs.
2.5 to 2.7) from Markle's method (1971) are presented in Figs.
2.4(a) and 2.4(b).
Based on the modified Mohr-Coulomb (M-C)
failure criteria (i.e.,
and
vary with failure plane inclin-
ations (a)), for a given value of
, failure in the soil occurs
when the point representing the state of stress at failure
(tff and aff) on the failure plane at orientation 6 lies on the
corresponding failure envelope (only one envelope exists for
given values of
and
f).
As a result, the failure envelope
can in the general case, intersect or be tangent to the Mohr
circle.
It is seen that the resulting equations are difficult
to use (since the variations of
and c with inclinations of
failure plane (a) must be known) and
they require
know-
ledge of pf for computing qf and the failure plane inclination
(a). As a result, they have little practical value for predicting the variation in qf with
, though the basic causes of ani-
sotropy in qf are hypothesized.
2.3.1(c)
Factors Affecting the Anisotropic Shear Strength
Behavior of Varved Clay
The variables that could affect 4 and c of varved
clays were studied by Townsend et al. (1965). They presented
the following general expression indicating the possible variables in the case where failure planes cut across the varves.
tan ~ = (1 -nc) tan ~s + nc tan ~c
21
(Eq. 2.8(a))
c = (1 - n c) C s +
and
(Eq. 2.8(b))
c Cs
p and c = M-C effective stress strength parameters
where
of bulk material
%s and cs= M-C effective stress strength parameters
of silt portion
c and cc= M-C effective stress strength parameters
of clay portion
n c = % of clay layer in the failure plane
The following are the assumptions used for developing Eq. 2.8:
1.
The failure surface is a plane that intersects clay
and silt layers at the same inclination.
2.
Both silt and clay layer exhibit isotropy with respect to the effective stress strength parameters 4
,
and c (i.e., c
c
s
,
s
Cc
and c s are independent
c
of failure plane inclinations).
3.
The same off acts in both clay and silt layers.
The above equation suggests that in the case where nc is
constant, values of
and c of varved clays will be constant.
Thus, for homogeneous varved clay, there are three possible
values of
and c.
These are:
(1) for failure within the clay
layer (i.e., n c = 1.0); (2) for failure within the silt layer
(i.e., n
= 0); and
(3) for failure across
the varves.
Data in Chapter 5 will show, provided samples have the
same failure mechanism (i.e., failure across the varves or
failure within the clay layer), that values of
and c at
(a1/a3)ma of OC samples are constant and independent of failure
22
plane inclination.
But for other conditions, i.e.,
(c=O)
at (al/a3)max of NC soil or ' and c at qf (when they are not
equal to those at (al/a3)
), it is not true.
max
2.3.2
Experimental Approach
The purpose of the literature review in this section is
to synthesize experimental evidence from several cohesive soils
regarding:
(1) the possible causes of anisotropy in Su , and
the factors influencing anisotropic behavior measured in laboratory tests.
The presentation of results will be organized
according to the types of laboratory tests, namely:
UUC* and
CIUC tests on samples with different inclinations; and CK0U
tests such as those described
in Table 2.1.
Experimental re-
sults from homogeneous clays and from varved clays are included.
2.3.2.(a)
Evidence from UUC Tests
Table 2.3 summarizes the data from UUC tests on
differently inclined homogeneous samples (presumably measuring
the inherent anisotropy) of six clays.
The results illustrate
the possible trends that may be encountered from UUC tests on
homogeneous samples, as follow:
(1)
The isotropic trend:
Jacobson (1955) reported results
of 34 unconfined compression tests performed at various inclinations on sarhples.of post-glacial marine clay fromiStockholm.
As shown in Table
2.3, the variation
in qf is within
the limit
*The unconfined compression test is treated as a UUC test
where
oc is equal to zero.
c
23
of the mean error.
This apparent isotropic trend could be the
result of sample disturbance which tends to mask undrained
strength anisotropy of homogeneous clay (Ladd et al., 1977).
(2) The anisotropic trend.
As shown in Table 2.3, with
the exception of test results from Kaolin (Mitchell, 1967; and
Duncan and Seed (1966(b))and from Little Belt clay (Hvorslev,
1960), the horizontal strength is the weakest.
Therefore,
these data suggest, for homogeneous clays, as shown for UUC
test data, that the lowest strength will be either in the inclined or horizontal loading direction.
(ii) Non-Homogeneous Clay
Non-homogeneous clays include stratified soils as
well as varved clays.
The London clay which has laminations
and fissures is also included in this category.
Table 2.4 and
Fig. 2.5 summarizes data on non-homogeneous clays from several
locations.
These data show that samples inclined with
to 45 to 60 degrees had the lowest strength.
equal
The UUC results
for Connecticut Valley varved clays, summarized in Fig. 2.5,
are of particular interest.
The main points worthy for mention
are:
1.
The qf (45)/qf(V) ratio ranges from 0.3 to 0.5
2.
The anisotropic effect is more pronounced in
samples at shallow depth (WNI 50%) compared to
those at greater depth.
qf(H)
(3)
) ratio is greater than one for varved clay
located near the surface, and less than oneforsample
24
located at greater depth.
This difference sug-
gests that the stress history influences anisotropic behavior, i.e., the sample at shallow
depth probably has a higher OCR.
Sample disturbance in non-homogeneous clay may amplify the
inherent anisotropic effect as measured with UUC tests.
Ward
et al. (1965) concluded that opening of the laminations in London clay (caused by disturbance during sampling) was likely to
have little effect on shear strength unless the failure surface
was nearly horizontal (generally observed in inclined loading)
in which case the shear strength could be appreciably reduced.
Disturbance due to separation of clay and silt layers in varved
clay during sampling and trimming may also cause a significant
reduction
in qf(4 5 0 ) while qf(V) is slightly
affected
(see
Chapter 6 and Section 2.3.2(b)).
2.3.2(b)
Evidence from CIUC Tests
The anisotropic behavior from CIUC tests on
samples with different inclinations was measured for both undisturbed samples (Bishop et al., 1965; Lo, 1966, and Perry and
Nadarajah, 1974) and laboratory prepared Kaolin samples (Duncan and Seed (1966(a)) and Mitchell (1972). Figures 2.6 to 2.10
summarize their results.
Test data from Connecticut Valley
varved clays (Long, 1975) are presented in Fig. 2.11.
For the
laboratory prepared Kaolin samples (Duncan and Seed, 1966(a);
and Mitchell, 1972), the test specimens were anisotropically
consolidated and then extruded so that the effective stress
25
immediately after sampling was due to a negative pore pressure
and was therefore isotropic.
The specimens were then isotropic-
ally consolidated to stresses less than the a
of the block
sample before they were sheared, with the exception of two
samples (vertical and horizontal) reported by Mitchell (1972)
where c was greater than
vc of the block sample.
The undis-
turbed soils are the OC Welland clay from Ontario (LO, 1966);
the heavily overconsolidated London clay at Ashford Common
Shaft (Bishop et al., 1965); an overconsolidated soft clayey
silt from Fulford, England; and the Connecticut Valley varved
clays.
The available index properties and the UUC test data
(only from Welland Clay (Lo, 1965) and Connecticut Valley
varved clay Long (1975)) are also presented.
These CIUC data, mostly for OC vertical and horizontal
samples, show the following:
1.
Data from London clay (Fig. 2.6) and laboratory
prepared Kaolin (Figs. 2.9 and 2.10) show that
the
failure
envelope in terms of effective
stress (i.e., the effective stress strength
parameters
and
) was essentially the same
for vertical and horizontal samples.
The differ-
ences in qf are therefore the result of the
differences in Af and hence pore pressure developed during shear.
Data from the soft clayey
silt from Fulford (Fig. 2.8) show only slight
differences in failure envelopes from vertical,
26
inclined, and horizontal samples (practically
insignificant considering the difficulty in
determining
and c from the scatter of lab-
oratory data).
For the East Windsor varved
clay, the anisotropy in qf is due to both the
anisotropy in effective stress strength parameters and the pore pressure (Fig. 2.11).
2.
The data from UUC and CIUC tests on Welland
clay (Fig. 2.7) and the East Windsor varved
clay (Fig. 2.11) show different trends with
respect to anisotropy
in qf.
It is believed,
however, that sample disturbance affected
the UUC test results so as to decrease the
anisotropy in qf for the Welland clay and
to increase it for the East Windsor varved
clay
3.
(see Chapter
6).
Based on test data from Welland clay (Fig. 2.7)
and laboratory prepared Kaolin (Fig. 2.10),
the excess pore pressure versus axial strain
relation is unique for samples tested with
various orientations, though Af
samples
4.
is different
of these
(since Ef is different).
For both homogeneous and non-homogeneous soils
(Figs. 2.6 to 2.11), the vertical
strength
(qf(V)) is larger than the horizontal strength
(qf(H)).
27
5.
An increase in the magnitude of isotropic consolidation stress ( ) in CIUC tests decreases
the anisotropic shear strength behavior.
Data
from Figs. 2.6 and 2.8 to 2.11 indicate that
qf(H)
qf(H)
qf(V)
Af(H)
andA
) approach unity as
Af (V)
c
reaches
vm for undisturbed soil or approach or exceeds
vc of the block for laboratory prepared
VC
Kaolin samples.
Test data from prior K
con-
solidated laboratory prepared Kaolin (Mitchell,
1972; in Fig. 2.10) in fact, showed that when
a
c
exceeds avc of the block, the Kaolin behaved
like an isotropic material.
The stress-strain-
pore pressure (a-e-u) behavior and thus the
stress paths (plotsof q versus
a
+
3
53
), of
horizontal and vertical samples (shown as the
dashed line in Fig. 2.9) are essentially identical.
Therefore, the data in Figs. 2.6 and
2.8 to 2.11 (especially
those in Fig. 2.9) sug-
gest that the structural anisotropy component
resulting from prior K
consolidation which
causes anisotropic behavior is reduced (or may
even be completely eliminated if a exceeds avm)
c
stresses.
vmconsolidation
by high isotropic consolidation stresses.
28
2.3.2(c)
Evidence from CK U Tests
The total effect of the combined anisotropy
(that resulting from the inherent and the stress system induced components) and the difference in the applied magnitudes of a2 (only when comparing data from TC and TE tests)
on variations of S
with
, is measured with CKoU tests.
This section will present the NC and OC anisotropic shear
strength behavior data obtained from Options 3, 4, and 5
of Table 2.1.
The variations of S
U
with loading direction,
of several soft clays from these options are presented in
Tables 2.5, 2.6, and 2.7 and Fig. 2.12.
The data in Table
2.6 and Fig. 2.12 were obtained from the RECOMPRESSION ( VC =
avo ) method of consolidation used
by NGI.. The SHANSEP
method of consolidation (avc > (1.5 to 2.0) a
) was used
to obtain test data in Tables 2.5 and 2.7 for CK U plane
strain and traixial test results.
Su(DSS)*
The strength ratios
qf(H)
and (V)
from Options 3 and 4 (excluding Option 5
because of problems in the interpretation of test data) are
presented in Figs. 2.13(a) and 2.13(b).
The data from
Tables 2.5, 2.6, and 2.7 and Figs. 2.13(a) and (b) show that:
1.
The anisotropy in S
is more pronounced in the
lean sensitive clay (having lower PI) than in
the plastic insensitive clays.
2.
For both plane strain and triaxial loading ,
*See the definition
in Tables
29
2.5 and 2.6.
SU(H) of homogeneous clays is lowest, partly
because of the effect of the rotation
of prin-
cipal planes and, for the triaxial tests, also
3.
due to the difference in the magnitude of a2.
qf(H)
qf(V) ratio is affected by the type of test
qf(v)
(i.e., plane strain versus triaxial) and (perhaps) the method of consolidation.
at same PI
(Fig. 2.13(b)),
Compared
SHANSEP CK U
o
strain (Option 4) tests yield higher
plane
qsf(H)
qf:(V
ratios than those from RECOMPRESSION (a =a )
VC VO
C--U triaxial (Option 3) tests,and to a lesser
0
extent, the SHANSEP CK U triaxial samples.
Su(DSS)
4.
Su (V)
ratio
is apparently
affected
little
by the type of test and the method of consolidation.
The test data from CK UDSS tests (Option 5, Fig. 2.12)
give the variation of qf with sample inclination rather
than the loading direction.
It is seen that the maximum
Su from CK oUDSS (C) tests occurs between sample inclinations
of 40 to 45 degrees,
and the minimum
S
from CK UDSS
(E)
tests occurs between sample inclinations of 50 to 60 degrees.
Thus, the use of sample inclinations of 450 (see
Chapter 5) should yield S
values close to S maxfrom
CK0 UDSS (C) tests and somewhat larger than Su min from
CK UDSS (E) tests.
30
Table 2.7 also presents a summary of effective stress
CKoUPSE, CKOUC and CK0 UE tests
strength data from CKoUPSC,
o
on SHANSEP normally consolidated samples of several soft
clays.
It is seen that the direction of loading can cause
differences in both
and Af.
It should be noted that, at
the same av , a smaller Af in CK UPSE tests compared to in
CKoUPSC tests, such as occurred with the San Francisco Bay
Mud and AGS CH clay,does not mean that the extension tests
had smaller pore pressure.
Because of the larger magnitude
of Aq required to cause failure, Pf/Ov
CK0UPSE tests (when K
value from the
<1) can be lower than that from
CK UPSC tests (i.e., the effect of stress system induced
anisotropy).
To illustrate this point,
Pf/vc
values from
San Francisco Bay Mud (Duncan and Seed, 1966(b))are given
below:
Af
qf/avc
Pf/vc
0
CK UPSC
1.12
0.12
0.60
380
CK UPSE
0.70
0.57
0.49
350
Type of Test
o
Based on test data in Table 2.7, the differences in Af, Yf*
and ~ from vertically and horizontally loaded samples are
more pronounced in lean sensitive clays than in plastic insensitive clays.
The stress history can also affect the combined
*Defined in Table 2.7, based on linear isotropic elastic theory.
31
anisotropic behavior.
Table 2.8 shows the effect of OCR on
the qf, Af and Ef ratios from plane strain compression and
extension tests on Boston Blue Clay.
anisotropy in qf and
increasing OCR.
It is seen that the
f decrease, but Af increases with
The fact that the effect of the rotation
of principal planes is more pronounced in NC samples than
in OC samples (since K
increases with increasing OCR) partly
explains such trends.
2.3.3.
Summary and Conclusions
Based on the review test data from UUC, CIU and CKoU
tests, the possible factors affecting the anisotropy in Su
measured
1.
in the laboratory
are:
Sample disturbance:
For homogeneous clay,
sample disturbance tends to mask the anisotropic effect measured with the UUC tests.
However, for non-homogeneous clay, sample
disturbance may
increase
the
anisotropic effect.
2.
The Magnitude
of
c:
The use of high a ,
especially using ac >avm in CIUC tests on
samples with different inclinations tends to
minimize, if not completely eliminate, the
structural anisotropy resulting from prior
in situ K
less K o
consolidation.
As a result, un-
o
is equal to 1, it is not possible
32
to truly separate the effect of stress system
induced anisotropy from the combined anisotropy by comparing test data from CIUC
tests
(avc = avo ) on samples with different inclinations (presumably measuring the inherent anisotropy), and from CKoU tests on vertical
samples (measuring combined anisotropy).
Anisotropy is found to be more pronounced in the lean
sensitive soils than in highly plastic clays.
Based on
limited data, the anisotropic behavior of cohesive soil is
also dependent upon the in situ stress history.
The magnitude
the q(V)
of the
2 used in the test can affect
ratio... Test data from CKoU plane strain and triqf(H)
axial tests show that the qf(v)
ratios from plane strain
tests are somewhat larger than those from triaxial tests.
Such differences are thought to be due to the differences
in the magnitudes of a2 between triaxial and plane strain
loading.
Data from both plane strain and triaxial tests on NC
sample show that the directional variation of S (qf) with
B
is due to the directional variations of both the effec-
tive stress strength parameter ~ at qf and the pore pressure.
The directional variation in pore pressure is in
turn the result of the directional variation of Af, and
the difference in the magnitudes of
Aqf.
The larger mag-
nitude of Aqf in a horizontally loaded sample, compared to
33
that in the vertically loaded sample, can cause larger pore
pressures and a lower effective stress at failure, even
though the Af for horizontally loaded sample is lower.
For OC samples, based on data from CIUC tests on RECOMPRESSION samples of several homogeneous cohesive soils with
different inclinations, the anisotropy in qf is due only to
the anisotropy in pore pressure.
strength parameters
The effective stress
and c at qf are practically unaffected
by sample inclination.
However, one should realize that
CIUC test do not cause a rotation of principal planes during
shear and thus, the directional variations in 4 and c could
occur in the combined anisotropic behavior of OC soil.
The interpretation of the anisotropic behavior of both
NC and OC Connecticut Valley varved clay measured using
SHANSEP method of consolidation will be presented in Chapter
7.
2.3.4
Effect of Certain Stress System Variables on Undrained
Shear Strength Behavior
This section
summarizesexperimental data showing
the general effects of intermediate principal stress and K
0
consolidation on undrained shear strength behavior.
The
effect of the intermediate principal stress is shown by
comparing test data from CK0 U triaxial and plane strain
tests.
The results from NC
and CKoUC (Ao>A
=A3)
T-rU PSC tests (A >a2>A
)3
tests are compared for the inter-
mediate principal' stress'effect in vertical loading.
34
For
horizontal loading, results from CK UPSE tests (Aa1 >Au2 >AG3 )
o CKUE
tests (A=Aa2>A
of K
3)
are compared.
(versus isotropic
consolidation
The influence
consolidation)
is
shown by comparing test data from CIUC and CKoUC tests, and
from CIUE and CK UE tests for vertical and horizontal load-
o
ing respectively.
2.3.4(a)
Effect of Intermediate Principal Stress
Tables
2.9(a) and 2.9(b) summarize data from
CK UPSC versus CK UC tests and from CK UPSE versus CK UE
o
o
o
o
tests on several normally consolidated undisturbed and remolded clays.
For San Francisco Bay Mud, Atchfalaya Clay,
and remolded Weald Clay, only data for vertical loading
are available.
These data show that shearing in plane
strain compared to that in triaxial:
1.
Increases qf for both vertical and horizontal
loadings.
Such increases are larger for hori-
zontal loading (up to 25%) thanfor the vertical loading
2.
(an increase
of about 10 to 15%).
Generally decreases Yf* for both vertical
and horizontal loadings.
Such decreases are,
however, larger for vertical loading than for
horizontal loading.
3.
Significantly increases
loading,
though
4 at qf for vertical
4 at (l/3)ma
x is only
*Defined according to Table 2.9 for triaxial and plane
strain tests based upon linear isotropic elasticity theory.
35
slightly increased.
the difference in
For horizontal loading,
at qf between TE and PSE
test is somewhat smaller.
4.
Does not consistently affect Af.
is not to say, however,
This
that increase
in a 2
will not increase Au.
For an ideal isotropic soil, the intermediate principal
stress will not affect the stress path in terms of octahedral stress (c
t
versus act)
ference in the magnitude of
For such a material, dif2 does not change the pore
pressure parameter "a" defined below:
Au =AcT
+ a Ar
oct
oct
(Eq. 2.9)
For vertical loading, test data from CK UC and CK UPSC
o
o
tests on remolded Weald clay and natural Hanney clay show
unique stress path in terms of octahedral stress.
For hori-
zontal loading, however, (i.e.,the data from CKoUPSE and
CKoUE test), the octahedral stress paths are far from unique.
2.3.4(b)
Effect of K
Consolidation
Ladd (1971) summarized the effect of isotropic versus K
consolidation undrained (CU) triaxial compres-
sion tests on several normally consolidated clays.
Accord-
ing to Ladd (1971), the results from NC clays generally
show that K
consolidation relative to the isotropic con-
solidation:
1.
Has a variable effect on qf/Ovc
36
the change
being + 10 to 15%;
2.
Generally decreases
3.
Causes a substantial change in stress strain
and Af at qf;
behavior, with a much smaller strain at failure,
and more strain softening; and
4.
Little
change in
at (/a3)max
For horizontal loading, NC test data on Boston Blue
clay (Ladd and Varallyay, 1965) indicate a significant decrease in qf and Af and a slight increase in
with K
con-
solidation.
For OC samples, test data from CIUC, CK UC, CIUE and
CK UE tests on Kaolin (Amerasinghe and Perry, 1975) indicate
essentially the same trends as reported by Ladd (1971) and
Ladd and Varallyay (1965) with the exception that for the
same loading direction (i.e., TC or TE tests), the isotropically consolidated specimens had the same failure envelope
(i.e., the same
and c).
37
-
>1
0(i
4-)
U)
.H 'Hl
,
H
rzl
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9
a)
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--
- --- - ---
38
---
---
Table 2 1
PREDICTION OF ANISOTROPY IN Su FROM HANSEN AND GIBSON'S
EQUATION FOR PLANE STRAIN LOADING
(DUNCAN AND SEED, 1966(a))'.
Assuming
Af = 1.0; Ko = 0.5; c = 38°;
f
o~~
Loading Direction
a
= 0
S u (a)/Cvc
Su)
S (a=64° )
Vertical
loading(i)
64
0.37
1.0
0
0.22
0.59
-26
0.20
0.54
Inclined loading(
Horizontal
loading(3)
.,
Notes:
.
.
.
(1) i.e., Data from PSC test
(2) i.e.,
Data from DSS (=0)
(3) i.e.,
Data from PSE test
test
Table 2.2
39
INFLUENCE OF INHERENT ANISOTROPY ON MEASURED UNDRAINED SHEAR
STRENGTHS OF HOMOGENEOUS CLAY FROM UUC TESTS
Relative Undrained Shear Strength
References
Clay
Vertical
akobson
(1955)
post glacial
1.0
450
1.13+
Horizontal
.02
1.05+ .05
marine clay
vorslev
(1960)
Vienna
1.0
0.92
0.87
Hvorslev (1960) Little Belt
1.0
1.08
1.20
Duncan and Seed
S.F. Bay Mud
1.0
0.86
0.81
Kaolin
1.0
0.75
0.87
itchell (1967) Kaolin
1.0
0.87
1.06
'Appolonia
Resediment
(1968) Boston Blue
Clay
1.0
0.9 +0.03
(1966(b))
Duncan and Seed
(1966(a))
0.82+0.03
Table 2.3
40
INFLUENCE OF INHERENT ANISOTROPY ON MEASURED UNDRAINED
SHEAR STRENGTHS OF NON HOMOGENEOUS CLAY FROM UUC TESTS
Investigators
Lo and Milligan
Relative Undrained Shear Strengi I
Vertical
450
Horizontal
Clay
Welland (Block F)
1.00
0.63
0.56
0.99
Heavily OC
London Clay
1.00
0.60
--
1.23
Varved clay from
Surte, Sweden
1.00
0.82
--
0.93
Varved clay from
Step Rock Lake
1.00
0.51
0.53
1.18
(1967)
Ward et al.
(1965)
Jakobson
(1952)
Eden (1955)
,~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~
,.
Table 2.4
41
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*)
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0
90
ANGLE
O0
gVrtical Direc
tion
-
?
e
DEFINITION
OF UNDRAINED
SHEAR STRENGTH ANISOTROPY
47
FIGURE
2.1
l
Q
.Mc
Fe
I.s
r 7 T'b--' ~
v
t,
%.UK I
QU
vK
\
4
n a%
0:
Q
Nrrr
l*.
IF
g l~~~~b-S[[LllL~·
'tpl
il
4.
0%
bbI.
·-
ani~~~~~~
1!11-4
tiii'-
q
0
±
s
I4
t
+
+
+
bq
±
C
I
1l3X[
g
0
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z
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!a ZCL))-
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o
jl4.
I
, -'--- /J
w
e z
CD
Cl)
W 2
z
w
Ew
Cfw
I
o0
a
J
LL.
0
0
.0
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z
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4n
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C-
0 <
<J
w
Z L
0
a
-
48
wJ
cnoo
0
IL
Q
a0 P(1Icrwfcr
x
O0
-
I-1:z
az
4
Ci
M:
0
w a cL
C
oCO
2
:3
z
CD
t r
oc
t
Cl)
w
.~~
X: Z
~ =:
w
_
CX
.0
FIGURE 2.2
I a --------
a ----
am
b s
404
f. (v)J
O.SClc)
s ()
____
?oszt(;;)3;c;
V)H
L
a= cos2l 4=a
where
A= s,() su(H)
L=
=
DAVIS-CHRISTIAN
2 /r3o
/+0)n2
ngyle between
and oaf
vartical
(1971) ELLIPTICAL
ANISOTROPIC
STRENGTH RELATIONSHIP
49
FIGURE 2.3
In Sitv VertIcal Directiotn
o(= -s
r. 2.5S
7f
alf -
a) DF/N/ITION OF fAILlRE
PLANE ORIENTAT/ON
rf
circle
*o-t
_ I
3f
from &eome try
cowi=I f f
in
r
si (26- )
Eq. 2.6
Fq. 2.7
b) STATE OF STRESS AT FAIL RE IN AN A/SOTROP/C
CLAYr
MODIFIED MOHR-COULOMB FAILURE CRITERIA FOR
CROSS ANISOTROPIC SOIL (After Merkle,1971)
FIGURE 2.4
50
Z5
SYM44. SOIL LOCATION
0
/lorthampton
0
East Windsor
/3
-
12
- I
REFERENCES
Ladd W,
fHealy,
/967
East Windsor Lon9 , 1975
/./
WN= SO%
-
depthr 7 ft iron surfrce
-- -. =60 / dA/th = -_,
/12 . IV,,,
frnm
10
1970
_N-\-
0.8
urfAro
__,
Ivx
ll
I
i)
7kt
I;
from
.
m
1"
07
su(V)
5,,
'v) 06
OS
0.4
0.5
m
0.2
m
0.1
0
0
I
/0
I
I
20. 30
I
40
I
I
50 60
I
70
I
80 90
P°
SUMMARY OF RESULTS FROM UUC TESTS ON
CONNECTICUT VALLEY VARVED CLAYS (Data from Healy,
1967; Ladd and
Wissa, 1970; and Long, 1975)
51
FIGURE
2.5
U.
oO
B
I
10
0..
S%
/
0
z
0
ox
I147
7/
'I'4
/8
a
'I
.41
OZ
I-0 2
i
Z:
0
Z ,¢~
4-
I
I
I
co
e.
(
o
() 0
rI
I*.-~
OU)
z U
0
_ot
Q<
O m
L L~~ II ~
LL
a
II
II
Ili
z cc~ U
0
0
FIGURE 2.
*_
ZIG
52
0
I
U)
w
e
a.
3
o0I
zj
0
_J
o E
0
~"
L
U.LL..
O
()Z
, I- Z
Iz
z
'I
LLL
4%
-
XDW(aUo)~
53
_
FIGURE
2.7
·
/. W
/.0
qf (v)
0.8
0.6_
40
0
/60
120
80
e,
kN/mn
A/ (4.5)
At(s)
al
Z2 _
i/
0
/.
Af a)
Af(V) 10
0Z0
"-
A (cV)
"0--
/
o
AtCH)
_
.tIN
rA
f j
-0
/I 80
I
II
40
I
i20
i , ktN/mZ
ANISOTROPY IN qf AND Af
OF FULFORD
/60
2OO
SOFT CLAYEY
SILT MEASURED FROM CIUC TESTS ON SAMPLES WITH
DIFFERENT INCLINATIONS (Data from Perry 8 Nadarajah,
54
54
1974)
FIGURE 2.8
-
:l
I
II!0)
zQ
tN
-
0
IIi
Crl-
0
'0
o
0
10
1b
1
I
I
1;
0
O- U
.
o
0
t- c
I
r1i
Cik
IION
Irs
OD
:kL
cn
C
o
I
all.
CD
(0
Z
0
z
ZY
r%
.
0
0MO.- 0
0in -J
0
w
.4
0
0
I-
r
LJ
111
-<
1o cn
coQ "
0
v<
Q
NoB)t
k
4%
0
o
k"
';
o
N
t
F
C
0
n
w
I
Z
11)
I
O
s
i
U
.
I
I
I
I
'6 eI
681e,
55
FIGURE 2.9
\I
4
LL
0
I'Z\s
NL
11
I)cn
w
4
Ca
z
J
0
he
-J
tvU 411
U
q'N
N
4
w
ti
w
z
i%
00,
C-
rl% 0-
z
N
4
I
a
I
I
'
'is
I
11
4
t
to
.N
t
N
z0
0 0wo
%4
IIcn
4
(IU)
tyI-.b
: 0)
wIM
Z%'
-
,
I
I
.4
e
I
N
4
II
0-
O O
llJ
'4
:
-
-i
IVI
4.
>
I-
.
I-
4
<cr
cr
C0
I' "
"
C
.I
'9
'.'1
a.
J-io
A.
Ik
4s
<
FIGURE 2.10
i
Aq0
-a
c
REMARK
2.6
22
28
0
0
0
C/i/c Tests
I
0
45
/2.0 I
90
.
l
!
._.-
cc
.-
_ .- -_--
o = 60 psi, OCR=/.0
/0.0
I
=8.5 psi
v
r
I
I.
8.0 '
&V," = 3-P pr
%
1
x
qf
5- =20,psi
I -
\~~elr _... m.mEU"°S
O.C
in psi
-011t=1.7
OCR
I
.
-1
I
0
ut/C T#'sts
2.0
0o
EFFECT
I
0
I.
20
OF &c ON THE
I.
i
I.
I.
40
,a
I
60
I
I
80
MI EASUREMENT OF UNDRAIINED
SHEAR STRENGTH ANISC)TROPY OF CONNECTICUT
VALLEY VARVED CLAY (After Long, 1975)
FIGURE 2.11
57
I-
ii E
L.
0--
z-J
00
CA L
W) cn
w
I- w
W
cn
0-ZC
cr
'6-
O
LL
-
ZL
-z
zz
icr
0 W.
LL
58
FIGURE
2.12
C
('a)
/.0
0.8
U
Jo.--
-l
(DSS)
sL (V)
0.6
-
~
0.4
a FronmTable 2.5
0 From 7able 2.6
-
0o.
0
I
0
I
o
,
40
I
I
60
80
/0
PLAS/C/ITY INPEX,PI %
O Plane Strain
(b)
o0 Trioial
1.
0
q (H)
q ()
0.u
·
0
0.6
0.4
04P
o.z
-_
13 From
O
*0
abl/e 2.7 (a
*· From Table 2.6
0 FromTble 2.7(h)
0
I
0
I
I
I
o0
80
PLA T/CITY INDEX, Pi %
20
40
1lg
UNDRAINED STRENGTH RATIO
VS PLASTICITY INDEX
ANISOTROPIC
59
FIGURE 2.13
3.
3.1
EXPERIMENTAL PROGRAM
GENERAL SCOPE AD
OBJECTIVES
This chapter provides a general description of varved
clays from four locations, namely:
East Windsor in Connecti-
cut, Amherst and Northampton in Massachusetts, and Hackensack
Valley in New Jersey.
The similarity in the general proper-
ties of these soils is discussed.
The experimental program,
and the details of the testing conditions used for the soil
samples at each location are presented.
The last portion of
the chapter describes the standard testing procedures used
for triaxial compression, triaxial extension, direct simple
shear, plane strain compression and plane strain extension
tests.
3.2
EXPERIMENTAL PROGRAM
The plan
of the laboratory testing program was to make
a comprehensive study of the anisotropic behavior of the four
varved clays.
Additional test results for the Northampton and
and Hackensack Valley varved clays are reported in Lacasse
et al. (1972) and Saxena et al. (1974).
Tables 3.1(a) and
3.1 (b) present a summary of the experimental program.
consisted
1.
of:
A series of unconsolidated-undrained triaxial
60
It
compression (UUC) tests run with different
orientations of the varves;
2.
A series of isotropically consolidated-undrained
triaxial compression (CIUC) tests with pore
pressure measurement run with different orientations of varves and at varying
3.
A series of K
OCR;
consolidated-undrained triaxial
compression (CK UC) tests with pore pressure
0
measurement at varying OCR;
4.
A series of isotropically consolidated-undrained
triaxial extension (CIUE) tests with pore pressure
measurement run on vertical and horizontal
samples at varying OCR;
5.
A series of K
o consolidated-undrained triaxial
extension (CKoUE) tests with pore pressure measurement at varying OCR;
6.
A series of Ko consolidated-undrained direct
simple shear (CKoUDSS)
7.
tests at varying OCR;
One special Ko consolidated-undrained direct
simple shear test run on a normally consolidated
sample where the normal and shear stresses were
applied during consolidation before the sample
was sheared by increasing the horizontal shear
stress to simulate plane strain compression
(CKoUDSS
8.
(C) test);
One special Ko consolidated-undrained direct
61
simple shear test run on a normally consolidated
sample where the normal and shear stresses were
applied during consolidation before the sample
was sheared by reversing the direction of the
horizontal shear stress to simulate plane strain
extension (CKoUDSS (E) test);
9.
One K
consolidated plane strain compression
(CKoUPSC) test run on a normally consolidated
sample from Northampton;
10.
Two Ko consolidated undrained plane strain
(CK0 UPSE) tests run on a normally consolidated
sample from Northampton;
11.
A series of consolidated drained direct shear
(CDDS) tests at different OCR on Northampton
varved clay;
12.
Two isotropically consolidated drained triaxial
unloading (CID (U)) tests on Northampton varved
clay.
Both SHANSEP and RECOMPRESSION* methods of consolidation were
used for items 3 and 5.
The locations and the testing condi-
tions of soil samples are presented in Section 3.3.
Chapter 4 will present the investigation of the normalized behavior of these soils, and a comparison of their NSP.
*RECOMPRESSION--the method of consolidation whereby the
sample is consolidated at the desired OCR with a vertical
consolidation stress that is less than the in situ maximum
past pressure.
62
The
testing program included CKoUC, CKoUE, and CKoUDSS tests
on varved clays from East Windsor, Amherst, Northampton, and
Hackensack Valley. NC CIUE test vertical samples and NC CIUC
test samples with different inclinations, of East Windsor and
Amherst varved clays, were also tested to compare their NSP.
All samples were consolidated according to the SHANSEP approach.
The effect of stress system on the NSP, obtained via the
SHANSEP approach, of Connecticut Valley varved clays is presented in Chapter 5.
1.
This includes:
The effect of initial shear stress (i.e., isotropic versus K
consolidation);
2.
The effect of the intermediate principal stress;
3.
The effect of testing method for shearing parallel
to the varves.
CK UC and CIUC tests on vertical samples which had OCR's of
1,2, and
4 were run to study the effect of initial shear
stress on NSP for vertical loading.
For horizontal loading,
CK UE and CIUE tests were run on vertical samples with OCR's
of 1, 2, and 4.
CIUC tests on vertical
samples,
and CIUE
tests on horizontal samples, at OCR's of 1, 2, and 4, were
run to study the effect of the magnitude of the intermediate
principal stress for vertical loading.
Similarly, CIUC
tests on horizontal samples and CIUE tests on vertical samples
at OCR's of 1,2, and 4, were performed to study the effect
of
2
on NSP for horizontal loading.
Results from CKoUPSC
and CK UPSE tests (Lacasse et al., 1972) were also used to
o
63
compare with data from CK oUC and CKo UE tests for the investigation of the effect of
2'
The special CK UDSS (C) and
CK UDSS (E) tests on inclined samples from East Windsor also
O
served this purpose.
and CK UDSS
CIUC tests on inclined samples (8=45),
tests on vertical samples at varying OCR, were
performed to study the effect of testing method for shearing
parallel to the varves.
Test results from drained direct
shear tests on samples from Northampton (Lacasse et al., 1972)
were also used for this purpose.
Chapter 6 will study the effect of "sample disturbance"
on NSP.
UUC tests on samples with different inclinations,
CIUC tests on samples with different inclinations consolidated
tests on verby the RECOMPRESSION method and CK 0UC and CK oUE
0
tical samples consolidated by the RECOMPRESSION method were
performed for this study. The NSP from these tests are then compared with the NSP from SHANSEP samples which have the same OCR.
Interpretation of the anisotropic undrained strength behavior of varved clays, which is based upon the results from
Chapter 5, is presented in Chapter 7, while Chapter 8 provides
the "best estimates" of soil parameters for use in the design
of structures on varved clays.
3.3
DESCRIPTION OF VARIOUS CLAYS
This section provides a general description of the soil
samples and their index properties.
64
Details of the geological
formation of the Connecticut Valley varved clays are presented
in Ladd and Wissa
(1970); the geological
formation
of the
Hackensack Valley varved clay is discussed in Saxena et al.,
(1974).
Briefly, the Connecticut Valley varved clays were deposited in glacial Lake Hitchcock during the late Pleistocene
time.
The bedrock of this area was covered by a widespread
but discontinuous blanket of till.
After deposition of the
varved clays, glacial Lake Hitchcock was drained when the natural dam at Rocky Hill, Connecticut breached.
As a result,
the varved clays along the Connecticut River were then deeply
eroded.
The eroded surface was subsequently covered by post
glacial sand deposits.
The upper present-day surface of the
varved clay is thus not necessarily the same as its original
surface under Lake Hitchcock.
The varve thicknesses vary
typically from 1/2 to 2 1/2 inch.
The thickness often de-
creases with increasing elevation because of the disappearance
of the glaciers with time.
The Hackensack Valley varved clay was obtained from the
Hackensack Meadows.
It is a tidal marsh occupying the area
of a glacial lake formed by the obstruction of a preglacial
valley with glacial drift in the vicinity of Perth Amboy.
In recent times, extensive areas of the meadows have been
covered with uncontrolled fill varying in thickness from 2
to 12 ft.
65
3.2.1
East Windsor Varved Clay
Two undisturbed one cubic foot block samples of East
Windsor varved clay from the same depth, were provided by Prof.
R.P. Long of the University of Connecticut.
They were ob-
tained from a clay pit located off Route 5 and owned by the
Kelsey-Ferguson Brick Co. in East Windsor, Connecticut.
The
clay is greyish in color and the varves show a gradual transition in grain size from top to bottom; but upon drying, the
"clay" and "silt" layer can be easily distinguished by their
dark and light colors, respectively.
The thickness of these
layers generally varies from 1/16 in. to 3/16 in.
The block samples were taken from the clay pit at a
depth of about 11 ft.
Five ft of medium sand and six ft of
brown varved clay overlie these samples.
They have an esti-
mated effective overburden pressure of 0.6 kg/cm 2.
The samples
generally had very uniform properties, based on torvane
strengths and natural bulk water contents.
Only portions of
these samples which had the same torvane strength and moisture
content were used for testing.
Table 3.2 shows the index and the engineering properties
of the East Windsor varved clay.
plasticity chart
According to Casagrande's
with the basis of limited data, the clay por-
tion can be classified as CH material and the silt portion
can be classified as ML material.
Fig. 3.1 shows results
from three standard consolidation tests on samples from these
two blocks.
The strains at the end of primary consolidation,
66
determined by the %'t~method, are used in the plot of the vertical strain versus log avc
A small variation in the values of
the maximum past pressure (avm) and the virgin compression
vm
ratio (CR) is noted in these tests.
clay is 2.4 kg/cm 2.
The average avm of this
The in situ OCR is therefore about four.
Table 3.3 gives the details of the testing conditions and the
types of tests performed on the samples.
3.2.2
Amherst Varved Clay
Under sponsorship of the Massachusetts Department of
Public Works and the US Department of Transportation Federal
Highway Administration for research on varved clays, the Department of Civil Engineering at MIT obtained 5 in. diameter
undisturbed samples from a boring near the Route 116 by-pass
and North Hadley Road at the University of Massachusetts, Amherst.
These samples were obtained by using an Osterberg
type fixed piston sampler.
Field vane tests using the Geonor
apparatus were also run adjacent to the boring.
The clay is generally grayish and brown in color, and
the varve thickness varies from 1/2 in. to 1 in.
The varve
thicknesses of the Amherst varved clay are larger than those
from East Windsor.
Figure 3.2 shows the soil profile, Atter-
berg limits and compressibility data.
The soil profile con-
sists of a thick deposit of varved clay overlain by 3 ft of
clayey sand.
The bulk moisture content and plasticity index
first increase and then decrease with depth (below El. 95),
67
but such changes are relatively small indicating a fairly uniform deposit of varved clay.
The bulk moisture content is
also larger than the bulk liquid limit, indicating that the
soil may be somewhat sensitive.
Figure 3.3 shows the general soil profile, stress history,
and undrained shear strength from triaxial UUC and field vane
tests.
The stress history shows that the soil is overconsoli-
dated at all depths.
of the maximum
It is presumed that the increasing values
past pressure
prior disiccation
(avm) above El. 110 is due to
(Ladd, 1975).
slightly overconsolidated.
Below El. 110, the soil is
This resulted from a lowering of
the ground water table and/or erosion of the overburden.
The
Geonor field vane strengths below the crust are almost constant with depth at 700 + 100 psf; the sensitivity from the
field vane test is about 7 above El. 100, and 5 to 6 below
that depth.
Triaxial UUC tests yield strengths fairly close
to those measured with the field vane in the upper portion
of the clay, whereas below El. 110 they are significantly
lower.
Table 3.4 shows the details of the testing conditions
and testing program.
3.2.3
Northampton Varved Clay
Four borings, B
through B4, were made at several loca-
tions near the intersection of I91 and Route 9 in Northampton,
Massachusetts (Ladd and Wissa, 1970).
Three in. diameter
undisturbed samples were obtained from borings B1 and B2 in
68
1967 and 5 in. diameter undisturbed Osterberg type samples
were obtained from Borings B3 and B4 in 1970.
Field vane tests
using the Geonor apparatus were run adjacent to Boring B3 and
B4.
Soil samples from Borings B3 and B4 were used for consoli-
dated - undrained strength tests.
See Ladd and Wissa (1970)
for further details.
Figure 3.4 shows the soil profile, Atterberg limits, and
virgin compression ratios from the borings.
The soil profile
generally consists of a thick deposit of varved clay which
is covered by medium sand.
The clay is grayish or gray brown
in color and has typical varve thicknesses of one-half to one
inch.
The Atterberg limits of the clay and silt portions of
the Northampton varved clay show no significant difference
from those of the Amherst and East Windsor varved clays.
Values of the virgin compression ratio range from 0.2 to 0.4.
The upper values are larger than those from Amherst and East
Windsor.
The varved clay deposit at Northampton is not as
uniform as that at Amherst.
Figure 3.5 shows stress history and undrained shear
strength data from triaxial UUC, unconfined compression, and
field vane tests.
The stress history from the "good quality"
undisturbed sample at Boring B3 and B4 shows that soil is
overconsolidated at all depths.
The stress history from Bor-
ings B1 and B2 is not as reliable because the samples were
more disturbed.
The results from the Geonor field vane tests
indicate that the clay is medium to stiff and that the strength
69
increases with depth from 1000 psf at El. 90 to 1700 psf at
El. 50.
Undrained strength data from unconfined compression
and triaxial UUC tests on vertical samples are significantly
lower, exhibit appreciable scatter and show no consistent
trend with depth.
This occurred in spite of the fact that
the tests were run on samples from 5 in. diameter fixed piston samples.
Table 3.5 shows the testing program of Northamp-
ton varved clay.
Other experimental results were also reported
in Lacasse et al.
(1972).
3.3.4
Hackensack Valley Varved Clay
Five in. diameter Osterberg type samples of varved clay
were obtained from the Hackensack Valley at Sacaucus, New
Jersey in cooperation with the Port Authority of New York and
New Jersey.
The boring was located immediately to the south-
west of the Palisades in the township of Secaucus (see Saxena
et al., 1974).
Figure 3.6 shows the general soil profile,
Atterberg limits, the stress history, and the values of the
virgin compression ratio.
The soil profile consists of fill,
peat, sand, and silty clay overlying about 35 ft of varved
clay.
The varved clay is red-brown in color.
nesses vary widely.
The varve thick-
They are about 3/8 in. at shallow depths
and about 1 in. below El. -40.
Though the bulk natural mois-
ture content is somewhat lower, the bulk Atterberg limits and
the values of the virgin compression ratio are generally
70
similar to those for the Connecticut Valley varved clays, at
least for samples between El. -35 to 50 ftwhich were used for
the consolidated-undrained strength tests.
The stress history at this site shows that the clay is
overconsolidated at all depths.
The causes of the overconsoli-
dation are mostly due to desiccation for soil above El. -40
and to the lowering of the ground water table for soil below
El. -40.
Table 3.6 shows the experimental program for the Hackensack Valley varved clay.
3.4
SUMPARY OF SOIL PROPERTIES
The following is the summary of the soil properties of
Northampton, Amherst, East Windsor, and Hackensack Valley
varved clays.
1.
The soil profiles of Amherst, Northampton and
Hackensack Valley have thick deposits of varved
clay, although the varve thicknesses varied.
2.
At Amherst, the deposit was quite uniform, as
were the thicknesses of the clay and silt layers.
On the other hand, the deposits at Northampton
and Hackensack Valley were more variable.
3.
The varve thicknesses of samples from Amherst,
Northampton, and East Windsor used in the testing program
(Tables 3.3 to 3.6) are different.
71
Northampton and Amherst samples have about the
same varve thicknesses.
East Windsor samples
have the smallest varve thickness.
The varve
thickness of Hackensack Valley samples are between those from Amherst and East Windsor.
4.
Despite the difference in varve thicknesses, the
samples used in the undrained strength tests from
East Windsor, Amherst, Northampton, and Hackensack Valley have essnetially the same bulk Atterberg limits and virgin recompression ratio (see
the summary in Table 3.7).
The Atterberg limits
of the individual clay and silt portions of the
Connecticut Valley varved clays (no data from
Hackensack Valley varved clay) are also essentially
the same.
However, the natural moisture contents
(listed below) of the Hackensack Valley varved
clay are significantly different from those of
the Connecticut Valley varved clays.
Soil Location
3.5
W
(%)
vm
(k/cm 2 )
M
Amherst
63+4
2.0+0.5
Northampton
59+3
3.0+0.5
East Windsor
57+2
2.4+0.2
Hackensack Valley
48+2
2.5+0.1
GENERAL 'ESTING PROCEDURE
All tests were performed at MIT using the general test
72
procedures described below.
3.5.1
Consolidated Undrained Triaxial Tests
Triaxial Compression Test
All triaxial compression tests were performed with Geo-
measurement cells.
consolidation.
Only bottom drainage was allowed during
One of the two drainage lines leading from the
bottom pedestal at the base of the triaxial cell was attached
to a burette for measuring the volume change.
An electrical
pressure transducer was attached to the other drainage line
for the measurement of pore pressure during shear.
The con-
fining pressure, applied to the triaxial cell water, and the
back pressure were obtained from adjustable mercury columns
(Bishop and Henkel, 1962).
After samples were cut to the de-
sired varved inclination, they were tested according to the
standard triaxial testing procedure presented in Ladd and
Varallyay (1965).
The experimental program includes both isotropic and Ko
consolidation.
For Ko consolidation, the axial consolidation
stress was applied to the samples with small load increments
by using dead weight.
The value of Kc , the ratio of the ef-
fective horizontal consolidation stress (hc)
tive vertical consolidation stress (vc ),
to the effec-
was adjusted with
each load increment to obtain essentially one-dimensional
strain.
Since an approximate relationship between Ko and OCR
was known in advance, Ko consolidation was closely followed
in most load steps.
Vertical filter strips are used for
73
accelerating the drainage during consolidation and equalizing
the pore pressure during shear.
The samples consolidated via
the SHANSEP approach were loaded in steps to a stress greater
than about 1.2 to 2 times the in situ maximum past pressure
(vm).
For testing samples at an OCR greater than one, the
samples were unloaded to the desired OCR.
The RECOMPRESSION
samples were loaded in steps to the vertical stress at the
desired OCR (--).
At the final consolidation stress, two
vc
days were used for normally consolidated SHANSEP samples to
allow a reasonable amount of secondary compression, and one
day was used for OC SHANSEP and RECOMPRESSION samples.
A
back pressure of 2.0 kg/cm 2 was used for all tests.
All but two of the overconsolidated SHANSEP CKoU test
samples were consolidated to a maximum vertical effective
2
stress of 3.6 kg/cm before they were unloaded.
The vertical
effective stress was increased in the following steps:
0.5
kg/cm 2 (isotropic), 0.95, 1.6, 2.0, 2.3, 2.75, 3.23 and 3.6.
The vertical effective stress was then rebounded to 1.8 kg/cm 2
in two steps and 0.9 kg/cm 2 in one step to obtain samples
which had OCR's of 2 and 4, respectively.
The other two OC
CKoUC test samples (OCR of four) were consolidated to 4.0
kg/cm
before being unloaded to 1.0 kg/cm 2
All NC SHANSEP samples used for CIU (CIUC and CIUE) tests
were consolidated to a maximum consolidation pressure of 4.0
kg/cm2.
The load steps were 0.3, 0.75, 1.5, 2.5 and 4.0
kg/cm 2
74
RECOMPRESSION samples from East Windsor were isotropically
consolidated to the overburden pressure of 0.6 kg/cm 2
two steps to obtain samples with an OCR of four.
in
Samples with
an OCR of two were anisotropically consolidated to a vertical
effective
stress of 1.2 kg/cm2
in three steps.
Wykeham-Ferrance loading frames wereused to produce controlled rates of strain during shear.
A strain .rate of 0.7%/hr
was used for SHANSEP samples, and 0.35%/hr for RECOMPRESSION
samples.
The samples were loaded (i.e., axial stress increased)
in compression tests and a proving ring was used for measuring
the load.
Readings of elapsed time, pore pressure, sample
length, and the vertical load from the proving ring were obtained from a data aquisition system with less frequent visual
checks.
The piston friction correction was obtained by measuring
the friction force required to move the piston through water
in a triaxial cell at the strain rate and cell pressure corresponding to the testing condition.
The correction was made
by subtracting this friction force from the total measured
axial force.
A filter paper correction was applied in CKoUC
tests according to the procedure outlined in Ladd and Varallyay
(1965).
Triaxial Extension Test
The testing procedure for TE test is different from that
for TC test in two aspects, as follows:
75
1.
During sample set-up: the filter strip paper is wrapped
spirally around the sample instead of being placed vertically
as in the TC test.
2.
Thus, there is no filter paper correction.
During shear:
the samples were unloaded (i.e., axial
stress decreased) in TE tests.
Also, suitable methods for
connecting the top cap to the loading piston and the loading
piston to the proving ring are needed.
Figure 3.7 shows a schematic diagram of the essential
elements in the TE test, including the connections between
the loading piston and the top cap, and between the loading
piston and the proving ring.
Two piston pins (1) are inserted
into the chamber (2) in the top cap (3) for connecting the
loading piston (4) to the top cap.
After lowering the load-
ing piston until it touches the ball bearing (5), the connection between the top cap and the piston was then made by rotating the pin across the slot.
As there is a small clearance
between these pins and the top cap, this arrangement will not
cause a torsional force to be applied to the top of the sample.
The proving ring and the loading piston were connected
by two pins (6) and two cross bars (7).
These two pins were
first put through the holes located in the piston block (8)
and the proving ring head before inserting the cross bars.
Because the diameter of these pins is smaller than the holes
in the cross bars, proving ring head, and the piston block,
the connection is very flexible.
Consequently, any small
deviation in alignment of the proving ring, loading piston
76
and top cap is not very important.
For CKoUE and CIUE tests, the load steps, the back pressure, the rate of shearing, and the time allowing for final
load increment during consolidation were identical to those
used for TC tests.
3.5.2
Testing Problems with Consolidated Undrained Triaxial
Tests
Due to the layered nature of varved clays, disturbance
due to trimming and setting up the sample, expecially with
TE tests, can severly affect the undrained shear strength.
Chapter 6 will discuss the effect of disturbance due to trimming
which causes separation of silt and clay portions in
horizontal and inclined samples.
This section will present
test data illustrating the importance at the details of testing on the undrained shear strength (su=qf) when disturbance
during setting up a TE test sample introduces a plane of
weakness.
In the TE test, the process of connecting the loading
piston to the top cap can introduce sample disturbance.
Table
3.8 shows qf data from NC and OC samples from Amherst, Northampton, and East Windsor, using two different testing procedures.
Upon using the "new" testing procedure described in
Section 3.5.1, qf/'vc (Table 3.7 (a)) from OC (OCR=4) samples
from Amherst and East Windsor are essentially identical (in
fact, it will be shown in Chapter 4 that NSP from Amherst,
Northampton, and East Windsor are identical).
77
Thus, the larger
qf/vc
(an increase of 30%) of NC East Windsor clay compared
to those for NC Amherst and Northampton (using "previous" testing procedure) was due to the difference in the testing procedures.
In the "previous" method, the tests (NC Amherst and
Northampton samples) were performed in a cell where the loading piston was screwed into the top cap after the samples had
been set up, and that probably caused a slight separation between one or more varves at the top of the specimen and hence
introduced planes of weakness.
This in turn probably caused
the decrease in qf for NC Amherst and Northampton samples.
The fact that necking occurred at the top of these samples
instead of at the middle of the sample as observed in the
East Windsor specimen seems to support this hypothesis.
3.5.3
The Direct Simple Shear Test
All tests were performed with a Geonor Model 4 direct
simple shear device.
Description of the equipment and the
step by step testing procedures for CKoU tests are outlined
in detail in Ladd and EdgerS(1972) and Saxena et al. (1974).
In this test, a cylindrical sample 8 cm in diameter and
approximately two cm high is confined in a wire-reinforced
rubber membrane.
This allows vertical deformations and hori-
zontal shear movement with little or no change in diameter.
During consolidation, there is very little lateral strain
so that Ko consolidation is closely approximated.
The con-
solidation process is the same as that of an ordinary consolidation test wherein a lever arm is used.
78
The direct simple
shear device is capable of applying an instantaneous load
which is achieved by locking the lever arm, putting on.a load
increment and then releasing the level arm.
The SHANSEP
samples were first loaded in steps to stresses greater than
about 1.5 to 2. times the in situ maximum past pressure for
the normally consolidated samples.
For the overconsolidated
samples, ths specimens were then unloaded to obtain the desired
overconsolidation ratio.
at the last load increment
A period of two days was allowed
for normally
consolidated
samples,
and one day was allowed for overconsolidated samples.
Shearing was achieved by applying a horizontal shear
force parallel to the top of the specimen.
The sample height
during shear was kept constant by a screw-controlled loading
mechanism to cause shearing under constant volume conditions.
This was done by regulating the vertical stress so that the
dial gage reading for vertical deformation remains constant,
after allowing for the compression or expansion of the porous
stones and the top and bottom caps.
The state.of stress dur-
ing shear can be considered to be close to a plane strain condition (Ladd and Edgers, 1972).
The rate of shear for the
tests was 5.5%/hr.
The special test procedures used for testing inclined
samples are described in Section 5.3.4.
3.5.4
The Plane Strain Tests
This section provides a very brief description of the
79
MIT plane strain device and the testing procedure used with the
Northampton varved clay.
Details of the device and testing
procedure are given in Bovee and Ladd (1970).
Two types of plane strain tests were performed, namely:
Ko consolidated undrained plane strain compression (CKoUPSC)
tests and plane strain extension (CK UPSE) tests on vertical
samples.
The CKoUPSC test sample was sheared by increasing the
vertical stress (i.e., alf = avf), while the CKoUPSE test sample
was sheared by decreasing the vertical stress.
Figure 3.8 shows a simplified diagram of the specimen and
the loading systems in the MIT plane strain device.
The sample
can be loaded both vertically and horizontally (the latter only
used during consolidation for the strain controlled test).
With the vertical loading system, during consolidation a vertical load is applied by a hanger placed with dead weights on
the loading piston.
During strained controlled shear, a load
cell is placed above the hanger and a variable speed load frame
is used to apply the vertical load.
For the horizontal loading system, a cell pressure (supplied by the self compensating mercury pot system) transmitted
to the sample via a cell membrane is used for the purposes of:
(1) Maintaining Ko consolidation.
Movable side platens
located within water-filled chambers enclosed by the cell
membranes and the side assembly plates are also used for restricting the horizontal deformation during consolidation.
80
(2) Applying (ah increments to the sample during shear in
stress controlled PS tests (not used for this study).
The pore pressures are measured from the transducer at the
bottom of the specimen.
The back end platen (Fig. 3.8) has a
flush diaphram transducer for measuring the horizontal normal
shear at that height.
Testing Procedure
The test specimen is 3.5 in by 3.5 in. by 1.4 in.
After
trimming, the entire specimen (after installing filter paper,
overlapping for CKoUPSE test and continuous for CKoUPSC test)
is enclosed in a 0.015 in thick membrane.
A watertight seal
is maintained by compression of latex membranes and an o-ring
between the platens located at the top and bottom of the specimen and the piston block or the base plate.
Prior to each consolidation increment (dead weight for
the vertical stress and cell pressure for the horizontal stress),
the side platens are brought into contact with the end platens
(Fig. 3.8) for restricting the horizontal deformation.
During
consolidation, adjustment of the cell pressure is required to
insure that the volume change resulting from axial movement
of the piston times the sample area equals the volume change
recorded on the volume change device.
At the later stages of
consolidation in each load increment, the side platens are
withdrawn.
After final consolidation, the CKoUPSC test sample
is vertically loaded by forcing the loading piston against
81
the load cell.
The CKoUPSE sample is sheared by pulling up a
loading piston and load cell.
The back pressure in the test was 2.0 kg/cm .
of shearing was 0.7%/hr.
The rate
The corrections for filter paper
(PSC only) and membrane resistance are given in Bovee and
Ladd
(1970).
82
SUMMARY OF THE EXPERIMENTAL PROGRAM
Type of
Test
UUC
(1)
8
Method(3 )
Consolidation
Consolidation
4
0, 45, 90
E
0, 90
E
1,2,4,20
S
45
E
1,4,20
S
45
A
1
S
0
H
1
S
E
4
R
CIUC
0,45,
CKoUC
0
_IUE
90
(1)
--
0
E-
1,2,4
S
0
A
1,2,4
S
0
H
1
S
0
E
2,4
R
0
A
1,6
R
0,
90
E
1,2,4
S
0,
90
A
1
S
E
4
R
90
Notes:
Soil(2)
Location
!
= angle between the varves and the in situ horizontal direction.
(2) E = East Windsor,
A = Amherst,
H = Hackensack.
(3) S = SHANSEP
R = RECOMPRESSION
Table 3.1(a)
83
SUMMARY OF THE EXPERIMENTAL PROGRAM
Type of
Soil
0
Test
Method
(1)
of
Consolidation
e
Location
0
E
1,4
S
A
2,4
S
E
2,4
R
A
1.6
R
0
E
1,2,4
S
0
A
S
0
N
1,3,6
1,2,4,8
0
H
1,2,4
S
0
A
1.6
R
CKoUDSS (C)
45
E
1
S
CK UDSS(E)
45
E
1
S
CK UPSC
0
N
1
CK UPSE
o
0
N
1
CDDS
0
CK UE
o
CK UDSS
o
(3)
OCR
S
I
S
I
0
N
1,1.8,3,5
Notes:
S,R
u
_
(1) E= East Windsor; A = Amherst; N = Northampton;
H= Hackensack
(2)
S = SHANSEP;
R = RECOMPRESSION
Table 3.1(b)
84
I
SUMMARY OF INDEX PROPERTIES AND ENGINEERING PROPERTIES
OF EAST WINDSOR VARVED CLAY
Physical Properties:
Description
Natural Water
Content
Bulk
Sample
(%)
55 -
60
Clay
Layer
Silt
Layer
63
45
Liquid Limit (%)
50.0
64.5
43.6
Plastic Limit (%)
28.9
31.4
30.5
Plasticity Index (%)
21.1
33.1
13.1
2.8
Specific Gravity
Total Density
104 - 105
--
--
.
.
(pcf)
Engineering Properties:
a
vm
= 2.4 + 0.1 kg/cm
CR
= 0.20 + 0.01
RR
= 0.035 + 0.003
2
qf from UUC test = 0.37 TSF (vertical sample)
Su from
Torvane
= 0.35 + 0.02 TSF
Table 3.2
85
LABORATORY TESTING PROGRAM OF EAST WINDSOR VARVED CLAY
( = varve inclination)
Type
of
Test
Test No.
to
EUCEUC-3
UUC
00
OCR
Method
of
Consolidation
4
--
'1,2,4,20
SHANSEP
0,45,90
EIC-i-1 to
EIC-I-4
0
EIC-2-1 to
EIC-2-3
CIUC
45
1,4,20
EIC-3-1 to
EIC-3-1
to
EIC-3-4
90
1,2,4,20
EIE--1 to
EIE-1-3
0
1,2,4
SHANSEP
CIUE
EIE-3-1 to
EIE-3-3
EIE-3-3 o90
to
EKC-5
EKE-1 t
1,2,4
CK UC
0
1,2,4
SHANSEP
CK UE
0
1,4
SHANSEP
CKoUDSS
o
0
1,4,7.38
SHANSEP
o
EKE-1 to
EKE-3
o
EDS-1
EDS-3to
ESDS-1
ESDS-2
CKoUDSS(C)
CK UDSS(E)
ERTC-1-1
ERIC-2-1
ERIC-3-1
45
45
1
1
0
4
CIUC
45
90
ERIE-1-1
CIUE
0
4
ERKC-1
CK UC
0
2
ERKE-1
CK UE
0
2
to
2,4
SHANSEP
4
ERIC-3-2
RECOMPRESSION
Table 3.3
86
LABORATORY TESTING PROGRAM OF AMHERST VARVED CLAY
Test No.
ACK- to(b)
Type
of
Test
OCR
Method
of
Consolidation
CKUC
0
1,2,4
SHANSEP
CK o UE
0
2,4
SHANSEP
CK o UDSS
0
1,3,6
SHANSEP
1
SHANSEP
AKE-1 to
AKEAKE-2to
ADS-1 to
ADS-4
ADS-4
AIC-1-1
0
AIC-2-1
CIUC
45
AIC-3-1
90
AIE-1-1
AIE-3-1
CUE
0
1
0
1
,9
ARKC-1
CK UC
0
ARDS-1
CK UDSS
0
~
=
varve
SHANSEP
1
0
1.6
RECOMPRESSION
0
1.7
RECOMPRESSION
.
inclination
Table 3.4
87
LABORATORY TESTING PROGRAM OF NORTHAMPTON VARVED CLAY
Test No.
NKC-1 and NKC-2
NDS-1 to NDS-8
NCD-1 to NCD-6
Type of Test
OCR
Method
of
Consolidation
SHANSEP
CK0 UC
0
1
CK UDSS
0
1,2,4,8
SHANSEP
0
1,1.83,3
SHANSEP
CDSS
NPC-1
CK .oUPSC
0
1
SHANSEP
NPE-1 and NPE-2
CK UPSE
0
1
SHANSEP
= varve
inclination
Table 3.5
88
LABORATORY TESTING PROGRAM OF HACKENSACK VALLEY
VARVED CLAY
Test No.
Type of
OCR
Method
of
Consolidation
HKC-1
CKoUC
0
1
SHANSEP
HIC-1
CIUC
0
1
SHANSEP
HDS-1 to HDS-4
CKoUDSS
0
1,1.98,4.12
SHANSEP
l
I ,
e = varve
.
I
inclination.
Table 3.6
89
SUMMARY OF THE INDEX PROPERTIES OF VARVED CLAY
SAMPLES FROM AMHERST, EAST WINDSOR,
NORTHAMPTON, AND HACKENSACK VALLEY
Typical
Varved
.
1/2 to 1 inch.
Thickness
.
-
_
----
WL
Wp
PI
Bulk
53 + 4
26 + 4
27 + 8
Clay*
70 + 5
35 + 5
35 + 10
ISilt*
40 + 4
26
_I_
;
. _-_
_
+ 4
14 + 8
_
CR = 0.2 to 0.3
*Average of available data where the strength
tests were made.
Table 3.7
90
COMPARISON OF THE UNDRAINED SHEAR STRENGTHS FROM
THE CK UE TESTS FROM DIFFERENT TESTING PROCEDURES
(a) Stregnth Data from New Testing Procedure:
Soil Location
Amherst
L
East Windsor
Uvm (L)
|
OCR
qf/Cvc
1.71
4
0.63
1.50
4
0.65
t
(b) Strength Data from Previous and New Testing Procedure( 3)
at OCR = 1.0
L
vm
Soil
Location
a
Amherst
1.76
Northampton
East Windsor
Note:
Testing
Procedure
Previous
q
f
vc
0.16(1)
Remark
From Ladd
(1975)
1.1 + 0.1C Previous 0.15 + 0.01(1 Ladd (1975)
1.50
Now
0.21(2)
2 Tests
(1) Failure at the top of the sample by necking.
(2) Failure at the middle of the sample by
necking across the varve.
(3) Described in Section 3.5.2.
Table 3.8
91
0
5'
/0
-
-
K4
A
LU
VYMA d
CR= 0.20 f O.0/
S4 8
£6.8
O
2s
!
, I1
RR =O.035+ O.003
5S.I
I
0.5L
/.0
VERTICAL CONOL IDA710N,
~iI
J0
vc, 7TSF
2.,.c
/0.O
I.O
0.75
Data for loadiny
c, determined 6y Vr nmethod
X.
ri
o
a0.s
o
0
D~eon /
I'
oodi ng
0Fisf
SI
A
U
v
0
·
0.I
0.f
.
·
zO
r
50
·
----
/0.0 .0.0
V6RTICAL CONSOLIDATION STRESS, 7SF
RESULTS OF CONSOLIDATION TESTS
ON EAST WINDSOR
92
VARVED CLAY
FIGURE 3.1
m,l
G.$.
El. M£l
13
water ontnt
40
0
Bulk
/JS!
/2S
p
X
0.2
I
t
-
OI
I
X
Grey
! I--
mnedium
i-
/arved
0.J
.
*
- loay
0.1
,
Sitf A
StifC,
o
w.
WN
6rey --
CR, RR
80
W
C/ayey Sand
Vorved
C/oy
./Q
M.)
i
1
soft
i
--
I
'4S
-
Jome
.stones
A
-4_
-44
X
roS - xeemr
I
more
si/ty
lal
i
I
6rey
/00 - Vorved
0---
!LI--*6
---
:-
i
i
I
I
I
Cloy
very oft
stones
x
-
---
95S
Ii-
Brown
cI
Varved
Cloy
8s
Pq{-
f
I
medium
soft
t
....
11
.
RR CR
I
min.
lwx.
,
A
I
7S
._
_
I
O-
---
SOIL PROFILE, ATTERBERG LIMITS, RECOMPRESSION
RATIO, AND VIRGIN COMPRESSION RATIO OF AMHERST
VARVED CLAY
93
FIGURE
3.2
a
w
r
I
0
I
zw
r.
1-
C,)
w
z
z:
r
-r
z
I
IrCl)
ct
r
I-
,)
w
-J
0
-J
94
FIGURE 3.3
so
CR
B/fB
ICu
.
-B4.
' ''
.
2
Sond
83
.
,-.
0o0nd.
iXlTER CO6NT£7
.20
C0
40
%
A0
COMPRES'SI/O
/00
RATIO
w
n2
0
v
-
_~_._
. __
.
---
4
T
i'slty
Kssli
Sand
Boring 83
N.-i
Grey
90
Grey
9
Grey
I-
Vared
o80F
Q-
ok-
v
Vurved
!
70o
c>_0
L-111-
Vrwd
I
-o
-4 8 estr
C/oy clay
v
I
I
J1
I'
-.4
60- F.
.
l
4J
-4 I!- I
I
C/oay
rI
0
L4j
i
Sand
Cloy
C/fw
IC7,,y
Cloy
Bulk
x ;-,
Silt
A r
C/ly
·*
E. (ft)
8/l 2
,o.
Sand
L
w
0-
R .S.
. .
nT/
'A.
V
tvwn
40O
iV
I
0
Grey
K
I
I
I
I
6rey-
-
85
84
/39
1/24
//
*max min.
OFDOMETER
t
CRSC
o
A
0
o
CR plotted
SUMMARY OF ATTERBERG LIMITS AND COMPRESSIBILITY OF
NORTHAMPTON VARVED CLAY (From Ladd & Wissa, 1970)
95
FIGURE 3.4
,
VI
0
I%
Mi
f%b
Cl)
WO
aw crE 0
z 0O Eof,
F
:
Z
La
O
LL.
0oa o
0
4
In
oc
D
_;z
L wz
aQ
,
_j
0
Sn>
en I
en
g
>
'Cla>
-4.
It
.6i
a
lb'
iU
L4J
1
I'
44
"I~4
e11ZIQ
l
(7S'W J
'OIl
96
VA373
FIGURE 3.5
w
cK
w
3
z
0D
w
lO
C)
(0
o
0Z.¢
Or c
aC E
m -o
L.
oLL
O
z
0I
0o
cn
z>
0
0
Er
1%~~~
~ ~~~~~
'9
ej
97
C')a
(63.
FGR
FIURE 3.6
I-
0
C
k
W
w
x-
h'W
-z
I...
-I
z
w
w
w
.I
CT
zw
Cl)
en
w
0
C
\3
1iS
98
:f
w
m
FIGURE 3.7
a., epi-ton)
VERTi/CAL
CV
LOADING
TrzAA
rend Ploten)
(From Ce/I
Membrorac)
HOR/ZONTAL
L OAD/IN&
SYSTEM
DIAGRAMATIC DRAWING OF PLANE STRAIN
SAMPLE AND LOADING SYSTEMS
99
FIGURE 3.8
4.
COMPARISON OF NORMALIZED BEHAVIOR OF
DIFFERENT VARVED CLAYS
4.1
INTRODUCTION AND BACKGROUND
Normalized soil parameters (NSP) such as qf/avc, Af, Eu/vc
¢
and c/Evm, obtained using the SHANSEP approach (Ladd and
Foott, 1974) have been used extensively for.studying the stressstrain-strength properties of the Connecticut Valley varved
clays(Ladd, 1975).
The objectives of this chapter are:
(1) to
demonstrate experimentally whether or not varved clays, particularly those from Connecticut Valley, exhibit normalized behavior; and (2) to compare the NSP of varved clays from several
locations.
The SHANSEP approach uses a maximum laboratory consolidation stress that is at least about 1.5 to 2.0 times the in situ
maximum past pressure ( vm).
Overconsolidated samples are ob-
tained by unloading from this maximum stress.
Thus, in general,
the SHANSEP samples which have the same laboratory OCR can
have different magnitudes of.the laboratory maximum past pressure [a (L)], and thus, the maximum past pressure ratio, (MPPR),
vm
a
vm
(L)
.
At a given laboratory OCR, a difference in the MPPR of
vm
SHANSEP samples can result from differences in the magnitude
of
vm (L) for samples with the same avm and/or from differences
of am
Vm of the samples used for testing.
100
If a soil exhibits normalized behavior, then at a given
laboratory OCR and for the same stress system (both during consolidation and shear), the MPPR will not have a significant
Thus, plots of NSP versus OCR obtained for a
Ovm(L)
. Since the
given stress system should be independent of
vm
soil structure (Lambe, 1958) prior to shear can change with
effect on NSP.
the magnitude of the laboratory applied consolidation stress,
SHANSEP samples at a given OCR, but having different MPPR, may
have different soil structure .
For a soil that exhibits near-
perfect normalized behavior, changes in soil structure with
MPPR are minimal, and NSP from SHANSEP samples which are consolidated and sheared with a stress system similar to that of
the in situ soil, should represent the in situ soil parameters.
The experiment program shown in Table 4.1 is divided into
two parts according to the objectives of the investigation , as
follows:
(1) Does varvedclayfran a givenlocation (Amherst, Northampton, East Windsor and Hackensack
behavior?
Valley) exhibit normalized
For this investigation, CKoUC and CKoUDSS tests were
performed on.NC and OC vertical samples at various MPPR (see
upper part of Table 4.1) for vertical and inclined loading respectively.
It should be noted, however, that the range in
MPPR is relatively small.
(2) Comparison of SHANSEP NSP from soils at different locations -- This testing program, shown in the bottom part of
Table 4.1, included several types of tests ranging from CIUC
101
tests on NC samples with different inclinations to CK UDSS
tests on both NC and OC samples, obtained from two to four locations.
Ideally, comparison of NSP from SHANSEP samples at
different locations should be made with samples which have the
same values of avm,' v(L)
and a
.
However, it was not always
possible to have such conditions, since some of the tests were
performed prior to start of this thesis.
Therefore, the com-
oarison is made with samples having approximately the same
a
vm
(L)
ratio, as shown in the lower part of Table 4.1.
The
vm
normalized soil parameters (NSP) studied are qf/avc, Af,
Yf, E/avc
(or G/avc), $ and c/avm.
f or
For the sake of complete-
ness, NSP data from the Connecticut Valley varved clays are compared to those from New Liskeard and Matagami in Canada.
4.2
EFFECT OF SAMPLE STRESS HISTORY (MPPR)
Test data from CKoUC and CKoUDSS tests used to investigate
the effect of sample stress history are summarized in Tables 4.2
through 4.4.
Figures 4.1 and 4.2 show normalized stress-strain
curves from CKoUC tests on OC (OCR=4) Amherst varved clay and
from CKoUDSS tests on NC Northampton varved clay at various
MPPR respectively.
Test data in Tables 4.1 through 4.4 are
samples which have a relatively small range in o
from 2.0 to 4.0 kg/cm2 ) and MPPR ( vm
2.35).
The following observations
102
(avm ranges
) ranges from 1.01 to
are made with regard to
the effect of MPPR on NSP (qf/Qvc' Af, Eu/Ovc, $ and c/avm)
from four locations.
(1) The MPPR has an insignificant effect on the normalized
undrained shear parameters (qf/vc' Af, and NC 4 (assuming c =
0)) from CK oUC tests on NC samples from Amherst, Northampton,
and East Windsor.
This is also true with respect
to qf/Cvc
and Af of OC samples from Amherst.
(2) The MPPR has an insignificant effect on
(v)at T
( v) athmax
rmCUS
Thmax
-
vc
from CKoUDSS tests on NC samples from
and
Amherst,
vc
Northampton, and Hackensack Valley.
OC samples from Northampton
also show the same trends.
(3) In general, Eu/cvc at
qf/Aqf = 1/2 from CKoUC tests
is practically unaffected by MPPR, considering the accuracy
with which soil modulus can be measured.
The variation of Eu/
c is also less than the variation of G/vcvc from DSS tests.
vc
(4) Based on limited data from the same soil and from MPPR
range
for OC samples (OCR=4), the MPPR has a significant ef-
fect on 3G/avc at 50% and 20% stress levels from CKoUDSS tests
(Fig. 4.3).
(5) For NC sample, Fig. 4.3 also shows the following:
a.
Based on Amherst data (two tests), the
increase in MPPR decreases G/avc at 50%
and 20% stress levels, while Northampton
data (3 tests) show no consistent trend.
But, collective data from East Windsor,
Hackensack Valley, and Northampton
103
arved
clay suggest a decrease in G/a
with invc
creasing MPPR:
b.
The fact that the NC Amherst sample has a
lower a
and a higher G/a
than at other
vm
vc
sites with the same MPPR suggests that the
magnitude of in situ a
per se affects
shear modulus (i.e., G/vm
vm decreases with
increasingam)
vm
(6) Though limited data are available the results from Amherst and Hackensack Valley varved clay suggest that the MPPR
has little effect on yf, though data from Northampton indicate
an increase in
f with increasing MPPR.
Based on results from NC and OC (limited data) samples
which have a small MPPR range and for two loading directions
(vertical and inclined loading), the important findings are:
1.
Varved clays from Amherst, Northampton, East
Windsor and Hackensack Valley exhibit normalized
behavior with respect to the undrained shear
parameters S/avc
Af,
*
(data only available
for TC), and
2.
v
from DSS tests;
avc
The normalized undrained modulus for vertical
loading is little affected by MPPR; but
3.
The normalized undrained shear modulus of NC
sample as measured from DSS tests most likely
decreases with increasing values of both the in
*Su
qf for TC and Su = Thmax for DSS.
104
vm
Based on the above, for any study of the effect of soil
location on G/avc, the samplesshould have the same values of
both
and a
vm
. However, this requirement can not always
vvm
be fulfilled since some of the sites have quite different values
of in situ maximum past pressure, avm
4.3
EFFECT OF SOIL LOCATION
The results presented in Section 4.3 are from samples at
four locations which have approximately the same MPPR, and avm
CIUC tests on samples with different inclinations, and CUE,
CKoUC, CKoUDSS and CKoUE tests on vertical samples are used for
this study.
4.3.1
Results from CIU Tests
A series of CIUC tests on samples with different varve
inclinations was performed on NC Amherst and East Windsor samples.
In addition, results from one vertical NC sample from Hackensack
Valley are available.
same magnitude of
These samples had approximately the
vm, ranging from 2.1 to 2.6 kg/cm
MPPR ranged from 1.5 to 1.9 (Table 4.5).
The
Thus, comparison of
the normalized undrained modulus data for all directions of
loading can be made.
The results from the CIUC tests are shown in Table 4.5
and Figs. 4.4 to 4.7.
1.
These data show that the soil location:
Does not have a significant effect on q/avc
either at the maximum shear stress or at the
105
maximum obliquity condition obtained from
=0,
45 and 900 tests.
2.
Shows no apparent effect on Af and
f (both at
conditions) for all three
the qf and ( 1 /a3 )
max
loading directions, though the Hackensack Valley
f at (a1 /a3 )
CIUC (8=0) test indicates a lower
max
[Such a difference may result from difficulty
in picking the strain at (l/a3)ma
x
since the
test had an almost horizontal (v/ah) versus
curve (see Fig. 4.5)].
The test on East Windsor
=13.5% before reaching (a1 /a3)
was stopped at
max
conditions.
3.
at qf (same as
Has no effect on
from CIUC (=90)
= 45 and
tests, though
*
from CIUC (08
0) tests.
(a1 /a 3 ))
at qf from CIUC
= 0 tests are affected.
effect of soil location on
4.
at
A significant
at qf is only observed
The reason is not clear.
Shows no practical effect on normalized stressstrain curves from tests with 8= 0, 45 and 900
(Figs. 4.4 to 4.7).
Variations in Eu/av
vc at e=0.5%
are not considered significant when considering
the difficulty
in measuring.Eu
(Table 4.5).
It should be noted that the variation in qf with loading
direction from the CIUC tests is much smaller than that from
UUC tests (see Chapter 2).
The explanation of such a differ-
ence will be presented in Chapter 6.
106
It was also observed
that the mode of failure varied with sample inclination.
vertical samples failed by bulging.
The
The inclined samples had
failure planes within clay layers and the horizontal samples
had failure planes across the varves.
Two CIUE (=0)
tests were performed on normally consoli-
dated samples from Amherst and East Windsor.. These samples
have approximately the same avm, avm (L), and thus
vm
ratio.
avm
The normalized stress-strain-obliquity curves from these two
tests are shown in Figs. 4.4 and 4.5 to be essentially identical.
In summary, the results from CIU tests on normally consolidated samples indicate that the Amherst and East Windsor varved
clays have essentially identical undrained stress-strain strength
parameters for vertical, inclined and horizontal loading.
Sec-
tion 4.3.2 will provide additional data from CKoU tests which
show generally identical undrained stress-strain strength properties from Amherst, East Windsor, Northampton and Hackensack
Valley.
4.3.2
Results from CK U Tests
Test results from CU
triaxial compression,extension, and
direct simple shear tests on NC and some OC vertical samples
from the four different locations, are presented in this section.
4.3.2(a)
Data from CKoUC Tests
Table 4.6 summarizes the test results from CKoUC
tests at the maximum shear stress condition.
107
Data are shown for
four locations at OCR=l and for two locations with OCR's equal
to 2 and 4.
These results were obtained from samples which
have approximately the same
vm
ratio so that a comparison
avm
of the normalized undrained modulus could be made. A comparison of NSP can not be made at maximum obliquity condition for
all tests, because some tests were stopped before reaching this
condition.
Test data in Table 4.6 and
igs. 4.8 to 4.14 show that,
for both NC and OC samples, the soil location has practically
no effect on:
1.
qf, Af,' f and E /vc at Aq/Aq = 0.5
u vc
2.
Normalized stress-strain curves (Figs. 4.8 and
(Table 4.6);
4.11, normalized stress paths (Fig. 4.9 for NC
samples), the failure envelope at qf for normally
consolidated samples (Fig. 4.10), and the normalized effective stress envelope (NESE), a plot
of qf/avm versus
3.
f/avm, in Fig. 4.13; and
Unique plots of Af,
f, and qf/3vc versus OCR
(Figs. 4.12 and 4.14).
The test data in Table 4.6 also show larger scatter in Af,
* and E/aV
f,
than in qf/av.
4.3.2(b)
Data from CK UDSS Tests
Detailed NC and OC (OCR=4) DSS test data having
about the same MPPR are compared to show the effect of soil
location.
Test results from samples having other OCR values
are also used for comparison of Thmax/avc versus
108
OCR and
· 1max/Qvm versus av/qvm
Table 4.7 presents a summary of results from CK o UDSS tests
on NC and OC (OCR=4) samples which have about the same MPPR.
The comparison is made at the maximum horizontal shear stress
(Thmax) condition.
The soil parameters are T/avc,
Yf, av/ovc
The results in Table 4.7 and Fig. 4.3 and Figs. 4.14 to 4.16
shows:
1.
For both NC and OC samples, the soil location has
practically no effect on:
(a) Thmax/avc and a vOvc
(Table 4.7); (b) the "unique" plot of
Tha
/avc
versus OCR (Fig. 4.14); (c) the normalized-stress
strain curves (Th/Uvc versus y) (Fig. 4.15); and
(d) the NESE, a plot of thmax/vom versus
v/ vm
(Fig. 4.16);
2.
For both NC and OC samples, G/O
vc at 50% stress
level from CKoUDSS tests are mostly affected by
the MPPR rather than the soil location (Fig. 4.3),
though the data from Amherst show some deviation
(probably because of a significant difference in
the magnitudes of avm),
4.3.2(c)
Data from CKoUE Tests
CK0 UE test data are only available for Amherst
and East Windsor.
From the results shown in Figs. 4.13, 4.14,
and 4.18, it is seen that the location had little effect on:
(1) plots of qf/avc, Af, and
at qf.
f versus OCR; and (2) the NESE
The difference in the normalized stress-strain curves
109
in Fig. 4.17 might
be the result of differences
of avm and MPPR rather than in soil location.
in the values
The higher modu-
lus of the Amherst sample compared to that of the East Windsor
sample
is due to the lower a
observed with G/o
higher.
vc
vm at Amherst
(same effect
as
data from DSS tests), though the MPPR is
Based on NC CIUE test data (Figs. 4.4 and 4.5) from
Amherst and East Windsor samples which had smaller differences
in MPPR compared to the CKoUE tests, one would expect the same
normalized stress-strain curves for these two clays if they had
similar values of a
vm
The normalized effective stress envelope (NESE) at qf
(Fig. 4.13) for vertical and horizontal loading are different.
For horizontal loading, it is interesting to note that the results from NC CIUE and CKoUE test samples and OC CKoUE test
samples plot on the same NESE.
Further discussion of this
point will be presented in Chapter 5.
In summary, CKoUC, CKoUDSS, and CKoUE test data show that
the samples from Amherst, Northampton, East Windsor, and Hackensack Valley (limited data) have very similar NSP (qf/vc
ef, Eu/avc'
,
and c/ vc).
from DSS tests and in E/avc
Af,
The differences in G/5vc values
from CKoUE tests are thought to be
due to the effect of variation in stress history rather than
the soil location
per se.
110
4.4
DISCUSSION
The experimental evidence in Section 4.2 demonstrated for
the range of MPPR under study that the varved clays from Amherst,
East Windsor, Northampton, and Hackensack Valley exhibit normalized behavior with respect to undrained shear parameters (qf/
avc' Af,
and C/3vm), and fairly consistent undrained modulus
data for vertical loading, i.e., from CKoUC tests.
However,
shear modulus data from DSS tests show a trend of decreasing
3G/avc with increasing MPPR and/or in situ a
Section 4.3 presented experimental evidence which indicated
that:
the varved clays from Northampton, Amherst, and East
Windsor, in the Connecticut Valley, have the same NSP.
Further-
more, limited data from the Hackensack Valley varved clay suggest that this deposit has essentially the same normalized properties as those obtained from the Connecticut Valley.
The above results suggest that varved clays which have
similar bulk index properties and composition, though not necessarily having the same Varve thickness, should have the same
NSP.
Thus, if the nature of the varves in other deposits varies
due to:
1.
Large differences in relative positions of "silt"
and "clay" layers;
2.
Large differences in Atterberg limits of the
silt and/or clay
3.
layers;
Variation in mineral composition of varves,
especially the presence of cementing agents
111
such as carbonates;
then, one might expect their NSP to differ significantly from
those of the Connecticut Valley varved clay.
An example of the possible effect of the difference in the
nature of varves on NSP is provided by data from the Matagami
varved clay located in Canada near James Bay (NGI, 1972).
The
two layers of this varve deposit are composed of a dark grey
layer of soft fissured clay and a light gray layer of soft silty
clay, generally with the thicknesses of 3/8 - 1/2 in. and less
than 1/5 in. respectively.
Compositional analyses show the
presence of large amounts of carbonates in both layers.
this soil is most likely naturally cemented.
Thus,
The soil profile
shows that the thickness of the silt layer increases with depth.
Table 4.8 shows the bulk Atterberg limits and undrained shear
strength data from samples at two depths.
These Su/avo values
are higher than those for the Connecticut Valley (compared Su
at the same OCR [see Fig. 4.14]).
The increase is thought be
be largely due to the difference in the nature of the varves,
though the method of consolidation (SHANSEP method for Connecticut Valley varved clay versus RECOMPRESSION (vc = avo) method
for Matagami varved clay) was also different (see Chapter 6
for general effect).
Based on FV data (Fig. 4.19) where the
method of consolidation is not a factor, S (FV)/avo
for the
Matagami varved clay is also significantly higher than that
from the Connecticut Valley.
As a result, the difference in
laboratory strengths are probably due to inherent differences
112
in the nature of the varves, and especially the carbonate cementation.
Basic differences in the nature of the varves are not always
shown by variation in the bulk Atterberg limits.
The best il-
lustration of this fact can be seen from a comparison of the
NSP of the varved clays from the Connecticut Valley and New
Liskeard in Canada.
A description of the general soil proper-
ties and undrained shear strength (Su ) data of the latter soil
are presented in Lacasse and Ladd (1973). Briefly, the varved
clay from New Liskeard has about 1/2 in. thick clay layers and
3/8 in. thick silt layers. The natural water content of the
silt portion varies between 24 and 30%, and varies between 60
and 80% for the clay portion.
On Casagrande's plasticity chart,
the CH "clay" layers plot above the A-line with WL = 70 + 10%
and PI = 47 + 13%, and the CL and ML "silt" layers plot above
and on the A-line with WL = 30 + 5% and PI = 10 + 6%.
Thus,
the silt and clay portions of New Liskeard varved clay have
similar Atterberg limits to those from the Connecticut Valley
(presented in Table 3.7), though composition analyses show large
amounts of carbonates in both layers of New Liskeard varved
clay (Quigley and Ogunbadejo, 1972).
Furthermore, comparison
of the bulk Atterberg limits of varved clays from Connecticut
Valley and New Liskeard indicates that the former is only
slightly more plastic.
Figure 4.19 shows plots of S/Oo
and
SU/avc versus OCR from field vane and the CKoUDSS tests (RECOMPRESSION consolidation for New Liskeard varved clay versus
113
SHANSEP method of consolidation for Connecticut Valley varved
clay).
The significantly higher strengths for the New Liskeard
varved clay compared to those for the Connecticut Valley are
probably the result of the presence of natural cementing agents,
e.g., calcite and carbonate.
(Both the Connecticut Valley and
New Liskeard varved clays have mainly illite and chlorite clay
minerals.)
These differences in the mineralogical composition, especially the presence of cementatin9 agents, seem to be a very
important factor affecting the NSP of varved clays.
of S(FV)
The plots
and Su(DSS)/ovo in Fig. 4.19 for Matagami and New
Liskeard, both having the same cementing agent (carbonates), are
essentially identical, despite having very different bulk Atterberg limits.
Large variations in the relative portions of silt and clay
layer may have a small influence on FV strengths.
Limited data
from the Lake Flushing varved clay in NY (Parsons, 1976) surprisingly suggest this.
Despite a much smaller percentage of
clay layers in the Lake Flushing varved clay compared with those
from the Connecticut Valley and Hackensack Valley, SU(FV)/Ovc
at an average OCR of four for the Lake Flushing varved clay is
roughly equal to that for the Connecticut Valley (0.6 for Lake
Flushing versus 0.74 for Connecticut Valley).
The clay layers
which have essentially the same Atterberg limits as those from
Connecticut Valley represent about 10% of total thickness compared to about 50% for the other two deposits.
114
Important conclusions obtained from a comparison of the
NSP of several varved clays are summarized, as follows:
Varved
clays which have the same bulk Atterberg limits, but coming
from different geological deposits, do not always have the same
NSP.
Thus, bulk index properties do not necessarily indicate
singificant differences in soil composition, especially the
presence of cementing agents.
Variations in the mineral compo-
sition of varves seems to be the most important factor that
affects NSP.
Varve thickness and differences in the relative
portions of "silt" and "clay" layer seem to have less influence
at least on FV strengths.
It should be clearly pointed out that the soil samples
which came from the Connecticut Valley have the same bulk index
properties and basic soil composition, and therefore should
have the same NSP.
4.5
SUMMARY AND CONCLUSIONS
Experimental evidence from this chapter has illustrated
that, besides having the same index properties (see Chapter 3),
varved clays from Connecticut Valley and Hackensack Valley have
the same normalized behavior with respect to undrained shear
strength parameters (qf/lvc' Af, $ and C/avm) .
for all directions of loading:
This is true
vertical,inclined, and horizontal.
Though the data show some differences in the normalized undrained shear modulus from the CKoUDSS tests, such differences
115
are due to the changes in the MPPR (
v
) and/or the in situ
vm
a
rather than the soil location per se.
vm
Analysis of the normalized undrained modulus of a given
varved clay shows that for samples which have the same OCR, the
vm(L)
sample's stress history (i.e.,
ratio) and magnitude of
vm
aVM had:
1.
Little effect on E /avc and
f obtained from the
CK0 UC test; and
2.
Some effect on G/avc obtained from the CKoUDSS
tests, i.e., decreasing modulus with increasing
MPPR and/or
vm
For the Connecticut Valley varved clay, the effect of sample
stress history on the normalized undrained modulus and failure
strains is expected to be more than the soil location.
a
In gene-
(L)
ral, the increase in the vm
ratio will result in a decrease
avm
in the normalized undrained shear modulus and an increase in yf
at
hmax
for SHANSEP samples which have the same
and OCR.
However, based on data for a small MPPR range, as used in the
study, the MPPR ratio has practically
qf/Uvc, Af,
no
effect on
and C/avm (i.e., these clays exhibit the norma-
lized behavior with respect to these parameters).
Consequently,
these normalized parameters from Amherst, Northampton, and
East Windsor samples can be used to study the effect of stress
system presented in Chapter 5.
The suitability of using the
SHANSEP method of consolidation for measuring the in situ normalized soil parameter is studied in Chapter 6.
116
The importance of soil composition was also illustrated.
Variation in the mineral composition of varves, especially the
presence of cementing agents, seems to be very important, this
leading to increased strength compared to uncemented varved
clay.
Also, bulk index properties do not necessarily indicate
changes in soil composition and thus soil properties.
117
EXPERIMENTAL PROGRAM FOR INVESTIGATING THE EFFECTS OF SOIL
SAMPLE'S STRESS HISTORY AND SOIL LOCATION ON NSP
Togic
Investigation
Type
Test
CK UC
Sample
.
-1
-E
N
0
N
CIUE
Soil Location:
1.12, 1.25, 1.50
1.07, 1.33
1
1.76, 2.35
1
1.01, 1.46, 1.52
2
1.25, 1.47
0,45,90
1
1.90
E
0,45,90
1
1.67
A
0
1
E
1.90
.
0
1
1.67
CK UC
N
o
A
E
A
E
E
H = Hackensack
Valley
CK UE
A
0
A
E
ton
1.71
0
1
1.50
1.40
0
2
0
4
1.71
1.50
1.90
1.67
1.50
1.33
0
2
1.71
0
4
2.20
1.67
0
1
A
CKoUDSS
0
1.76
N
E
H.
KOUDSS
1.60
A
E
H
Windsor
N= Northamp-
1
1
A
A= Amherst
E= East
1.71, 1.81
1.90,
.2-5
12H',
T.
.
C I.UC
vm (L)
Om
A
KoUDSS
Values of
0
A
A
Stress History
of
Soil
Soil
Location
1.52
1.67
__
E
H
4
1.60
1.67
1.60
Table 4.1
118
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a,
Table 4.3
H
I
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I
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120
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i
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0
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Table 4.6
r-
II
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0:0 N N
W
d O
en
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b
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Table 4.7
N
C
,.,.,CO
C
124
UNDRAINED STRENGTH DATA ON SAMPLES FROM TWO DEPTHS OF
MATAGAMI VARVED CLAY
Data from Matagami Varved Clay from NGI Report No. 71307-3
WL = 94.1 + 3.0%; W = 40.9 + 3.2%; PI = 53.2 + 6.2%
I
*
2
Depth = 23 ft; WN = 98.7 + 3.5%; aO
Type
of
Test
K
OCR
CK UC(1
0.6
1.8
0.72+0.01
CK UDSS
0.6
1.8
0.39+0.01
0.6
1.8
0.48+0.05
CK UE
2
Su/a
SU=qf
S =f
CKUE
Kc
OCR
0.60
1.7
.60
_ 0.60
Notes:
(av
of 4 tests)
(avg. of 3 tests)
U.
Depth = 28 ft; WN = 87.8 + 0.8%;
CK 0 UDSS
2
vm
vm :1000 lb/ft
(avg. of 2 tests)
hmax
WL = 76.1 + 0.7%; Wp = 36.9 + 0.9%;
CKUC1)
;
%
Remarks
·
Type
of
Test
= 550 lb/ft
t
ao =760 lb/ft2 ;
S /avo
v(3)=1300 lb/ft2
Remarks
0.57
1.7 0.34+
1.7
PI = 39.2 + 1.6%
Su
.05
0.43
Su
qf
hmax (avg. of 3 tests)
hmax
S,
(1) Failure across the varves was observed in
CKoUC tests.
(2) Failure by bulging of silt was observed
in CKoUE tests.
(3) Interpolated to these particular depths.
Table 4.8
125
/F
.6
1.4
1.2
/.o
0.a
0.6
0.4
0.2
0
o
!
Z
3
9
s
6
7
9
9
10
/2
I!
/3
AXIAL STRAIN, %
------
(°/aY)mox
3.O0
- -
2.25
s
=2.2
/
/~~~~~~~~~A
°h
2to
IA
I.
W
|~, ,= 2.i K9/c,_
I
0
_..._
/
I
Z
I
3
I
*
v,m L)190
' '
I
s
6
I
7
9
I
9 /0
I
//
I
/2
/3
AYIAL STRAIN, %
NORMALIZED STRESS-STRAIN
CKoUC TESTS
CURVES FROM
ON OVERCONSOLIDATED
AMHERST
CLAY AT TWO vm(L) RATIOS
"vm
126
VARVED
FIGURE 4.1
rh;
avc
l
U.a
0.6
bu
#;- 0.4
0.2
C
rl
-
!0
I
NORMALIZED STRESS-STRAIN
CURVES FROM
CKoUDSS TESTS ON NORMALLY CONSOLIDATED NORTHAMPTON VARVED CLAY
FIGURE 4.2
127
J~
-
!
W~ r
At
h//hmlx
-1
L OCA T/ON
A C.
0 I0
E
A
A
0o 0
/o0
Fvc
SO/IL
i
\'vm= /.7
3C
SYM.
n
4
90
H
A
U
0
OCR' 4
80
J4
*rrn ronges from 2. S t
70
60
s0
%I
i
I
"'zo
.
I
.0
.5
3.5
-. 0
2.5
vin4)
v/n ?7
u~
= 0.2
At
too
36
oic
Th Ma
0- \
0
E = East Windsor
A
mb7-h-er;t
J L
L\
/50
/-/ = Haockensac
V/alley
OCR-I\
N = Aorthampton
n
w, 0
I
A
!
I
3.0
2..5
t.O
I
3.5
O'vm(z
Oto
n
3G
EFFECT OF MPPR ON VALUES OF rv c FROM
CKoUDSS TESTS ON NORTHAMPTON, AMHERST, EAST
WINDSOR AND HACKENSACK VALLEY VARVED CLAYS
128
FIGURE
4.3
COMPARISON OF STRESS-STRAIN
CIUC AND ClUE TESTS ON NORMALLY
CURVES FROM
CONSOLIDATED VERTICAL SAMPLES FROM AMHERST, EAST WINDSOR,
AND HACKENSACK VALLEY
129
FIGURE
4.4
r,
.
l
MPPR
YM. L OCA TION
40C
40 -.
o
E. Windsv.r
1.0
1.67
57. /
Amherst
9..0
1 90
65S5
Clue TESTS
Jw - _;7
3.0
2.0
Stress in Ag/cm2
IU
T
I
0
I
I
2
4
I
I
I
I
I
8
6
I
/0
I
I
I
/f
J2
AXIAL STRAIN, Q %
.ro
i
ii
LOCAT/ON
0
4.0
4.0
A mherst
HacKensocir 4.0
.vc
E. Windsor
A
h
,P
;
YM.
MPPR I
0
1.67 56.0
1.90
.54
CIUC TESTS
58.4
.-
J3.O
C=-
6-
-..
II2.0
'
Strres in ko/cm a
1A
,.F_ ..... __
r_
I
I
2
. I
I'
4
1_
I.
6
I
AXI/AL STRAI#, 6.
.3
CURVES
I
/0
8
I
I
/.
%
FROM CIUC AND CIUE TESTS ON NOR-
MALLY CONSOLIDATED VERTICAL SAMPLES FROM
AMHERST, EAST WINDSOR, AND HACKENSACK VALLEY
130
FIGURE 4.5
0.8
.
;
lU
.
SYM.
wN
LOCATION
Ovm
E. Windsor 58.7 1.67
0.6
O
An herst
A
60.3 1.90
I
--
0.4
LI !
0.2
II
I
2
t
I
I
Mi
I
I
! I'
I
4
I,.
I iI
I
I
I
a
6
I.
.
12
10
14
AXIAL STRAIN, %
4.0
.
l
LOCA TION
YA.
/0
358.7
E. Windsor
0
30
A mherrt
--
1
60.3
---
67
190
_~~
_
&3
,I
I.U _ I
0
I
I
2
I
I
4
I
I
6
I
I
8
I
I
o0
I
I
12
I
'4-
AXIAL STRAIN,%
COMPARISON OF STRESS-STRAIN
CURVES FROM
CIUC TESTS ON INCLINED SAMPLES (45 ° ) OF NORMALLY
CONSOLIDATED AMHERST AND E. WINDSOR VARVED CLAYS
131
FIGURE 4.6
LU
w
1
I
LOCAT/ON
SYM.
0.8
E. Windsor
_
0
Amnherst
Wn
/
Cr.
¢
4.0
167 S3.7
1.90 61.
4.0
rm
0.6
-,
_
5
O~~~r
0
tr.
0.a
\/
I
'I
1
0
I
I
2
c
I
Ii
I
I
II
I
I
I
6
4
0
I
t
I
/4-
/2
AXIAL STRAIN, %
4.0
3.0
SYM.
LOCATION
*
E. Windror
4.0
0
Amherst
4.0
vcm
8,C,,
.67
1. 7
1.90 6.2
3
2.0
r.
I
0
r
I
I
2
I
q.
I
II
6
I
I
U
.
i
I
/0
I
IZ
I
AX/AL STRAIN, %
COMPARISON OF STRESS-STRAIN
CURVES
FROM
CIUC TESTS ON HORIZONTAL SAMPLES OF NORMALLY
CONSOLIDATED
AMHERST AND E.WINDSOR
132
VARVED CLAYS
FIGURE 4.7
1/
0.6
-
0.5
_-
I
1[
0
F
(r
0
0
*0
C3
V *
0
Ia
I
ed5
A
A
0
0.4
A
I
r
lrvc
0.3
0.5
.
u
.
I
SYM.
0.2
_
0.1
A L
"'00
¢
|MPPR
'vm
A
4.0 J3
v
3.4- 3.3
O
3.6
2.4
3.o
2.4- 2
O
*
*
LOCATION
12/ Northompton
.03
vorthompton
E. Windxor
/.50
E. Windsor
2. 6Y 2.4 1.12 E. Windsor
2.99 1.6 1.87 Amherst
II
0.
I
I
II
1.0
I
[
I
I
I
I!
I.
I.S
2.
4
I.
6
I.
8
I
!.
0
/2
AXIAL STRAIN, %
NORMALIZED STRESS-STRAIN CURVES FROM
CKoUC TESTS ON NORMALLY CONSOLI DATEI[) VERTICAL
SAMPLES FROM CONNECTICUT VALLEY
133
FIGURE 4.8
/14
c%
1-.
-J
J
-
cn
C-D
Igm
I
/
cn
I-
I
w
zZ
Y
L
z0
0
O
L.
O
C-)
n
w
C)
I0 a.J
$i~
b,
11t
to
--Z
a-.
'6
0
0\o os \os S\oOs V,~o
%O
lbo
cn
U)
w
o ei
i
v
a<c
X
lb'b
I-
cs t- o a
w
)
It
aZ
n
C
Cs
hj ~p:jic
LL
W 0I
t + t C4
^ 2 0 4 LL
_ ~Y4 N4
.
W
N
z
0
e0
cr
<I4
I '
I'
,<lia
I
i
z0
LI <H tII
4q
,, u of
0
0
'T 2t W
0
I
!~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~
0
1
q4
0
I
I
Nb
134
FIGURE 4.9
Z
wX
LL
I-
WJ
e4
..
v-.
lcu
U0
)
L
aJ
0
z
z0
,%&C
%4. 11
cr
a
wCrr
135
FIGURE 4.10
a: -cr,
lrvc
AXIAL STRAIN,
/@
4.0
h
%Fr
2.0
AXIAL
COMPARISON
STRAIN, %
OF NORMALIZED STRESS-STRAIN
CURVES FROM CKoUC TESTS
ON OVERCONSOLIDATED
AMHERST AND EAST WINDSOR VARVED CLAYS
136
FIGURE
4.11
14t'~SYM.
SOIL
L
/
0
C,
0
A
·
I E
X\
.,[/..)
TIONIV
1.7/-1.9
15'-167
H
.8s-
,,m
v
lI.
N
.33
Af
0.6
0.4
I
o
I
3
2
(a)
I
4
I
12
I
Amherst
E =
East
E--
/0
8
Cf6
P
6
OCR
_
Windsor
H = H4acXensocK Valley
N = Northornpton
6
1'
C
K0 1/
4
2
uA
(b)
d
/
"'"
I
I
I
2
3
4
s
6
OCR
Af AND Ef VS OCR FROM CKoUC TESTS
ON AMHERST AND EAST WINDSOR VARVED CLAYS
137
FIGURE
4.12
_
|
.~_
-I
0
==
E
I
N
I-
t
99
;
-. :
ONO
E
0
.
c
cn
Ibl b'
ON:
.
00
II i*,
0th
6
w
00
a:
Ge
ct
%, %I -*
Y zaI
ou
E
oth t4%i
os
pi
K
-,
11
-4 -4.1
"C1q
'p
IT
N
.
0I<^
tiQ
0
x
r
0
Q!
2:
Joz140t
-~
LyK
I
w
t {Q
O
.Q
_Z
--
() a
o
C
'C<
Ow
11
c
Ow
fi
3
n
'4-
WL9
_
o"
0'
LL
W
Ill
Ill
w
Ici
0 t
llOC
I .
N Iv
N
1
/
w
/
Q l 'IN
/
0W
I
O
N
/
/
1I)J
<5
<2/
/
/
I
Cr5
/
,=
,·
,
Ii
C1
I
Ie
16
Z
I
O
(CJ
ts~
r:
O
I
)l''
138
FIGURE 4.13
SE'
&VC
OCR
NORMALIZED UNDRAINED SHEAR STRENGTH
OF CONNECTICUT VALLEY AND HACKENSACK VALLEY
VARVED CLAYS
FIGURE 4.14
139
0.2
0
d, %
AU
qvc
ar' yo
-1
'ch
&IV
0
4
8
/2
/16
Jo
24
(', %
NORMALIZED STRESS-STRAIN CURVES FROM CKoUDSS
TESTS ON NORMALLY CONSOLIDATED NORTHAMPTON, AMHERST, EAST WINDSOR, AND HACKENSACK VALLEY
VARVED CLAYS
FIGURE 4.15
140
-
'.4
6>7
I' .iI
cn
Q
I-
I
-JU)
-J
10
Io
1%i
N
f6
t;·
v
Or
zW LL
Zcn
0>
.-
I
t.,
*b
o
rZt
Q
P
R
(zi
4
(4
waJ
Z
E
NO
k
.
sJ
t1
A1
.O
I
*
oo'
<
-----
I
B
<I
.I
-
ti
141
FIGURE 4.16
A
&.C .
2
0
4
/0
8
6
14
AXIAL STRAIN, %
4.0
, '1
I
I
I
I
I
I
....
I I
V
I
I ':
LOCATION
20
.i
"'0
Jsn
I
I
I
I
4
I
I
6
i
,•l) wN
m,, . %
*
Eat Windsor/.67 58.7
0
Amherst
2.d.O
I
63.1
I
8
/0
/2
/1
AXIAL STRAIN, %
NORMALIZED STRESS-OBLIQUITY-STRAIN
FROM CKoUE
TESTS ON OVERCONSOLIDATED AMHERST AND EAST WINDSOR VARVED CLAYS
142
FIGURE 4.17
1
/.4
.
SYM.
.2
1.0
4
0
\
SOIL
all
LOCAT/ON
E. Windsor
/.671
Amherst
Z.9-2.2
0.8
0.6
0.4
0
m (L) = 2.
o.:I
L
_
I
I
I.
2
3
I.
4
I. aPt
JS
I.
I
6
7
_
OCR
Is
/9
13
/2
12
I
'C
/c
OCR
VS OCR FROM CKoUE
OF EAST WINDSOR AND AMHERST VARVED CLAYS
Af AND Ef
TESTS
14 3
FIGURE
4.18
: -
l
r~~~~~~~~~~~
SYhlM.
So/L
TYPE OF TES T
1OCA TION
New Lisleard
/.o
x
New LisKeoard
0
rl
Moa 9 0mi
I/DS
CKo
rv
C K I DSS
Malagami
FY
i
o0.
-
/
FV of the
FV from Motogami x/
and ew Liseord
Su
°vo
._ Connecticut Volley
I/
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Yaorved Cloy
Loodd Foott (/977)
/
0.6
/
-
T/
/
/J
0.4
XO
ee
T°
½VC X
Pvc
AK1;0
'rop
C VDS
tests of the
Connecticvt
l hey
V/ored C/oys
a2 I
I
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I
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6
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7 8 9
OCR
COMPARISON OF THE NORMALIZED UNDRAINED
SHEAR STRENGTHS OF CONINECTICUT VALLEY, MATAGAMI,
AND NEW LISKEARD VARVED CLAYS
144
FIGURE 4.19
5.
EFFECT OF STRESS SYSTEM ON UNDRAINED BEHAVIOR
FROM SHANSEP SAMPLES
5.1
INTRODUCTION
The general requirements for accurately measuring the undrained stress-strain-strength behavior of varved clay and its
anisotropy are:
1.
Performing tests on samples having the same soil
structure and physical conditions as the in situ
varved clay;
2.
Performing tests on samples in a manner to insure
that the stress system, time, and environment
(temperature, pore fluid characteristics, etc.)
are.the same as those which will be imposed in
the field.
From these requirements, the major factors that could affect the
measured behavior are;
(1) sample disturbance; (2) the stress
system before and during shear; (3) time prior to and during
shear; and
(4) the environment.
The effects of stress system on undrained behavior are
studied in this chapter.
All data are obtained by the SHANSEP
method of consolidation and time and environment are treated
as constants.
The effects of sample disturbance and alternate
methods of consolidation will be covered in Chapter 6.
The stress system, both prior to shear and that applied
during shear means the direction and relative magnitudes of the
145
three principal stresses.
The stress system prior to shear is
therefore that resulting from the relative magnitude and the
direction of
lc' a2c' and a3c'
Laboratory samples are frequently
isotropically consolidated whereas the in situ stress condition
is generally anisotropic, with a K o stress state often being
of greatest interest.
The applied stress system during shear
is that resulting from the applied total stress path, which in
turn specifies the relative magnitude
and Aa 3 during shear.
and direction ofAol1 A
Thus, the direction of alf versus
2
alc
(i.e,, rotation of principal stresses) and the magnitude of the
intermediate principal stress are the variables for the applied
stress system during shear.
The' study in this chapter concentrates on three major topiics, as follows:
1,
Effect of Ko versus isotropic consolidation, i.e.,
the study of the effect of stress system prior to
shear;
2.
Effect of the intermediate principal stress (a2 );
3.
Effect of different stress system used for causing
shear parallel to the varves.
The section presents data from tests having different directions
of alf relative to the orientation of the:s-%rves and from tests
with rotation of the principal stresses.
Such effects.are in-
vestigated by comparing the resultant NSP (qovc' Af, E/avc,
4 and c/vm).
Table 5.1 presents a summary of the experimental program.
It consisted
of:
146
1.
CIU and CKoU tests on vertical and horizontal
samples of OCR's of 1, 2, and 4 for investigating
the effect of K o versus isotropic consolidation;
2.
CIUC and CIUE tests on vertical and horizontal
samples at OCR's of 1, 2, and 4 for investigating
the effect of
3.
a2 in CIU tests;
CKoUC, CKoUE, CKoUPSC, CKoUPSE, CKoUDSS(C),
CKoUDSS(E)
tests on normally consolidated samples
for investigating
4.
and
the effect of a 2 in CK 0 U tests;
CIUC tests on inclined samples at OCR's of 1, 4,
and 20, CKoUDSS and CDDS tests on vertical samples
with various OCR's for investigating the effect
of testing method for shearing parallel to the
varves.
As noted previously, the SHANSEP method of consolidation
was used in the study.
Experimental results from K
consolidated
SHANSEP samples in Chapter 4 indicated that Connecticut Valley
varved clays exhibit normalized behavior over the stress range
of interest (except perhaps for modulus) and samples from different locations also had the same NSP.
Therefore, the NSP con-
cept and test results from three varved clays (Amherst, Northampton, and East Windsor) are used for this study.
Appendix C presents details regarding the test results and
testing conditions.
The general testing procedures for triax-
ial, direct simple shear, and plane strain tests were presented
in Section 3.5.
The special method of consolidation used for
147
CKoUDSS(C) and CKoUDSS(E) tests is discussed in Section 5.3.4.
It is emphasized that, as a result of using the SHANSEP
method of consolidation, these samples do not necessarily have
the same soil structure as the in situ soil, as will be discussed
in Chapter 6.
In addition, the structure of isotropically con-
solidated laboratory specimens will obviously be different from
that of th( in situ soil.
Becau.;e the soil structure of SHANSEP samples is different
from that )f the in situ soil, the effect of stress systems, as
shown from the studies presented in this chapter, may not truely
reflect
the>
in situ soil behavior.
However, the SHANSEP method
of consoli¢lation was chosen because disturbance due to separations of tiie varve , especially with horizontal and inclined
samples,
c()uld
seriously affect the results, if the RECOMPRESSION
method of :onsolidation had been used (see Chapter 6).
In ad-
dition,
th ~ effect of material anisotropy component (defined in
Chapter
2)
could be mostly separated from that due to the struc-
tural and Lnduced anisotropy (see Chapter 7), therefore helping
to provide a basic understanding of anisotropic behavior of
the varved clay.
Finally, the results are applicable for a
given soil structure even if this material does not necessarily
have the sttme NSP as the in situ soil.
5.2
5.2.1
EFFEC! OF K
VERSUS ISOTROPIC CONSOLIDATION
Met] od of Investigation
Tab,Le 5.1 lists the experimental program that is used
148
for this study.
For vertical loading, the effect of Ko consoli-
dation is shown by comparing NSP from CKoUC and CIUC tests,
while CK0 UE versus CIUE tests show the effect for horizontal
loading.
The laboratory maximum past pressure o
(L) was either
avm
Volume Change Data from CIU and CK U Tests
5.2.2
AV
Figure 5.1 shows plots of volumetric strain, Vo , versus
the vertical consolidation stress, avc' for samples from East
Windsor.
Data from CIUC and CIUE tests are compared with those
from CK UC and CKOUE tests.
0
Figure 5.2 presents Ko versus OCR
data from a Ko cell and Ladd's (1975) recommended relationship.
Measured Ko values from CKoUC and E tests agreed quite well
with the recommended relationship.
For simplicity, since Ko
is nearly equal to unity at an OCR of 4, CKoU tests with an OCR
of 4 started shear with Kc= 1.0
It is seen in Fig. 5.1 that isotropically consolidated East
Windsor samples compress only slightly more than K o consolidated
samples.
Based upon these rather limited data, one can conclude
that the stress system during consolidation (i.e., K o versus isotropical consolidation) has little effect on the volume change
during consolidation, at least in NC range.
Since there was
little variation in the natural moisture content (WN) of the
East Windsor samples, both
= 1 and Ko consolidated specimens
have approximately the same water content prior to shear.
Thus,
water content per se is not a factor in comparing CIU versus
149
CKoU test data.
However, the same final water content does not necessarily
mean that CIU and CKoU test samples have the same soil structure.
One would expect Ko consolidation to retain and/or produce a
more parallel orientation of particles perpendicular to the vertical direction.
Hence, it is likely that the soil structure
of these samples will be different.
Moreover, this difference
should still exist for overconsolidated samples of an OCR of 4
even though 1Kis equal to 1.0 (Martin and Ladd, 1975).
5.2.3
Re'sultsfrom CI-UCand CK UC Tests
Test results from CKoUC and CIUC tests are summarized in
Table 5.2 and Figs. 5.3 through 5.10.
Before discussing the re-
sults, some discussion of "normalized effective stress envelope
(NESE)," a plot in terms of p /"
f
1.
vm versus qf/m
is in order.
In order to compare NC* and OC samples and to account
for variation in avm, all results are divided by avm
2.
Results of triaxial tests are plotted for convenience in terms of q and p (or q/3vm vs. p4m
yield an intercept a (or avm
3.
)
)
which
and a slope a.
In order to convert the above parameters to normalized Mohr-Coulomb (M-C) strength parameters, i.e.,
to plots of Tff/avm
versus aff/avm to give
/avm
and ~, the following relationships were used.
*Note that p and q divided by a
of all NC samples will
end up at the same point in the plotvgf NESE if the soil exhibits
normalized behavior.
150
sin 4 = tan
C/avm = a/cos
Tff/Vm=
avm
af
qf/avm cos
=
/
Pfvmqf/vm
f/vm
sin
These assume that the failure plane is given by the
point of tangency between the normalized (with respect to avm ) Mohr circle at failure and the normalized M-C envelope, whereas, in fact, the inclination
of the actual failure plnae is not known.
4.
Reference to Fig. 5.8 shows that there are several
normalized envelopes when NC and OC samplesa are
considered collectively.
represents the NESE at (
also at qf when Pf/vm
Note, however, that Line A
1 /a3 )
for all samples and
max
is less than about 0.3.
These envelopes were used to obtain the corresponding values of
and c/vm
vm except in the case of NC
samples when 4 was computed from a assuming a=c=O.
5.
To avoid confusion, a NESE with a bar on top (i.e.,
NESE) refers to Mohr-Coulomb normalized effective
stress envelope (i.e., a plot of aff/avm versus
Tff/avm), whereas the term NESE designates an envelope
in terms of Pf/vm
versus qf/3vm
Data in Figs. 5.3 through 5.10 and the results in Table 5.2
show that K o consolidation (relative to isotropic consolidation):
1.
Did not significantly change qf/avc at OCR's of 1,2,
and
4.
151
2.
Did not significantly change
at
Tff/a c at (al/a3)max
an OCR of four, but decreased it at OCR's of
1 and 2 (Table 5.2).
3.
Had little effect on Af at qf, but increased A at
(a1 / 3)
4.
at low OCR (Table 5.2).
max
Significantly decreased f at qf (Fig. 5.6(a)).
5.
Did not change
f at (
1 /a3 )
of samples at OCR's
max
of one and two, but decreased f of the sample with
an OCR of four.
6.
At low OCR, caused a significant lowering of the NESE
at qf (Fig, 5.8).
At an OCR of one,
(c=O) of NC
sample decreases from 270 + 1.5° to only 21
°.
+ 0.70
This reduction in NESE becomes less pronounced as
the OCR approaches four.
7.
Did not change the NESE at (a1 /a 3)
except per-
max
haps only slightly changing
(=0)
for NC samples
because of the difference in pf.
8,
Increased the secant Eu/qf at the same applied shear
stress level (Aq/A qf) for OC samples (OCR=2 and 4),
based on limited data, whereas K
consolidated
apparently did not significantly affect E /qf for
NC samples (Fig. 5.9).
9.
Increased Eu/ vc at
a
= 0.2% for OC samples and
decreased EU/avc at the same strain with NC samples
(Fig. 5.10).
The compensating effects of the stress system during
152
consolidation on the NESE at qf and the pore pressure response
caused qf(V) to be nearly identical.
CK 0 UC samples have a lower
NESE at qf (Fig. 5.8), but higher values of pf than CIUC samples
(Figs. 5.7 and 5.8).
The decrease in NESE at qf with Ko consol-
idation is consistent with data on other clays (Ladd, 1965).
The lowering of the NESE at qf from CKoUC tests might be
explained by a difference in the degree of mobilization of friction and cohesion in the clay and silt portions.
According to
the concepts of Schmertmann and Osterberg (1960), the mobiliza
tion of cohesion and friction and a potential failure plane
are strain dependent.
The friction component only becomes fully
mobilized at large strain, whereas the cohesion reaches its
peak at low strains.
Since CIUC test samples have a higher
f at qf
than CK0 UC test samples, the amount of friction mobilized at qf
is more in isotropically consolidated samples than in K o consolidated samples at low OCR and hence K o values.
At (a 1/
3)
, the NESE is hypothesized to represent the
fully mobilized frictional strength of both clay and silt layers.
The fact that,.despite having different failure mechanisms
(CIUC test samples failed by bulging while CKoUC test samples
failed by developing distinct failure planes across the varves
at the end of the test), NESE at (/a3)max
of CIUC and CK UC
test samples are identical (Fig. 5.8) suggests that in the undrained tests the frictional strengths of both clay and silt
layers in the CIUC and CKoUC samples are fully mobilized.
should be noted in Fig. 5.8 that the same NESE at (
It
1 /a3 )
max
153
also holds for both NC and OC samples from CIUC and CKoUC tests.
Consequently, NESE at (l1/a3)max for vertical loading is independent of the stress history and stress system during consolidation.
At low OCR, the decrease in Tff/avc at (a1 /a3 )
with
max
CKoUC tests is due to the fact that CKoUC samples exhibit more
strain softening.
This can probably be explained by the fact
that distinct failure planes formed in the K o consolidated
samples
The "soil structure" effect on NSP might be inferred by
comparing NSP of samples from CIUC and CKoUC tests at an OCR of
four, since-in both tests shear started at the same stresses.
Table 5.2 shows that qf(V) and Af values of these samples are
practically identical.
These results, therefore, indicate little
effect of soil structure on qf(V) and Af.
However, soil struc-
ture apparently does have an effect on the undrained modulus,
as indicated
in Fig. 5.9.
In conclusion, the stress system during consolidation does
not significantly affect qf(V) and NESE at (a1 /a3 )
for vermax
tical loading, but it has an effect on the envelope at qf and
the stress strain and pore pressure characteristics.
5.2.4
Results from CIUE and CK UE Tests
Test results from the CKoUE and CIUE tests are summarized
in Table 5.3 and Figs. 5.5, 5.6(b), 5.7(b) and Figs. 5.10 to
5.13.
These data show that K o consolidation relative to iso-
tropic consolidation:
154
1.
Decreased qf/ vc by about 15% practically independent of OCR (Table 5.3).
2.
Decreased Af at qf and ( 1/3)
of normally
consolidated samples, but did not significantly
change Af of overconsolidated samples (OCR=2 and
4), shown in Fig. 5.5.b.
Unlike the samples from
triaxial compression tests, samples from CIUE and
condition
max
Thus, for CUE and CKoUE
CK UE tests reached qf and (a/a 3 )
at the same strain.
test samples, Af and NESE at qf are identical to
those at (
1 /a3 )
. However, ~ (c=0) of NC sample
max
is increased due to lower Pf/3vc (Table 5.2).
3.
Did not change ef at qf or (a1/3)
. The f
max
is also practically independent of OCR (Fig.
5,6(b)),
4.
Decreased Eu/-vc at the same strain of both NC
and OC samples (Fig. 5.10).
5.
Had little effect on Eu/qf at OCR of one, but OC
CKoUE samples have lower E/qf
samples.
values than CIUE
This reduction increases with increas-
ing OCR (Fig. 5.13).
The fact that samples from both CIUE and CKoUE tests had
distinct failure planes across the varves (though less visible
in NC CKoUE test sample) probably partly explains why the NESE
at ( 1 /; 3)
or at qf from CIUE and CKoUE tests are identical.
The lower pf in the NC CKoUE test sample relative to that in
155
the CIUE test sample explains why
(c=0) of the NC
W
E test
sample is larger, since NESE are identical.
Since NESE is not affected by the stress system during
consolidation, the decrease in qf(H)/vc
and Eu/3vc of CKoUE
test samples relative to those from CIUE test samples is caused
by the development of larger excess pore pressures in CKoUE
test samples.
These are two possible reasons to explain this
different excess pore pressure at failure (Auf).
1.
The difference in the magnitudes of Aqf due to the
rotation of principal planes:
The principal planes
are rotated by 90 degrees in CKoUE test samples
which have K o less than one.
In these samples,
shearing first causes q to decrease to zero and
then to increase in the horizontal direction.
OCR=1.0, when q = 0, the K
At
consolidated sample has
a much lower p than the isotropically consolidated
sample (Fig. 57).
Thus, for this NC CKoJE sample,
in spite of having a lower Af, pf is smaller.
2.
A difference in soil structure:
The CKoUE test
samples may have a more preferred particle orientation compared to that of the CfIUEtest samples,
in spite of insignificant differences in the final
water contents.
It is difficult to assess the magnitude of the effect of
soil structure on the difference in *qf. However, the fact that
the K o consolidated sample developed larger pore pressure during
156
shear at OCR = 4 (Ko=1l),suggests that there is a soil structure
That is, a more preferred particle orientation leads
effect.
to an increase in the pore pressure developed in CKoUE samples,
i.e., when alf acts in the horizontal direction.
This hypothe-
sis is supported by the results of Duncan and Seed (1966(a)) on
Kaolin.
In summary, the decrease in qf/vc
and EU/Uvc of NC CKoUE
samples are thought to be due to both the effects of the differ
ence in Aqf and in the soil structure, while the decrease in
and E
qf/vc
v
of OC samples are mainly due to the differences
in the soil structure.
Test data in this section, therefore, show that the stress
system during consolidation has a considerable effect on qf/v
Af, and E/avc
from horizontal loading.
(or at qf) nor perhaps
effect
on NESE at (
5.2.5
Summary and Discussion
A.
1 /a3 )
However, it has no
f.
Summary of Principal Findings from the Study of K.
Versus Isotropic Consolidation
(a) Behavior from Vertical Loading (Triaxial
Compression)
Test data from CIUC and CKoUC tests show that:
1.
The stress system during consolidation has an insignificant effect on qf at all OCR values studied,
due to the compensating effect of Ko on developed
pore pressure and NESE at qf.
The CKoUC test
samples (1 < OCR< 4) have a lower NESE at qf, but
have larger pf values than the CIUC test samples.
157
At an OCR of four, CKoUC and CIUC test samples
have essentially the same qf and Af (also Ko=l).
2.
The CKoUC test samples, especially those at OCR
of one and two, exhibit considerable more strain
softening behavior than CIUC test samples. That
is probably due to the formation of failure planes
during shear in the CKoUC test samples.
3.
Because
f at qf of CoUC
test samples, especially
that of the NC sample, is smaller than ef of CIUC
test samples, NESE at qf from CKoUC tests is lower.
4.
For both CIUC and CK UC test samples, the difference between NESE at qf and (/a3)max
decreases
with increasing OCR, as the difference in axial
strains at qf and ( 1 /a3)
of both tests also
max
decreases with increasing OCR.
5.
Despite having different failure mechanisms (CIUC
test samples having failed by bulging while (CKoUC
test samples developed distinct failure planes
across the varves), NESE at (a /a 3 )
from CK0 UC
max
and CIUC tests are identical.
6.
Eu/qf from OC CIUC test samples are lower than those
from OC CKoUC test samples, though Eu/qf of NC CIUC
and CKoUC test samples are practically identical.
(b) Behavior from Horizontal Loading (Triaxial Extension)
Test data from CKoUE and CIUE tests lead to the follow-
ing findings:
1.
CKoUE test samples have about 15% lower qf than
158
CIUE test samples at OCR = 1 to 4.
2.
The decrease in qf is due to the lower pf (increased
pore pressure), while the NESE is unchanged.
For
normally consolidated CKoU test samples, the increase
in
Auf is thought to be due to both the increase in
Aqf and to differences in the soil structure.
For
overconsolidated samples, such a decrease is mainly
due to the difference in the soil structure.
3.
For both CIUE and CKoUE tests, the qf and (1/03)
conditions occur at the same strain and they have
the same failure mechanism.
(Both tests develop
a failure plane across the varves.)
Thus, their
NESE at (al/a3)max which are also the same as those
at qf are identical.
B.
Uniqueness of NESE at (Ol/a3)max
It is seen in Figs. 5.7 and 5.8 that the same NESE
at (/3)max
holds for both NC and OC samples from CIUC, CKoUC,
CIUE and CK UE tests.
Furthermore, upon comparing the stress
systems of CIUC, CKoUC, CIUE and CKoUE tests, one can conclude
that NESE at (
3)max for failure across the varves (since
both TC and TE test samples have potential failure planes across
the varves)
is independent
of:
1.
The stress system during consolidation;
2.
The magnitude of
2
and the rotation of principal
planes (hence, the stress system during shear);
3.
The magnitude
of
v
(m
of CIU tests is 4.0 kg/cm 2
159vm
159
while av
vm
is 3.6 kg/cm2 for some of the CKoU tests;
0
additional data from RECOMPRESSION samples (presented in Chapter 6) which have aVm ranging from
1,6 to 2.4 also indicate this independency).
The fact that despite bulging type of failure observed at the
end of the test, the NESE at (al/a3max from CIUC tests are identical to that from CKoUC, CUE
and CKoUE tests, in which their
samples developed distinct failure plane across the varves at
the end of the test, suggests that CTUC test samples also failed
across the varves.
Additional data confirming the independency of the NESE at
(al/ 3 )max from the magnitude of
2 from CIUC (G=90) and CIUE
(8=90) tests on NC and OC samples and from a CKoUPSC test and
from two CKoUPSE tests on NC soil will be presented in Section
5.3.
The uniqueness of NESE at (l/a3)max suggests that it can
be considered as a material property.
On the other hand, NESE at qf, when it is not equal to that
at (al/a3)max (e.g., NESE at qf for CIUC and CKoUC tests when
> 0.3) can be greatly affected by the stress system during
Pf/vm
consolidation and the direction of loading.
5.3
EFFECT OF INTERMEDIATE PRINCIPAL STRESS
5.3.1
Method of Investigation
The investigation of the effect of
2 on NSP was conducted
on both K o and isotropically consolidated SHANSEP samples
(Table 5.1).
160
The effect of a 2 with CIU tests is shown by comparing NSP
from CIUC and CIUE tests.
CIUC tests on vertical samples (=0)
and CIUE tests on horizontal samples (=90)
study the effect for vertical loading.
veritical samples (=0)
samples (=90)
are performed to
NSP from CIUE tests on
and those from CIUC tests on horizontal
are compared for the case of horizontal loading.
Figure 5.14 presents the state of stress during shear of these
samples.
It is seen in Fig, 5.14 that, when using horizontal
triaxial samples, one of the a2 directions is perpendicular to
the varves instead of being parallel as should occur in situ.
For CKoU tests, the effect of a 2 is shown by comparing NSP
from triaxial and plane strain condition.
NSP from a CKoUPSC
test on a normally consolidated sample from Northampton (Lacasse
et al., 1972) and that from a CKoUC test on a normally consolidated sample from Amherst are compared for the case of vertical
loading.
For horizontal loading, NSP from two CKoUPSE tests
(Lacasse et al., 1972) and those from a CKoUE test are compared,
all being normally consolidated.
Since the DSS device is thought
to have a state of stress during shear closer to.that of plane
strain (Ladd and Edger, 1972),it was also used for tests on two
normally consolidated inclined (=45)
samples.
These CKoUDSS(C)
and CKoUDSS(E) tests simulate CK0 UPSC and CKoUPSE tests, respectively.
Section 5.3.4 will present the testing procedures and
results as well as the problems of interpreting test data from
these special DSS tests,
5.3.2
Results from CIU Tests
Test results from U-UC and C-IUEtests on vertical and hori161
zontal samples are summarized in Tables 5.4 and 5.5 and Figs.
These test results are from samples which have
5.15 to 5.19.
approximately the same moisture content.
Thus, the differences
in NSP from CIUC and CIUE test samples are mostly the result of
the difference in the magnitudes of 02 and perhaps also in the
direction of
2 relative to the varve structure.
Test data
show that for vertical loading, shear in extension is relative to
shear in compression:
1.
Slightly increased (though considered insignificant
for practical purposes)
(~1/~3)
max
2.
qf(V)/avc and
Tff/vc
at
for both NC and OC samples;
Caused an increase in pore pressure parameter A for
OC samples, but had no effect on A for NC soil
(Fig. 5.19);
3.
Increased NESE at qf, though NESE at (a1/3)max
is
not affected (Fig. 5.17);
4.
Increased Eu/avc at e =0.5%, though the maximum increase was less than 30% (Table 5.4);
5.
Increased
(c = 0) of NC samples (Table 5.4).
For horizontal loading, shear in extension relative to shear
in compression:
1.
Decreased qf(H)/vc
and Tff/avc at (/f3)max
by
5 to 15%, this effect also decreasing with increasing OCR (Table 5.5);
2.
Increased the pore pressure and hence A (Fig. 5.19);
3.
Increased
(assuming c = 0) of NC soil
162
(Table 5.5)
because of lower Pf/Vm;
4.
Did not change NESE at qf or (al/a3)max (same
envelope), Figs. 5.17 and 5.19;
5.
Slightly decreased E /0vC at
= 0.5% but the maxi-
mum decrease is only 30% or less (Table 5.5).
An analysis of the effect of a2 on E/avc
can not be made
without assuming that varved clay is an isotropic material.
The
vertical strains were measured in 0=0 samples, whereas the "horizontal" strains (with respect to the in situ orientation) were
measured in 0=90 samples.
shown by comparing Eu/vc
Since the effect of a 2 on E/Ov
is
from CIUC (.=0) and CIUE (8-=90)tests
for vertical loading, and those from CIUC (=90)
and CIUE (0=0)
tests for horizontal loading, it is necessary to assume that the
material behaves as if it were isotropic.
With this assumption,
a2 has relatively little effect on Eu/avc for both vertical and
horizontal loading
within 30%.
(Tables 54
and 5.5), the variation being
Though a difference of 30% is not considered signi-
ficant (considering the limitation in the method of analysis),
it can not be concluded that a2 has no effect on E/
vc because
of the unknown direction of the error due to the assumption of
isotropy.
The fact that the modulus from
than for 0
=90 samples is always larger
samples, independent of the type of test (i.e.,
when comparing TC (=90) versus TE (8=0) tests or TC (=0)
sus TE (=90)
ver-
tests) suggests that the use of samples with dif-
ferent inclinations affects the outcome of the results at least
as much as the value of the intermediate principal stress.
163
For an isotropic "ideal" material, interpretation of test
data in terms of Toct and aoct' rather than p and q, and use of
pore pressure parameter, a, should account for the changes in
the magnitude of a22O
.
octversus
ct
different.
However, Fig. 5.18 shows that the
ct Aoct)
Oct
Oct m
[( oct ) m = vm ) from TC and TE tests are
For both vertical and horizontal loading, the pore
pressure parameter a from CIUC test samples is always greater
than that from CIUE tests (Tables 5.4 and 5.5).
Therefore, the
increase in the magnitude of a2 decreases the pore pressure parameter a.
The effect is more pronounced with vertical loading
than with horizontal loading.
The decrease in a is also larger
in OC samples than in NC samples for both loading directions.
Figure 5,17 shows the NESE at qf and (a/a 3 )max from the
test series.
These data and data in Fig. 5.7 show that NESE at
(al/3)ma
x for
ent of:
failure across the varves is practically independ-
(1) the stress system during consolidation; (2) the
intermediate principal stress; (3) rotation of principal planes;
and (4) the magnitude of avm
vm (limited data only; more data being
presented in Chapter 6).
It is stressed that the effects of the differences in the
magnitudes of a2 and in the directions of Aa2 relative to the
varve structure (perpendicular or parallel to the varves) shown
from comparing TC and TE test data on vertical and horizontal
samples can not be separated.
be as important as the a2 .
In fact, sample orientation might
Thus, no definite conclusion regard-
ing a2 can be drawn from this test series.
164
5.3.3
Results from K
Consolidated Plane Strain and Triaxial Tests
Figures 5.20 and 5.21 show normalized stress-strain curves
of normally consolidated samples from CKoUC and CKoUPSE tests,
and those from CKoUE and CKoUPSE tests respectively.
The samples
from the CKoUC and CKoUPSC tests reached the qf condition at a
very low strain and the (1/a3)max condition at a large strain
(Fig. 5.20).
The CKoUE and CK o UPSE tests, on the other hand,
reached qf and (al/a3)max simultaneously at much larger strains
especially the TE test (Fig. 5.21).
Normalized effective stress
paths for these tests are plotted in Fig. 5.22.
These data and
the data in Table 5.6 and Fig. 5.23 show that shear in plane
strain relative to shear in triaxial tests:
1.
Slightly increased qf(V)/avc by 10%;
2.
Increased qf(H)/avc by 20%;
3.
Decreased Af for both vertical and horizontal
loading ;
4.
Decreased the shear strain at failure (Yf) computed from the linear elasticity theory (Table 5.6);
5.
Increased E/v
6.
Increased
u
vc
(Fig. 5.23);
(assuming c = 0) at qf for vertical
loading (i.e,,
PSC >
TC), but decreased f
c = 0) at qf for horizontal loading
larger Pf/am
(i.e., a TE> i PSE).
loading, 5 at (1/03)max
because of
For vertical
from CKoUC and CK UPSC
tests are almost identical.
It should be mentioned that because of the inherent problems
with friction forces in the plane strain device, the results
165
from the plane strain tests may be questionable.
Any frictional
forces along the vertical boundaries of the rectangular specimen
will affect the measured vertical force.
Thus, the qf/avc and
EU/avc measured from the plane strain test may be too high.
In
plane strain compression (PSC) tests, the errors in measuring
qf/avc and perhaps Eu/avc probably are small because the direction of the major principal stress during consolidation and shear
are the same.
Because of the difference in loading direction
during consolidation and shear, the errors in measuring qf/vc
and Eu/avc from plane strain extension (PSE) tests will be more
important.
Despite having friction problems in the apparatus, the increases in qf from triaxial to plane strain tests for both vertical and horizontal loading agree with the other test data reviewed in Chapter 2.
As shown in Table 2.7, data from Boston Blue
clay (Ladd and Edger, 1972) show an increase in qf/avc from TE
to PSE of about 20%.
Data from Haney clay (Vaid and Campannella,
1974) also show the same increase of about 20%.
Data from
CKoUPSC and CKoUC on these two soils also show an increase in qf
from TC to PSC similar to that for the Connecticut Valley varved
clay.
Regarding the measurement of undrained modulus, it is difficult, if not impossible, to make any assessment of the effect
of friction forces.
However, Boston Blue clay data obtained
from using the same plane strain equipment show no practical
difference in E/
c
from NC CKoUPSE and CK UE tests, despite
166
having friction problems in the apparatus (Ladd et al., 1971).
Since the increases in Eu/avc (Fig, 5,23) from TC to PSC and
from TE to PSE are very large, there may be a significant effect
of
2
on E/vC,
though probably not as large as measured.
The plane strain test sample also failed across the varves
(though only the CK0 UPSC test developed distinct failure planes).
Thus, based on the result of Section 5.3.2, the NESE at (al/a3)max
for failure across the varves (shown to be independent of the
stress system during consolidation and of the magnitude of a2)
shown in Fig. 5.17 should be identical to that from the plane
strain tests (Fig. 5,22),
Figure 5.22 shows the NESE at qf and
(al/a3 )max conditions from CK0 UC and CR0 UE tests together with
normalized stress paths from CKoUPSC and PSE tests on NC samples.
This figure shows that;
1.
The magnitude of a 2 has little effect on NESE at qf
and no effect on NESE at (A1/
3 )max
for vertical
loading.
2.
Test results from the CKoUPSE tests fall slightly
from CKoUE tests
max
(also see Fig. 5.26) and from CIUE (=0) and CIUC
below the NESE at (/a
(8=90) tests.
3)
This might be due to the fact that
the envelope at high pf may not be a perfect straight
line* and/or due to friction problems in the plane
strain device.
*Note that data from CIUC (=90)
fall slightly below the envelope.
167
tests on NC samples also
Since qf values for OC plane strain tests are not available,
they were estimated as follows.
qf at avc/qf at
Figure 5.24 shows a plot of
vm (equal to SR shown in Fig. 5.24) versus OCR
from several types of tests,
It is seen that for the same load-
ing direction (i.e., either horizontal or vertical loading),
the ratio qf at
SHANSEP test.
vc/qf at
vm is independent of the type of
Therefore, qf values for a particular loading di-
rection at a given OCR for a plane strain test can be reasonably
estimated by the relation given below.
qf/vc
qf/aVC
of the NC sample x SRx OcR
=
(Eq. 5.3)
Using values of SR for horizontal and vertical loading, Fig.
5.2.4 presents the calculated plane strain test qf/Uvc values
from Eq. 5.3 and also the measured data from triaxial tests.
Use of triaxial tests to estimate S
will obviously lead to lower
strengths than plane strain tests and thus presumably yield somewhat conservative results for field situations involving plane
strain loading conditions.
If one wants to obtain the estimates of plane strain qf
values from triaxial tests, CIUC and CIUE tests on vertical
samples seem to be the appropriate tests to perform.
As shown
below, the agreement in qf between CKoUPSE and CIUE tests is
fairly good, while qf from CIUC tests are about 10% lower than
those from CK UPSC tests,
168
.f/avc
f
from Vertical
OCR
-o
CK UPSC
1
2
4
c from Horizontal
Loading
Loading
--
CIUC
0.28
0.47
0.76
CK UPSE
.25
.46
.75
.25
.44
.72
0.245
0.42
0.68
CIUE
Note, however. that this approach may not be equally applicable
to all varved clay deposits.
Since NESE at (1/3)ma
3
max
for failure across the varves is
a material property (see Section 5.2.5 and later
Section 5.3.4),
it is believed that NESE at qf which is the same as that at
(l/3)max
from CKoUE tests should also hold for OC samples
from CK UPSE tests.
Furthermore, based on NC test data, NESE
at qf when pf/sVm> 0.3 for NC and OC CKoUPSC tests may be slightly
above the NESE at qf from CKoUC tests (though by very little).
The estimated values of
and c/vm
vm for NESE at qf from CKoUPSC
0
tests are 160 and 0.080 respectively (versus
= 0.088 from CKoUC tests).
*=
14.50 and c/am
Vm
These NESE values will be used for
the interpretation of anisotropic behavior in plane strain loading presented
in Chapter
7.
In conclusion, the comparison of NC NSP from CK o U plane
strain and triaxial tests on Connecticut Valley varved claysgenerally indicates trends similar to those for other homogeneous
clays.
169
5.3.4
Interpretation and Results from Special DSS Tests
Problems with the DSS Test
Since the state of stress of the sample in the DSS test
is rather complex and only the normal and shear stresses on a
horizontal plane are measured, the problems that must be considered in the analysis of DSS test results are:
1.
The complete state of stress during shear is unknown;
2.
The state of stress in the sample is not uniform.
A uniform stress distribution is expected in the middle of the
samples, but high stress concentrations occur along the edges
of the samples (Duncan and Dunlop, 1969; and Lucks et al., 1972).
Based on the above mentioned problems, uncertainty regarding
the state of stress at failure is a major problem.
Ladd and
Edgers (1972) analyzed four possible assumptions which might be
used for estimating the state of stress at failure.*
These are
obtained by assuming:
a.
That the applied stress system is one of pur shear,
i.e., A
b.
=
Aa 3 = ATh
That the horizontal plane is a failure surface, i.e.,
Th =
C.
1
Tff and Ovf =
ff
That the horizontal shear stress is the maximum shear
stress in the sample, i.e., Th = qf and avf . = pf.
d.
The location of Mohr-Coulomb failure envelope.
The
smallest Mohr circle representing the state of stress
at failure must then pass through Th and
*Arbitrarily defined either at Thmax or at
170
f and be
(Th/av)max
tangent to the assumed Mohr-Coulomb envelope.
Provided the failure envelope is known, assumption (d), where
the location of Mohr-Coulomb failure envelope is used, is most
reasonable.
Based on the analyses performed on NC, CH and CL
clays, using DSS (=0) tests by Ladd and Edgers(1972), the pure
shear assumption yields qf values which are significantly larger
than those from assumptions (b), (c) and (d).
The pure shear
assumption is only reasonable when the soil is assumed to behave
like an elastic isotropic material (Duncan and Dunlop, 1969; and
Lucks et al., 1972).
Thus, this assumption is questionable and
probably should only be used at low strains where the behavior
of a soil may be nearly elastic.
Interpretation of the Special DSS Data
For this portion of the study, the effect of a 2 on NSP (qf/
avc' Af and Eu/Ovc) are shown by comparing NSP from special DSS,
PS, and triaxial tests.
NSP from CKoUC and CKoUPSC tests and f-
rom a CKoUDSS(C) test on a normally consolidated inclined sample
are compared for the study of the effect of intermediate principal on stress-strain-strength behavior in the vertical loading.
For horizontal loading, NSP from CKoUE, CKoUPSE and CKoUDSS(E)
tests are compared.
The assumptions employed for the interpretation of special
DSS test data are as follows:
(1) During consolidation:
The normal and shear stresses
are applied to an inclined sample so that "Ko consolidation"*
*True K consolidation can not be maintained because the
stresses on he vertical sides of the inclined specimen can not
be controlled.
171
is maintained during consolidation.
Such a condition requires
that the complimentary shear and normal stress be developed on
the vertical sides of the inclined specimen.
As a result, the
consolidation shear stress on the inclined (a) plane, Tac
avc (-Ko)sinecosO;
and the consolidation normal stress on the
inclined (a) plane, sc'
(vc(Cose+K
sin2 8); note that
=
is
450 in this case.
Two separate assumptions are used, namely:
(2) During Shear:
the pure shear sssumption (i.e., assumption (a)) and that of the
location of the failure envelope (assumption (d)). The first assumpResults.
tion was mainly used to illustrate that it yieldsunreasonable
For the pure shear assumption, the states of stress during
and CK0 UDSS(E) test samples
consolidation and shear of CKoUDSS(C)
0
are shown in Fig. 5.25.
As a result of assumption (1), the
origin of planes, i.e., the pole, of the Mohr circle representing the state of stress during consolidation is located at the
point defined by the stresses Tac and a
ac
of the varves is 450.
CT
defined as pc
, since the inclination
a0
This point (-Fig.525(a:):)cou-ldalso be
+a
-
and q
q
vc
.
For a compres-
sion test, simulating a CKoUPSC condition, the sample is sheared
by increasing
Ta
.
Figure 5.25(b) shows Mohr circles represent-
ing the states of stress prior to shear and those at failure.
Upon assuming that the applied shear stress is that of pure
shear, the origins of planes for the total stress circle (TSC)
and the effective stress circle (ESC) at failure are located at
a point defined by pf = Pc and qf for the TSC and at a point defined by pf and qf for the ESC.
Thus, alf in the CKoUDSS(C)
test acts perpendicular to the varves and the principal planes
172
are not rotated during shear.
Therefore, based on these assump-
tions, a CK0 UDSS test on an inclined sample should measure qf(V).
In the CKoUDSS(E) test (Fig. 5.25(c)), the sample is sheared by
reversing the direction of T .
As a result of using the pure
shear assumption, the pole of the TSC moves from the point defined by Pc and qc before shear, to the point defined by p =
pf and q = -qf at failure.
Thus, the principal planes in the
CKoUDSS(E) test sample are rotated by 900 and alf is in the direction parallel to the varves.
presumably measured qf(H).
Therefore, the CKoUDSS(E) test
From the above discussion, it is
seen that the varved inclination of 450 is selected so that qf(V)
and qf(H) could be measured, assuming a pure shear state of
stress during shear.
For analyses based on an assumed location of the failure
envelope, a NESE for failure across the varves was selected
(referred to as NESE assumption).
Because of the uniqueness of
NESE at (l/a 3)max for failure across the varves (i.e., it is
a material property), it can be used with reasonable confidence
for the interpretation of CKoUPSS(E) test data on a NC inclined
sample which reaches (Ta)max' the maximum shear stress on the
inclined plane, and (T /ao)max
condition at the same strain.
The experimental evidence supporting the uniqueness of NESE at
(al/U3)max and thus NESE is given below.
All the test data [pf/vm versus qf/avm at ( 1/a 3)max]
pre-
viously presented in Sections 5.2 and 5.3 are plotted in Fig.
5.26.
It also includes data from two stress-controlled, iso-
tropically consolidated drained triaxial compression unloading
173
tests (CIDC(U)) on heavily OC RECOMPRESSION, Northampton samples,
developing distinct failure planes across the varves observing
at the end of the tests (Lacasse et al., 1972).
Figure 5.26
shows the following.
1.
Results from NC samples from CIUC (=0),
CIUE (0=90), CIUE (=0),
CIUC (=90),
CK UC, CK UPSC, CKoUE and,
to a less extent, CK UPSE tests lie on the same NESE
at (a-/a3)max for failure across the varves.
Further-
more, the same NESE at (1/a3)max holds for OC
samples from the above types of tests.
2.
Thus NESE at (a1 /a3 )
for failure across the
max
varves is independent of the stress system during
consolidation and during shear.
This is one of
the two necessary requirements; the other is the
fact that the same NESE holds for both NC and OC
samples.
3.
Though data are limited, NESE at (a1 /a 3)max for
failure across the varves is also independent of
drainage conditions (based on CIDC(U) tests) and
the magnitude of
Chapter
vm (see additional data in
6).
Since NESE at qf when Pf/vm>
0.3 from CKoUC and CKoUPSC
tests are practically identical (and therefore, practically independent of the magnitude of A
2 ),
the NESE at qf from CKoUC
tests is thought to hold for practical purposes for both NC and
174
OC triaxial and DSS samples.
This NESE will be used* for the
interpretation of DSS test data at (T )ma x from the CKoUDSS(C)
tests.
At (T/o) max, NESE at ( 1/G 3)
for failure across
max
the varves will be used.
The suitability of the pure shear assumption can be evaluated by comparing, for both vertical and horizontal loading,
qf/avc values estimated from the pure shear assumption to those
from the NESE assumption and to those from the triaxial tests,
keeping in mind that qf/avc from the CK UE test should be the
lowest value (since A
1=
A 2>Aa 3 in TE tests).
Experimental Procedure
The only portion of the special DSS experimental procedure
J
which differs from that presented in Section 3.5 is the method
In this case, normal and shear stresses were
of consolidation.
applied to inclined samples.
Upon using assumption (1) (the
state of stress during consolidation), the magnitude of T c and
ac were computed as follows:
acC
a=
For
=
vc(c° s 2
Vc
vc(
+ K
o
sin 2 8)
- K o) sinOcosO
= 450; avc
=c
he
(IC
+
(C
ac -Tac
Eq. 5.4
Eq. 5.5
Eq. 5.6
I
tzq.
A -.. r- ID.4-,fq.
D .D)
Eq. 5.7
(Eq. 5.4-Eq. 5.5)
Since both avm and K o of East Windsor varved clay were known
in advance, it is possible to consolidate so that the "assumed"
*Since the principal planes are not significantly rotated
in the CKOUDSS(C) test, the independence of NESE from the magnitude of A 2 is the major requirement.
175
Ko consolidation could be maintained in every load step.
For
the purpose of preventing the sample from failing at low stresses,
only ac
was applied during the first three load increments
(i.e., tac = 0).
weight.
Tac is applied to the samples by using dead
Table 5.7 shows the load increments (and the computed
values of
vc and ahc) that were applied to the samples from the
CK0 UDSS(C) and CKoUDSS(E) tests.
It should be noted that small
load increments were used to prevent excessive deformation in
the sample.
Experimental Results
The normalized stress-strain curves from CKoUDSS(C) and
CKoUDSS(E) tests on normally consolidated inclined samples are
shown in Fig. 5.27.
The CKoUDSS(C) test sample reached Tamax'
presumably equivalent to the qf condition, at a fairly low strain
and (T /a )max
presumably equivalent to the (l/a3)max
condi-
The CKoUDSS(E) test sample, on
tion, at a much larger strain.
the other hand, reached the Tamax and (T /
x )ma
the same very large shear strain.
condition at
Hence, the CK UDSS(C) test
behaved like a CKoUC test and the CKoUDSS(E) test had a behavior
similar to that of a CKoUE test.
Therefore, this similar be-
havior suggests that the uses of NESE at qf for (Ta)ma
tion and of NESE at (a1/a3)
for
(Ia/a)ma
x
x
condi-
are appropriate.
Figure 5.28 shows Mohr cirlces representing the state of
stress at qf/Ovc from the two tests determined from the NESE
assumption.
Values of qf/avc and Pf/avc from CK0 UC, CKoUE,
CKoUPSC, and CKoUPSE are also plotted for comparison.
The
points defined by qf/avc and pf/vc from CKoUC and CK0 UE tests
176
tests fall almost exactly on the Mohr circles from the CKoUDSS(C)
and CK UDSS(E) tests respectively determined from NESE assumption.
O
The pore pressure parameter, Af, for the DSS tests was estimated by assuming
dated samples.
a value of K
(K
= 0.68) for normally
Table 5.8 shows how the values of Af are obtained.
Table 5.6 presents the comparisons of qf/ vc (or
Af
and
consoli-
at the qfand(l/a 3 )
r ff/svc),
Yf,
condition from the triaxial, plane
strain, and special DSS tests (determined from NESE assumption).
These data show that for both vertical and horizontal loading:
1.
qf/avc values from the DSS and triaxial tests are
identical.
2.
Af at qfand
at (c 1/Ca3)
from the DSS are roughly
the same as those from triaxial tests, especially
in the horizontal loading.
3.
qf/vc
values from the plane strain tests are
higher than those from the DSS tests; Af values
are also lower.
/
_
4.
_
~ (assuming c = 0) at qf from the DSS and triaxial
tests are equal (since NESE and qf are identical).
However,
from NC CK UPSC test is larger than
that from NC CKoUC test, and
from NC CKoUPSE
test is lower than that from NC CKoUE test.
B angles determined from Fig. 5.28 are 9
test and 1110 from the CK UDSS(E) test.
B
from the CKoUDSS(C)
The small deviation in
from 0° and 90° respectively indicates that qf values deter-
mined from CKoUDSS(C) and CKOUDSS(E) tests could logically be
177
compared with those from CKoU triaxial and CK U plane strain test.
The agreement between the special DSS and CKoU triaxial
data for both vertical and horizontal loading
does not mean
that the states of stress in the DSS and triaxial test are identical and that the
2
has no
effect on these soil parameters
for the following two reasons:
1.
avc from the DSS test may not be correct.
The com-
plementary shear and normal stress cannot develop
on the vertical sides of the specimen, especially
at the edges of the specimen.
Consequently, the
"computed" avc may be too large compared to that
computed from considering no development of shear
stress on the vertical sides of the sample, and
vc "measured" from DSS tests might be
thus, qf/vc
f
too small.
2.
The use of 450 varved inclination is somewhat approximate, though it generally yields a larger (T)max
to using 8 = 45+ Fj/2 for vertical
compared
and using
= 45 -
loading
/2 for horizontal loading.
(See data from two homogeneous clays reported in
Chapter
2.)
As a result, the agreements between Af and qf from the special
CKoUDSS and triaxial tests may be fortuitous.
For both vertical and horizontal loading, the decrease in
qf/Ovc from CK0 U plane strain to special CK0 UDSS test could
partly be the result of:
178
1.
qf/0vc from the plane strain test, especially
that from the CKoUPSE test, might be too high
because of the friction problem in the plane
strain device;
2.
qf/avc measured from the special
SS tests is too
low because of problems in the interpretation of
DSS test data.
Mohr circles representing the state of stresses at qf obtained from the pure shear assumption are compared to the previous results and shown in Fig. 5.29.
Although qf/fvc derived
from the NESE assumption may be too low for both vertical and
horizontal loading, qf/avc obtained from using the pure shear
assumption is even lower.
(Note that the same avc was used for
computing qf/avc values for both assumptions.)
For horizontal
loading, the pure shear assumption gives a qf/avc equal to 0.19
compared to 0.21 from the CKoUE test.
Since the latter value
represents the minimum strength, the pure shear assumption is
not reasonable.
Eu/vc
values at various applied shear stress levels (the
applied shear stress level being equal to Aq/Aqf for triaxial
and plane strain tests and equal to ATa/(Aa)ma
x
for the special
DSS test) from CKoU triaxial, plane strain, and special DSS
tests shown in Fig. 5.23 are based on the assumption that the
soil is a linear isotropic elastic material.
It is seen that
only test data from CKoUE and CKoUDSS(E) tests show good agreement with respect to modulus.
Such agreement, however, may be
179
fortuitous.
For vertical loading, the special DSS test gave a
much lower value of modulus than that from the triaxial compression test.
In summary, the major findings from this section are as
follows:
1.
The agreements in qf/avc, Af and
(c=O) of NC
soil from the CKoU triaxial and DSS tests, for
both vertical and horizontal loading, are probably
fortuitous.
2.
(since it is a material
(a1 /a3 )
at
The use of NESE, especially
property),
max
for determining
the state of stress in DSS tests is apparently
fairly reasonable, at least compared to the pure
shear assumption.
In general, it is thought that
NESE at qf, if not equal to that at (a1 /a3 ) max
max
is dependent upon the stress system.
For the
interpretation of the CKoUDSS(C) test, it is fortunate that NESE at qf from CKoU vertical loading
tests appears to be little affected by the magnitude
of Aa2 and that the same NESE at qf holds for both
NC and OC samples.
(Thus, NESE at qf, when Pf/avm>
0.3 from CK0 UC tests is thought to hold for NC
CK0 UDSS(C) test samples.)
These requirements are
sufficient for an interpretation of CKoUDSS(C)
test data, since the principal planes are not expected
to be significantly rotated.
180
5.3.5
Summary and Discussion
Tables 5.9 and 5.10 summarize the general effect of Aa 2
presented in Sections 5.3.2, 5.3.3, and 5.3.4.
The important
points that need to be further mentioned are as follows:
1.
For both vertical and horizontal loading, the agreement of qf/avc and Af from CKoU triaxial and special
DSS (8=450) tests are probably fortuitous.
This is
due mostly to inherent problems in interpreting DSS
test data, especially those arising from the assump-
2.
tion for computing ac
vc
The difference in the values of qf/avc from the
CK UPSE and CKoUE tests on NC soil is somewhat ampli0
fied by inherent problems (friction forces) in the
plane strain device.
This error is, however, less
significant when test data from CKoUC and CKoUPSC
tests are compared.
Despite the friction problem in
the plane strain device, for both vertical and horizontal loading, the trends observed from the plane
strain and triaxial tests are similar to those observed from other homogeneous clays (see Chapter 2),
i.e., the effect of
2 is more pronounced in the ex-
tension than in the compression test.
3.
The behavior observed from the CIU tests is probably
not mainly due to effect of the intermediate principal stress.
Because of using horizontal (=90)
samples, the effect of inherent anisotropy (defined
181
in Chapter 2) is included.
of A
2
One of the directions
acting on the horizontal sample is perpendic-
ular to the varves instead of being, in every direction, parallel to the varves, as for the vertical
sample.
Although for both vertical and horizontal
loading , the trends of qf from NC CIU tests are similar to those observed from comparing CKoU plane
strain and triaxial test data, these trends might
equally be explained by using different sample
orientations.
5.4
EFFECT OF TESTING METHOD FOR SHEARING PARALLEL TO VARVES
5.4.1
Introduction and Method of Investigation
For shearing across the varves, the effects of the stress
system during consolidation and the magnitude of the intermediate
principal stress on NSP were studied in Sections 5.2 and 5.3.
In this section, the effect of stress system on NSP for shearing
parallel to the varves is studied.
The influence is shown by
comparing results from two types of tests, namely:
CIUC tests
on inclined samples and CKoUDSS tests on vertical samples.
The stress system differences between CIUC and CKoUDSS
tests are as follows:
1.
For the stress system during consolidation, the
samples from CIUC tests are isotropically consolidated, while those from DSS tests are Ko consolidated.
182
(Note that one direction of the consolidation stress
(" ) in the CIUC test acts at 450 to the varve strucc
ture, while
hc in the CK UDSS test always acts
parallel to the varve structure.)
2.
For the stress system during shear, CIUC test samples
are sheared in triaxial compression, i.e., Aal> A
Aa 3.
In the
2
=
DSS test, the state of stress during
shear is close to that of the plane strain, i.e.,
A1 >AC 2 >Aa 3 , and principal planes are continuously
rotated during shear
Any difference between NSP obtained from DSS and CIUC tests
(besides the inherent problem of interpreting DSS test data)
could be the result of:
1.
The
influence of initial shear stress (Ko consoli-
dation);
2.
The effect of intermediate principal stress;
3.
The effect of rotation of principal planes; and
4.
The difference in the relative direction of
respect to the varve structure.
2 with
One of the A
2
di-
rections in the CIUC tests on inclined samples is
450 to the varve structure instead of parallel to
the varves as in the DSS test on vertical samples.
These factors could affect qf though their effects on the pore
pressure response and/or on the failure envelope.
The direct simple shear test has a problem with respect to
interpretation of the data.
However, the study in Section 5.3.4
showed that a NESE could be used to determine with reasonable
183
accuracy the state of stress at failure in DSS tests.
Thus,
value of NESE at qf and (al/a3)
for failure parallel to the
max
varves are needed (since DSS samples at OCR's between 1 and 4
do not reach
Thmax and (h/v)ma
x
at the same strain).
The
correct selection of a NESE requires that it -be a material property.
Based on the undrained test results from shearing across
the varves, this is only true for the NESE at (a1l/3)max or for
NESE at qf when it is identical to that at (a1/3)max
qf, when it is not
identical
ent upon the stress system.
to that at (a1 /a 3)maM
NESE at
is independ-
Therefore, the location of the NESE
at qf for a conventional DSS test is not known for sure.
interpreting thest data at Thmax
For
the NESE at qf for the appro-
priate range of Pf/avm from CIUC (=45)
tests is first used.
The sensitivity of qf to variations in the NESE is then studied.
For interpretation of DSS test data at (h/v)max'
the
uniqueness of NESE at (a1 /a 3 )max for failure parallel to the
varves should be established.
For this purpose, NESE at (a1/a3)ax
from CDDS tests (the envelope being determined assuming Th=Tff)
and from CIUC tests on inclined samples are compared.
It can
probably be assumed that, apart from the difference in the
drainage conditions, the stress system in a CDDS test is similar
to that in DSS tests, though the latter has a more uniform
stress distribution.
5.4.2
Results from CDDS and CIUC Tests
NSP from CIUC tests are summarized in Table 5.11.
The
test data from CDDS tests on samples from Northampton (Lacasse
et al., 1972) where failures take place in the clay layer are
184
presented in Table 5.12.
qf and pf were computed by assuming
that the horizontal plane in the direct shear test is the failure plane.
Consequently, qf is equal to Th/CosT and pf is
equal to avc + Th tanT.
The
value in Table 5.12 are those re-
ported in Lacasse et al. (1972).
For shearing parallel to the varves, the samples could fail
either in a clay or a silt layer.
However, for the range of
OCR's under study (1<OCR<20), observations of CIUC (0=45) test
samples at the end of the tests indicated that the failure in
these samples took place within a clay layer.
(Hence, the fail-
ures in the DSS samples are also assumed to be within the clay
layer.)
Consequently, the NESE from CDDS test samples which
have failure within the clay layer is compared to that from
CIUC tests.
Figure 5.30 shows the NESE at qf from CIUC tests and NESE
at (l/a3)max from CDDS and CIUC tests. Figure 5.31 shows NESE
at ( 1 /; 3)
of both NC and OC samples -from CDDS and CIUC
(8=45) tests at high stress.
(al/a3)max
from
It can be seen that NESE at
these tests are identical (Fig. 5.31).
Thus,
provided the assumed state of stress at failure in the CDDS test
is correct, NESE at (al/'3)max (Fig. 5.31) for failure within
the clay portion of both OC and NC samples is independent of:
1.
The stress system during consolidation;
2.
The stress system during shear;
3.
The drainage conditions;
4.
The magnitude of a
(samples from CDDS tests have
vm different among themselves and from CIUC (=45)
185
test samples).
Consequently, it is a material property.
Therefore, it will be
used to calculate the state of stress at (Th/ v)max in direct
simple shear tests.
For interpreting the DSS test data at Th,
both NC and
OC, the NESE obtained from CIUC (8=45) tests at OCR's equal to
1 and 4 will first be used.
This is line D in Fig. 5.30.
is assumed to be linear for Pf/ma
CIUC
(=0)
5.4.3
It
> 0.3, based on data from
tests at OCR = 1,2, and 4 (Fig. 5.8).
Results from CKoUDSS Tests and Their Interpretation
Three representative normalized stress-strain curves from
CK0 UDSS tests having OCR's of one, two, and four, are shown in
Fig. 5.32.
The determination of the state of stress at Thmax
of these samples is shown in Figs. 5.33, 5.34, and 5.35.
Mohr-
circles representing the state of stress at qf of inclined samples
from CIUC tests are also shown for comparison.
plot at (T /ov)max are shown in Fig. 5.36.
The corresponding
Table 5.13 presents
a summary of the results from Figs. 5.33 to 5.36.
These results
show that:
1.
At Thm
, as the OCR increases, Thmax/Tff increases
from 0.9 to unity at an OCR of four;
2.
At (h//v)max
Tff is equal to Th;
3.
The horizontal plane in the DSS test is not a
failure plane (this being true at both the qf
and (/ 3)ma x conditions);
4.
At Thmax, 8 increases with increasing OCR from
about 300 to 370 versus
186
450 if Th = qf; and
5.
At (h/av)max,
8 is approximately constant
(370 + 20) and is thus independent of OCR.
The angle 8 corresponds to the computed rotation of principal stresses.
Since the principal stresses are continuously ro-
tating during shear, the amount of rotation will depend upon
the magnitude of shear strain.
Hence, samples that fail at
larger strains should have a larger
.
This only applies to
the'hmax condition where in Yf increases with increasing OCR
(Fig. 5.27).
Consequently, $ follows the same general trend. At
(Th/av)max Y is quite variable, but also very large and thus
OCR has little effect on 8.
The sensitivity of qf values to the location of NESE is
evaluated and presented in Fig. 5.38.
Three locations of NESE
(arbitrarily selected X and Y NESE and that from assuming a horizontal failure plane) were used to compute qf/avc' Tff/avc and
8 for DSS samples at OCR's of one and two.
These computed
values are compared with those from using the CIUC (8=45) NESE,
also presented in Fig. 5.38.
As shown in Fig. 5.38, it is seen
that:
1.
The differences in qf are within 15%, the lowest
value of qf being computed from the lower limit
of NESE (i.e., that assuming a horizontal failure
plane).
qf values from NESE X and Y are within 10%
of those determined from the CIUC (=45)
2.
The values of
NESE.
vary with different assumed states
of stress in the DSS test.
187
The fact that'
c=O0)for
NC soil determined from this horizontal failure
plane assumption is unrealistically low (=15
and that the computed
°)
value is significantly
different from the X, Y, and CIUC (8=45) NESE
shows that the use of a horizontal failure plane
assumption is not suitable.
3.
The small variations in qf and 8 determined from using NESE X and NESE Y suggest that the use of NESE
from CIUC (e=45) test data for interpreting the
DSS test is suitable and much better than the use
of horizontal failure plane assumption.
In general, Th is not equal to either
ff or qf.
Ladd and
Edgers(1972) concluded that they would expect Tff<Thmax<qf.
Based on the results presented in Table 5.13, Thmax can be up
to 12% less than
Tff for samples which have OCR's of 1 and 2,
and is approximately equal to Tff at an OCR of four.
data suggest that the use of
hmax as the value of
undrained total stress analysis
will
These
ff in an
be conservative,
especially at low OCR values.
In summary, the use of NESE from CIUC (=45)
tests leads
to a reasonable interpretation of DSS data at the Thmax and
(Th/av) condition. Thmax values of samples at low OCR are
generally lower than
ff.
As the OCR increases, the ratio
Thmax/Tff approaches unity.
5.4.4
Comparison of Results from CIUC (8=45) and CKoUDSS Tests
Table 5.14 presents a summary of NSP obtained from CIUC
188
tests on inclined samples and CKoUDSS tests on vertical samples.
qf and Af from DSS tests were determined from using NESE at qf
from CIUC (=450) tests.
The results in Table 5.14 show that
shearing in a triaxial condition relative to a DSS condition:
1.
Increased qf/avc by about 10 and 20% at OCR's of
one to four, respectively;
2.
Increased Tff/Ovc at (/a3)max
by about 20+5% at
both OCR;
3.
Increased Af at both qf and (l/a3)ma
for nor-
mally consolidated soil, but decreased them at
OCR = 4.
As previously stated, the differences between NSP from CIUC
(8=45) and CKoUDSS tests could be the result of several factors.
Based on the comparison of the stress systems in these tests,
it can be hypothesized that such differences are due to the
rotation of principal planes, the intermediate principal stress,
and the initial shear stress effects, as well as inherent problems due to the interpretation of DSS test data.
In summary, the use of SHANSEP CIUC (80=45)test samples
for measuring the in situ qf for inclined loading will not be
conservative.
The analyses presented in this section also
show that the use of NESE at qf from CIUC (=45)
tests leads
to a reasonable interpretation of DSS test data, at least
better than using the assumption that the horizontal plane is
the failure plane.
Because it is a material property, NESE
at (al/a3)max for failure within the clay layer can be used
189
with confidence for interpreting DSS test data when the Th
and (Th/)max
conditions occur at the same shear strain (e.g.,
in heavily OC SHANSEP samples or OC RECOMPRESSION samples.
5.5
GENERAL DISCUSSION
One outcome from the studies of the effect of stress sys-
tem on NSP is the opportunity to investigate the possibility
of using a simplified testing method for evaluating anisotropic
Su values for total stress analyses.
The CIUC test is the
simplest test and could be performed on vertical, inclined, and
horizontal samples.
Therefore, data from CKoUC, CKoUPSC,
CKOUDSS, CKoUPSE, and CKoUE tests will be compared with those
obtained from CIUC tests on samples with different varve
in-
clinations.
Figure 5.39 shows plots of qf/avc versus OCR obtained from
these tests.
The data from CIUE tests are also included for
the sake of completeness.
It can be seen that the stress sys-
tem during consolidation and the intermediate principal stress
have practically no effect on qf(V).
Therefore, CIUC tests
on vertical samples could be used in practice for measuring
qf.(qf from
UC
tests are about 10% lower than those from
CKOUPSC tests.)
CIUC tests on inclined and horizontal samples, however, can
not be used in place of CKoUDSS and CKoUE or CKoUPSE tests,
since qf values from inclined and horizontal loading are significantly affected by the stress system and/or sample
190
orientation.
The combined effects cause qf(H) from CIUC (8=90)
tests to be up to 35% higher than those from CKoUE (=0)
and about 17% higher than those from CK0 UPSE tests.
ferences are rather large and, hence, CIUC (=90)
not
tests
The dif-
tests are
suitable to use in practice for measuring the horizontal
undrained shear strength.
For inclined loading, qf from CIUC
(0=45) tests can be 10 to 20% higher than that from DSS tests
(Fig. 5.39).
Such a difference is also somewhat too large.
Thus, in practice, CIUC tests on inclined samples overestimate
the strength for the measurement of Su with inclined failure
surface.
The above conclusions are based on the test results from
SHANSEP samples.
For RECOMPRESSION samples (see data in Chapter
6), CIUC tests yield a reasonable value of qf(V), but qf(H) is
too low at the in situ OCR of 4 because of disturbance due to
separation of the varves.
On the other hand, qf(4 50 ) from a
RECOMPRESSION CIUC test is practically equal to qf from SHANSEP
CK0 UDSS test, though this may be fortuitous.
5.6
SUMMARY AND CONCLUSIONS
The principal findings from the study of the effect of
stress system on NSP with different loading directions using
the SHANSEP method of consolidation are as follows:
1.
For vertical loading, the stress system during
consolidation and the intermediate principal
stress have little effect on qf(V).
191
However,
they affect Af,
f, Eu/avc,
*
(c=0) of NC samples,
and generally NESE at qf, though NESE at (a /a3)max
is not affected.
(See Section 5.2.5 for a de-
tailed summary of the effect of stress system during consolidation and Table 5.9 for the effect of a2 .)
2.
For horizontal loading, the stress system during
consolidation and the intermediate principal stress
affect qf, Af,
f, E/avc,
4 (c=0) of NC samples,
but have no effect on NESE at qf (same envelope as
that at (/a3)max)
(See Section 5.2.5 for a de-
tailed summary of the effect of the stress system
during consolidation and Table 5.10 for the effect
of a 2 .)
3.
Limited data show that the stress systems during
consolidation and shear affect NSP from inclined
loading more than fromvertical loading.
4.
The stress system during consolidation and the
intermediate principal stress have more effect on
the horizontal undrained shear strength behavior
than on the vertical undrained shear strength
behavior.
5.
For vertical and horizontal loading, the NC test
data from CKoU triaxial and plane strain tests,
show trends similar to those observed with homogeneous natural clays (reviewed in Chapter 2).
Due to inherent problems in interpreting DSS test
data, especially the assumption used for estimating
192
avc value, qf, Af, and
4
(c=O) for NC samples
from the special DSS tests on inclined samples
may be fortuitously identical to those from
CKoU triaxial tests.
6.
The use of strength data from CKoU triaxial tests
rather than plane strain tests for the evaluation
of aniostropic S u values for total stress stability
analyses in plane strain loading should be conservative.
Also, the use of Thmaxfrom
tests as Tff
7.
seems to be conservative
CIUC tests on samples cut at
DSS (=0)
as well.
=45 ° and 900 will
yield unconservative values of qf for undrained
total stress stability analyses, though qf(V)
from CIUC tests are slightly lower than those
from CKoUPSC tests.
8.
Reasonable
estimates
of qf(V) and qf(H) for PS
conditions can be obtained from CIUC and CIUE
tests on vertical samples.
The results in this chapter also provide insight into the
basic understanding of the anisotropic behavior of varved clay,
as presented below.
1.
There are four possible failure mechanisms in
varved clays, namely:
failure by developing
distinct failure planes across the varves, failure
by bulging, failure within the clay layer, and
failure within the silt layer.,
193
2.
Despite failure by bulging in CIUC (=0)
test
samples, their NESE at (al/c3 )max is identical to
that from CK UC test samples which develop disO
tinct failure planes across the varves.
This
indicates that the basic failure mechanism (i.e.,
failure across the varves) is the same even though
bulging occurs.
3.
The NESE at (a1 /a3) ax for failure across the varves
and for failure within the clay layer are shown
to be independent of:
a.
The stress system during consolidation;
b.
The stress system during shear;
c.
The magnitude of av
vm (see additional data
in Chapter
6);
d.
The drainage conditions (limited data); and
e.
The OCR of the sample.
Therefore, these envelopes are considered to be material properThe values of
ties.
and c/avm of these envelopes are given
below:
5.
c/avm
0
Type of Failure
Across the varves
30
Within the clay layer
22
0.017
0.027
The NESE at qf, if not equal to that at (l/a3)ma
x
(e.g., for vertical and inclined loading), is
dependent upon the stress system.
For vertical
for failure across the varves is much larger
*Note that
than that for failure within the clay layer.
194
loading, the NESE at qf, when pf/a
> 0.3 from
CKoUC tests,is lower than that from CIUC tests
(Fig. 5.8).
Results in Sections 5.3.4 and 5.4.3 suggest that appropriate values of NESE at qf and (a1 /a3 )max can be used to interpret the results of DSS tests, and that is much better than
using the pure shear assumption or assuming a horizontal failure
plane.
195
EXPERIMENTAL PROGRAM FOR INVESTIGATION THE EFFECT OF
STRESS SYSTE-MON SHANSEP NSP
Stress in kg/cm2
Variable
Type
of Tes0
Test
CIUC
0
EA
1,2,4
4.0
CK UC
0
E,A (2)
1,2,4
3.6 to 4.0
CIUE
0
E,A ( 2)
1,2,4
4.0
CK UE
0
E,A (2 )
1,2,4
3.6
CIUC
0
E
1,2,4
4.0
Intermediate
CIUE
90
E
1,2,4
4.0
Principal
CIUC
90
E
1,2,4
4.0
Stress
CIUE
0
E
1,2,4
4.0
CK 0 UC
0
A
1
CKoUPSC
0
N
1
45
A
1
CK 0 UE
0
E
1
CK UPSE
0
N
1
45
E
1
45
E
K o versus
Isotropic
OCR
OCR
vm ()
L)
_-
Consolidatio
CK UDSS(C)
CKoUDSS(E)
Effect of
Testing Method CIUC
Testing Mt:'d
for Shearing
Parallel to
theVarves
Notes:
Soil
S
Location
(1)
(2)
2.99 to 4.0
3.6 to 4.0
1,4,20
4.0
CK UDSS
o
0 A,N,E
1,2,4
3.0 to 4.0
CDDS
0
1,1.83,3.4
3.4 to 6.0
N
E = East'Windsor; A = Amherst; N = Northampton
Only soil from East Windsor that are tested
at indicated OCR.
Table 5.1
196
00
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4)
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4
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r
0
ii.
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I
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Table 5.4
0
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v)
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H
EU
(d
44
U
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a 0
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Table 5.5
__
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-
o
N
CN
U)
(0
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U)
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w
(n
4-.)
tqU)
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U4)
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HI I
U
X
U
P
4
0
H
1
Table 5.6
aU
201
LOAD INCREMENTS USED IN CKoUDSS(C) and CK UDSS(E) TESTS
ON NORMALLY CONSOLIDATED INCLINED SAMPLES
ac
2
= aV (cose + K
VC(1 - K
TC=
)
2
sine)
sinecose
Stress in kg/cm2
OCR
OCR
Ko
|
T
Tac
I:
ac
|
ahc
vc
23
1.0
0.104
0
0.104
0.104
11
1.0
0.211
0
0.211
0.211
4.6
1.0
0.520
0
0.520
0.520
2.36
0.91
0.971
0.049
1.016
0.927
1.61
0.80
1.342
0.146
1.488
1.196
1.133
0.74
1.838
0.281
2.119
1.557
1.0
0.68
2.097
0.399
2.496
1.698
1.0
0.68
2.505
0.479
2.984
2.026
1.0
0.68
3.020
0.575
3.595
2.445
1.0
0.68
3.395
0.639
4.034
2.756
For e = 450:
avc
=
sa c
+
c and
hc
ac
-
t
eac
Table 5.7
202
ESTIMATION OF Af FROM DSS TEST SAMPLE (8= 45)
NESE ASSUMPTION
USING
o-f
oc AC'
\\ -
at
Qo
rm
'i
_
_
_,
_
'A
&,,
_
4--
from e¢u
cation
Assuming
two
dimensonalo/
applied stress
systemn
= A %f-A af
, f - c
O.
I
At,
,,
'o, -
2
f
/(
=,)
/
{ a .- 2 AUf: c
=
- 1C
203
2
E4.S.9
Eq. .//
Eq. 5?/O
*(zIC)
2
A:/f
Z
III
Eq.5/2
Eq.5./3
io7 DSs test
TABLE 5.8
_____ L
______ ___
4.4
x
N
x
0
x
a
0
m
.4
E
__
m
103
Z
H
X
X
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a)
4
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o
0
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0
Cd
a
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Cd
I 11
I
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tr
0
. rF-1
z
0
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>
H
t-
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> 0
C--
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*,
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P4UU)
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~~ ~ ~ ~ ~
a)C
a 3
a)
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o U0
U.r-f--
z0
m0k
U
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204
A
Table 5.9
4-4
0
0
to
C
ro
0
Cd
44
44
4.)
0
to
0
01
4-4
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H
tdX
m
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W m
t
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al
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54
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cU)
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to
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205
0
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Table 5.10
dO
(A)
w
00
o
r
o
a\
O
0
·J
o
(N
o0
L)
Ln
0
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LA
r4
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0
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10
11
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U
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r
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00
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mo
0
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0
L
LA
0
un
mr
L
LA
or
(N
n
Ln
m
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0;
0
H
o
rl
LA
H
44 0
1
.
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10.1 1
b
v
01
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O
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n
r-
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U)
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C0
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.*
H
r4
r-i
(n
o 0) 0
(N
(U)
112
E4
3
0
.C*
H
e
C4
0
(N
0
C,)
a)
IC
En
'r
W
UA
H
Table 5.11
I
(N
E-4
uI
U
H
F1
IIC)(
(N
N
(N
I
I
I
U
H
N
U
H
w
C)
H
r
206
RESULTS FROM DRAINED DIRECT SHEAR TEST ON NORTHAMPTON
VARVED CLAY FOR FAILURE WITHIN CLAY LAYER
(LADD AND WISSA, 1970 AND LACASSE ET AL., 19-72)
Test N.
Test
__
aT
ah
TSF
TSF
(1)
vc
(2)
qf
(3)
q(1
ciqvc
_vmvm
NCD-5
3.65
2.00
0.380
19.50 0,410
0.220
0.620
NCD-6
3.4
1.01
0.495
19.50 0.525
0.160
0.350
NCD-1
4.0(4)
4.0
0.385
21.00 0.410
0.410
1.150
NCD-2
5.0(4)
5.0
0.390
21.30 0.420
0.420
1.150
5.14
0.450
24.1° 0.490
0.490
1.220
6.00( 4 ) 6.00
0.340
18. 8 0.360
0.3.60
1.120
NCD-314
NDC-4
Notes:
(1)
From Lacasse
et al.
(1972) for overconsolidated
samples:
= tan 1
h
for normally consoli-
ovc dated samples.
(2) qf/avc =(Th/vc) (1/cos)
(3)
Pf/vm
= Th/a vm tan
+ °vc
vc
vm
(4) Laboratory maximum past pressure (a m(L))
Table 5.12
207
0
cc
S4
r
HU
.cl
0)
U
H)
4-)
o
o
z
o
r
oN
0
LH
0
H
H
0
Lr)
A
c
Ln
0O
o
00o
O
m
H
o
Cd
-
CNI
Hf
O
LO
x
rH
.H
010
0
U)
o4
a)
10
n
EU)
4)
E-
41 U
p4 IO>
U)
cn
u
4tlc>
a:
m
o
0
0
0
0
c7
(v-)
r-
D
H
N
o
U
Clu
0
N-
i4
c
01I
¢· UltT3
LA
N-
x
Cd
-i
tlD
0
0
00
o
0
0
L;
LA
d
0
c;
oN
o
HI.
0
o
Lo
LO
H
0
o
En
U)
4a)
4-)
z
0
*,{
m
{.o
4-4
0
t
..........
E-
U
co
X-
Q
a
0U)
44
Cd
U)
WH
W U)
.
z rd
.a)
U)
r.,
-r1
1-~4
)
a)
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m
P·ltC
Ln
O
O
M
Iti
rd
.f-e -H0
·-
kir
4 1,C
I >3
U)
Wt
r-
-
--
H
r-q
o
v
o
cn
0
(N
C4
_
0
M
V)
a)
4)
U
0
H
N
R
O
Z
Table 5.13
208
0H
I
.-
x
C)
I t,
1tHm
N)
'o
H
Im
-
l
cd
H
N
HI
H:1
I
o
0
m
4t V)
I....
O
N
.
o
0
x
H
ib
5
U)
W>0
Igu
W
'1bt
Co
H
.
E
4
o
rl
44-4
1U
Cd
0
-4r-i
'u
UL
m
O
0
o
00
o
O
O
U)
.
O
o
H
0
N
H
o
8W
\D
4-4
zH
Ul
44
U
E4
I-r E-1
IEb
XI
B
_cn
W
3
0
bU
'INI
U)
rI
H
Q) 4U.4J
4
a
u]-'
0
U u
O
U
0
0 Cl)
z
j
-
O
U
N
H
H
N
I
U
U)
U
U)
H
Q
N
N
4.)
U)
E-I
z Q)
U
H U
U
H
v
N
H
0
H
Table 5.14
209
0
2
4
6
8
/0
AV
V
/2
_
/0
--
Dat rom 3 C/; N =-56.9 .0'%
/4 _
Daoto from
/6
CKo U tests; wN '=
.3
O.7 %
_
/8 _
20 _
I4 _
'21L
O.I
-~~~~~~~~~
i
O.Z
I
I
0.4
I
B
0.6
I
I
0.
I
B
/.0
I
I
2.0
o
&'vc, kg/em2
VOLUME CHANGE DATA FROM I clu AND CKoU
01I SAMPLES
TESTS ON HORIZONTAL VARVE
(8: =
FROM EAST WINDSOR
210
FIGURE
5.1
.0o
L
a~~~~~~~~~~~~~~~~J
a
rl
TEST K -/I, AMHERS7 MA.
/Ist
Unload from 8 2 7tf
to OCR=32
18 -
0 Reload to 20.5 Tss
O 2nd
Unload from
20.S TSF
wN = 60.4%,El.
/07 ft
/
/
/
/
/
/
/
//
/
1/
12
/
/
/,
/
/
_
/
/
/
/
/
1.0 _
/
/
S/
/
/
_^ _
,
Relatio.s*i
0/
-·i
-irecomn endea
p
/
~f
/0
0.
0.1
I
I
A
I
3
I
4
I
I
I
I
Ii
5 6 7 8 9 /0
20
OCR
Ko VERSUS
OCR FOR CONNECTICUT VALLEY
VARVED
CLAYS
211
FIGURE 5.2
0.8
OGR /
0.6
_'
4.
- - -a
s
0
)
v
*v c
C
TTs
NO
/
0
I
0
2
I
4
I
SYM. TYPEOF ar yc
TES
oY | rEs
T rT g/co .
A/C-I -/
0
C1IC
AKC-/
*
CK
0L/C2.99 187
I
6
I
i,
I
8
/o
AX/AL S TRA/N,
4.0
I
.90
I
12
%
h
NORMALI ZED STRESS-STRAIN DATA FRO M CIUC
AND CKoUC TESTS ON NORMALLY CONSOLIDATED AMHERST VARVED CLAY
212
FIGURE 5.3
/'
r.
1.6
pr
5vc 0.8
vc LAB.
' OF
-ST
7c /.0 4.o
0.4+
'oc I.0 40
,I
0,
&;m
K9/a
1
I
2
I
I
I
4
I
I
6
I
8
AXIAL STRAIN,
I
I
,
I
I
/0
I
I
I
12
qa%
A
-
AXIAL STRAIN, Ea %
NORMALIZED STRESS-STRAIN
CURVES
CIUC AND
CKoUC TESTS ON OVERCONSOLIDATED (OCR.=4)
WINDSOR VARVED CLAY
213
EAST
FIGURE 5.4
14
Af
OCR
(a)
VERTICAL
LOADING
OCR
(b) HORIZONTAL LOADING
Af VERSUS OCR FROM CIUC, CKoUC, CUE, AND CKoUE
TESTS ON CONNECTICUT VALLEY VARVED CLAYS FIGURE 5.5
214
o
/s
o10
%
,5
0
/
2
(ao)
VRTICAL
'5
4
s
6
7
OCR
LOADIN6
I ()/"a1
ore identical
t fI"-c)mzox and (a&
~~-,/ o
Ef
/0
3
1--
E%I
0
o
C/uE
*
CKo IvE
I
i·
3
2
I·
4
I·
J
·I
6
7
OCR
(b ) HORIZONTAL LOADING
Et
VERSUS
OCR FROM CIUC, CKoUC, CIUE,
AND CKoUE
TESTS ON CONNECTICUT VALLEY VARVED CLAYS
215
FIGURE 5.6
0
I-
ROZ
ZO
,)
O -J
I U)z
O
I
_4
Lz
b
4.
0z
I,-
L.I
-o
0
w
z
aO --j
-.I
0
x 0
2o
3:
0
zw
C:3
Z
,b'l4
216
Ib" 'l
FIGURE 5.7
C/)
IL O
w
cn >
> >
b
W .
-Jw
en
ib
W
__z
oWO
.
w OW. M
Nw
w
C
Z
N
zTo
P)
a,
'
e
217
l
0
FIGURE 5.8
4000
-
I
TYPE OF OCR NO. OF
S YM
3000 _
0
2000
I
l w
TES T
TES TS
clUc
/
2
4
2
I
2
2
Ck'ouc
-
0
Uo
CkoUC
/
/000 _
700 _
500
E,,
9f
_
400
.00
_
O\~
200 _
)
OCR from CIU tests
(
/00 _
£)
OCR from C/4ot tests
J,
0
an
"
I
I
I
a6
0.4
CI
0.8
/.0
a 9
Eu
VERSUS
qf
ON SAMPLES
Aq
Z-qfFROM CIUC AND CKoUC TESTS
L~q 1
FROM CONNECTICUT VALLEY
218
FIGURE 5.9
/3O I
f,
SYM. Type of Tejt
(9o
0
0
CKUoC
1.20
0
1/0
El
CKo C
I-,
CIUE
0
c(-[E
0o
sO
sO
sO
/00
/1b
90
80
/f
C/~/I/ ond C/iL
-
70
/
I
Wf
60o
r
-
. .
m-
.
0
40
-
1
J01 Ik
1
3
OCR
?I
Eu
vc
AT
"i~c,
C= 0.2%
CIUE,
I
4
I
s
I
6
I
I
7 8
VERSUS OCR FROM CIUC,
AND CKoUE TESTS
ON
SAMPLES FROM EAST WINDSOR
2 19
FIGURE 5.10
I---- -~
0.2z
3
0
S5T
~~~~~~~~~~~~~7~
\
~~~~~NO.
-0.4
vc
m
4.0
/.67
CKc/VE3.6
50
CIR7
*_
EKE-i
C
7.TSM. K9 l,,'
o
-0.2
_~~~~~~~~~~~~~~~~~
~,.,.=..
.
-0.6
TYPE OF
sYM
I
I
2
-CI)
I
I
I
4
I
I
6
.
--
.
I
8
AXIAL STRAIN,
/4
/0
a %
.;
3.0
2$
zo -
*
CK UE
A
O.
-
0
_.
I
a.
I
I
4
I
i
6
I
J
I
8
AXIAL STRAIN,
(a
i
/0
i
I
12,
I
%
CURVES FROM CIUE
NORMALIZED STRESS-STRAIN
AND CKoUE TESTS ON NORMALLY CONSOLIDATED EAST
WINDSOR VARVED CLAY
FIGURE 5.11
220
/.
f.
0.
-o
Ir.O.;
Li
AXIAL
STRAINV, 6a %
¥1
h
&V
AXIAL STRAIN, Ea %
NORMALIZED STRESS-STRAIN CURVE' i
CKoUE TESTS ON OVERCONSOLIDATED
CONNECTICUT VALLEY VARVED CL
221
FROM
CIUI E
(OCR = 4.0)
WYS
FIGURE 5.12
scLIU
SYtlBOL TYPE OF TEST OCR
3000
0
CIK LE
Eu-= Secant Modulus
1.
IO
2000 -
O
U
C/LIE
CKo0E
2.0
C/lUE
2.0
4.0
4.0
Ceo
C
k, UE
E
/000
_
(-<
Eu
Su
OCR from CIU tests
ts
5$00
400
300
200
_
/00
_
2 tests
(Eoch from Amhert
and East Windsor)
A^
I
0--C
)
0.2
·-
I
0.4
0.6
I
0.w
10
FROM
VERSUS
Af FROMCIUE AND CKoUE TESTS
ON SAMPLES
FROM EAST WINDSOR
222
FIGURE
5.13
One of A c
directions
acts perpendicu Jar
to
the varves
-
Cl/, (9 = 90) a
I
tJN
cr
a
(a c,
,T=O
-0o)
ERTICAL LOADIN& (0=0)
(0)
One of tAc
directions
j acts perpendiclar
to the vorvcs
A
aa0
7
Cl
CIUc
)E A o0, =a S2 > a3
{e =o)
(b)
(4J
(e -90)
= 0)
> a( C = o
)= o
(&C25. = aSr.Y)
O
HORIZONTAL LOADIG& (3=90)
STATES
OF STRESS
FOR CIUC (=O0),
CIUE (=O0),
CIUE (- 90), AND CIUC( :90) TEST SAMPLES
223
FIGURE 5.14
0.4
0.3
_
7EE
-w
________
C
./------0----
0.2
I-
l
//=_
0.1
0
CIJE e =90)i
----
I
0
I
iII
Cl-C(eA=
:0);
i
2
6
4
I
Test Ala. E/E-3-!
TestNo. EIC-I-I
I
0
I
/0
I
2
4
.3
A
.__
-
--
2
I
I
O
I
II
I
4
,I
I
6
8
I
/0
/2
NORMALIZED STRESS-STRAIN CURVES FOR
NORMALLY CONSOLIDATED CIUE (= 90) AND CIUC
(6- O) TEST SAMPLES (VERTICAL
224
LOADING)
FIGURE 5.15(a)
0.8
T1
_ _
____
___
I
-
0.6
v-
0.4
CUE (e =9o): TeJst No. EIE-3-3
---
0.2
CIUC (e=
/
0
l
2
0
4
l
l
l
est No. EIC-1-3
):
l
l
/2
/0
8
6
£0, %
----
TE
4.0
TC
3.'
J.0
_-0-
_-
/7
Ir.
2S
/
2.0
OC
_
///
/.5
,. _
ltl
I
I
I
I
4
I
I
I
6
I
I
I
/0
I
/2
NORMALIZED STRESS-STRAIN CURVES FOR OVERCONSOLIDATED CIUE (8= 90) AND CIUC (8= O) TEST
SAMPLES
(OCR = 4.0)
225
FIGURE 5.15 (b)
0.4
0.3 _
TC
i
TE
a2
z
.
VC
F/ic e 90); Test No: EC-3-I
--
0/
---- e-)
O
n
I
4To
I
2
CI
u1
t (e- o
I
I
I
4
6
I
Test No: EIE-I-/
I
8
/0
12
£q, %
I
lz
h
o
%
Ea, %
NORMALIZED STRESS-STRAIN
CURVES FOR NORMALLY
CONSOL_IDATED CIUE (e=O) AND CIUC (e = 90)
TEST SAMPLES (HORIZONTAL LOADING)
FIGURE 5.16 (a)
226
"cri
£q, %
4.0I
I
.0TE
'0--
3.0
Rz
cr
"
2.0
PA
lId
L.
p
I
I
I
I
4
6
8
I
I
I
/0
/2
So, %
NORMALIZED SIrRESS-STRAIN CURVES FOR
OVERCONSOLI DATED CIUE (8=;) AND CIUC (8= 90)
TEST SAMPLES (OCR = 4.0)
227
FIGURE 5.16(b)
W
-J
p-
Z O
o 0
w
D Z
O Z
0
L
0
W W
Oa-
w
Z
n-Q
fW
WO
w m
> ZoU
W O
N W
r
on?
Z Pp-
w
1l
228
FIGURE 5.17
z
O
O3
-Jon
II
1a
<
-Jl
IU
D
LI-O
LL
T
a-
C)
N
O_ z
Io--
ul
)3"
Q
0
229
FIGURE 5.18
0
Cf,
Dz
O-
Oh
LL w
ALC
(n
0
nL
wz
.J <
Ib~
rl6ft l
zo::
>
0
Cl
:
Ow
WI-
Zt
z
0N>
0
wn
L
hlU
(3
_o
4s
IZ)
6CS~~~~~~
:i
!5
z,:
230
FIGURE 5.19
O6
-
!
_ -
0.4
av- h
- U/SC ,
CiyYPSC, Test Ala. A'P-I
0.2 _
---
a
c)
cKO Uc ;
I
0o.s
0
-
-
o. APC-
4
irt.
- ---
TestNo.
i
at q
0
ct
9
0 at
AKC-I
I
I
I
I
I
I
/.s
2
4
6
H
/0
I
/.5
2
4
6
8
/0
2
I
I
I
I
II
.2 /4
/2
/4
Ct.1o
A
F
NORMALIZED
STRESS-STRAIN CURVES OF NORMALLY
CONSOLIDATED SAMPLES
FROM CKoUC AND CKoUPSC
231
FIGURE
TESTS
5.20
Y
lrvg
,0
J.3
~~~~~~~-
I-,~
-
2.5
hi
abv
2.0
-_--Cl cdYE; Tt
1.s
-
No.EKE-I
CKo.LPSE; Trt No. NPE-2
-.
CKKoPSE;
7t
7No. NPE-I
Io
vrt .
0
I
I
2
I
I
4
I
I
6
I
£a, °/
I
8
I
I
/0
I
I
A*
I
14
A
NORMALIZED STRESS-STRAIN CURVES OF NORMALLY
CONSOLIDATED SAMPLES FROM CKoUE AND CKoUPSE TESTS
232
FIGURE 5.21
W
-3
J
w
WZ
<
Q:
w
Z Zz
-Jo
a0
or
IJ-
()9
o-z
(n
w
Uf)
aJ
0
0
W
?0
WLL
LAO
IL
w cr
Ow
WI
c>o
0 <O
Ct
I
I
I
I~ I iI:
,:Ihk1
233
FIGURE 5.22
4L
TIrc
APPLIED
COMPARISON OF
SHEAR STRESS LEVEL,
vc
%
FROM CKoU PLANE STRAIN,
TRIAXIAL, ANI) SPECIAL DSS TESTS
234
FIGURE 5.23
___
9-f/er-"
I
.
_
OCR
CUi PsC
wP
CKoVE
/Uc
CK
028
0.22 0.2S 0.21
0.47 0.41 0.44 0.36'
0.76 0.67 0.72 0.65'
2
I.I
_
_
""N
_
~ ~ ~ ~ ~
TYPE OF
TEST
SYM
e
Ck lUE
0
0
0
0
--
cl-E ( =o)
C/I/g 990) 0
90
A
0
A
90
C/I/C((
-- ".
9'o)
O
90
0.9
oc
"I'll\
11
E 1E
* 1
,, t3,
\ .~
~~-
0.7
Verticai
o Horizonto
P/
H/
to
.s
Looding
~
Wjd
a3
I
I
1
I
2
3
4
I
I
s 6 7
I
I
I
.9 /0
I
/II
I
I
20
JO 40
_
_
OCR
ESTIMATION
OF
qf
avc OF
OVERCONSOLIDATED
PLANE STRAIN HORIZONTAL AND VERTICAL LOADING
SAMPLES FROM CONNECTICUT VALLEY VARVED CLAYS
235
FIGURE 5.24
2E
0
w
O-
ZO
1_Y)
V
0C
cn en
@0
of_
c
Lio
Z
_
w.OU z
I-O
Xr z
Q0
I
I
_o
o
z o
w
u
Q~
n c
,
V)
c]
nO
O
cn
W
c
z
236
FIGURE 5.25
w
w
m
Uf)
U)
0
m
Ir
-i
cr:
w
.0
LiL
.x
a
I-E
16-110
w
U
w
z
iI
237
FIGURE 5.26
0.3
NAlo.ESDS-I
r I Oonr. zor
Bn(L)
Tes
0.2
z~~~~~~~~~~~~~
t~>
0./
v-.c
I
I
4
I
I
I
I
I
/2.
8
2-,
I
16
I
I
I
20
2 4
20
5.2T
2
%
-o./
No.£$D5-2
__,_
0.6
TESTS
CKOUDSS( CKDUDSS(E)
~0.2~FIGURE
4
8
3
16
NORMALIZED STRESS-STRAIN CURVES OF
NORMALLY CONSOLIDATED INCLINED SAMPLES FROM
CKoUDSS(C) AND CKOUDSS(E) TESTS
FIGURE 5.27
238
rr
O.
0
--
Wnz
O
ow
F-
I
LL Iai
0 w
z 3
0 L
E/za
L
0
L
0 w
NI1J
239
FIGURE 5.28
z
ZJ
w:
{
> C
.J
w
O i<
LL a
OZQ
,I::{
A
N
^
C*U
FIGURE 5.29
FIGURE 5.29
a
z
z
0
oz
n
Ld
LL
O
U
oa
EZ
C)
U
O n
<
[
zo
241
FIGURE 5.30
o
'o
w
C)
lJ
U
0
I
ri4
o'b
I
4bI,$
00 *
ci
3
11
o
L
w
z
*
vs
242
FIGURE 5.31
0
-
-
-
0,
5!
OCR = 4, Test No: EDS-3
---
_---at
o
---.. OCR = 2, Test No: ADS-4
OCR - , Tt No: ADS- I
0.6
P'hn1,
rvC
o , IX
Zh
-
-
0.4
cb1
-
-
-
--
-
_
--
·-
·
-?
-
-
·
10
·--
*-.-
/
0.2
.i
00
V
I
II
I
I
2
II
I
I
6
4
I
I
II
8
II
II
/o
I!
I
I
14
/Z
a, %
0.4
AL4
~
vc
0
-
-0.4
.
'~~~~~~~~~c
l'~~~~~~~~~~~~~~~
_
_. -
I
_
_>,~-. _
I_
., ,
_r
-0.+
l
0.c
_---I1.
I
0.3
''
r
=,
l
I .I1-
0.2
0.i
0
... .___....-'. - _--.-
,//
_///
I/~
0
4
I
2
I
I
4
I
I
6
I
I
8
7hZ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~
I
I
/0
i
I
/2
I
t~~~~
TYPICAL NORMALIZED STRESS-STRAIN CURVES FROM
CKoUDSS TESTS ON NC AND OC CONNECTICUT VALLEY
VARVED CLAYS
243
FIGURE 5.32
PS
(A
OZ
E
0
o
V
v!
d
bilb
C)
E
0Z
L IE
a
0E
0 cr
w
0
o
E
Z v
:
IO
11
O
J
I
__
244
FIGURE 5.33
,<
x
LL a
w a
0C
Qa
0C->.'-
~
"ot,b
o
DWW
I.
o0
oe- A
I .1
u)
owFZ0Z 0
FIGURE 5.34
C%
bn
0
E
L:JW
~ > 0.1
0 o> .
O
n~ c
0
E
vsLJ
z q
O
E
C)
Cab
0C)~ <
£
,.ilJII
246
FIGURE 5.35
Iw
n
0
o0
XL
K
Co
0r
z
w
w
CL
(f
N ,$
247
FIGURE 5.36
/.o
2.0
3.0
4.0
So
OCR
FAILURE
CKoUDSS TESTS
STRAINS
AT
hmax
AND(v)max
ON CONNECTICUT VALLEY
FROM
VARVED
CLAYS
248
FIGURE 5.37
O
Q
>
0
zZ
in
0
UZ
n
w
z
t
O
o
--
O
n0
0
O
Z0
o
,s
0
249
FIGURE 5.38
NN.
~q<% ~ *j
°
IIts~~~~~~~~~~~~~L
~
ig
\Z
fr1 OO
1I-.
CNob ,
e"Fz
o~~~v
L
z
~u
-
vS
v
0c
~rb-c
u0
I-
c6J
O
C3
U
`
O
IJ-
uQ)
ao) \X
,
-
'
e At-
\I
II
Ie
.'
'
.!
(nc
tZ
\us
oR04
t
51e
N
o
t
w
0
''0
K)(-)
4
3
'q.
-!
IC~~·~
WOg
LL
'1
0
to.~'~~~
z:
uJ
u W
u
W
j)11~ ~LL
,., It
I a
Z I
0
I
>oC
2s50
FIGURE 5.39
6.
6.1
CONSIDERATION OF SAMPLE DISTURBANCE
INTRODUCTION AND BACKGROUND
6.1.1
Purpose
Sample disturbance in homogeneous clays has been studied
fairly extensively (e.g., Ladd and Lambe, 1963; Davis and Poulos, 1967; Bozozuk, 1970; and Milovic, 1970), but disturbance
in varved clay has received little attention.
The special
problems with varved clays include formation of transverse
cracks located at the interface between clay and silt portions
due to sampling technique (Kenney and Chan, 1972) and separation
of varves during trimming of samples (Saxena et al.,
1974).
In this chapter, methods for minimizing disturbance in
varved clay are studied for the purpose of providing:
parison of NSP (qf/vm, Af, Eu/ vc,
testing methods:
(1) com-
4 and c/avm) from these
the UUC test, the RECOMPRESSION method, and
the SHANSEP method; and (2) recommended methods for measuring
the in situ strength-deformation properties of varved clay.
6.1.2
Types and Causes of Sample Disturbance
There are two basic types of sample disturbance in varved
clay, namely:
1.
Disturbance due to the removal of the in situ
stresses which caused a change from anisotropic
stresses to an equivalent isotropic stress, aps; and
2.
Disturbance due to sampling, handling, trimming,
and setting up a test, all of which may cause
251
permanent damage to the soil structure, and with
varved clays, cause separations or transverse
cracks located at the interfaces between silt and
clay layers.
The separations are most likely to
occur between the varves.
The formation of cracks or separations are the result of sampling techniques (Kenney and Chan, 1972), and trimming, especially
with inclined and horizontal samples (Saxena et al., 1974).
Kenney and Chan (1972) concluded that the formation of transverse cracks due to sampling of varved clay samples was primarily the result of the large friction force created between the
coarser-grained soil (silt portion) and the sampling tube.
The
formations of cracks or separations resulting from trimming
samples are difficult to avoid, even when a thin wire is used.
Such an effect will be very pronounced in horizontal samples,
slightly less pronounced in inclined samples, and will be negligible in vertical samples.
6.1.3
General Effect of Sample Disturbance
During sampling, the change in effective stress due to
the removal of the in situ stress is unavoidable.
Even for the
"perfect sample," defined as a sample which undergoes only an
undrained release of the in situ stresses to end up with an isotropic effective stress equal to aps (Ladd and Lambe, 1963),
the strength and undrained modulus of the soil are affected.
As a result of structural damage (including both soil structure
damage and separation
of varves), and expansion (swelling) of
252
the sample, during sampling, extruding, trimming, and setting
up of a test, the isotropic effective stress in a disturbed
sample is decreased from aps to as (the effective stress of
the sample prior to a test).
Such a reduction
in effective
stress in turn causes an additional decrease in undrained shear
strength (qf) and Young's modulus (E). In addition, planes of
weakness at the interfaces between clay and silt portions, which
are created by the transverse cracks or separations (i.e.,
part of structural damage) may cause further decreases in undrained shear strength and modulus, especially when
lf acts
parallel to the varves and when shearing is parallel to the
varves.
This would
produce a directional dependence of
sample disturbance.
6.1.4
Methods for Minimizing Sample Disturbance
Methods for minimizing sample disturbance to be studied
are the UUC test, the RECOMPRESSION method, and the SHANSEP
method.
Since the cracks or separations can be closed by pres-
sure, the UUC test (where a cell pressure is used to confine
the sample) is considered
as a method
for minimizing
sample
disturbance compared to an unconfined compression test.
In the
RECOMPRESSION method, the sample is Kx consolidated to the in
situ vertical and horizontal stresses (Bjerrum, 1973).
It
therefore minimizes disturbance by reapplying the field effective stresses,
i.e.,
vc =
vo and ahc =ho
Because
of the
resultant decrease in moisture content and any prior structural
damage (soil structure damage and formations of cracks or
253
separations located at the interfaces between clay and silt
portions), the RECOMPRESSION method may yield a sample having
a different soil structure and hence
different soil proper-
ties compared to the in situ soil.
Research reported by Ladd and Foott (1974) and Ladd et
al. (1977) suggest that, if a soil exhibits normalized behavior
when av > am , K
consolidated SHANSEP samples (especially NC
samples) should yield reasonable qf/avc values compared to those
of the in situ soil.
The basis of the SHANSEP method proposed
by Ladd and Foott (1974) as a means for minimizing sample disturbance effects on strength-deformation properties measured
in the laboratory
is as follows.
As indicated
in Fig. 6.1(a),
the laboratory curve typically approaches the in situ virgin
= v(L) > (1.5-2) avm)
vm
vm
vc
Thus, CK U test samples consolidated to Point D should have a
compression curve at Point C (
soil structure similar to the in situ normally consolidated
clay and hence yield reasonable qf/avc values if a soil exhibits
normalized behavior.
For overconsolidated samples (A and B in
Fig. 6.1(a)), the samples are first consolidated into the virgin compression range to obtain a normally consolidated clay
(i.e., at Point D), and unloaded prior to shear (shown as
Point A and B in Fig. 6.1(a)).
These OC samples then should
yield reasonable qf/avc values at the required OCR's.
It is seen in Fig. 6.1(a) that the SHANSEP and RECOMPRESSION
samples at the same OCR have different magnitudes of
compare
vc of Sample A and A' in Fig. 6.1(a).
254
vc' e.g.,
Therefore, the
comparison of the measured soil parameters of these samples
must be in normalized form.
For the purpose of making exten-
sive comparison of NSP from SHANSEP and RECOMPRESSION methods
of consolidation, RECOMPRESSION samples considered in this
study will also include those which are consolidated above
but below
.vm NSP (qf/avm Aft Eu/avc,
and C/vm)
vo
of a
RECOMPRESSION sample such as that indicated by B' (Fig. 6.1(a))
will therefore be compared with those of a SHANSEP sample which
has the same OCR.
qf/Cvm versus
Figure 6.1(b) shows a hypothetical plot of
vc/vm(L)
from SHANSEP samples A, B, C, and D.
Since samples C and D are normally consolidated, their qf/vm
are identical
if the normalized behavior concept applies.
Figure 6.1(c) shows hypothetical plots of qf/cvm versus avc/avm
from RECOMPRESSION and UUC test samples and qf/%vm(L) versus
avca vm(L) from SHANSEP samples.
Because of the difference in
the capabilities for minimizing disturbance effects and the
use of different methods of consolidation, qf/vm
of the RE-
COMPRESSION sample at the in situ OCR can be higher or lower
than that of the SHANSEP sample.
preshear effective stress as
As a result of the smaller
qf/avm of the UUC test sample
will be lower than those from the SHANSEP and RECOMPRESSION
samples at the in situ OCR.
This chapter compares data obtained from the SHANSEP and
RECOMPRESSION methods in an attempt to ascertain their relative
advantages and disadvantages.
Results from both isotropic and
K O consolidated tests are included since CIUC tests on samples
cut at different inclinations have been used in practice.
255
6.1.5
Experimental Program
Table 6.1 presents the experimental program used in the
study.
Since sample disturbance can affect the undrained shear
strength behavior differently for various loading directions,
the investigation is made for three loading directions:
cal, inclined (450), and horizontal.
and E/
NSP (qf/avm' Af,
verti, C/avm
vc) of K o consolidated OC SHANSEP samples, RECOMPRESSION
OC samples (Ko and isotropically consolidated), and UUC test
samples are compared.
High quality block samples (in situ
OCR ~ 4) from East Windsor were primarily used for the UUC and
RECOMPRESSION test programs, supplemented by some data from
5 in. diameter tube samples from Amherst and Northampton.
Re-
sults from SHANSEP samples (presented in Chapters 4 and 5) are
used for comparisons shown in Table 6.1.
It should be noted that K o consolidated RECOMPRESSION East
Windsor samples at an OCR of four have K o equal to unity.
For
simplicity, these samples were, therefore, isotropically consolidated in the laboratory prior to shear.
Data on overconsolidated samples with 8 = 0, 450 and 900
are used to study the method of consolidation in CIUC tests.
6.2
EFFECT OF METHODS FOR MINIMIZING DISTURBANCE ON UNDRAINED
STRENGTH BEHAVIOR
It is emphasized that the soil samples used in the study,
especially those from block samples, are of much better quality
than normally encountered in practice.
256
Thus, results of this
study may not apply to the poorer quality "undisturbed" samples
obtained from 3 in. diameter tube samples or obtained from 5
in. diameter tube samples with poor sampling techniques.
6.2.1
Behavior from Vertical Loading
The upper half of Table 6.2 summarizes test data from
CKoUC tests on SHANSEP and RECOMPRESSION samples.
Figures 6.2
through 6.5 present plots of test data from the UUC test, and
from the CU tests at OCR's of two and four.
These data show
the following results.
1.
At OCR's of two and four, CKoUC tests on SHANSEP
and RECOMPRESSION samples at the same OCR yield
essentially the same qf/Ovc within experimental
accuracy.
However, RECOMPRESSION samples have
a higher NESE at qf, and a higher pore pressure
parameter A than the SHANSEP samples (Figs. 6.2 to
6.5).
Therefore, the behavior in terms of effec-
tive stress is really quite different.
Further-
more, the stress-strain behavior is also somewhat
different
(Figs. 6.2 and 6.3).
samples have a lower
RECOMPRESSION
f and higher EU/vc
high strain, though E/vc
at
at low strain of both
SHANSEP and RECOMPRESSION samples are nearly identical.
Based on a comparison of
- E -u behavior,
it is seen that RECOMPRESSION samples exhibit a
more sensitive and brittle behavior.
The fact
that RECOMPRESSION samples have larger normalized
257
pore pressures, and also a higher NESE at qf
explains why qf/%vc from SHANSEP and RECOMPRESSION
samples turned out to be practically identical.
2.
qf/avm from the UUC (0=0) test is also essentially
the same as that from CKoUC tests on SHANSEP and
RECOMPRESSION samples at OCR of four, though the
UUC test sample has the largest
Eu/avm.
f and the smallest
The agreement in qf values indicates that
disturbance due to structural damage and removal
of in situ stress is relatively small, though the
modulus is significantly reduced.
(Note, however,
that disturbance due to separation of varves, if
there is any, shown from TC tests will be small
since shearing would tend to close the separations
or cracks).
However, as will be shown in Section
6.2.2,
thedisturbance of this block sample is mainly
due to separation of varves.
This has a pronounced
effect with horizontal and inclined samples, but
little effect
on
vertical samples.
It is emphasized that the good agreement among UUC and the
SHANSEP and RECOMPRESSION CKoUC test qf/avm values is based on
data obtained with very good quality block samples from East
Windsor.
With 5 in. diameter tube samples (such as those from
Amherst and Northampton), where the effect of disturbance is
greater, especially from samples at great depths, qf(V)/avc from
UUC tests are generally much lower than those from SHANSEP
samples, Fig. 6.6.
These data also show that, because of
258
variable sample disturbance, UUC qf values from tube samples
are very scattered.
qf(V)/avc from a RECOMPRESSION CK UC test on 5 in. diameter
Amherst tube sample is compared with the values from SHANSEP
and RECOMPRESSION block samples in Fig. 6.7.
These data show
that qf(V)/Fvc of an Amherst sample with OCR of 1.6 is only
slightly higher than that of the SHANSEP samples.
noted that qf(V)/vc
from a CIUC (=0)
It should be
test on a RECOMPRESSION
sample (OCR=1.7) from East Windsor (Long, 1975) is approximately
equal to that of the Ko consolidated RECOMPRESSION samples.
0
These data indicate an insignificant effect on stress system
during consolidation on qf(V)/Zvc obtained from the RECOMPRESSION
method.
(The same trend is also observed with SHANSEP samples;
see Chapter 5.)
At an OCR of four, it is also interesting to compare data
from CIUC tests on SHANSEP and RECOMPRESSION samples.
(Note
that at OCR=4, the "CKoU" RECOMPRESSION sample was isotropically
consolidated.)
qf/avc and NESE at (a1/a 3)
from these tests
max
are roughly the same (see bottom part of Table 6.3), though Af
and NESE at qf of the RECOMPRESSION sample are higher.
In conclusion, for vertical loading, the SHANSEP method
of consolidation yields reasonable values of qf and NESE at
(al/O3)ma , even though the magnitude of a
is significantly
larger (3.6 versus 2.4 kg/cm 2) than the in situ value.
However,
it yields a lower Eu/avc and quite different behavior in terms
of effective stress compared to the RECOMPRESSION method.
259
6.2.2
Behavior from Horizontal Loading
Test data in Figs. 6.4, 6.5 and 6.8 to 6.11 and the sum-
mary in Table 6.2 show the following results.
(1) At OCR's of two and four, the RECOMPRESSION CKoUE
test samples have a lower qf/avc and
f, and a higher Af and
Eu/avc than SHANSEP samples at the same OCR (Figs. 6.8 and 6.9),
though the NESE at qf (also same envelope as at (/a
these samples are identical.
3
)
of
max
Therefore, the decrease in qf/avc
from SHANSEP to RECOMPRESSION results from larger normalized
pore pressures in the RECOMPRESSION samples.
of
The lower values
f, and the higher values of Eu/avc, (Fig. 6.10), Af and
Au/a
from RECOMPRESSION samples show that RECOMPRESSION samples
exhibit a more brittle and sensitive behavior and thus have a
more sensitive soil structure.
These results and results in
Section 6.2.1, therefore, lead to the conclusion that the SHANSEP
method of consolidation causes the soil to have a less sensitive
and less brittle structure
than the in situ soil since the high
quality undisturbed RECOMPRESSION block sample should have a
soil structure close to that of the in situ soil.
(2) At the in situ OCR (OCR=4), qf(H)/avm from the UUC
(6=90) test is about 80% of the RECOMPRESSION CIUC (=90)
and 62% of the SHANSEP CKoUE value (Fig. 6.8).
value
The lower UUC
qf and the large decrease in strength after failure is thought
to be mainly caused by disturbance due to varve separation,
even though the sample developed a failure plane across the
varves.
However, it is not known whether the separation occurred
during sampling or mostly as the result of trimming.
260
(Note
that the effect on vertical samples would be much smaller since
shearing would tend to close the cracks.)
(3) Further evidence that disturbance (presumably due to
varve separation)is most important with
=90 sample is partly
provided (since.CIUC.SHANSEP and RECOMPRESSION samples have dif.ferentsoil structure) by
comparing CIUC (=90)
RECOMPRESSION and SHANSEP samples.
test data on
The strength of the RECOM-
PRESSION sample is only about two-thirds of that from the SHANSEP sample at OCR's of 2 and 4 (Table 6.3).
The moduli from
RECOMPRESSION tests are also lower (Figs. 6.11 to 6.13).
This
is in contrast to the CKoUE test where the RECOMPRESSION strengths
were 21% less at OCR = 4 and 6% less at OCR = 2.
(Note, however,
that the CKoU RECOMPRESSION samples had a much higher undrained
modulus and larger pore pressure, i.e., they behave in more sensitive and brittle fashion).
This suggests that the RECOMPRES-
SION method of consolidation only partially minimizes
the ef-
fect of varve separation with horizontal samples.
The fact that the strength in a CIUC (=90)
test is lower
than that in a CKOUE (0=0) test, both performed on RECOMPRESSION
samples at OCR= 4, despite the larger magnitude of a2 applied
in the extension test also suggests that disturbance (presumably
due to varve separation) is most important with 0=90 samples.
In summary, the two most important findings appear to be:
1.
Disturbance due to varve separation is most important with horizontally trimmed samples and can be
only partially overcome by the RECOMPRESSION method.
2.
The SHANSEP method causes the soil to have less
261
sensitive and brittle behavior.
Finally, varve separation may have affected the RECOMPRESSION CKoUE test at the in situ OCR of four, perhaps causing a
lowering of the strength, since the difference in the RECOMPRESSION and SHANSEP strengths at the OCR is much greater than an
OCR of 2.
6.2.3
Behavior from Inclined Loading
Test data from the CIUC (=45)
and UUC (=
45 ) tests on
overconsolidated samples (OCR=4) from East Windsor, presented
in Fig. 6.14 and Table 6.3 (lower part), show the following results.
qf/avm from the UUC (=45)
test is only 64% of that from
the SHANSEP CIUC (0=45) test and 75% of that from the RECOMPRESSION CIUC (=45)
test.
The decrease in qf/vc
from SHANSEP to
RECOMPRESSION samples is due to larger normalized pore pressures
while the NESE at (al/a3)max of these samples is probably identical (Fig. 6.15).
The fact that UUC tests and the RECOMPRES-
SION CIUC (8=45) test samples failed along a plane of weakness
(located at the interface between the clay and silt portions),
while the SHANSEP sample failed within the clay layer partly
explains their lower qf/vm.
However, because of the use of a
high isotropic consolidation stress,the SHANSEP strength is
probably too high compared to the in situ soil.
Comparison of test data (Figs. 6.16 and 6.17) from SHANSEP
CKoUDSS tests and from one RECOMPRESSION sample also indicates
a trend similar to that from CIUC (=45)
RECOMPRESSION (vc
Vc =
tests:
h/avc of the
)
vo
Vo Amherst sample is lower than that of
the SHANSEP sample (Fig. 6.16).
262
The value of G/i
at Y = 1%
from RECOMPRESSION test is also somewhat lower (Fig. 6.17).
6.3
DISCUSSION
Disturbance in Varved Clay
As stated in Section 6.1.1, disturbance in varved clay is
due to the removal of the in situ stresses, soil structure damage, and varve separation.
For block samples, unless
alf
acts perpendicular to the varves, disturbance due to varve separation, which probably occurs mostly during trimming, is an important factor.
As shown from UUC tests, it is most pronounced
in inclined and horizontal samples, but perhaps small in vertical samples.
dependent.
Thus, disturbance in varved clay is directional
The following experimental evidence illustrates the
existence of disturbance due to varve separations in block samples.
(1) As shown in Fig. 6.18, for an OCR = 4, the increase in
qf from a UUC ( = 45) test to a RECOMPRESSION CIUC (=
45) test
is about the same as that from a UUC (0= 90) test to a RECOMPRESSION CIUC ( = 90) test, while qf(V) from a UUC (=0) test is
essentially identical to that from a RECOMPRESSION CKoUC (=0)
test at OCR = 4.
(2) The qf in a CIUC ( = 90) test is lower than that in
a CKoUE (6=0) test, both performed on RECOMPRESSION samples at
OCR = 4, despite the larger magnitude of a 2 applied in the extension test.
(See Section 6.2.2.)
(3) The fact that a failure plane developed at the interface between varves in the UUC (0=45) test instead of within the
263
clay layer as observed from SHANSEP samples indicates the existence of a plane of weakness at this location (Section 6.2.3).
With 5 in. diameter tube samples, the effect of disturbance
is greater (see test data from Amherst and Northampton, Section
6.2.1) than for block samples, probably because of increased soil
structure damage.
With tube samples, qf(V) from UUC (=
0)
tests are generally much lower than would be obtained from RECOMPRESSION CKoUC tests.
Thus, based on datafrm
vertical load-
ing, one expects that disturbance in tube samples will cause
greater reductions in qf in all directions, compared to block
samples.
SHANSEP Versus RECOMPRESSION as a Method for Minimizing
Sample Disturbance
The increase in qf from UUC tests to RECOMPRESSION (vc =OO)
CIUC tests on inclined and horizontal samples indicates that
disturbance due to varve separations can be minimized by reconsolidating the samples. SHANSEP method of consolidation appears
to be more effective than the RECOMPRESSION method in minimizing this type of disturbance, especially with inclined and horizontal samples, based on the following data from CIUC (0=45)
tests on block samples.
Despite reconsolidation to avo, the
failure plane which developed in the RECOMPRESSION CIUC (=45)
test occurred at the interface between silt and clay portions,
whereas it occurred within a clay layer for the SHANSEP sample.
In addition, the difference in qf(H)/avc between the RECOMPRESSION and SHANSEP CKoUE (6=0) tests increases with increasing
OCR, i.e., 0.375/0.355 = 1.06 at OCR = 2, and 0.64'5/0.535=1.21
264
at OCR = 4 (see Table 6.2).
Since both SHANSEP samples had
been consolidated beyond the in situ avm, this suggests, but
does not prove, that SHANSEP may minimize disturbance due to
varve separation better than the RECOMPRESSION method.
However,
it is not certain if the 20% difference in strength at OCR= 4
is mainly due to the influence of the varve separations.
Though the SHANSEP method of consolidation can effectively
minimize disturbance due to separation of varves, it also has
a drawback in that it causes significant changes in soil structure, especially when isotropic consolidation is used.
effect of these changes
below.
.
in
The
the soil structure is discussed
-
With CKOU tests, except for possible effects of varve separations (when
alf does not act perpendicular to the varves), it
is thought that OC RECOMPRESSION data on block samples give the
best indication of the in situ behavior of varved clay, since
the structure in the RECOMPRESSION sample should be close to
that of the in situ soil.
For both vertical and horizontal
loading, OC RECOMPRESSION samples compared to OC SHANSEP samples
generally have significantly higher moduli, pore pressures and
NESE at qf (when it is not equal to that at (ol/a3)max
in the vertical loading).
e.g.,
These general trends show that OC
RECOMPRESSION samples act in a more brittle and sensitive manner.
This in turn strongly suggests that the SHANSEP method of consolidation causes the test samples to have a less brittle and
sensitive behavior than that of the in situ soil.
265
Thus, though
data (mostly from NC soil) over a small MPPR range on Connecticut Valley varved clays show normalized behavior with'respect
to qf, Af,
and c/avm when consolidated above avm' based on
OC East Windsor block sample test data one concludes that stressstrain-pore pressure parameters from OC ( 2 < OCR < 4) SHANSEP
samples do not represent the true in situ soil behavior.
How-
ever, SHANSEP samples apparently yield reasonable estimates of
qf and NESE at ( 1 /a
3 )
max
for both vertical and horizontal
loading.
With CIU tests, SHANSEP yields much larger qf values than the
RECOMPRESSION
method, except for vertical loading on 8 = 0 samples,
as shown in Fig. 6.18.
This may result both from the influence
of disturbance due to varve separation which is important with
inclined and horizontal samples, and from the fact that high
isotropic consolidation may produce a less parallel orientation
of clay particles perpendicular to the in situ vertical direction.
6.4
SUMMARY AND CONCLUSIONS
Block sample test data presented in this chapter show that
the UUC test is not suitable for measuring anisotropic strength
and modulus.
separation
Despite using block samples, disturbance due to
of varves causes the value of qf from inclined and
horizontal samples to be too low compared to more reliable data
obtained from CIUC RECOMPRESSION tests.
For typical tube samples,
increased soil structure damage will probably cause a further
reduction in qf and undrained modulus in all directions.
266
OC CKoUC block sample test data show that the SHANSEP
method of consolidation causes a change in soil structure to
produce samples that are less brittle and sensitive than the
in situ clay.
On the other hand, the RECOMPRESSION method bet-
ter preserves the in situ structure.
However, SHANSEP may yield
more reasonable estimates of qf in cases where disturbance due
to separation, of varves affects the results.
With CKoU tube samples, which are likely to be much more
disturbed than block samples, little data exist for comparison
of the SHANSEP and RECOMPRESSION methods of consolidation.
But
in the view of the author, SHANSEP method is probably the better
approach if the in situ soil has a low OCR, since RECOMPRESSION
yields the maximum discrepancy between the in situ and laboratory compression curves in the vicinity of the maximum past
pressure (e.g., see Fig. 6.1(a)), while use-of the RECOMPRESSION
method in tube sample is probably better if the clay is heavily
overconsolidated and there is little disturbance due to separation of varve.
In any case, SHANSEP has the advantage of mini-
mizing adverse effects of varve separation whereas RECOMPRESSION
has the advantage of better preserving the in situ structure of
OC samples.
For both block and tube samples, CIUC tests on samples with
different inclinations consolidated by either the SHANSEP or
RECOMPRESSION methods are not recommended for obtaining anisotropic undrained parameters.
In general, the pros and cons of the SHANSEP and RECOMPRESSION methods of testing for use with good quality "undisturbed"
267
Connecticut varved clay samples are as follows:
(A) RECOMPRESSION Method
Advantages
(1) OC RECOMPRESSION samples (OCR > 2) probably yield
reasonable estimates of in situ Su/avc, though we can not really
be sure whether it is too high or too low and the error probably
depends upon the sample quality and direction of loading.
(2) RECOMPRESSION samples probably better represent
the in situ pre-failure a-e-u behavior of OC samples, especially
for modes of failure involving rotation of principal stresses.
Disadvantages
(1)
If soil parameters are not related to stress his-
tory, i.e., OCR, each new job will need CKoU testing to obtain
data versus depth.
(2) For NC and perhaps slightly OC deposits (OCR < 2)
the RECOMPRESSION method of consolidation (avc
vc
overestimate the in situ qf.
=
a)
vo
could
(B) 'SHANSEP Method
Advantages:
(1) For normally consolidated and slightly OC deposits,
where good quality undisturbed samples are especially difficult
to obtain, the SHANSEP method probably yields better estimates
of the in situ qf and -e-u
behavior.
(2) NSP from many types of SHANSEP strength tests
have been shown to be the same for several locations within the
Connecticut Valley (see Chapter 4).
(3)' Thus, since soil parameters have been developed
268
as a function of OCR, the testing program can concentrate on consolidation tests for determining the stress history, plots of
NSP versus OCR already having been established.
(4) NSP versus OCR data provide information to predict changes in in situ soil parameters due to changes in consolidation stress, such as occur during consolidation under embankments.
Disadvantages
(1)
Because of changes in soil structure, the SHANSEP
method of consolidation may not provide very reliable a-e-u data
for OC samples.
(2) Because NSP are expressed as a function of OCR, the
SHANSEP method is not suitable to use at a site where the stress
history is very scattered.
In such cases, tests such as the
field vane are recommended.
In summary, the RECOMPRESSION method is recommended if the
deposit is heavily overconsolidated (OCR) 4) and good quality
samples can be obtained, though it would still be desirable to
correlate the results with values of in situ OCR and with those
from SHANSEP method.
In such cases, when predictions of de-
formations and pore pressures are especially important, the
RECOMPRESSION procedure will probably yield much better parameters.
For cases where stability is of prime'concern, and if
the stress history can be established with reasonable certainty,
then the SHANSEP method has a distinct advantage since extensive
Su/avo versus OCR data exist for the Connecticut Valley varved
269
clay.
Finally, for normally consolidated and slightly OC
deposits, use of the SHANSEP method is recommended.
270
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o
O
NH
H:n
o
o
O
H.
O
N
0
O
LA
,
o
N
LA
N
N
0
*
O
H
00
O
0
'
O'
O
0
S
ro
U)
0
l-
0H
U)
MU
U)
q0I
U1
H
U
o
0
IP
0
0)
M:
N
18
13 u
I
II
rM
n
E
Po
N
40
tc
Ei
i
N
I
M
UZ
O
r-4
I
en'
M
I
I
EQ)
H
w
I
ti:
w
n
I
I
)
I
I
N
I
')
I
U
I
H
A
U
H
.
U
U
H
H
W
-. -
l
273
i
,-I
I
i
,-i
I
L)
rH
U
H
W
H
r
C)
4.)
0
Table 6.3
5'JJ
In
Mrve
c7ryc
0vc
-
In
1%
°-vm (L)
ot A-
aC
-4
I),
Somp/es of A
/9
/0 9
Fron
6r. ()
C
nd A' have the same OCR
'(fao)
O
No
6:y"
10
SHANSEP Method
D
BC
A
/c/
O
(b)
)
bc/
/-/r
Al
or
qf/
from
i/
SHJANSEP
A
)Raneof S/.
m ()
rom RECOMPRESS10N method
Ul test
/.
cr,,,
/,7 51tv
or
Cr..
v
1
(c)
rvon
, a)
GENERAL EFFECTS OF DISTURBANCE ON UUC
TEST, RECOMPRESSION, AND SHANSEP SAMPLES
274
FIGURE 6.1
- --
0.35
--.
C--,'
0.30
m
0.2
//
0.20
RECOMPRESS/ON ({rvo) Test Alo. ERIC-/)
-- SHANSEP (/est No. EKC-5)
.UUL/, (Test No. EUC-/)
.- *
/
/
IN
I/
61"
0./0
OCR
0aos
I
I
2
co
I
]~
I
I
,
6
4-
I
I
I
I
...
/0
8
AXIAL STRAIN, eo
I
I
I
%
6.u r
so
h 40
&v
3.0
/
2.0
i
4
2
.-- ¢_
2
6
8
/0
/2
/4
AXIAL STRA IN, 64 %
r=
r
0.7
A
r
0.5
0.3
-
nI
L-
I
I
_
_
I
,
i
I
I
6
8
/o
/
/4
AXIAL STRAIN, a %
EFFECT OF METHOD OF CONSOLIDATION ON VERTICAL UNBEHAVIOR FROM TRIAXIAL COMPRESDRAINED -g-STRENGTH
SION TESTS ON EAST WINDSOR VARVED CLAY OCR = 4
FIGURE 6.2
275
0
.2.
4-
/2
/.o
_
0.6
PAAAP.'J('AI
/rF
r\_yc
1-111,
b
IN
0.4 It
=
vyC
--
4'
-,
.-
v/m7
(Test No. ERKC-/)
SHANSEP (Test No. AKC-2)
--0.2
0
I
I
I
2
0
I
I
4
I
I
I
I
/0
8
6
AXIAL STRAIN,
I
I
/2
I
.,.I
,%
AXIAL STRAIN, a, %
0.7
A
I-
0.6
s -r~~~~~~~-
0.5
\
\I
0
//
n
2
Er
I
I
I
I
I
I
4
6
I
I
I
8
AXIAL STRAIN,
I
17
I
I
I
I
I
/2
/0
i
I
/4.
(0,%/
EFFECT OF METHOD OF CONSOLIDATION ON VERTICAL
BEHAVIOR FROM TRIAXIAL COMUNDRAINED o--STRENGTH
PRESSION TESTS ON EAST WINDSOR VARVED CLAY AT
OCR=2
276
FIGURE 6.3
ILJ^
a.
C)
zz
0
OT
LLcrn
-Jo
C-
IW
C>1
L '
Z
LLW
> w
0oz IE
I
CLL
F
z
kf
IN
277
FIGURE 6.4
a
>-
z I
I I)
I-
0
Oo
o
00W
O
LL
I
o .JO
.I
-r 2
e o
a M
O0
o
Z <
6
-
278
_
I
FIGURE 6.5
cr-·rr
C /3
a*~~~~~~~
U
(*~~~~~~~~~~~~~~~~
I
1
TYPE O
LOCA TiONA
ast
TEST
CKAN
065 Windsor
SYM.
P)
0
(SH4NVSEP)
£ost
Windsor C,oVC (RECOM)
Amrnherst
UU(V)
North -
amp ton
Ut v)
W/indsor
A
±
UL(v)
- East
0.S
0 -
qf (v)
ofvc
Data from
or
V~O
Amherst
A/
A
0.35
_0_
S = qf
·
_,
0.,
z(
U
_ -
/ I
'
I. 0
a .=I
I
2.0
OCR.
4.0
I
S0
I
I
6.0 70 8.0
OCR
qf(V)
COMPARISON
OF
FROM
Vc
CKoUC TEST
SHANSEP
SAMPLES AND UU TEST SAMPLES FROM AMHERST
AND NORTHAMPTON
279
FIGURE 6.6
0.7
0.6
C Test
0 UA-CK
tesults from
Long, /97S
9q(V)
'vc
K
solidated
SHANSEP s
E/
ornpIes
From F 9
0.4
.
4., /4
Notes: () R= PE(COMPRES510N
S SH
-2) E =Ea.rst Windsor
0.3
Ar n herst
A
I ,A
r-
--
A
_
TEST
A
CKOUC
0
0
l
0
I
l
E
E
-7-Uc
v7T7c
E
.
.
3
2
4-
TYPE OF
SAMPLE
t----------
E
0
.
MOC
.
0
0
-
SOIL
LOCATION
TYPE OF
!._
.
__
=
e
SYM.
0.2
4 NSEP
.
R
R
Block
R
Bloc
S
Block
I
5
'Bloq/;:
6
7 8
OCR
COMPARISON
qf (V)
OF =VERSUS
CIUC, AND UU(v) TESTS
SAMPLES
OCR FROM CKoUC,
ON SHANSEP AND RECOMPRESSION
OF AMHERST AND EAST WINDSOR VARVED CLAYS
280
FIGURE 6.7
C-(
(T
7
AXIAL STRAIN, E,, %
6.0
..'ERIC-3-2
7est No. ERIE-/-/
No. Kf-3
--- Test
s0
4.0
4
.-2.
0
/O
8
6
AXIAL STRAIN,
/2
/4
I
1
E, %
/o
0.8
A
06
-
0.+
n.2
-0
-"
-,
I
..
-
I
2
'
'
4
I
6
I
I
88
I
/
I
I
AXIAL STRAIN, a,%
EFFECT OF METHOD OF CONSOLIDATION ON HORIZONTAL UNDRAINED cr-E-STRENGTH BEHAVIOR FROM TRIAXIAL TESTS
FIGURE 6.8
OF EAST WINDSOR VARVED CLAY AT OCR = 4
281
,,
(r,y,
5.3
,
Test No. ERIC-3-/
o..*
Test No. ERKE-I
AKE - I
Test ANo.
45
0
0S··
2..
I
-S
*
*
*
,
-
......
-.
. -.
-r
--
.
-.
_-
/._
I
I
I
I
2
I
4
I
I
I
6
I
I
I
I
I
I
t2
8
O
0.7
A
6*
O..
_-
j,
.
I -- _-
-
. -. .-
-
.z
C
0
I
I
2
II
II
I
I
6
8
I
I
/0
I
I
I
/2
AX/AL STRAIN, £o, %,
EFFECT OF METHOD OF CONSOLIDATION
BEHAVIOR
0--STRENGTH
ON HORIZONTAL UNDRAINED
FROM TRIAXIAL TESTS OF CONNECTICUT VARVED CLAYS
AT OCR = 2.O
FIGURE 6.9
282
/4
Ei/.
at
29S
2
/
3
4
S
6
7
OCR
Eu
I Aq
6vc AT Aqf FROM CKoUC AND CKoUE TESTS
ON SHANSEP AND RECOMPRESSION SAMPLES FROM
CONNECTICUT VALLEY
283
FIGURE 6.10
0.8
A_
s
;
SYM
0.7 -
i
t
-
e°
90
90
|
SOIL
LOCA TION
TYPE OF
TEST
CItc
CILC
CK0UE
TYPE OF
SA/MPLE
Av0C
I
E. Windsor
S
blocK
E. Windsor
E. Windsor
R
block
R
blocK
bloc
/U (H)
90
E. Windsor
No/es: () S- SHANSEP
(2) R
A'ECOMP'EStY
0.6
From CKEE
/
SHA NSEP
Samples
qf rH)
6
vc
(IfC.
A
4.14)
Os
4,0
0. 4
4;
--2 T.ets
xu,,
/
.3rests
A
firor,
P .Con,,
0.3
02
V.j
I
/
COMPARISON
I
3
OCR
2
q (H)
OF vc VERSUS
I
4
OCR FROM
I
I
I
S
6
7
8
CKoUE, CIUC
(8= 90) AND UU (8= 90) TESTS ON SHANSEP AND RECOMPRESSION SAMPLES
OF AMHERST AND EAST
VARVED CLAYS
284
WINDSOR
FIGURE 6.11
a-v
OCvtit
AXIAL STRAIN,
J.
,,%
t·
,
30
.o.. ' ,
._
00,
2.5
000
2.0
1.5
0
I
I
=~~~~~~~~~~~~~~~~~
I
I
!
I
I
__
L
I
B
4
6
II I
8
I
/0
12
AXIAL STRAIN, 'o,%
0.6 Aw_
,_
0.45
A
0.2S
e, e
_...0
_L_
-
2
_
tw_ - r -
4
-I - -
6
-I
I
6
-
I
/0
I
I
/a
AXIAL SRAIN> , %
NORMALIZED
STRESS-STRAIN
CURVES
FROM CI-C (
90) TESTS ON SHANSEP AND RECOMPRESSION SAMPLES((OCR=4) FROM EAST WINDSOR
285
FIGURE6.12
0.6
0.5
r 0.4
I
I
[
I
I
IT
I
MVO.
c
-iI
0.3
-
0.2
I
0.I
-
a
I
aw
I
I
A.
2
r
I
A
I
I
a
I
I
I
I
1
I
1
w vw
-
iAd
SHANSEP
---
Recornpression
3.0
.
i0g
__0--
/7
2.o
__
-.
/I
/.0 k
,
V
A.
4O
-
T
^
IQP
#r
/o
/2
O
L
0.7
O./
0o
2.
4
6
8
AXIAL STRAIN, Ea, %
NORMALIZED
(8 = 90)
STRESS-STRAIN
CURVES FROM CIUC
TESTS ON SHANSEP AND RECOMPRESSION
SAMPLES
(OCR 2) FROM EAST WINDSOR
286
FIGURE 6.13
/,
_" - -0---
0.3
No.EC-2-3
/fTest
Test /o.ERIC-2-/
T-1
/
slppge
,est No.El/C-3-'-indcAted
W., ¢c pc",vor
O.2
J.
1;-weancsKss
_/
S,)
UN4
.' --.-CIC
0.1
---CVC SANSEP e =-s)
i
n
I
6
C
rv = r,,(e =+s)
I
2
)
I
I
4
I
I
I
8
6
AXIAL STRAIN,
I
I
/0
/2.
/4
"2
J4
.
%
A
AX/AL STRAIN, %
I
0.1 &
x
Ir
0.S
A
0.4
0.3
-ilo
w.-
0.2
0.
I
I
I
2
C
I
4%
I
I
I
6
I
I
8
AXIAL STRAIN, 4, %
I
/0
NORMALIZED STRESS-STRAIN
C IUC (8 = 45)
SAMPLES
TEST
(OCR
=
4)
CURVES FROM
ON SHANSEP AND RECOMPRESSION
FROM EAST WINDSOR
287
FIGURE 6.14
0Z
n)
w
a-
0
w
zao
w
z
I
en
O w
-J
I-C
en
It
C-)
w
w
z
288
FIGURE 6.15
0
O
0.
qf
rve
0.
or
-rbag
Irr,,
0.
0.
0
OCR
COMPARISON OF
2uc
VERSUIS OCR FROM CIUC
(8 45), CKoUDSS, AND UU (45) TESTS ON SHANSEP
AND RECOMPRESSION SAMPLES FR( )M AMHERST AND EXkST
WINDSOR
289
FIGURE
6.16
so
-
-
l
-
l
SYM.
MOC
0
SHA NSEP
TO
40
RECOMPRESSION
Amherst Data
i
SNANSEP samples had the some
9mf"L) = /.76 X a:m
@6
tc =/ dy
c =/ day
0c
2'zo
0
Pcotc
- a
7nd
crvrr= / 7 kg/cm
t = 3 days
c
7
I/
days
2
OCR
3
I
I
5-
6
I
7
G
COMPARISON OF avc AT
=
1.0% FROM CKoUDSS
TESTS ON SHANSEP AND RECOMPRESSION
290
SAMPLES
FIGURE 6.17
0.32
5YM.1 TYPE OF TEST
o U/C test
0.28 _ V
RECOMPRESSION/ C/C
tests
SHANSEP C/UC tests
V
RECOMPRESS/ON Ce/
iC
-_
_
/ C6U£IE
tests
SHANSEP CKUC / CUIE tests
0.24
0.20
m
Ff
/
SHA NSEP
,t_
°'vm
·II
/
0.16
I'
I
OCR =4
m
',
0./2
sts
X
,- _ !
L-JrK = At '
0.08
0.o0
IUt/ic tests
_
,_
,,
C)
I.
/0
.
I
20
I
I
I.
I
I
I
0
O
I
l
30SO60
I
I
70
I
80
90
O
qf
COMPARISON OF °c
FROM CIUC AND CKoU
TESTS ON SHANSEP AND RECOMPRESSION SAMPLES
291
FIGURE 6.18
7.
INTERPRETATION OF ANISOTROPIC BEHAVIOR OF CONNECTICUT
VALLEY VARVED CLAY
7.1
INTRODUCTION
As presented in Chapter 2, the term anisotropy means that
the soil properties (e.g., qf. A,
, and C/Ovm) vary with
loading direction (defined by angle
; see Fig. 2.1) as they
are measured in the laboratory using controlled tests with
different directions of alf.
It is convenient to divide aniso-
tropy as measured in the laboratory into three components, as
follows:
1.
Inherent Anisotropy--For varved clay, the inherent
anisotropy includes:
a.
Structural Anisotropy:
This component is a
result of the mode of deposition and subsequent K o consolidation that produces a preferred orientation of particles.
K o consoli-
dation causes plately shaped particles to
have a tendency to align themselves at right
angles to the direction of the major principal stress.
b.
Material Anisotropy:
This component of aniso-
tropy results only from the layered nature of
varved clay, the individual layers being considered as isotropic material.
The two components of inherent anisotropy would
292
be expected to lead to changes in effective stress
envelope, pore pressure parameter, and modulus
with direction of loading.
2.
Stress System Induced Anisotropy--This component
of anisotropy results from the fact that different increments of shear stress are required to produce failure as the major principal stress varies
between the vertical and horizontal directionswhen
K o is not equal to unity.
In essence, stress sys-
tem induced anisotropy results in changes in effective stress as al varies in direction during
shear.
This component occurs even if there is
no change in properties due to inherent anisotropy.
3.
Combined Anisotropy--Thisterm (used in place of
"in situ" anisotropy because the anisotropic behavior presented in this Chapter is measured via
the SHANSEP method of consolidation) includes the
total effects of inherent and stress system induced
anisotropy.
The study presented in this chapter hopes to provide an
understanding of the fundamentals controlling anisotropic undrained shear strength behavior.
The study concentrates on
the anisotropy in qf interpreted in terms of effective stress,
the anisotropy in NESE (both at qf and (al/3)max), and the
anisotropy in pore pressure parameter A (considered at a particular strain).
Some consideration is given to the anisotropy
in undrained modulus.
293
7.2
METHOD OF INVESTIGATION
To provide this understanding, an attempt was made:
1.
To separate the material anisotropy component from
the structural and stress system induced anisotropy
components so that the anisotropy in qf, A, NESE
and E
primarily due to the layered nature of varved
clay (i.e., the material anisotropy component)
could be studied; and
2.
To illustrate the general effects of structural
anisotropy.
Combined anisotropy is studied in terms of effective stress
from results of CKoUPSC, DSS and PSE test data, i.e., for plane
strain test conditions.
Although more extensive data exist
for shear in TC and TE, interpretation of the results are complicated by effects due to changes in a2 which are difficult
to separate out.
Details of the types of tests used for measuring material
and combined anisotropy are presented below.
Measurement of Material Anisotropy
Conceptually, for plane strain loading, CIU PSC tests on
samples with different inclinations that have been consolidated
to extremely high stresses (
>>>vm
)
to minimize the in situ
structural anisotropy should be used for this purpose.
How-
ever, since this was not possible, SHANSEP CIUC tests on samples
at different
inclinations
(=0,
45
and 90 ) at OCR's of 1, 4,
and 20 were used to study material anisotropy (i.e., Option 2
294
in Table 2.1).
These CIUC tests have some drawbacks.
The magnitude of
a2 is controlled in tests of this type, but the direction of
02
relative to the varve orientation changes with sample
inclination.
This complicates interpretation of the results.
Also, the values of
c used were only 4.0 kg/cm 2
xipared
tovm=
2.4 kg/cm 2 for the East Windsor varved clay which is used in
the study),which is not high enough to eliminate the effects
Therefore, the word
of structural anisotropy.
material
is
put in quotes, i.e., "material" anisotropy is used to emphasize the fact that the results are only an approximate measurement of the true material anisotropy.
Measurement of Combined Anisotropy
Since the anisotropic fabric is oriented in the same direction as the in situ soil and since shear starts from a K o condition, SHANSEP CKoUPSC, CKoUDSS and CKoUPSE tests were performed for this purpose.
Note, however, that results for OC
CKRUPSC, PSE tests had to be estimated based on trends measured in CKoUC and CKoUE tests.
Effect of Structural Anisotropy Component
An attempt was made to approximately determine the structural anisotropy component.
Conceptually, this component of
anisotropy should be shown by comparing qf values from isotropically consolidated PSC tests run on samples with different inclinations with those from CKoU
0 PSC, DSS and PSE tests
295
at OCR = 4 where Ko = 1.0 so that shearing of the CIU and CKoU
samples starts from the same conditions.
Since this type of
plane strain testing was not possible, SHANSEP CIUC test data
were used for this purpose.
7.3
INTERPRETATION OF "MATERIAL" ANISOTROPIC BEHAVIOR
Test data from CIUC (8=0, 45, 90) tests are plotted in
Figs. 7.1 through 7.6 and summarized in Table 7.1.
Interpre-
tation of the anisotropic behavior arising primarily from the
layered nature of varved clay for both NC and OC samples
is
as follows:
(1) The "material" anisotropy in qf/avc of varved clay
results from anisotropy with respect to the normalized effective stress-strength parameters (Fig. 7.4), and to the pore
pressure parameter Af (Fig. 7.3 (c)), and as a result of this,
to the excess pore pressure at failure (Fig. 7.4).
seen that the variation in qf with
It is
is within 20% (Fig. 7.3(a))
and the ratio qf(B)/qf(V) is independent of OCR (Fig. 7.3(b)).
qf/Ovc from the horizontal samples are the largest and qf/avc
from the inclined samples are the smallest (Fig. 7.3(a)).
The inclined (450) samples, which are sheared parallel to the
varves, have the lowest NESE at qf (Fig. 7.4) and for NC soil,
also the largest Af.
Comparing vertical and horizontal samples
(4> OCR >1), both being sheared across the varves, vertical
loading always develops
larger pore pressures and causes a
lower NESE at qf (Fig. 7.4).
These both result in a lower
296
f
vc
(2) The anisotropy in NESE at qf when Pf/vm>
7.4) or in
0.3 (Fig.
(Table 7.1) is due to "material"
(c = O) of NC soil
anisotropy and variation in the degree of mobilization of friction and perhaps cohesion during shear.
plete mobilization of friction
Because of the incom-
for vertical loading, NESE at
qf is lower than that at (al/a3)max' For horizontal loading,
since NESE at qf is equal to that at ( 1 /53)max' it is thought
that the frictional component is fully mobilized at qf (Bjerrum, 1973).
As a result, NESE at qf for horizontal loading is
higher than for vertical loading when Pf/Ovm
0.3.
Note
that
at (al/a3)max' NESE for vertical and horizontal samples are
identical.
Because failure takes place within a clay layer,
NESE at qf for inclined loading is the lowest.
(3) The anisotropy in NESE at ( 1/3)max
is only the re-
sult of whether failure occurs across the varves or within
when shearing is parallel to the varves.
a clay layer
The independency
from inclinations of the failure plane (a = 600 in vertical
sample and
= 300 in horizontal sample) of NESE at (al/3)max
for failure across the varves suggests that
(al/a3 )max of the clay and silt
and c/avm at
portions are isotropic.
That
is, structuralanisotropy does not cause a change in the NESE
at (l/a3)max.
It then follows that the increase in NESE for
failure across the varves (i.e., vertical and horizontal samples) compared to failure within a clay layer (=450
sample)
is mainly the result of the increased strength of the silt
layer.
[CD direct shear test data (see Fig. 2.3 in Lacasse
297
1972) show a higher envelope
et al.,
for failure
in a silt
layer than in a clay layer, except at very high OCR values.]
Furthermore, the NESE at (al/a3)max' for failure across the
varves and for failure within a clay layer are considered to be
material properties.
(4) The anisotropy in pore pressure A at
in Fig. 7.5.
At the same strain (=
10%),
= 1.0% is shown
the horizontal samples
at OCR's of one and four have smaller A and Au than vertical
samples, while results from inclined samples fall between these
values.
It is difficult to assess whether this behavior is
mainly due to material or structural anisotropy, or a combination of both effects.
(5)
In Fig. 7.6, the increase in E(H)
compared to that
from vertical and inclined samples is probably due to a lower
Au and a larger horizontal drained modulus.
An indication of
a larger horizontal drained modulus compared to the vertical
modulus is obtained from CRSC (constant rate of strain compression test) data on Hackensack Valley varved clay (Saxena et al.,
1974).
CR and RR of-a vertical sample were larger than those
from a horizontal sample at the same depth.
In summary, the 'material anisotropy causes:
1.
Anisotropy in qf due to anisotropy in both NESE
at qf and the pore pressure at failure (qf(H)>
qf(V)> qf(4 50 )).
The qf( 8 )/qf(V) ratio was
found to be independent of OCR.
2.
Anisotropy in NESE at qf, though this behavior
298
is also due to variation of friction and cohesion
during shear (horizontal NESE > vertical NESE
>
inclined NESE).
3.
Anisotropy in NESE at ia1 /a3)max.
This anisotropy,
considered to be a true material property, depends
upon whether
failure occurs within the clay layer
or across the varves, the latter leading to a significantly higher failure envelope.
4.
Anisotropy in excess pore pressures, these generally
being largest in vertical samples and lowest in
horizontal samples.
5.
Anisotropy
in
(Eu(H) > E(V
7.4
E u at the same strain
(
= 1.0%)
EU(450) .
)
EFFECT OF STRUCTURAL ANISOTROPY COMPONENT
Figure 7.7 shows a
plot
of qf/vm
versus
at OCR's
equal to 1 and 4 from CIUC tests on samples with different
inclinations and from CK UPSC, CKoUDSS and CKoUPSE tests (results presented in Table 7.2).
The shaded area for the OCR=4
data represents the combined effects of structural anisotropy
and variations in
2.
For OCR = 1.0, the shaded area also
includes the effects of stress induced anisotropy and of shearing samples which have different initial shear stresses, since
K
is less than 1.
In Fig. 7.8, an estimation
of the struc-
0
tural anisotropy component is shown.
versus
Plotting qf(a)/qf(V)
, where qf(V) is from CIUC ( = 0) and CKoUPSC
tests,
0
299
is an approximate method for adjusting CIUC test qf values to
equivalent CIUPSC test conditions as discussed below.
As previously stated, the structural anisotropy component
effect, at least conceptually, would be shown by data at OCR=
4 (i.e., when K
= 1) from CIUPSC tests on samples with dif-
ferent inclinations and from CKoUPSC, CKoUDSS, and CKoUPSE
However, it is reasonable to assume that qf(V) from
tests.
"CIUC PSC (=0)"
since K
and CK UPSC tests would be essentially equal
versus isotropic consolidation showed little effect
on CU TC strengths (Section 53.1).
In essence, Fig. 7.8
ob-
tains the equivalent qf from the plane strain condition by increasing
and 90
qf values
for CIUC samples at inclinations
of 0, 45,
by a factor equal to
qf(V) from CKoUPSC te.st
qf(V) from CIUC
(=0)
test
[Therefore, for example, the estimated qf from a CIU PSC
(8=45) test is equal to qf from CIUC (8=450) test
qf from
qf(V) from CXKUPSC test
qf(V)
from FmIc
(0=0)
qf from CIUC (8=45 ) test
(=0) test
qf(V) from cIc
and
]
CPSC.
x
.).test
(..=.45
test
Also, there is no stress sys-
tem induced anisotropy when K
= 1.
Thus, Fig. 7.8 represents
an estimate of the effect of structural anisotropy on undrained
strength.
As shown in Fig. 7.8, it is seen that structural anisotropy significantly decreases qf for inclined and horizontal
loading and therefore increases the anisotropy in qf over
300
that due to material anisotropy.
The increased anisotropy in
qf is probably due to changes in anisotropy in NESE at qf and
the pore pressure response,but these effects are difficult to
estimate.
7.5
INTERPRETATION OF COMBINED ANISOTROPY
The combined anisotropy in qf, Af and NESE from CUPSC,
'CK UDSS, and CK UPSE data (4 > OCR > 1) are presented in Figs.
0
0
7.9 and 7.10 and Table 7.2.
The method used to estimate the
OC qf values was presented in Section 5.3.3 and Fig. 5.24.
Af
was computed* from Ko values (.seeFig. 5.2) and estimated locations of NESE at qf
(presented in Section 5.3.3).
In Fig.
7.10, the locations of NESE at qf and at (l/a 3 )max for NC and
OC CKoUPSC and CKoUPSE samples are estimated as follows:
1.
As shown in Fig. 5.22, the NESE at qf when
f/avc>
0.3 for NC and OC CKoUPSC samples was estimated
to be slightly above the NESE at qf for NC and OC
CKoUC samples based on NC CKoUPSC test data.
At
(a1 /a3)max' the NESE values for failure across
the varves (a material property) is believed to
hold for NC and OC CKoUPSC samples.
2.
The NESE at ( 1 /O3 )max for failure across the
varves is believed to be the same as the failure
envelopes at qf and at (1l/'3)max for NC and OC
*The values are sensitive to the assumed location of the
NESE, especially at OCR=4.
301
CKoUPSE samples, i.e., for these samples, the envelopes at qf
and ( l/ 3)max are identical.
With help from the study presented in Sections 7.3 and 7.4,
interpretations of the combined anisotropy in qf, A, and NESE
of NC and OC samples (4 > OCR > 1) are as follows:
1.
The combined ansiotropy in qf/avc (Fig. 7.9)
results from:
a.
Anisotropy with respect to NESE at qf
(or the anisotropy in
(c=0) for NC
samples, Table 7.2);
b.
Anisotropy with respect to the pore pressure parameter Af (Fig. 7.9), and
c.
Differences in the magnitudes of Aqf due
to the rotation of principal planes (i.e.,
the stress system induced anisotropy component) when K
is less than 1.
The anisotropy in Af and the changes in Aqf in turn
lead to further anisotropy in the pore pressure at
failure.
qf(H) is slightly smaller than qf(V)
even though the NESE at qf is much higher for horizontal loading.
The decrease results from signi-
ficantly lower effective stress at failure.
This
occurs even though Af for horizontal loading is
lower, as illustrated below for data on NC soil.
0
-m
Type of Test
Pf/a
f -ve
A
-cf
CK UPSC
0.70
1.09
23.5
CKoUPSE
0.55
0.85
27
302
The significantly lower qf/avc for inclined loading
in DSS tests results both from a lower NESE at qf
(Fig. 7.10) and probably from a lower value of
Pf/avc
For NC soil, the DSS pfavc
is 0.49 (Fig.
5.33) compared to 0.55 for PSE.
2.
As previously discussed, the anisotropy in A is
due to both the material and structural components
(i.e., due to inherent anisotropy).
A at
= 0.3%
and Aq/Aqf = 1/2 for NC CKoUPSE are smaller than
for CK0 UPSC as shown below.
Type of Test
3.
A ate=
0.3%
A at Aq/Aqf=1/2
CK UPSC
1.02
0.80
CKoUPSE
0.45
0.52
The anisotropy in NESE at qf, when NESE at qf is not
equal to that at (al/a3)max, or the anisotropy in
Mohr-Coulomb
at qf of NC soil are the result of
inherent anisotropy and of variation in the degree
of mobilization of friction and cohesion.
For fail-
ure across the varves, since the mobilization of
friction and cohesion are strain dependent, the much
larger
f of a NC CK UPSE test compared to that of
a NC CKoUPSC test (f = 4.0%versus0.4%)partly explains why
4 at qf from a horizontally loaded sample is
larger.
[Note that NC
~ at qf and at (l/a3)max
from TUR7PSE tests are identical (Table 7.2).]
Because failure occurs within a clay layer instead
303
of across the varves (i.e., the material anisotropy
component), the NESE at qf from CKoUDSS tests is
the lowest.
The fact that
a
larger difference
between NESE at qf from vertical and horizontal
loading is obtained from CKU
plane strain compres-
sion and extension tests than from CIUC tests on
vertical and horizontal samples (comparing Fig. 7.4
to Fig. 7.10) might suggest that the structural
anisotropy could cause the anisotropy in NESE.
4.
The anisotropy is NESE at (al/a3 )max is only due to
the material anisotropy component, i.e., failure
across the varves versus failure within a clay
layer.
Based on several types of tests reported
in Chapter 5, NESE at (al/a3)max for vertically and
horizontally loaded samples are identical.
Thus,
the lower f at (/a3)mx
from NC CK UPSE compared
1 3 max
o
to CK0 UPSC (Table 7.2) is due to larger p/avc
of the CK UPSE Sample.
For the sake of completeness, the combined anisotropy in
undrained shear modulus is shown in Figs. 7.11 and 7.12 by comparing test data from CK UDSS(C), CK UDSS, and CK UDSS(E)
tests on NC soil.
At low stress levels, inclined loading re-
sults in a lower modulus than for either vertical or horizontal loading.
But at the higher shear stress level, little
modulus anisotropy is measured in this type of testing.
304
This
contradicts the results shown in Fig. 5.27 from PSC and PSE
tests, but these data may be influenced by apparatus friction.
7.6
SUMMARY AND CONCLUSIONS
The factors that control anisotropic strength behavior of
varved clay are the layered nature of the soil (the material
anisotropy component), the prior K o consolidation (the structural anistropy component) and the stress system induced anisotropy, especially important when Ko is low.
of the combined anisotropy in qf, A, Au,
Coulomb
Interpretation
and c/avm (or Mohr-
and c) of NC and OC samples ( 1<
OCR < 4) as shown
fro the SHANSEP method of consolidation are as follows.
1.
The anisotropy in qf results from anisotropy in
effective stress at failure and in NESE at qf.
Due to material anisotropy component, qf(H) of NC
and OC samples is the largest.
With the additional
effects of structural and stress system induced
anisotropy components, the anisotropy in qf
increases from that due to the material anisotropy
component and qf(V) from CK-UPSC tests is larger
than qf(H) from CKoUPSE tests.
The lower qf(H)
compared to qf(V) from CKoU plane strain tests
results from the lower pf of the CKoUPSE sample,
though the NESE at qf is higher.
Mostly because
of having the lowest NESE at qf, qf for inclined
loading is the smallest.
305
2.
The anisotropy in NESE at qf, when NESE at qf is
not equal to that at (l/a3)max' is due to inherent
anisotropy component and variation in the degree
of mobilization of friction and cohesion.
For
vertical loading, because of the incomplete mobilization of the friction and cohesion components of
strength, NESE at qf is lower than that at (1/a3)max
with NC soil.
For shearing across the varves, be-
cause of differences in the degree of mobilization
of friction and cohesion, NESE at qf for vertically
loaded samples is lower than that for horizontally
loaded samples.
This lowering is also due to the
structural anisotropy component.
Because failure
takes place within a clay layer, NESE at qf for
inclined loading is the lowest.
3.
The anisotropy in NESE at (al/aC3)maxis only dependent upon whether failure occurs within a clay
layer or across the varves.
Thus, the anisotropy in
NESE at (a1/a3)max is only due to the material
anisotropy component.
Despite having the same NESE
at ( 1 /a3 )max ,
at (
(=0)
1 /a 3)
ax for NC CK 0 UPSC
tests is larger than that from CK0 UPSE tests, because P/vc
4.
at (1/C3)max
is smaller.
The anisotropy in pore pressure is due to anisotropy in A resulting from the inherent anisotropy
and from differences in the magnitudes of Aq.
306
Due to the material anisotropy component, at the
same strain, A for vertical loading is generally
the largest and A for horizontal loading is the
lowest.
However, at the same strain, despite
having a lower A, a CKoUPSE test can have higher
pore pressure and a lower p/avc than a CKRUPSC
test because of the much larger Aq.
307
-11
-q
to
C
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z
Table 7.2
I
309
·_
/. O
a08
'4
I,
IP
I4
0.6
0.4
o0.
--0
2
4
6
8
o/
AXIAL SRAIN, 6 %
/4
I, A
Y..J
3.0
cr
3
.2.0
AXIAL STRAIN,
,%
EFFECT OF SAMPL.E ORIENTATION ON STRESS
VS STRAIN FROM CIUC TESTS ON SAMPLES OF NORMALLY CONSOLIDATED EAST WINDSOR VARVED CLAY
310
FIGURE 7.1
I
AXIAL STRAIN, 4
%
,
-
3
AX,'AL STRAIN, F,
%
EFFECT OF SAMPLE ()RIENTATION ON STRESS VS
STRAIN FROM CIUC TESTS ON SAMPLES OF OVERCONSOLIDATED (OCR 4.0)
EAST WINDSOR VARVED CLAY
311
FIGURE 7.2
A
U.9
0.3
0.
O.
0
0
(a)
14
/.2
q~yg)
,?f(y
#?depe17de7t,
Of OCR
/.0
0.8
I
7o
/0 20
30
I
40
I
50
60
I
70
I
I
80 90 /
/9
(b)
I..
,
Af,
9 ),
'OR =/
/.0
Af (g. 0
II
n4 0 I
cI)
-,
I
I
II
I
/0
I
i
20 30
I
I
I
I
0
0
I
--
laa
ANISOTROPY IN
AND Af
VARVED CLAY RESULTING FROM
COMPONENT
312
--- -II
60
O CR = 4
I
I
70
I
I
80
90
/00
(C)
OF CONNECTICUT VALLEY
MATERIAL"
ANISOTROPY
FIGURE 7. 3
z
0z
cn
z
uZ
_
--
UC)Z
L
N
ZL Q
0_ >.
·.
LiJ
~J
ZL
~b
'N
313
FIGURE 7.4
0.6
1
0.*
OCR =
OC:
o'vm
at
.1
-Iv
=o/.0%
0.2
$
OCR 4
I
ii
0 o
/I
/o
I
o
Iv
I&
,
3o 40
Ii
5
II
60
I--
-
I
I
70
8o 90
o0
(Q)
I v
1. .
/.0
A
at
E- 10%
0.5
-
/0
-
20
30
40
SO
60
7
WJhS0
f(b)
"MATERIAL'
PORE
ANISOTROPY
PRE'SSURE
IN POR E PRESSURE
PARAMETER A AT
314
AND
:1.0%
FIGURE 7.5
450
A
3SS
o From
C/ZC
D From
cTC r-e5)
EH -
Ev =
EX
at &
5%
From C//C (eo=90) test
=
(e =
o)
test
/A
test
Horizontal Undrained Modvuus
Verticol t/ndroined Modulus
/
Inc//ned Moduis
2at
.
/
EH (e= 90)
/
E(,, - o)
ISO
I
,/
I,{C
20
so
OCR
ANISOTROPY
IN UNDRAINED
CIUC TESTS
315
MODULUS FROM
FIGURE 7.6
I
,o
rnfMDA^DIK
A llnTDDYV
Vl.vlr
VII
r l I. r(VI PI1IV
SI IV I Il
Im
FROM PLANE
n
L-LV
%,f MF'AI
IVL-I , VIICr
STRAIN AND CIUC TESTS
FIGURE 7.7
_
_
2.0
/.8
1.6
/.4
f (V)
/.0
0.8
0.6
0.4
1o
EFFECT OF STRUCTURAL ANISOTROPY COMPONENT
ON ANISOTROPY IN qf
317
FIGURE 7.8
0.4
----
0.3
I
OCR =/
-*----
OCR
OCR 2
-
OCR=4
estimated
}
I
Er.m
V.,
II,
/o
20o
I
3
,
40
SO
la
60
70
80
I
90
1.0
,
qf(v)
a0.
0.7
0.6
8.5
6.5
AsF(v)
Af(v)
4.5
I-,
\~~
2.5
II
-
I
I1
20
I
I°
40
.o
L
60
I
I
0o
/00
COMBINED ANISOTROPY IN qf OF CONNECTICUT
VALLEY VARVED CLAY MEASURED FROM CKoU PLANE
STRAIN TESTS
318
FIGURE 7.9
z
,[
(X)
w
z
0.
0
LL
w
U)
w
z
z
>-
0
(nI
0Z
W
Z
0O
X
O
0
w
U
bon Eo
.
IG
319
_
FIGURE
()
W
7.10
.
b6
50
40
G (e)
rrvc
30
20
/0
4.5comprssion)
0
Ad O
Li 29
-fr'extension)
(
90
G
ANISOTROPY WITH RESPECT TO avc OF NORMALLY
CONSOLIDATED EAST WINDSOR VARVED CLAY AT VARIOUS
APPLIED SHEAR STRESS LEVELS
320
FIGURE
7.11
60
Initia/ tongent
mo
duls /
so KtN%.-- ,
1·
I
-
ace__0
40
ve)
X-a/i
30
I
20
From
CouIDSS tests
Jr
_-
45 0(cmpeessioon)
o
0
___
-5° fe xtension)
B= 90 °
(3=29°
G
ANISOTROPY WITH RESPECT TO
c OF NORMALLY
CONSOLIDATED EAST WINDSOR VARVED CLAY AT 6 .1%
321
FIGURE 7.12
8.
EVALUATION OF UNDRAINED STRESS-STRAIN-STRENGTH PROPERTIES OF CVVC FOR USE IN DESIGN
8.1
OBJECTIVES AND SCOPE
The purposes of this chapter are:
(1) to present best
estimates of the normalized stress-strain-strength parameters
for use in the design of structures on Connecticut Valley varved
clays, with the background showing how these parameters were
developed; and (2) to recommend suitable testing procedures for
developing
anisotropic S /a
u
v
parameters for other varved clay
deposits.
The normalized stress-strain strength parameters are for
the following analyses that might be included as part of the
design process of structures ,Specifically an embankment) constructed on Connecticut Valley varved clays.
1.
Total stress stability analyses to determine the
factor of safety for undrained and partially
drained conditions.
Table 8.1 summarizes recom-
mended uses for total stress analyses.
2.
Analyses to predict stress, pore pressure, and
deformation due to undrained loading.
Recommended parameters such as A- and E -are presented to use
with available chart solutions and with MIT's finite element
program FEECON.
meters.)
(See Table 8.2 for required input soil para-
The parameters for estimation of initial settlement
based on D'Appolonia
et al.
(1971) method
presented
in
Table 8.3. Soil parameters for effective stress stability
322
analyses for drained and undrained conditions, and those for
predicting the amount and rate of long term
not presented here.
settlements are
However, such information can be obtained
from Ladd (1975).
The material presented in this chapter will concentrate on
the background for the development of the soil parameters which
is based upon the study presented in Chapters 4 through 7
(including the use of normalized FV data) rather than presenting the details of methods of analyses.
Since the background
for the development of soil parameters of Connecticut Valley
varved clay is fully discussed, this approach can also be extended to develop normalized anisotropic Su parameters for
other varved clay deposits, presented in Section 8.6.2.
8.2
DEVELOPMENT OF UNDRAINED STRENGTH PARAMETERS OF CVVC FOR
TOTAL STRESS STABILITY ANALYSES USING SHANSEP
Two requirements for obtaining reliable estimates of factor of safety from total stress analyses are:
1.
The use of a suitable method of analysis; and
2.
The selection of
suitable undrained shear
strength values which are compatible with the
method of analysis and reflect in situ soil
behavior.
Since the undrained shear strength behavior of varved clays is
highly anisotropic, the method of analyses should be capable
of taking this anisotropy into consideration.
The Morgenstern
and Price (1965) method of analysis (M-P) or similar "generalized"
323
methods, is recommended for. all cases involving the input of
anisotropic undrained strength data.
Strength data obtained from the SHANSEP method of consolidation will be used for developing the design Su parameters
for the reasons given below.
1.
It is thought that the SHANSEP method should yield
better Su values for NC clays than the RECOMPRESSION method, and also give reasonable estimates
of Su for OC deposits.
2.
NSP for different varved clays located in the Connecticut Valley as measured in several types of
tests using the SHANSEP method of consolidation
are practically identical.
This section presents:
(1) a recommended method for ob-
taining suitable anisotropic undrained shear strength parameters
measured with the SHANSEP method for use with M-P analyses;
and (2) recommended values of SHANSEP anisotropic undrained
shear strength parameters to use in the design of structures
on Connecticut Valley varved clays.
8.2.1
Criteria for Developing Design Undrained Strength Parameters
Values of undrained shear strength parameters should be
theoretically compatible with the method of analysis (e.g.,
Morganstern-Price (M-P) analyses as recommended in this thesis)
and also reflect in situ behavior.
The various factors that
need to be considered in developing appropriate design values
of Su for total stress analyses are as follows:
324
1.
The need to consider the anisotropy in qf--As
shown in Chapter 7, the values of peak strength
and stress-strain characteristics are directional
dependent.
Though tnese properties can be mea-
sured for various loading directions, a practical
method (e.g., that proposed by Ladd (1975) in Fig.
8.1) for considering anisotropy and other factors
is needed.
2.
The suitable definition of undrained shear strength
for M-P analyses--M-P analyses accurately analyze
failure surfaces composed of straight lines.
Therefore, the un-
drained shear strength for such analyses is defined
as the shear strengthon the failure plane.
In order
to reflect in situ behavior and since effective
rather than total stresses control this behavior,
it is appropriate to adopt the inclination of failure planes from the effective stress Mohr-Coulomb
failure criteria (i.e., the inclination of failure
plane with the major principal plane is(45
+
/2).*
Consequently, for failure at the peak strength (qf),
the shear stress on the failure plane, Tff, for a
given mode of failure (PSC mode for vertical loading, DSS mode for inclined loading and PSE mode for
horizontal loading), is equal to qf cos
( at the
*Observations in the laboratory indicate that the inclination of failure planes, if visible, with the direction of the
major principal plane is more than 450 (predicted by a total
stress Mohr-Coulomb failure criteria) and is close to 45 + /2
( at (al/a3)max)
325
qf condition.
3.
The strain rate effect--The stress-strain characteristics and hence qf are dependent upon the
strain rate, a decrease in the rate of strain
decreasing qf.
uncertain.
The magnitude of this effect is
However, based on information from
Ladd et al. (1977), it is thought that the slower
in situ strain rate compared to that used in these
CKoU laboratory testing programs should cause relatively little reduction in qf.
4.
A correction for strain compatibility--Because of
anisotropy in stress-strain-strength properties of
soil, it is unlikely that the peak strength (qf)
for varying mode of failure can be simultaneously
mobilized in strain softening material since these
peaks occur at different strains.
formation
of a distinct failure
In fact, the
(slip) surface
throughout the foundation soil is thought to be
primarily a function of the strains which occur.
Ladd (1975) proposed a method (Strain Compatibility Method) to account for strain compatibility
assuming that the shear strains along a potential
failure surface are equal.
Note, however, that
this assumption is not based on theory or observation.
326
8.2.2
Anisotropic S
Parameters Developed by Ladd (1975)
This section presents Ladd's (1975) basic approach for
obtaining anisotropic Su design values.
This was done before
all test data (especially for horizontally loaded samples)
presented in this thesis were available.
The basic approach
is as follows:
1.
For simplicity, the three types of failure surfaces
and corresponding strength shown in Fig. 8.1 are
recommended for considering Su anisotropy for use
with the M-P method of analysis.
2.
The definition of undrained shear strength is that
previously defined in Section 8.2.1.
qf cosT for PSC and PSE modes and
Tff
Thus,
Tff
in the DSS
mode is assumed to be equal to Thmax.
3.
The strain rate effect is not considered.
4.
The strain compatibility effect is considered
by using the strain compatibility method described
below.
The recommended anisotropic Su values from three modes of
failure are considered to be equal to the laboratory measured
shear stresses which yield the maximum average shearing resistance that could be mobilized at the same shear strain for the
three modes of failure.
The shear stresses and shear strains
are determined as follows.
(a) The value of shear stress on the potential
failure plane at a particular strain for either
PSC or PSE modes of failure is equal to q cosT.
327
The values of cosT selected were:
Mode of Failure
OCR
cosT
PSC
1
>2
0.92
0.90
PSE
>1
0.90
(b) The shear stress on the potential failure plane
at a particular shear strain for the DSS mode
is equal to Th.
(c) To obtain shear strains, the measured vertical
strains for PSC and PSE samples is multiplied
by 2 (i.e.,y=2
e vertical).
In DSS tests, Y is
equal to the horizontal displacement divided by
the height of the sample.
Table 8.4 shows how SHANSEP anisotropic Su values at OCR1.0 were determined.
It is seen that the average maximum mobi-
lized resistance, which occurs at a shear strain of 6%, is 8%
less than the average of the peak shear strengths.
At this
strain, Su for PSC is significantly less than the peak vertical
shear strength (0.258), due to the strain softening.
Su is at Tff (= T hmax =0.16).
The DSS
The 6% shear strain is not large
enough to fully mobilize the peak horizontal shear strength
(0.225). In essence, for considering strain compatibility,
ff
values for PSC and PSE were reduced by 16% and 6% respectively,
i.e., the strain compatibility factor (SCF) have values of 0.86
and 0.94 respectively.
For the DSS modes, the SCF is equal to
1.0.
For OC soil, since the laboratory plane strain stress328
strain-strength data were not available, OC qf values for PSC
and PSE were estimated from very limited TC and TE results
(using a method similar to that presented in Fig. 5.24).
SR
values for OC PSC samples used by Ladd (1975) are essentially
the same as those presented in Fig. 5.24.
However, qf(H) for
OC PSE samples was estimated by assuming that K
was equal to 0.89 independent of OCR.
(Ks = qf(H)/qf(V))
This means that the same*
relationship of SR versus OCR was used for vertical and horizontal loading.
To consider strain compatibility, the follow-
ing values of SCF were used based on stress-strain behavior observed with other clays.
SCF
5.
Avg. SCF
OCR
PSC
DSS
PSE
2
0.925
1.0
0.975
0.967
4
1.0
1.0
1.0
1.0
Finally, Ladd (1975) arbitrarily reduced the design strength by 5%, i.e., Su=qf cosT x SCR x 0.95.
8.2.3
Design Parameters Developed by the Author
The basic approach used by the author for developing ani-
sotropic Su design parameters generally follows that proposed
by Ladd (1975), with the exceptions that:
1.
The method for estimating OC SCR for each mode of
failure is different.
The author used results
from triaxial tests (details given below).
2.
A strain rate effect is considered.
A strain
*As shown in Fig. 5.24, SR versus OCR relationships for
PSC and PSE are not the same.
329
rate factor of 0.95 is arbitrarily selected for
Connecticut Valley varved clays. (Note that based
on Ladd et al. (1977), Aqf/Alog tf
A< 0.5%/hr.)
3.
Thus,
10+5% at
<
u=qf cost* x SCF x 0.95.
The OC qf values for PSC and PSE were estimated
from Fig. 5.24.
This gives the same qf(V) but
different values of qf(H) compared to Ladd (1975).
SCF values for NC soil are those from Table 8.4.
The
basis for estimating SCF for OC (OCR=2 and 4) soil is as follows.
Comparing Tables 8.4 and 8.5, it is seen that the average
SCF at OCR = 1.0 for plane strain and triaxial tests are essentially identical, i.e., by using the stress-strain-strength data
from CKoUC, CKoUDSS, and CK UE tests.
tests,
is equal to 1.5
(For CKoUC and CK UE
vertical and the shear stress on the
potential failure plane at a particular strain is equal to
cost.)
q
Note, however, that SCF values from PSC and PSE are difFurthermore, for
ferent respectively from those from TC and TE.
both triaxial and plane strain loadings, SCF for the DSS mode
is always equal to unity.
On the basis of the NC data, the esti-
mated OC average SCF values for plane strain loadings at OCR's
of 2 and 4 are assumed to be equal to those from the triaxial
loadings, shown in Tables 8.6(a) and 8.6(b).
The OC SCF for
the DSS mode was assumed to be equal to unity.
Details of how SCF for OC (OCR=2 and 4) PSC and PSE modes,
and hence Su values (Table 8.7), were evaluated are given below.
*cost values are identical to those proposed by Ladd (1975).
330
1.
At OCR=4, since the average SCF is practically
equal to 1, and since it is thought that the OC
CKoUPSC sample should not exhibit significant
strain softening, SCF for the
were selected
2.
PSC and PSE modes
to be 1.0.
At OCR = 2, SCF values for PSC and PSE modes are
determined from the average SCF value from triaxial loading (avg. SCF = 0.96) and from the estimated SU(H)/Su(V) ratio.
As shown in Table 8.7,
at OCR's of 1 and 4, this ratio is equal to 1.05.
Thus, at an OCR of two, it is reasonable to adopt
Su (H)/ S u(V) = 1.05 for the determinations of Su (V)
and S (H).
Table 8.8 compares
u values recommended by the author and
by Ladd (1975). It is seen that the only difference is in Su(H)
at OCR's of 2 and 4 '(the author's S(H)
being 9% larger at
OCR = 4 and 6% larger at OCR = 2).
8.3
COMPARISON OF THE USE OF SHANSEP ANISOTROPIC Su VERSUS CORRECTED FV STRENGTH DATA
8.3.1
Introduction
The objectives of this section are to present:
1.
The background of how the correction factor p is
estimated; and
2.
Checking of the compatibility of using recommended
SHANSEP strength data with M-P analyses for undrained
total stress stability analyses.
331
This is done by comparing factors of safety from the simplified
Bishop circular arc analyses with corrected field vane strengths
and from M-P analyses with Ladd's (1975) SHANSEP anisotropic Su
parameters.
(Note that since che author's recommended anisotro-
pic Su values are only slightly different from those recommended
by Ladd (1975), the conclusions drawn from using Ladd's (1975)
recommendation should also hold with the author's anisotropic
Su parameters.)
8.3.2
Recommended Correction Factor () For Connecticut Valle
Varved Clay
According to Bjerrum (1972), the correction factor
em-
pirically relates the field vane strength data to the value of
the average in situ undrained shear strength on an assumed circular arc failure surface obtained from total stress analyses
with
= 0 and c = field vane strength.
nature, reliable values of
Because of its empirical
can only be determined from case
studies.
The basis for recommending two
soil and
values ( = 0.85 for OC
= 1.0 for NC soil) for CVVC is as follows.
(1). The recommended i= 0.85 for OC clay is based on case
studies at Amherst and Northampton.
The average OCR's of the
foundation clay within the critical zone at these two locations
were approximately 3.5 at Amherst (ranging from 1.5 to 8) and
about 2 at Northampton.
Thus, the correction factor of 0.85
was developed for moderately to heavily OC clay.
(2) The recommended
=- 1.0 for NC clay is based upon the
trend established from a comparison of the average SHANSEP anisotropic design shear strength parameters with corrected and
332
uncorrected field vane strengths (Table 8.8).
At OCR's of 2
and 4, (i.e., within the OCR range of the case studies), the
average SHANSEP design S/Zvc
from both Ladd (1975) and the
author is close to the corrected FV strength using p = 0.85.
However, at OCR = 1.0, the average SHANSEP Su/vc
at 0.19 is
equal to measured field vane strength (Su/Uvc = 0.19).
fore, the value of
There-
for NC soil is taken to be equal to 1.0.
Since the use of corrected field vane strengths with Simplified Bishop circular arc analyses is an empirical approach
and since the value of
is dependent on several factors (e.g.,
strain rate, anisotropy, strain compatibility, etc.), its use
has limitations which are discussed below.
(1) The value of
is dependent upon the geometry of the
failure surface and hence the geometry of the embankment. Because of the anisotropy in Su, the average shear strength along
a long, shallow wedge-type surface will be different from that
along a deep circular surface.
Thus, an embankment with wide
berms would require a different correction factor U from one
without berms.
(2) The value of u is dependent upon the stress history of
the site.
The anisotropy, the strain rate and the strain com-
patibility effects are dependent upon OCR.
Plane strain data
from Boston Blue clay (see Chapter 2) show that qf(H)/qf(V) increases with increasing OCR, though a much smaller increase was
observed from estimated data for the Connecticut Valley varved
clays (see Chapter 7).
Furthermore, the use of circular arc analyses with corrected
333
field vane strengths for total stress stability analyses for the
partially drained case is not reliable.
8.3.3
Compatibility of Using SHANSEP Anisotropic S
Analyses
with M-P
To check the compatibility of using SHANSEP anisotropic
Su with M-P analyses, it is necessary to see if the M-P analyses with SHANSEP strengths yield about the same factor of safety
as that determined from the Simplified Bishop analyses with
corrected field vane strengths.
Note that the correction fac-
tor was determined from case studies.
Figure 8.3 compares values of the factor of safety and the
locations of the critical surface of an embankment, used in a
design problem presented in Ladd and Foott (1977), from M-P
analyses with Ladd's (1975) SHANSEP anisotropic strength parameters and from Simplified Bishop circular arc analyses with
corrected field vane strengths.
equal to 0.85.
The correction factor
was
Since the stress history, soil profile, and the
geometry of the embankment used in the design problem are somewhat similar to those from the case studies, the fact that the
factor of safety from Ladd's (1975) SHANSEP S
is slightly
lower than that from the corrected field vane strengths, and
that these values of factor of safety are equal to or very close
to 1,suggests that Ladd's (1975) SHANSEP Su values used with
M-P analyses are suitable to determine the factor of safety from
total stress stability analyses for Connecticut Valley varved
clays.
Since the difference between Ladd's (1975) and the author's
334
Su values is only the value of S(H)
OC S (H) is about
10% larger),
for OC soil (the author's
the factor of safety using the
author's recommended Su values with M-P analyses should also be
suitable.
Comparing Su values from triaxial loading, Tables 8.5,
8.6(a), and 8.6(b) to those from Ladd (1975) (Table 8.8), the
use of Su values from triaxial testing with M-P analyses will
be conservative.
8.4
EVALUATION OF UNDRAINED SHEAR STRENGTH FOR TOTAL STRESS
ANALYSES IN PRACTICE
The objective of this section is to recommend practical
methods for evaluating the undrained shear strength of Connecticut Valley varved clays for total stress analyses used in undrained and partially drained cases (Table 8.1).
This informa-
tion is partly extracted from Ladd (1975).
8.4.1
Initial Undrained Shear Strength for Total Stress Analyses
(Undrained Case)
The factor of safety for the undrained case determined
by total stress analyses can be obtained from two approaches:
1.
Corrected Geonor field vane strengths with Simplifie.d Bishop circular arc failure analyses.
2.
SHANSEP anisotropic shear strength parameters determined from Section 8.2 (Fig. 8.4) with M-P
analyses.
Since the correction factor
for Connecticut Valley varved clays
335
(presented in Section 8.3) is believed to be reasonably reliable
and since the critical failure surface from Simplified Bishop
analyses can easily be determined using computer programs, such
as LEASE (Bailey, W.A. and Christian, J.T., 1969), the first
approach is generally recommended for determining the initial
factor of safety.
However, as previously discussed
ments with long berms, since the values of
for embank-
can be affected by
the geometry of the slip surface, the use of the
values re-
commended in Section 8.3 may be unsafe.
Since field vane strength data generally show some scatter,
sufficient Geonor field vane tests should be performed to determine the change in Su(FV) with depth, and to ascertain whether
or not there is a significant lateral variation.
Upon knowing
the stress history of the site, the plot of S (FV)/avc versus
OCR (Fig. 8.4) can also be checked.
Section 8.6 presents re-
commended types of laboratory tests and also Geonor field vane
tests for checking pertinent soil parameters.
8.4.2
Undrained Shear Strength for Total Stress Analysis in a
Partially Drained Case
Total stress M-P analyses with SHANSEP anisotropic un-
drained shear strengths are recommended for determining the factor of safety for the partially drained case (Table 8.1) for
three reasons.
1.
SHANSEP undrained strength data can reasonably be
used for estimating the gain in Su with consolidation for stability analyses, even though consolidation of the foundation clay under the embankment
336
will not conform to K
consolidation.
The error
in vertical undrained shear strength should be
minor (Section 5.2.3); Su(DSS) should increase
when consolidated with a horizontal shear stress
(Ladd and Edger, 1972); and the horizontal undrained
shear strength should be unaffected since this portion of the failure plane will usually lie beyond
the influence of consolidation from the embankment.
2.
A method of analysis whereby the difference in Su
with different failure surface inclinations can be
taken into consideration is needed.
3.
M-P total stress analysis is relatively simple
compared with effective stress analyses.
Further-
more, because of the difficulty in predicting the
pore pressure at failure, effective stress analyses are not recommended.
For the evaluation of Su for M-P analyses of the partially
drained case, the following
1.
steps are recommended.
Determine the initial stress history, i.e., a
vo
and a
prior to construction.
vm
2.
Compute the increase in total vertical stress Aav
versus depth at representative locations using
elastic theory.
3.
Determine the degree of consolidation, U, versus
depth corresponding to the time at which the
factor of safety is required.
337
4.
Compute the increase in the vertical stress, AOvc'
versus depth at representative offsets from Aa
UAav
5.
+ Avc versus depth (i.e.,
vc =
vc
vo
vc
assume that o
does not change during undrained
v
Compute
loading).
Though this is not absolutely correct,
the error is unimportant compared to other uncertainties in the analyses.
6.
Divide the foundation clay into zones having the
same
7.
vm and
vc versus depth.
Compute values of the anisotropic Su for the appropriate modes of failure (i.e., PSC, DSS, PSE) within
each zone using the Su/vc
versus OCR = avm/vc
curves plotted in Fig. 8.4.
Ladd (975)
presents an example problem using the above proce-
dure for computing the factor of safety for the partially drained
case.
The evaluation of S u from SHANSEP anisotropic shear strength
parameters (Fig. 8.4) requires an accurate evaluation of the
stress history, i.e., a
and
. The value of the initial in
vo
vm
situ vertical effective stress ( vo) is determined from measurement of total unit weights and the pore pressure conditions.
The latter requires a knowledge of the elevation of the water
table and a check to determine whether or not the pore pressures
are hydrostatic.
The maximum past pressure, avm' is commonly
evaluated from the results of consolidometer tests.
discussed appropriate methods for estimating a
vm
sion curves.
338
Ladd (1975)
from compres-
8.5
RECOMMENDED PARAMETERS FOR UNDRAINED ANALYSES OF PORE PRESSURE AND DEFORMATION
The objective of this section is to present soil parameters
to be used in making predictions of pore pressure and undrained
deformations.
Either one of the following two approaches, dis-
tinctly different in their level of sophistication and amount of
information yielded, is recommended for such predictions.
The
first approach (Table 8.3) employs hand calculations and the
theory of elasticity corrected for effects of local yielding
(D'Appolonia, Poulos, and Ladd, 1971).
meters for this approach are:
The required soil para-
(1) E u , initial shear stress
ratio (f), and the undrained factor of safety for the predictions
of the initial settlement, pi; and (2) Skempton's (1954) pore
pressure parameter, A, for pore pressure predictions.
The second
approach, still being developed, employs the finite element program FEECON for estimating the undrained deformations, stresses,
and pore pressures throughout the foundation clay.
This approach
requires hyperbolic undrained stress-strain-strength parameters
(including anisotropic effects), and the pore pressure parameter,
a(S).
Table 8.2 lists the required soil parameters for the
finite element program, FEECON.
The presentation will include background for the development
of these soil parameters and a brief outline of the procedures.
8.5.1
Parameters for Predictions Using Elastic Theory
(A) Prediction of Initial Settlement
The'relationship for prediction of initial settle-
ment (D'Appolonia, Poulos, and Ladd, 1971) at the center of the
339
loaded area is given below.
A
P.
1
B I
v
IEq.
P
Eu(field)SR
8.1
where
Aav = applied vertical stress;
I
= influence
B = width
factor;
of the loaded area;
Eu(field) = in situ undrained
modulus;
SR = correction factor for local yielding
initial settlement (Pi)
elastic settlement (Pe)
The values of SR can be obtained from Fig. 8.7.
of:
(1) the initial
shear stress ratio
SR is a function
(f) (Fig. 8.6);
(2) the
applied shear stress ratio, Aq/Aqf, = 1/FS; and (3) the ratio of
the depth of the foundation (H) to the width of the loaded area
(B). The coefficient, Ip, can be obtained from charts in Poulos
and Davis (1974).
For the evaluation of Eu(field)' a correction factor needs
to be applied to the measured Eu values given in Fig. 8.5 for
the following three reasons:
(1) The laboratory SHANSEP Eu is generally lower than that
from the RECOMPRESSION method, though the difference is small
for vertical loading (Chapter 6).
The in situ Eu may also be
somewhat different from that obtained from the RECOMPRESSION
method.
(2)
situ E
Eq. 8.1 requires
a single value of E u , while the in
varies with loading direction.
Though E/S
u
values from
SHANSEP CKoUC tests are only somewhat larger than those from
340
SHANSEP CKoUDSS tests (Fig. 8.5), the larger value of S
from
CK0 UC tests compared to that from CKoUDSS tests (qf(V)>Thmax)
causes a significant increase in E.
Eu from CKoUE tests on
overconsolidated RECOMPRESSION samples are also higher than
those from CKoUC tests (Chapter 6), though the NC value measured
using SHANSEP is lower.
(3) Although a correction is made for local yielding (Eq.
8.1), an error in computing
a straight line to curved
theory.
Pi still exists because of fitting
-e data used in the linear elasticity
The values of secant modulus decrease with decreasing
factor of safety.
The correction factors presented below are from Ladd (1975)
based on a case study of am embankment at Northampton, Mass.
(Connell et al., 1973).
E
Factor of Safety
u (field)
E(field)
>2.5
2 - 3
1.25 - 2.5
1.5 - 2.0
<1.25
1.0 - 1.5
The factor of safety is determined from the method recommended
in Table 8.1.
The following are the five steps used in the eval-
uation of Eu(field)
1.
Determine the average OCR, average corrected field
vane strengths, and the undrained factor of safety.
The maximum depth for determining the average OCR
and average field vane strength is recommended to
be between 1.5 to 2.0 B.
2.
Equate the applied stress ratio to the inverse of
341
the factor of safety.
3.
Determine E/S
u
from Fig. 8.5 at the appropriate
OCR and Th/Su = 1/FS.
4.
Compute an average E(DSS)
using the above E/S
u
value and the average field vane strength.
5.
Select the upper and lower estimates of Eu(field)
from the correction factors given above and compute Eu(field) from:
Eu(field) = correction fac-
tor x E(DSS)
(B) Prediction of Excess Pore Pressure
Pore pressure predictions by hand-calculations are obtained using Skempton's A parameter and Eq. 8.2 (for B = 1.0)
given below.
Au = A
3
+ A(B)(Aa 1 - Aa) for Aq/Aqf < 1.0
The magnitudes of A
elasticity.
1
and A
3
Eq. 8.2
are computed from the theory of
(See Poulos and Davis (1974) for comprehensive
collection of graphs, tables and formulas.)
Figure 8.8
pre-
sents the recommended A parameter for vertical, inclined and
horizontal loading at Aq/Aqf = 0.5 and 1.0 versus OCR.
In Fig. 8.8, the recommended A values for NC soil are obtained from SHANSEP CK PSC and CKoUPSE tests since it is be0
0
lieved that the SHANSEP method of consolidation should yield
the better value of A than the RECOMPRESSION method.
However,
for OC samples, since the RECOMPRESSION method yields more
realistic values of A than the SHANSEP method (see Chapter 6),
OC Af for plane strain loading are estimated from the OC
342
-
I
RECOMPRESSION data.
For both vertical and horizontal loading,
OC Af values for plane strain loading were estimated from the
value of NESE at qf for RECOMPRESSION CKoUC and CKoUE tests,
which are also equal to NESE at (a1/~3)max (the material property), from K
values shown in Fig. 5.2, and from the estimated
SHANSEP OC qf values in Fig. 5.24.
At
q/Aqf = 1/2, based on
RECOMPRESSION CKoUC and CKoUE sample behavior for both vertical and horizontal loading, OC A values are equal to those from
NC soil.
Parameters for Finite Element Program FEECON
8.5.2
The finite element program FEECON was developed at MIT
for plane strain analyses of stresses and deformations caused by
construction of embankments on soft clay foundation.
Briefly,
FEECON is capable of:
1.
Considering initial in situ shear stress;
2.
Simulating the accumulation of material (e.g., construction of an embankment, excavation of material,
and incremental loading);
3.' Using hyperbolic and bilinear stress-strain models; and
Considering anisotropy in qf using elliptical re-
4.
lationships (Davis and Christian, 1971) for local
yielding.
The aniostropic stress-strain behavior and pore pressure response
can be taken into account by dividing the finite element grid
into three zones:
vertical, inclined, and horizontal.
In each
zone, the soil is assumed to be a piece-wise linear isotropic
material.
Therefore, three sets of soil parameters for stress343
strain and pore pressure are needed.
Table 8.2 lists the required soil parameters for the fiThey are the initial undrained
nite element program FEECON.
modulus (Gi), the factor relating the actual undrained shear
strength (qf) to the strength computed at infinite strain (Rf),
a/b ratio, K s and the pore pressure parameter a.
The parameters Rf and G i for the hyperbolic stress-strain
model are obtained from the relationship given below:
For triaxial and plane strain tests:
q/ =
/(/v t
Eq. 8.4
(9f/avc)
Aq = 0.5(Aa
1
-
Aa 3 )
Aqf = 0.5(Aalf - A
=
1 .5
3 f)
(axial) for triaxial
loading
and
y = 2
(vertical) for plane strain loading.
For direct simple shear tests:
T
/c
h/vc
Eq. 8.5
l1/i/7vc)+ RfY
(Thmax)
Recommended values of Gi/Ovc and Rf are shown in Fig. 8.9 along
with the test data used to develop these recommendations.
The recommended values of K
5.2 and 8.10, respectively.
and qf(V) are shown in Figs.
The values of qf for vertical
loading and horizontal loading were taken from Fig. 5.24. Values
of Ks and b/a ratio versus OCR for plane strain and triaxial
344
qf from CKoUDSS and the
loading are presented in Fig. 8.11.
measured and estimated qf values from plane strain vertical and
horizontal loading were used to determine
qf(4 5 )
b/a =
¢qf (V)qf (H)
The recommended pore pressure parameter a versus OCR at applied
shear stress ratios equal to 0.5 and 1.0 (Fig. 8.12) were determined from Skempton's A parameters (Fig. 8.8) with the equation
given below
(assuming A
2
= 0.5 (A
1
+ A 3 )).
aa = (3A - 1.5)
8.6
Eq. 8.6
RECOMMENDED PROCEDURES FOR CHECKING CVVC NSP AND FOR DEVELOPING NORMALIZED SHEAR STRENGTH PARAMETERS FOR OTHER
VARVED CLAYS
The objective of this section is to recommend suitable
testing procedures for checking the design soil parameters of
Connecticut Valley varved clays and for developing anisotropic
Su parameters used in the design of structures on other varved
clay deposits,
Section 8.6.1 will present the recommended
testing methods for checking CVVC soil parameters.
If the NSP
presented in Sections 8.2 through 8.5 cannot be used, suitable
laboratory testing procedures for developing other design NSP,
specifically the anisotropic Su values, are presented in Section 8.6.2.
345
8.6.1
Recommended Method for Checking Soil Parameters
The soil parameters presented in Section 8.2 through 8.5
were developed to use in the design of structures on Connecticut Valley varved clays.
They are based on test data from
Northampton, Amherst, and East Windsor varved clays.
Therefore,
besides determining the stress history of the site, and the
Atterberg limits of the clay, silt and bulk portions for comparison with values given in Table 3.7, the testing program
should include the following addition tests:
1.
FV tests for indicating spacial variations in
strength.
The FV data are used for three purposes:
(a)
to obtain S (FV) for TSA;
(b)
for use in conjunction with results from
oedometer tests to check whether the measured
value of normalized FV strength [SU(FV)/ vo]
at a given OCR agree with that in Fig. 8.4; and
(c) if they agree, to use the relationship in
Fig. 8.4 to help for obtaining a better defined stress history profile.
2.
SHANSEP CIUC (=0) tests or CKoUC tests and constant volume direct shear or CKoUDSS tests at
various OCR's because of the need to check the
shear strength in two out of three directions.
8.6.2
Recommended Laboratory Testing Procedure for Developing
Normalized Soil Parameters for Other Varved Clay Deposits
(A) Non-Cemented Soil
Based on the study presented in Chapters 4, 5, and 6,
346
and on the background used for developing the design normalized
soil parameters presented in this chapter, the SHANSEP CKoUC
CKoUDSS, and CKoUE tests are recommended for developing anisotropic Su .
It is believed that for tube samples, the SHANSEP
method of consolidation should yield reasonable estimates of
the NC and OC in situ qf values, especially when the soil has
a low OCR, though it causes the sample to behave in a less
brittle and sensitive manner compared to the in situ soil.
The use of SHANSEP CKoUC and CK UE tests for the evaluation
of Su for plane strain loading will yield conservative values.
As a result, an adjustment to convert triaxial S
strain Su values is recommended.
to plane
This can be done by assuming
that Su(V)/qf from CKoUC test and Su(H)/qf from CKoUE test
ratiosfrom the Connecticut Valley varved clay also hold for
other deposits.
That is, for a given mode of failure, the
design:
SU
qf (measured from other deposits)x design Su values -or
Conn. Valley
qf (measured from Conn. Valley deposit)
qf values from CKoUC and CKoUE tests for Connecticut Valley varved
clays are presented in Fig. 4.14 and Su values for PSC and PSE
failure modes for Connecticut Valley varved clays are presented
in Fig. 8,4. Thmax from the CKoUDSS test can be used directly as
Su value for the DSS mode.
Though conventionally used in practice, UUC tests or RECOMPRESSION (avc = Ovo) CIUC tests on samples with different
inclinations are not recommended because of disturbance due to
separation
of varves and the fact that the principal planes
are not rotated during shear (i.e., the stress system anisotropy
347
component is not included in the measurement).
SHANSEP CIUC
tests on smaples with different inclinations are also not recommended partly because they neglect the importance of stres
system induced components of anisotropy.
qf(H) and qf(45) from
these tests are too large compared to those from CKoUPSE and
CKoUDSS tests, though qf(V) from CIUC (=0) tests will probably
be only slightly lower than that from CKoUPSC tests.
(B) Naturally Cemented Soil
)
The RECOMPRESSION method of consolidation ( vc = avo
vo
is recommended with the types of tests similar to those for the
non-cemented soil.
348
rn
0
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349
Table 8.1
IL
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Table 8.2
i
350
SOIL PARAMETERS FOR THE ESTIMATION OF INITIAL SETTLEMENT
(D'APPOLONIA ET AL., 1971)-AND FORE PRESSURE
Soil Parameter
Item
u(field) (Fig
Initial Settlement:
Aav B I
p
v
85)
f (initial shear stress
p
ratio)
versus OCR:
E(field)SR
SR:
(Fig. 8.6)
(Fig. 8.7)
at i of loaded area
Pore Pressure Parameter A
Pore Pressure:
(B = 1.0)
at Aq/Aqf = 0.5 and 1.0
u = Ac 3 + A(al -
A
(Fig. 8.8)
3
Table 8.3
351
DETERMINATION OF S
AT OCR =;1.0 FROM PLANE STRAIN LOADING
USING STRAIN COMPATIBILITY METHOD (AFTER LADD, 1975)
Avg. Normalized Peak Shear Strength = 1/3 (qf(V)cosf+Thmax+qf(H)cosf)
= 1/3 (0.28x0.92+0.16+0.25x0.90)
= 1/3(0.258+0.16+0.225)
= 0.214
(%)
(%)
PSC
DSS
q:(V)
cost
Th
PSE
q(H)cosV Avg.
% of Avg. Peak
Shear Strength
1.0
2
0.248
0.150
0.153
0.184
86
2.0
4
0.235
0.160
0.194
0.196
91
3.0
6
0.221
0.160
0.212
0.198
92 (max)
4.0
8
0.21602
202
.0.16.0
025
0.196
91
From above results.
scF
Notes:
(1)
(2)
cos
SCF
Avg. SCF
PSC
1
DSS
I
PSE
0.857
1
1.0
I
0.942
= 0.92 for PSC, and cos = 0.90 for PSE
= qcos~ at max % of Avg. Peak Shear Strength
qfcos
(3) Avg. SCF = 1/3 (SCF(PSC) + SCF(DSS) +
(4)
0.935
SCF(PSE))
% of Avg. Peak Shear Strength =
1/3(q(V)cosp + Th + q(H)cos4)
Ave. normalized peak shear strength
(5) Stress-strain data are presented in Chapter 5.
Table 8.4
I
352
AT OCR = 1.0 FROM TRIAXIAL LOADING USING
DETERMINATION OF S
STRAIN COMPATIBILITY METHOD
%
q()cos(l1
5.33
8
0.212(2)
0.1662)
0171(2)
.18
11.33
17
0.193
0.150
0.189
0.17
%.
Th
Strength
q(H)cos(l)
+
94 (max)
92
Avg. normalized peak shear strength = 1/3(qf(V)cosT
Thmax
= 1/3
+ qf(H)cos4)
(0.25x0.92+0.16+0.2x0.9)
= 0.193
From above results:
Avg. SCE (4)
A.
CKUC
0.922
Notes:
(1)
I
CKUDSS
1.0
I
CK^UE
- jm
0.864
0.930
cos4 = 0.92 for CK UC and cos4 = 0.90 for CK 0 UE test.
(2) Equal to Su (no adjustment for strain rate effect).
qcos% at max. % of Avg. Peak Shear Strength
(3) -SCF =
(4)
qfcos
5
Stress-strain data are presented in Chapter 5.
Table 8.5
353
DETERMINATION (1)'OF Su BY US;INGSTRAIN COMPATIBILITY METHOD FROM
TRIAXIAL LOADING AT OCR = 2
At OCR = 2; avg. normalized peak shear strength
= 1/3 (0.40 x 0.90 + 0.28 + 0.38 x 0.90)
= 1/3 (0.360 + 0.28 + 0.342) = 0.327
Vertical
Inclined
E y
:5.33
8
142)
%h
0.351
· q.V~cs()
(·%
q(H)cosp
0.280
0.315
0.280(1)
o.331
3 )
Avg.
Shear
Strength
0.315
96
0.318
97 (ax)
8.00
12
0.3423 (
10.00
15
0.333
0.280
0.342
0.318
97
12.00
18
0.324
0..270
0.342.
0..312
95
Avg SCF =
Notes:
)
% of Ag.
Horizontal
0:342
0.28
1/3 ( 360 + 0.28
0.318
+ 0i342) = 0.96
(1) Tabulated stressstrain data from CKoUC, CKoUDSS
and CKoUE are presented in Appendix D.
(2) cos4 = 0.9 in all cases.
(3) Equal to S u (no strain rate adjustment).
Table 8.6(a)
354
DETERMINATION (1
)
OF Su BY USING STRAIN COMPATIBILITY METHOD FROM
TRIAXIAL LOADING AT OCR = 4
Avg. normalized peak shear strength = 1/3 (0.66x0.9+0.46+0.64x0.9)
= 1/3 (0.594 + 0.46 + 0.576)
= 0.543
%%
8.00
Vertical
Inclined
-'2'
2
q (V)h
12
0.594
10.67
16
0.594
12.00
18
0.585
(3 )
Horizontal
Notes:
(H)cos (2 ) vg.
0.450
0.531
0.460 (3 )
0.538
0.380
0.576
Avg.SCF = 1/3 (0.594
0.594 + 0.46 + 0.576
of Avg.
eak Shear
(3 )
Strength
0.525
97
0.537
99 (max)
0.514
95
= 0.991.00
.
(1)
Tabulated stress strain data from CK UC, CKoUDSS,
and CKoUE are presented in Appendix i.
(2)
cosT = 0.9 in all cases.
(3)
Equal to S
(no strain rate adjustment).
Table 8.6(b)
355
DETERMINATION OF S
U.
ORPSC
BY THE AUTHOR'S APPROACH
DSS
PSE
OCR
U(H)
qfcos (2
(
SSCF
S
j
h
)CF
COS
u
S(
u
1
0.258
0.86 0.210 0.160 .00.15
0.225
0.940.20
1.05
2
0.423
0.94 0.38
0.2801.0 0.27
0.396
0.950 .36
1.05
4
0.864
1.0
0.4601.00.44
0.648
1.00 .62
1.05
Notes:
0.65
(1) Su = qfcos
(2)
cos
x SCF x 0.95 (strain rate factor)
= 0.92 for NC PSC mode and cosl = 0.90 for
NC PSE Mode. For OC soil, for both PSC and
PSE mode cos = 0.90.
Table 8.7
COMPARISON OF S
VALUES FROM LADD (1975), AUTHOR AND CORRECTED
FIELD VANE STRENGTHS
Ladd (1975)
OCR
S (PSC)
(DSS) S(PSE)Avg. P
Author
Sc
orrect
Su(DSS)(PSE
Avg trengths
1
0.21
0.15
0.20 3.187 0.21C 0.15
0.20
0.187
0.19
2
0.38
0.27
0.34 3.330 0.38
0.27
0.36 0.337
0.34
4
0.65
0.44
0.57 ).555 0.65
0.44
0.62 0.570
0.63
Note:
(1) FV versus OCR from Ladd and Foott (1977); see Fig. 8.4.
p = 0.85 for OC soil;
p = 1.0
for NC soil.
Table 8.8
356
I
zw
Ln
-1i J
-
rI%
a-
ZGI
W
z
0
(Z LLJ
iO
W
cn aI
W
i t.qj
K
o Z
~s
qj
I.
I
Z .Z
O0
V)
357
FIGURE 8. 1
m~
sU
a-c
OCR
SU
RECOMMENDED
avc FOR TOTAL STRESS
BY LADD (1975)
358
ANALYSES
FIGURE 8.2
zOr
Z i
t LZ
0"LLJ
`O ,
10
q
h,
1
0)
:
Z
C)
CA QC
1
v
?·
-
0IJ
C)
Or
C4
0Eo
Ld
O0
a
LL
()n
/
(
0L
0O
a.
00
C)
Ub
,U 'NO1 VA377
359
FIGURE 8.3
Al
U.5
I
Vria ion of
S.
I
fro7, UUt/C testvo
of Ambhers
/
and Northampton
0.7
i
varved c/oy.
0.6
/
I
/
/
0.5
_
Su
.
///
/
I
/
VC
/
/
0.3
u -qf from 1kC tests
zI
0.2
5/
0./
/1977), bosad o
i
I
/
£fV) fron Lodd ' Foott
z
Amherst
ond Vorthampton data
I
s~~~~~~~
3
I
I
I
4-
s
6
I---
I
7
OCR
COMPARISON OF NORMALIZED UNDRAINED SHEAR
STRENGTH DATA FROM UUV, FV AND THESIS'
RECOMMENDED UNDRAINED SHEAR STRENGTH PARAMETERS FOR
DESIGN
360
FIGURE 8.4
OCR
SHANSEP
Eu
Su
VERSUS OCR FOR CONNECTICUT VALLEY
VARVED CLAYS
361
FIGURE
8.5
0.
0.4
0.,
0.,
C
- 0. ,
OCR
RECOMMENDED INITIAL STRESS RATIO VS OCR FOR
CONNECTICUT VALLEY VARVED CLAYS FOR PLANE STRAIN
LOADING
362
FIGURE 8.6
1.0
0.8
0.6
0.4
0.2
In
ci
O
0.2
0.4
0.6
0.8
z0
0.2
0.4
0.6
0.8
t.O
.
II
k
LIU
LA
-1
k
KL
I
I
/.>
0.8
0.6
0.4
0.2
0
-0
APPLIED STRESS RATIO, q/quit
SETTLEMENT
RATIO VS APPLIED STRESS RATIO FOR
STRIP LOAD ON ISOTROPIC HOMOGENEOUS FOUNDATION
(From D'Appolonia, Poulos and Ladd, 1971)
363
FIGURE8.7
1
.I(IJ
I-
09
04
0
%l
i a
LJ "IC
s~
t
11
--
~
ooe *
-+j00
.
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0
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W t4
li
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I
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kn
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wo
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364
FIGURE 8.8
OCR
---
OCR
RECOMMENDED UNDRAINED HYPERBOLIC STRESS-STRAIN
PARAMETERS FOR USE IN FINITE ELEMENT
ANALYSES
WITH CONNECTICUT VALLEY VARVED CLAYS
365
FIGURE 8.9
0.8
0.7
0.6
qs(v)
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as
0.4
0.3
0.2
/
2
3
4
6
OCR
RECOMMENDED VERTICAL UNDRAINED STRENGTH
FOR USE IN FINITE ELEMENT ANALYSES WITH CONN.
VALLEY VARVED CLAYS IN PLANE STRAIN LOADING
366
FIGURE 8. 10
Clta
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OCR
RECOMMENDED ELLIPTICAL ANISOTROPIC UNDRAINED
STRENGTH PARAMETERS FOR USE IN FINITE ELEMENT
ANALYSES WITH CONNECTICUT VALLEY VARVED CLAYS
367
FIGURE
8.11
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FIGURE 8.12
9
SUMMARY AND CONCLUSIONS
The objectives of this thesis are:
1.
To contribute to a fundamental understanding of
the nature of undrained shear strength anisotropy
of varved clays;
2.
To present best estimates of normalized stressstrain strength parameters for undrained total
stress analyses and for predictions of pore pressures and undrained deformations used in the design of structures on Connecticut Valley varved
clays, with background information showing how
these parameters were developed; and
3.
To recommend suitable testing procedures for developing normalized soil parameters for other
varved clay deposits.
9.1
THE NATURE OF UNDRAINED SHEAR STRENGTH ANISOTROPY OF VARVED
CLAYS
The factors that control the anisotropic strength behavior
of varved clay are:
the inherent anisotropy, which includes
anisotropy due to the layered nature of varved clay (material
anisotropy) and that due to the prior K o consolidation (structural anisotropy); and when K o is not equal to one, the stress
system induced anisotropy.
The latter results frox the fact
that different increments of shear stress are required to
369
produce failure as the major principal stress varies between
the vertical and horizontal directions.
The major result of
stress system induced anisotropy is a change in the effective
stress at failure as ol varies in direction during shear.
The anisotropy in qf/avm of normally consolidated (NC)
and overconsolidated (OC) varved clays results from anisotropy
in the pore pressure response and in the normalized effective
stress envelope* (NESE) at qf.
As shown in Figure 9.1, due to
the material anisotropy component, qf(H) is larger than qf(V)
and qf from inclined loading is the smallest.
With the addi-
tional effects of structure and stress system induced anisotropy,
i.e., combined anisotropy, qf(V) is larger than qf(H) while qf
from inclined loading is still the smallest.
(Note, however,
that in Figure 9.1 only the effect of the structural anisotropy
component is shown since K o is equal to one).
The fact that qf
is smallest in inclined loading is partly due to the material
anisotropy component the failure envelope for inclined loading
being the lowest because failure takes place within a clay layer
(Curve D, Fig. 9.2) whereas for vertical and horizontal loading
failure takes place across the varves.
With combined anisotropy,
i.e., total effects of inherent and stress induced anisotropy,
qf(H) is slightly lower than qf(V) even though the NESE at qf
is much higher for horizontal loading
in Fig. 9.2).
(Curve A vs. Curve B'
The smaller qf(H) compared to qf(V) results from
significantly lower effective stresses at failure.
The anisotropy in NESE at qf shown in Fig. 9.2 when it is
*The plot of q/vm
versus 370
Pvm.
370
not equal to that at (
3)max' is due to both inherent aniso-
tropy and to variation in the degree of mobilization of friction and cohesion during shear.
Because of the latter, the
NESE at qf for vertical loading (Curve
than that at (1/
',
Fig. 9.2) is lower
3)max' i.e., Curve A in Fig. 9.2, which holds
for both vertical and horizontal loading.
The anisotropy in developed pore pressures is the result
of anisotropy in the pore pressure parameter A due to inherent
anisotropy (see Fig. 7.9) and differences in the magnitude of
Aq due to the stress system induced component.
The anisotropy in NESE at (a1/a3 )max is found to be due to
material anisotropy.
Thus, there are three possible values of
NESE at ( l/a
1
3 )max for varved clay which depend only on the
shearing direction., i.e., across the varves or parallel to
the varves.
These are: (l) for failure across the varves (Fig.
9.2, Curve A);
(2) for failure within
a silt layer
(
> 240
based on CD direct shear tests); and (3) for failure within a
clay layer (Figure 9.2, Curve E), the latter two occurring when
shearing is parallel to the varves.
Based on several types of strength tests (see Chapter 5),
the NESE at (1l/3)max
for failure across the varves (Curve A
in Fig. 9.2 and Fig. 5.26) and for failure within a clay layer
(Curve E in Fig. 9.2 and Fig. 5.31) are considered to be "basic
material" properties since they were shown to be independent of:
1.
The stress system during consolidation;
2.
The stress system during shear;
371
9.2
3.
The magnitude of ov;
4.
The drainage conditions; and
5.
The OCR of the samples.
DEVELOPMENT OF NORMALIZED STRESS-STRAIN-STRENGTH PARAMETERS
FOR USE IN THE DESIGN OF STRUCTURES ON CONNECTICUT VALLEY
VARVED CLAYS
Best estimates of normalized stress-strain-strength para-
meters for use in the design of structures on Connecticut Valley
varved clays, and recommended testing methods for checking
these design parameters, were presented in Chapter 8.
The back-
ground for developing the normalized soil parameters for Connecticut Valley varved clays is as follows:
1.
Anisotropic S u parameters for plane strain loading-Tff at the qf condition from SHANSEP Ko consolidated
plane strain compression, direct simple shear, and
plane strain extension (CKoUPSC, CKoUDSS, CKoUPSE)
tests corrected for strain compatibility (Section
8.2.2 are recommended as design values
(Fig. 9.3) for
Morgenstern and Price undrained total stress analyses.
The reasons for considering these values to be best
estimates are:
a.
K o consolidated plane strain and direct simple
shear tests simulate representative in situ
modes of fialure, i.e., the PSC, DSS, and PSE
elements shown in Fig. 9.3.
b.
Values of qf/vc
from several types of SHANSEP
372
strength tests (e.g., CKoUC, CKoUDSS, and
CKoUE tests) have been shown to be the same
for three locations (Amherst, Northampton,
and East Windsor) within the Connecticut
Valley (e.g., Fig. 4.14).
c. On the basis of CKoU triaxial data on block
samples, the SHANSEP method of consolidation
is thought to yield reasonable estimates of NC
and OC in situ qf values, especially at low
OCR, though it causes more heavily OC samples
to behave in a less brittle and sensitive manner compared to the in situ soil (see Fig. 9.4
and 6.5).
Note that in Fig. 6.5 the lower qf
value from the RECOMPRESSION*
CKoUE OC (OCR=4
the in situ OCR) sample compared to that from
the SHANSEP sample probably results mainly from
disturbance due to separation of varves, though
when
lf acts perpendicular to the varves or
when the effect of disturbance due to separation
of varves is small, it is thought that OC
RE-
COMPRESSION data on block samples give the best
indication of the in situ soil behavior.
d. In order to reflect the in situ behavior, and
since the effective stress rather than total
*RECOMPRESSION--the method of consolidation whereby the
sample is consolidated of the desired OCR with the vertical consolidation stress that is less than the in situ maximum past
pressure.
373
stresses control the soil strength behavior,
the undrained shear strength is defined as
Tff rather than qf.
NOTE:
In Fig. 9.3, the Su values for NC ver-
tical and horizontal loading are based on plane
strain ccmpression(PSC) and plane strain extension (PSE) tests.
However, the variation in
strength with OCR was obtained from trends
measured in triaxial compression and extension
test (see
2.
Fig. 5.23 and Section 8.2.3).
The recommended values of pre-failure pore pressure
parameter A (Fig. 8.8) were based upon SHANSEP plane
strain loading tests for NC soil (since they are believed to yield better results than RECOMPRESSION
tests) and on estimated RECOMPRESSION tests data for
OC soil since RECOMPRESSION tests on block samples are
thought to yield realistic pre-failure stress-strain
pore pressure behavior for overconsolidated deposits
(Figs. 9.4 and 6.5).
The other recommended normalized soil parameters plotted
versus OCR are Eu/Su (Fig. 8.5), the initial shear stress ratio,
f, (Fig. 8.6) and hyperbolic-.stres-strain parameters (Gi and
Rf (Fig. 8.9).
9.3
RECOMMENDED TESTING PROCEDURES FOR DEVELOPING NORMALIZED
SOIL PARAMETERS FOR OTHER VARVED CLAY DEPOSITS
For undrained total stress stability analyses, SHANSEP Ko
374
consolidated undrained triaxial compression, direct simple shear
and triaxial extension (CKoUC, CKoUDSS and CKoUE) tests were
recommended for obtaining anisotropic plane strain S u values,
with an adjustment to converttriaxial Su to plane strain S u
values.
(Note that use of triaxial tests will lead to lower
strength thus yielding somewhat conservative results for field
situations involving plane strain loading (see Section 8.3.3)).
The adjustment can be done by assuming that the ratios Su(V)/qf
from CKoUC tests and Su(H)/qf from CKoUE tests for the Connecticut Valley varved clay also hold for other deposits.
for a given mode of failure
That is,
(PSC or PSE), the design:
triaxialq measured for other depsit)x
design
value for Conn. Valley
triaxial qf (measured frm Cnn. Valley deposit)
%7-
Su values for PSC and PSE failure modes for Connecticut Valley
varved clays are presented in Fig. 9.3.
qf values from CKoUC
and CKoUE tests on Connecticut Valley varved clays are presented
in Fig. 4.14.
Thmax from the CKoUDSS test can be used directly
as S u value for DSS mode.
Due to distrubance from structural damage, especially due
to separation of varves, UUC tests and RECOMPRESSION CIUC tests
on samples with different inclinations are not recommended (see
Chapter 6).
SHANSEP CIUC tests on samples with different in-
clinations are not recommended because they yield unconservative qf values for plane strain horizontal and inclined loadings
(see Chapter
5).
In cases where predictions of pore pressures and deformations
375
are important, in OC deposits, RECOMPRESSION CK0 UC, CKoUDSS
and CKoUE tests are recommended, rather than the SHANSEP method
if the samples are of good quality.
376
I
l..
qf ('3)
0.4
0.,
o.
/O° Angle Between crf
MATERIAL
'
YerticO/ DiectIn)
AND COMBINED ANISOTROPY
CONNECTICUT VALLEY
VARVED
377
CLAYS
IN
qf
OF
AT OCR = 4
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378
FIGURE 9.2
0.
0.
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RECOMMENDED
3
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4.
5
6
7
8 9
ANISOTROPIC UNDRAINED STRENGTH
PARAMETERS FOR USE IN TOTAL STRESS MORGENSTERNPRICE STABILITY
ANALYSES WITH CONNECTICUT VALLEY
VARVED CLAYS
379
FIGURE 9.3
.j
CK
o 1C test (t5 )
-
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SHA
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EFFECT OF METHOD OF CONSOLIDATION ON VERTICAL AND
HORIZONTAL o--STRENGTH
BEHAVIOR FROM CKJJC AND
CJOUE TESTS ON CONNECTICUT VALLEY VARVED CLAYS AT
OCR = 2
380
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SECTION D.2-4
TABULATION OF TEST DATA FROM CKoUDSS TESTS ON
VERTICAL EAST WINDSOR SAMPLES
493
Sheet I of 2
SHEAR TEST
DIRECT - SIMPLE
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OF TECHNOLOGY
494
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Sheet 2 of 2
SHEAR (continued)
DIRECT - SIMPLE
PROJECT 8012
TIME
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ENGINEERING
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SHEAR (continued)
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REMARKS:
LABORATORY
MECHANICS
SOIL
DEPT. OF CIVIL ENGINEERING
OF TECHNOLOGY
INSTITUTE
MASSACHUSETTS
497
_7
Sheet I of 2
SHEAR TEST
DIRECT - SIMPLE
D65-3
PROJECT 804z
TYPE
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TESTED BY '3L
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ENGINEERING
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MASSACHUSETTS
INSTITUTE
498
REMARKS:
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INSTITUTE
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TECHNOLOGY
499
SECTION D.2-5
TABULATION OF TEST DATA FROM CKoUDSS TESTS ON
INCLINED EAST WINDSOR SAMPLES
500
10
Note:
REFERENCES
ASCE
ASTM
= American Society of Civil Engineers
= American Society of Testing and
Materials
JGED
= Journal of Geotechnical Engineering
Division
JSMFD = Journal of the Soil Mechanics and
Foundation Division
ICSMFE = International Conference on Soil
Mechanics and Foundation Engineering
Amerasinghe, S.F. and Perry, R.H.G (1975), "The Anisotropy in
Heavily Overconsolidated Kaolin," JGED, ASCE, Vol. 100,
No. GT 12, pp. 1277-1291.
Bailey, W.A. and Christian, J.T. (1969), "ICES-LEASE-I, A Problem Oriented Language for Slope Stability Analyses,"
Res. Report R-69-22, Dept. of Civil Eng., MIT, 56 p.
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386
BIOGRAPHY
Surachat Sambhandharaksa was born to Mr. Chanai and Mrs.
Sabha Sambhandharaksa on Feb. 24, 1944 in Bangkok, Thailand.
He attended high school in Bangkok, Thailand.
In 1962, he got
a Colombo plan scholarship to study at the University of New
South Wales in Sydney, Australia and in 1967 was awarded the
degree of Bachelor of Engineering.
While an undergradaute, he
was employed for two 4-month periods by the Department of Main
Roads in New South Wales, Australia.
The writer went to work as an instructor at Chulalongkorn
University for one year, before continuing his study under a
scholarship from the Asian Institute of Technology where in 1970
he was awarded the Master of Engineering.
He again lectured at
Chulalongkorn University.
In 1971, he and Dr, Za-Chieh Moh, his thesis advisor at
the Asian Institute of Technology, wrote a paper entitled, "Effects of Salt Content on Thixotropic Behavior of a Compacted
Clay,"~ published in the Proceedings of the First Australia-New
Zealand Conference on Geomechanics, Melbourne, 1971.
On leave of absence from Chulalongkorn University since
1971, he went to study at Northwestern University for a year
before transferring to MIT to study for his Doctor of Science.
387
APPENDIX A
Prefix
indicates a change
Suffix f indicates a final or failure condition
A bar over a stress indicates an effective stress
A bar over a property indicates value in terms of effective stress.
INDEX AND CLASSIFICATION PROPERTIES
eo
-
Initial void ratio
Gs
-
Specific gravity
LI
-
Liquidity index
PI
-
Plasticity index
w
-
Water content
WL
-
Liquid limit
WN
-
Natural water content
W
-
Plastic limit
P
Bouyant unit weight
Y~b
Yt
¥w
-
Total unit weight
-
-w
~
Unit weight of water
STRESSES, STRAINS, MODULUS, AND STRENGTH PATRAMTERS
a
a
a/-a
-
Pore pressure parameter (Eq. 2.9)
-
Intercept of qf vs.
-
f failure envelope
Normalized intercept of qf/avm vs. Pf/vm
envelope
388
failure
A
-
Skempton's pore pressure parameter A
Af(V)
-
Skempton's pore pressure parameter A at failure
for vertical loading
Af(PSC)
-
Skempton's pore pressure parameter A at failure
from K consolidated plane strain compression
test
Af(H)
-
Skempton's pore pressure parameter A at failure
for horizontal loading
Af(PSE)
-
Skempton's pore pressure parameter A at failure
from K o consolidated plane strain extension test
Af(45 °)
-
Skempton's pore pressure parameter A at failure
for inclined (45° ) loading
Af(DSS)
-
Skempton's pore pressure parameter A at failure
from Ko consolidated direct simple shear test
B
-
Skempton's pore pressure parameter B = Au/ac
b/a
-
Axes of Davis-Christian (1971) elliptical strength
relationship
c,
-
Intercept of Mohr-Coulomb failure envelope
0~~~~~~~~~~
c/a,c
vm
6vi
Normalized intercept of normalized Mohr-Coulomb
failure envelope
E, E
-
Young's modulus
Eu
-
Undrained secant E
EU (H)
Eu (V)
-
Undrained secant E from horizontal loading
-
Undrained secant E from vertical loading
E u (45°)
Undrained secant E from inclined (45°) loading
f
-
Initial shear stress ratio = (-Ko)/[2Su(V)/Avo)
0 u ~vo
G
-
Shear Modulus
Gi
-
Initial G for hyperbolic stress-strain model
Gy
-
Yielded G for hyperbolic stress-strain model
K
-
Bulk modulus
Gy
Kc
c/ a
~vch~c
389
Ko
Coefficient of earth pressure at rest
-
Ks
Anisotropic strength ratio = qf(H)/qf(V)
NESE
Normalized effective stress envelope, the plot
in terms of qf/avm versus f/;vm
NESE
-
Normalized effective stress envelope, the plot
in terms of aff/vm versus Tff/avm
OCR
-
Overconsolidation ratio = avm/avo, avm(L)/avc
0.5(o
P, P
+ aoh) 1/ 2 (
+ ah)
- 0.5(av - oh) or 1/2 (
q
- a3)
qf
-
q at failure
qf(V)
-
qf from vertical loading
qf(H)
-
qf from horizontal loading
qf(4 5 °)
-
qf from inclined (45'°) loading
Parameter for hyperbolic stress strain model
R.
Rf
St
-
-
Su(V)
Su(45°)
Su (undisturbed)/Su (remolded)
Undrained shear strength = qf or
Su
Su (H)
Sensitivity
Su from horizontal loading
Su from vertical loading
Su from inclined (45° loading
-
-
Su (DSS)
Tff = qfcos~
Su from direct simple shear test
from plane strain compression test
Su (PSC)
-
S
Su (PSE)
-
S u from plane strain
SCF
-
Strain compatibility factor (see Table 8.4)
SR
-
Strength ratio (see Fig. 5.24)
t
-
Time at failure
f
u
extension
Pore water pressure
-
u
m
ud
u for undrained condition
u from field measurement
Udn
390
test
-
u at undrained failure
a5
-
Slope of qf vs. pf failure envelope
a
-
Angle between the failure plane and the horizontal plane
6
-
Angle between the normal to the failure plane
and the major principal stress axis
B
-
Angle between
-
Varve inclination
-
Shear strain
-
Linear strain
-
Axial strain
Uu
e
Y,
f
C
a
a
lf and vertical direction
sv
Vertical strain
Svol
Volumetric strain
Field vane correction factor
V, v
-
Poisson's ratio
a, a
-
Normal total stress, normal effective stress
°1'
2' :3
(a 1/3)max
c
ac
a
c
Principal stress
Maximum obliquity
-
Confining pressure (isotropic)
-
Consolidation pressure (isotropic)
Offi aff-
Normal stress on failure plane at failure
ah
-
Horizontal normal stress
ah
-
Norizontal consolidation stress
-
Octahedral stress
a
Ps
as
-
Isotropic stress of perfect sample
-
Pre-shear effective stress of UUC sample
v
-
Vertical consolidation stress
-
Final vertical effective stress
~hc
0 oct
5
avf
391
-
vo
Initial vertical effective stress
a-
Maximum past pressure
aVm (L)
Laboratory maximum past pressure
vm
Ovm(L
a
-
Normal stress on a plane with inclinations to
the horizontal plane
-
Initial
a
,a
a
ao
a0
~~T
-Shear
stress
-T
T
on failure plane at failure
ff
Th
-
Maximum
Thm
Thmax
T
oct
Ta
T on horizontal plane (direct-simple shear test)
Th
-
Octahedral shear stress
-
T
on a plane with inclination
plane
T
Initial T
Tf
T
¢, ~T
to the horizontal
at failure
Slope of Mohr-Coulomb or normalized Mohr-Coulomb
failure envelope
CONSOLIDATION, STABILITY AND SETTLEMENT
CONSOLIDATION, STABILITY AND SETTLEMENT
B
-
Width of loaded area
Coefficient of consolidation for vertical flow
Cv
CR
-
Virgin compression ratio
FS
-
Factor of Safety
H
-
Depth of foundation clay or height of clay layer
I_
Influence factor for elastic settlement (Eq. 8.1)
P
PC-
Quasi-consolidation pressure
RR
-
Recompression ratio
SR
-
Settlement ratio = P
-Pe
392
1
time
t
t required for primary consolidation
t
p
Degree of consolidation
U
Average degree of consolidation
UA
Anglebetween failure surface and horizontal
direction for plane strain compression failure
surface
A
Angle between failure surface and horizontal direction for plane strain extension failure surface
0
p
"Elastic" settlement
Pe
Initial settlement
Pi
CONSOLIDATION AND STRENGTH TESTS
CD
Consolidated-drained shear test
CDDS
-
CD direct shear test
CRSC
-
Constant rate of strain consolidation test
CU
Consolidated-undrained shear test
CIU
Isotropically consolidated-undrained shear test
CIUC
CIU triaxial compression test
CIUE
-
Ko consolidated-undrained shear test
CK U
0o
CKoUC
CIU triaxial extension test
-
CKoU triaxial compression test
CKoUDSS
CKoU direct simple shear test
CKoUDSS (C)
CK U direct simple shear test run on inclined
sample and sheared by increasing shear stress
CK UDSS(E)
CK U direct simple shear test run on inclined
sample and sheared by decreasing shear stress
CKoUE
CKoUE
-
CKoU triaxial extension test
393
CKoUPSC
CKoU plane strain compression test
CKoUPSE
CKoU plane strain extension test
DSS
Direct-simple shear
FV
Field vane test
PSC
Plane strain compression (alf = Oavf)
PSE
Plane strain extension
TC
Triaxial compression
TE
Triaxial extension
UC
Unconfined compression
UU
Unconsolidated-undrained shear test
UUC
UU triaxial compression test.
(lf
=
hf)
MISCELLANEOUS
CVVC
Connecticut Valley varved clays
FV
Field vane
M-C
Mohr-Coulomb
M-P
Morgenstern and Price
MPPR
Maximum Past Pressure Ratio =
MOC
Method of Consolidation
NC
Normally consolidated
OC
Over consolidated
PS
Plane strain
a
394
(L)
vm
APPENDIX B
BEST ESTIMATES* OF STRESS-STRAIN-STRENGTH DATA
OF CONNECTICUT VALLEY VARVED CLAYS
Summary of best estimates of test data from SHANSEP CKoUC,
CK0 UE and CKoUDSS tests are presented in Tables B-1 through
B-3.
Figures B-1 through B-6 show best estimates of normalized
stress-strain curves and normalized stress paths from CKoUC,
CKoUE and CKoUDSS tests on NC and OC samples.
The normalized
stress-strain curves from CKoUPSC and CKUPSE tests on NC samples
0
~0
are shown in Figs. 5.20 amd 5.21.
The following are the lists
of tables and figures
LIST OF TABLES
B-1
Best Estimates of Test Data from CKoUC Tests on NC
and OC Samples from the Connecticut Valley
B-2
Best Estimates of Test Data from CKoUDSS Tests on
NC and OC Samples from the Connecticut Valley
B-3
Best Estimates of Test Data from CKoUE Tests on NC
and OC Samples from the Connecticut Valley
LIST OF FIGURES
B-1
Normalized Stress-Strain Curves from CKOU Triaxial
Compression and Extension Tests on Normally Consolidated Connecticut Valley Varved Clays
B-2
Normalized Stress-Strain Curves from CKoU Triaxial
Compression and Extension Tests on Overconsolidated
Connecticut Valley Varved Clays: OCR=2
*Determined from average data of samples from Amherst, Northampton and East Windsor.
395
LIST OF FIGURES
(Continued)
B-3
Normalized Stress-Strain Curves from CK 0 U Triaxial
Compression and Extension Tests on Overconsolidated
Connecticut Valley Varved Clays: OCR=4
B-4
Normalized Stress Paths from CKoU Triaxial Compression and Extension Tests on Connecticut Valley Varved
Clays
B-5
Normalized Stress-Strain Curves from CKoUDSS Tests
on Normally Consolidated and Overconsolidated Connecticut Valley Varved Clays
B-6
Normalized Stress Paths from CKoUDSS Tests on Connecticut Valley Varved Clays
396
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CONNECTICUT VALLEY VARVED CLAYS
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14
COMPRESSION
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VARVED CLAYS: OCR=4
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APPENDIX C
SUMMARY OF TEST RESULTS FROM CONNECTICUT VALLEY AND
HACKENSACK VALLEY VARVED CLAYS
Appendix C contains a summary of test results from UUC tests
,CIU and
CKOU tests on NC and OC samples from Amherst, East
Windsor, Northampton and Hackensack Valley with details of the
testing conditions.
The following are the lists of tables
which are presented in this appendix.
SUMMARY OF TEST RESULTS FROM AMHERST VARVED CLAY
C.1
LIST OF TABLES
C.l-l
Locations of Amherst Samples Used for Undrained
Strength Tests
C.1-2
Summary of UUC Tests on Amherst Samples
C.1-3
Summary of CIUC Tests on NC Amherst Samples at DifUE
ferent Inclinations (e=0, 45, and 90) and
Tests on NC Vertical and Horizontal Samples
C.1-4
Summary of CKoUC Tests on NC and OC Amherst Samples
C.1-5
Summary of CKoUE Tests on OC Amherst Samples
C.1-6
Summary of CKoUDSS Tests on NC and OC Amherst Samples
SUMMARY OF TESTS RESULTS FROM EAST WINDSOR VARVED CLAY
C.2
LIST OF TABLES
C.2-1
Summary of UUC Tests on East Windsor Samples at Different Inclinations
C.2-2a
Summary of CIUC Tests on NC and OC SHANSEP East Windsor Samples at Different Inclinations
406
C.2.2b
Summary of CIUC Tests on OC SHANSEP East Windsor
Samples at Different Inclinations
C.2-3
Summary of CIUE Tests on NC and OC SHANSEP East Windsor (=0 and 90) Samples
C.2-4
Summary of CKoUC Tests on NC and OC SHANSEP East Windsor Vertical Samples
C.2-5
Summary of CKoUE Tests on NC and OC SHANSEP East Windsor Vertical Samples
C.2-6
Summary of CKoUC, CKoUE, and CIUC (=0, 45, 90) Tests
on OC RECOMPRESSION East Windsor Samples
C.2-7
Summary of CKoUDSS Tests on NC and OC East Windsor
SHANSEP Samples
C.2-8
Summary of CKoUDSS Tests on NC SHANSEP Inclined East
Windsor Samples
SUMMARY OF TEST RESULTS FROM NORTHAMPTON VARVED CLAY
C.3
LIST OF TABLES
C.3-1
Summary of CKoUC, CKoUPSC, and CKoUPSE Tests on NC
Northampton Varved Cay
C.3-2
Summary of CKOUDSS Tests on NC and OC SHANSEP Northampton Varved Clay
C.3-3
Summary of CDDS Tests on NC and OC Northampton
Varved Clay
SUMMARY OF TEST RESULTS FROM HACKENSACK VALLEY VARVED CLAY
C.4
NOTE: Additional data from Hackensack Valley varved clay can
be obtained from Saxena et al. (1974).
LIST OF TABLES
C.4-1
Summary of CIUC, CKoUC Tests on NC and of CKoUDSS
Tests on NC and OC SHANSEP Hackensack Valley Varved
Clay Samples
407
SUMMARY OF TEST DATA FROM CIU AND CKoU TESTS
ON SHANSEP AND RECOMPRESSION AMHERST
SAMPLES
408
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APPENDIX D
TABULATION OF TEST DATA
Consolidated undrained strength test data from Amherst and
East Windsor varved clays are tabulated in Sections D.1 and D.2.
Strength data from Northampton varved clay were presented in
Lacasse et al. (1972). Saxena et al. (1974) presented the test
data from Hackensack Valley varved clay.
The organization of
the tabulated test data is as follows:
Section D.1-l:
Tabulation of Test Data from CIUC
and CIUE Tests on NC SHANSEP Amherst
Samples
Section D.1-2:
Tabulation of Test Data from CK UC
and CKoUE Tests on NC and OC SHINSEP
Amherst Samples
Section D.1-3:
Tabulation of Test Data from CKoUDSS
Tests on NC and OC SHANSEP AND RECOMPRESSION Amherst Samples
Section D.2-1:
Tabulation of Test Data from CIUC and
CIUE Tests on NC and OC SHANSEP East
Windsor Samples
Section D.2-2:
Tabulation of Test Data from CKoUC and
CK0 UE Tests on NC and OC SHANSEP East
Windsor Samples
Section D.2-3:
Tabulation of Test Data from CKoUC,
CIUC, and CKoUE Tests on RECOMPRESSION
East Windsor Samples
Section D.2-4:
Tabulation of Test Data from CK UDSS
Tests on Vertical East Windsor amples
Section D.2-5:
Tabulation of Test Data from CKoUDSS
Tests on Inclined East Windsor Samples
431
SECTION D.1-1
TABULATION OF TEST DATA FROM CIUC AND CIUE TESTS ON
NC SHANSEP AMHERST SAMPLES
432
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SECTION D.1-2
TABULATION OF TEST DATA FROM CK UC AND CK UE TESTS
ON NCOC
ANDSHANSEP AMHERST SAMPLES
ON NC AND OC SHANSEP
438
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I 4455
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