to m.e. and marta
TRANSCRIPT
A METHOD
PARTIALLY
TO PREDICT DEFORMATIONS FOR
DRAINED CONDITIONS IN BRACED
EXCAVATIONS
by
Theodore von Rosenvinge IV
B.S. Eng. Northeastern University
(1978)
Submitted in part
of the requirements
Master of
ial fulfillment
for the degree of
Science
at the
Massachusetts Institute of
(August 1980)
Q Theodore von Rosenvinge
Technology
IV
Signature of Author
Department of Civil Engineering, August 14, 1980
Certified by
Thesis Supervisor
Accepted by
Chairman, Departmental Committee on Graduate
Students of the Department of Civil Engineering.
MASSACHUSETTS INSTIIUT:OF TECH ,VOG,
OC T 9 1980
2
A METHOD TO PREDICT DEFORMATONS FOR
PARTIALLY DRAINED CONDITIONS IN
BRACED EXCAVATIONS
by
Theodore von Rosenvinge IV
Submitted to the Department of Civil Engineeringon August 14, 1980 in partial fulfillment of the
requirements for the degree of Master of Science inCivil Engineering
ABSTRACT
The purpose of this thesis is to develop a method topredict deformations in braced excavations, which includesthe effects of excess pore pressure dissipation. The timevariables considered are time for construction compared withtime required for pore pressure equilization due to alteredflow conditions and shear. The method will apply the StressPath Method to determine appropriate soil parameters to usein the existing finite element program, BRACE III in a waythat accounts for time effects.
The method will be applied to an actual case todemonstrate its applicability and illustrate its use.
Thesis Supervisor: Dr. William Allen Marr
Title: Research Associate
3
To M.E. and Marta
4
ACKNOWLEDGMENTS
The author is grateful to Dr. William Allen Marr for
his interest and guidance during the work on this thesis. The
author is also grateful to Professor T.W. Lambe for his support,
providing information and soil specimens tested for the case
study.
The author also wishes to acknowledge the generous advice
and discussion provided by Dr. Walter E. Jaworski of Northeast-
ern University during the early stages of this thesis.
Special thanks is due to my wife, M.E., and young daughter,
Marta for their patience during my education. M.E.'s tireless
hours on the word processor are gratefully acknowledged. The
sustaining humor was supplied by Marta's often welcome distrac-
tion while her father was working on his "thethiss". And a
note of thanks to my entire family for their continued support
and encouragement during the last several years.
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TABLE OF CONTENTS
Page
TITLE PAGE
ABSTRACT
DEDICATION
ACKNOWLEDGEMENTS
TABLE OF CONTENTS
LIST OF TABLES
LIST OF FIGURES
CHAPTER ONE INTRODUCTION
1.0 INTRODUCTION
1.1 THESIS OBJECTIVE
1.2 THESIS SCOPE
CHAPTER TWO THE ROLE OF TIME
2.0 INTRODUCTION
2.1 STRESS PATHS
2.1.1 Undrained Shear
2.1.2 Dissipation of Pore Pressure Following
Shear
2.1.3 Stress Paths - Summary
2.2 THE PERFORMANCE OF BRACED EXCAVATIONS VS. TIME
2.2.1 Braced Excavations in Cohesionless Sand
2.2.2 Piping, Suffosion
2.2.3 Effects of Frost Action
2.2.4 Braced Excavations In Normally
Consolidated Clay
1
2
3
4
5
9
10
14
19
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21
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25
25
25
26
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Page
CHAP
CHAP
2.2.5 Braced Excavations in Overconsolidated
Clay
2.2.6 Braced Excavation in Stiff Fissured Clay
2.2.7 Braced Excavation in Silt
2.3 METHODS OF ANALYSIS AND DESIGN
2.3.1 Semi - Empirical Methods
2.3.2 Finite Element Analysis
2.3.3 The Stress Path Method
2.4 CONCLUSIONS
TER THREE PROPOSED PRECEDURE FOR INCORPORATING TIME
IN THE PREDICTION OF THE PERFORMANCE OF A
BRACED EXCAVATION
3.0 INTRODUCTION
3.1 OUTLINE OF THE PROPOSED METHOD
TER FOUR CASE STUDY - PREDICTION OF THE PERFORMANCE
OF A BRACED EXCAVATION
4.0 INTRODUCTION
4.1 TYPE OF PREDICTION
4.2 PROJECT DESCRIPTION
4.2.1 General
4.2.2 Construction Schedule
4.3 SUBSURFACE CONDITIONS
4.3.1 Soil Conditions
4.3.2 Groundwater
4.4 NAKAGAWA PREDICTION
4.4.1 Initial Model of Field Situation
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46
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84
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Page
4.4.2 Stress Paths from Initial BRACE III
Analysis 87
4.4.3 Laboratory Tests and Revised Soil
Profile 89
A. Soil Classification 89
B. Stress Path Tests 90
C. Coefficient of Consolidation 91
D. Revised Soil Profile 91
E. Contours of Equal Vertical Strain 95
4.4.4 Revised BRACE III Analysis for Partially
Drained Conditions 95
A. General 95
B. Computation of Partially Drained
Moduli 96
1. Undrained Shear Strains 96
2. Consolidation Strains Due to
Dissipation of Excess Pore Pressure
Within the Excavation 97
3. Consolidation Strains Due to
Dissipation of Excess Pore Pressure
Behind the Excavation 98
4. Contours of Moduli 100
5. Ko 102
6. Anisotropy 102
C. Predicted Loads, Deformations and
Stab i 1 ity 103
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1. General
2. Deformations
3. Loads
4. Stability
CHAPTER FIVE SUMMARY AND CONCLUSIONS
5.0 GENERAL
Page
103
103
106
107
152
153
154
156
161
163
5.1 LIMITATIONS
5.2 SUGGESTIONSFORFURTHERRESEARCH
5.3 CONCLUSIONS
LIST OF SYMBOLS
REFERENCES
APPENDIX A GRAINSIZE PLOTS AND SQUARE ROOT OF TIME
CONSOLIDATION PLOTS - NAKAGAWA SOIL
SPECIMENS 167
APPENDIX B VERTICAL EFFECTIVE STRESS VERSUS DEPTH- 182
NAKAGAWA
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LIST OF TABLES
Table No.
3.1
3.2
4.1
4.2
4.3
4.4
4.5
Title
GEOTECHNICAL PERFORMANCE
STRESS PATH METHOD
SOIL INDEX PROPERTIES, AND CLASSIFICATION
SUMMARY OF COEFFICIENT OF CONSOLIDATIONDATA FROM STRESS PATH TESTS
SUMMARY OF SHEAR STRENGTH DATA
SUMMARY OF DETERMINATION OF PARTIALLY
DRAINED STRESS STRAIN MODULI
SUMMARY OF STRENGTHS FOR AVERAGE ELEMENTS
IN STABILITY ANALYSIS
Page
81
82
115
121
124
129
151
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LIST OF FIGURES
Figure No. Title Page
1.1 EFFECT OF FORCED DELAY ON A CRITICALCONSTRUCTION EVENT 20
2.1.1 EXAMPLE BRACED EXCAVATION GEOMETRY 48
2.1.2 STRESS PATHS FOR UNDRAINED UNLOADING OF
NORMALLY AND OVER CONSOLIDATED SOIL
ELEMENTS 49
2.1.3a STRESS PATHS FOR UNDRAINED UNLOADING
FOLLOWED BY PORE PRESSURE DISSIPATION FOR
NORMALLY CONSOLIDATED SOIL-NO DEWATERING 50
2.1.3b STRESS PATHS FOR UNDRAINED UNLOADING
FOLLOWED BY PORE PRESSURE DISSIPATION FOR
NORMALLY CONSOLIDATED SOIL-WITH DEWATERING 51
2.1.3c STRESS PATHS FOR UNDRAINED UNLOADING
FOLLOWED BY PORE PRESSURE DISSIPATION FOR
OVER CONSOLIDATED SOIL-NO DEWATERING 52
2.1.3d STRESS PATHS FOR UNDRAINED UNLOADING
FOLLOWED BY PORE PRESSURE DISSIPATION
FOR OVERCONSOLIDATED SOIL-WITH DEWATERING 53
2.1.4 DETERMINATION OF TOTAL HEAD SUBJECT TO
EXCAVATION AND/OR DEWATERING 54
2.1.5 STRESS PATHS FOR ACTIVE AND PASSIVE
CONDITIONS FOR NORMALLY AND
OVERCONSOLIDATED SOIL 55
2.2.1 STRESS PATHS FOR EXCAVATION IN COHESIONLESS
SAND 56
2.2.2 PIPING/SUFFOSION MECHANISMS 57
2.2.3 INCREASE IN STRUT LOAD VS. TIME FOR A
BRACED CUT IN STIFF FISSURED CLAY 58
2.2.4 VARIATION IN COEFFICIENT OF CONSOLIDATION 59
2.2.5 SOIL PROFILE, EXCAVATION, ANDINSTRUMENTATION DETAILS FOR A BRACED SLURRY
WALL IN SOFT CLAY, OSLO 60
2.2.6 SUMMARY OF OBSERVATIONS FOR BRACED CUT IN
SOFT CLAY, OSLO 61
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Page2.2.7 STRESS DATA FOR PIEZOMETER P-i, BELOW THE
CAES FOUNDATION EXCAVATION 62
2.2.8 SOUTH COVE SUBWAY EXCAVATION TEST SECTIONPIEZOMETER P-4 63
2.2.9 STRUT LOAD AND PORE WATER PRESSURE VS. TIME-SOUTH COVE TEST SECTION 64
2.2.10 MEASURED AND PREDICTED STEADY STATE POREPRESSURES-SOUTH COVE TEST SECTION 65
2.2.11 TIME DEPENDENT BEHAVIOR OF TEST CUT INSTIFF FISSURED CLAY 66
2.2.12 EFFECT OF VARIATION OF TIME TO FAILURE ONSTRENGTH IN UNDRAINED TRIAXIAL TESTS 67
2.2.13 EXCAVATION PENETRATING SILT: MBTA TESTSECTIONS A & B 68
2.2.14 MEASURED AND PREDICTED PORE PRESSUREMBTA TEST SECTIONS A AND B 69
2.3.la TERZAGHI & PECK DESIGN ENVELOPES 70
2.3.1b TERZAGHI & PECK DESIGN ENVELOPES 71
2.3.2 CONSOLIDATION AT THE END OF ONE-DIMENSIONALEXCAVATION 72
2.3.3a CONTOURS OF NEGATIVE EXCESS PORE PRESSUREFOR TWO-DIMENSIONAL EXCAVATION 73
2.3.3b CONTOURS OF NEGATIVE EXCESS PORE PRESSUREFOR TWO-DIMENSIONAL EXCAVATION 74
3.1 CONTOURS OF EQUAL VERTICAL STRAIN 83
4.2.1 PLAN VIEW AND PROFILE OF NAKAGAWA SEWAGETREATMENT PLANT EXCAVATION - SITE B 110
4.2.2 SITE B CONSTRUCTION SCHEDULE 111
4.4.1 INITIAL PROFILE AND FINITE ELEMENT MESH 112
4.4.2 DETAIL OF FINITE ELEMENT MESH IN ZONE OFEXCAVATION 113
4.4.3 TYPICAL TOTAL STRESS PATHS INITIAL BRACE IIIANALYSIS LINEARLY ELASTIC MODULUS 114
12
PLASTICITY CHART
STRESS PATH TEST TC-1
STRESS PATH TEST TC-2
STRESS PATH TEST TE-1
STRESS PATH TEST TE-2
REVISED SOIL PROFILE
4.4.4
4.4. 5a
4.4. 5b
4.4. 5c
4.4. 5d
4.4.6
4.4.7
4.4. 8a
4.4 .8b
4.4. 8c
4.4.9
4.4.10
4.4.11
4.4. 12a
4.4. 12b
4.4.13
4.4. 14a
4.4. 14b
4.4.15
4.4.16
4.4.17
TEST RESULTS
STRAIN -
STRAIN -
STRAIN -
SIXTH EXCAVATION STAGE SIMPLIFIED PROFILE 128
SUPERPOSITION OF TOTAL STRESS PATHS ON
CONTOUR PLOT OF EQUAL STRAIN 131
DEPTH OF EXCAVATION VS. TIME 132
TIME FACTOR VS. AVERAGE DEGREE OF
CONSOLIDATION 133
NORMALIZED DEPDTH VS. CONSOLIDATION RATIO 133
NORMALIZED DEPTH VS. CONSOLIDATION RATIO FORTRIANGULAR INITIAL EXCESS PORE PRESSURE 134
VOLUMETRIC STRAIN VS DISSIPATION OF POREPRESSURE 135
VOLUMETRIC STRAIN VS. LOGARITHMIC DISSIPATIONOF PORE PRESSURE 1
CONTOURS OF MODULI - BRACE III ANALYSIS -REVISED RUN B
CONTOURS OF MODULI - BRACE III ANALYSIS -FINAL RUN C
DEFORMED FINITE ELEMENT MESH
Page116
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126
127
UNDRAINED STRENGTH TRIAXIALVERSUS SAMPLE DEPTH
ESTIMATED CONTOURS OF EQUALCOMPRESSION
ESTIMATED CONTOURS OF EQUALCOMPRESSION
ESTIMATED CONTOURS OF EQUAL
EXTENSION
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137
138
139
13
4.4.18
4.4.19
4.4.20
4.4.21
4.4.22
4.4.23
4.4.24
4.4.25
4.4.26
4.4.27
4.4.28
5.1
5.2
A. 1-A. 6
A.7-A. 14
B.1
SELECTED LINES OF LATERAL DISPLACEMENT
SELECTED LINES OF VERTICAL HEAVE
CONTOURS OF SHEAR STRESS
CONTOURS OF TOTAL SIGMA Z
CONTOURS OF TOTAL SIGMA X
CONTOURS SHOWING TRENDS OF NEGATIVE EXCESSPORE PRESSURE
COMPARISON OF PECK'S CHART WITH SETTLEMENTBEHIND THE NAKAGAWA EXCAVATION
ESTIMATED TOTAL STRESS DISTRIBUTION BEHINDTHE STEEL PILE WALL
TERZAGHI BASAL HEAVE ANALYSIS
SIXTH EXCAVATION STAGE - STABILITYANALYSIS
STRESS PATHS FOR STABILITY ANALYSIS
LINEAR VS. NON LINEAR STRESS STRAIN
HYPERBOLIC STRESS STRAIN CURVES
COMBINED GRAINSIZE ANALYSIS PLOTS
SQUARE ROOT OF TIME CONSOLIDATION PLOTS
VERTICAL EFFECTIVE STRESS VS. DEPTH,NAKAGAWA
Page140
141
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159
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174
183
14
CHAPTER ONE
INTRODUCTION
1.0 INTRODUCTION
An engineer should carefully consider the role of time
and specifically the role of excess pore pressure
dissipation, in designing and constructing "deep" or major
excavations. Design economy, safety, and the potential for
damage to the excavation and adjacent facilities can be
highly influenced by the effect of time on the response of
the soil mass.
Presently the usual practice of engineers is to idealize
the excavation or "unloading" case and assume "undrained" or
"drained" conditions. For excavations in clay the most
common approach is to use "undrained" strength and
deformation parameters for design. This approach assumes
the following:
(1) Construction occurs rapidly.
(2) No dissipation of excess pore pressure occurs
during construction (i.e., no drainage occurs).
(3) The soil strength is the in-situ undrained shear
strength.
The assumptions above are applicable only for an
"instantaneous" excavation. Actual excavations in cohesive
soils are not "instantaneous" and are partially drained.
15
Although for some situations (i.e. rapid temporary
excavation in soil with a very low permeability) the
undrained assumption is nearly valid at least for design
purposes, many major braced excavations are open for months
or even years. Significant pore pressure dissipation
following shear stressing of the soil or altered flow
conditions may occur. Undetected layers of pervious soil
may facilitate rapid drainage. As a consequence unpredicted
and significant increases or decreases in soil strength may
occur as well as variation in the stress/strain
characteristics of the soil with time.
Although relatively few studies of instrumented
excavations have been made to observe the effect of time on
the performance of braced cuts, some significant
observations have been made. To briefly summarize several
cases:
(1) Bjerrum & DiBiagio (1956) reported on an
experimental 4 meter deep trench in stiff-fissured
clay instrumented for the purpose of studying the
influence of time on earth pressure. They observed
the average force per meter as measured in the
struts increase from 2.3 tons/meter to 7.17
tons/meter over several months (Sept. to Dec.).
(2) Measurements from a braced slurry wall in stiff
Boston Blue Clay reported by Jaworski (1973)
indicate a gradual and significant increase in
16
earth pressure after excavation to the design
depth.
(3) DiBiagio & Roti (1972) reported earth and pore
pressure measurements for a braced slurry wall
excavation in soft clay. Significant decreases and
increases in earth and pore pressures during and
after completion of excavation to the final depth
were observed.
(4) Lambe (1968) reported piezometer data for an
excavation 22 feet deep penetrating an
overconsolidated medium Boston Blue Clay.
Negative excess pore pressures in the clay at the
bottom of the excavation decreased during the 24 day
period the cut was open. This indicated an overall
decrease in the strength of the soil as well as a
decreasing factor of safety with time.
The above observations reflect the fact that time is an
important variable to be recognized by the engineer.
Recently Clough & Osaimi (1979) investigated pore
pressure dissipation for excavations, including braced
excavations using a finite element model to determine how
long the undrained case was applicable. From the results
the authors concluded that pore pressure dissipation for
excavation in clay is likely to occur to a greater degree
than previously believed and that the undrained assumption
should be used with care.
17
Determining the degree to which dissipation of pore
pressure occurs for a given braced cut can be a complex and
difficult task. For sands time may be a factor with respect
to other phenomena such as soil transport (piping).
The engineer should attempt to assess the importance of
several variables prior to design:
(1) The rate of excavation and total time for which the
design support system is to be in service (i.e.
temporary or permanent).
(2) The character of altered groundwater conditions due
to the imposition of new boundary conditions and
groundwater control.
(3) The changes in total stress distribution due to
excavation.
(4) The effect of excess pore pressure dissipation
on the soil strength and deformation parameters.
(5) The effect of forced delay-(i.e. labor strikes etc).
Although construction time may be accurately estimated
for reasonable weather and contractor conditions unforseen
events such as labor strikes may unexpectably extend
construction time and hence influence performance.
Critical excavation events such as placing support or
bracing after an excavation stage can be most influenced by
such delays. Clough & Davidson (1977) report one case where
a final phase of construction (just prior to casting
and backfilling pile caps and slab) at the bottom of a
foundation excavation for a major office building, was
18
adversely affected by an unexpected delay. The slab was to
be used to support raker bracing for the sheet pile wall. A
carpenter strike delayed pile cap formwork for six days
during which the excavation stayed open and the walls crept
inward. After 20 centimeters of movement had occurred some
remedial support was applied by temporary filling of the
pile cap excavation to retard further movement. Eventually
over 30 centimeters of wall movement occurred. A large
crack appeared behind the wall. Figure 1.1 shows movement
versus time and the pertinent events for this excavation.
Corbet, Davies and Langford (1974) reported on a braced
excavation in stiff London Clay where a national building
workers strike led to delay of a critical period in the
excavation. Although the specific results of this delay are
not reported, the authors state that the delay was a cause
of concern for the stability of the excavation.
It is apparent to the author that the profession would
benefit from additional research and measurements of
excavation performance vs. time. The behavior of
excavations penetrating intermediate soils between the clay
and sand types such as silts, clayey silts and clayey sands
etc. are of particular interest. The demand for the
engineer to design with economy and safety warrant a better
understanding of the role of time and drainage on behavior of
soil subject to excavation unloading.
19
1.1 THESIS OBJECTIVE
The objective of this thesis is to formulate and
apply a method which includes the effects of pore pressure
change with time to predict deformations of an excavation.
1.2 THESIS SCOPE
Following the discussion of the effect of time for
several simple cases, the significance of measurements
made over time for several braced excavations will be
discussed. Several current design and analysis approaches
will be discussed. A Stress Path Method approach will be
outlined and applied for a type A prediction of a case
study. This method will include a composite of finite
element analysis, and stress path testing of soil specimens
in the lab to predict the role of time in the study of a
deep braced excavation.
Strike
1_3
Poured Pile Caps-
Pile Caps Partially Filled With Soil
o East Sheetpile Wall
3 West Sheetpile Wall
Excavation of Pile Caps(-Ilm)
Excavation to Subgrade (-9.5m)
5
0.
30-
25:
20-
10
5-
w2Li
0
-J
ELAPSED TIME SINCE START OF EXCAVATION (DAYS)
EFFECT OF FORCED DELAY ON -A CRITICAL CONSTRUCTION E
Elevation Om
. 32
DAY 6
Elevation Om
DAY. 3
DAY 7-14FrOM CLOt&-CL" j CA'IDSON (9-17)
VENT FIG. 1.1
5.
0-
' ' ' 5 ' ' ' ' 'O ' ' ' '2 ' 1150' ' ' ' '
0
21
CHAPTER TWO
THE ROLE OF TIME
2.0 INTRODUCTION
In this chapter the effect of time on strength and
deformation for several simple examples will be discussed to
outline the importance of excess pore pressure dissipation
for braced excavations. Related observations and
measurements of pressure and time from the literature will
be discussed. Various current methods of analysis and
design will be reviewed on the basis of how the methods
incorporate time effects. Limitations and assumptions made
for each method are discussed.
2.1 STRESS PATHS
Use of stress paths and the Stress Path Method (Lambe
1967, Lambe & Marr 1979) provides an organized framework
within which the field problem can be studied. Stress paths
for field situations can be modeled using standard
laboratory tests, such as the three dimensional triaxial
test.
2.1.1 Undrained Shear
At this point it is convenient to refer to Figures 2.1.1
and 2.1.2 constructed by the writer. Shown in the figures
are an example of braced excavation geometry and total and
effective stress paths for normally and overconsolidated
clay soil elements adjacent to the excavation.
22
The total stress paths between points 1A and 1B and
between 2A and 2B represent the type of unloading typically
experienced by soil elements 1 and 2. The mode of total
stress change for element 2 at the center line of the
excavation is extension unloading. Soil element 1 behind
the wall experiences compression unloading or an "active"
stress condition.
Typical effective stress paths for the undrained shear
can be traced between points 1A and 1C and points 2A and 2C
on each p-q plot. The horizontal width of the shaded area
at any given value of q is the value of excess pore pressure
generated by shear.
The shaded region also represents the region within
which the effective stress paths lie for partially drained
conditions. There are an infinite number of possible
effective stress paths in this region. Rapid excavation in
a material such as silt for example, may exhibit partially
drained behavior.
2.1.2 Dissipation Of Pore Pressure Following Shear
Strength and stress-strain modulus of the soil are
dependent on the effective stress path. Since changes in
pore pressure directly affect the effective stress path,
these changes are important in the field and when estimating
soil parameters for design.
Figure 2.1.3a-d shows stress paths for the example
geometry of Figure 2.1.1 for typical overconsolidated and
23
normally consolidated clays. The plots in the figure
exhibit a range of situations bracketing the type of pore
pressure behavior likely to occur during and after
excavation to the design depth. A simplifying assumption is
made in order to draw these plots; undrained shear due to
excavation occurs first and then is followed by dissipation
of excess pore pressure.
Excess pore pressure results from the total stress
release due to excavation and/or from decrease in pore
pressure due to altered flow conditions (i.e. construction
dewatering). With respect to dissipation of excess pore
pressures the following qualitative conclusions can be made
after construction of the Figure 2.1.3 plots:
Normally Consolidated Clay
(1) 2.1.3a--This soil will undergo a minor decrease in
strength (i.e. move towards the failure envelope)
following dissipation of excess pore pressure to
the original pore water pressure. This occurs at
both the bottom of the excavation and behind the
excavation wall. Consolidation heave will occur.
(2) 2.1.3b-- Consolidation due to dissipation of
positive excess pore pressure caused by dewatering
will result in settlement. The soil will experience
an increase in strength with consolidation. Net
vertical strain will be higher with dewatering for
element 1 due to additional compressive strains at
24
constant shear stress. However the net vertical
strain for element 2 will be less than the strain
associated with no dewatering. If two dimensional
steady state flow conditions are attained the
steady-state pore pressure can be determined from
a flow net. Although it may be unlikely that full
steady state flow conditions will be reached, this
condition represents an outside bound.
(3) 2.1.3c--For an increasing over consolidation ratio
the effective stress path will shift to the right
due to the generation of large negative excess
pore pressures. Subsequent consolidation will
result in a substantial decrease in strength
and a consolidation heave.
(4) 2.1.3d--If the absolute value of change in pore
pressure due to altered flow conditions is equal
to or greater than the absolute value of the excess
pore pressure due to shear, the strength will
increase. Otherwise the strength will decrease
with time.
For a given excavation with excess pore pressures
generated by shear and dewatering the net excess pore
pressure can be estimated by super-imposing or adding the
two. Lambe (1968) discusses this. Figure 2.1.4
schematically illustrates the procedure for simplifying the
field situation.
252.1.3 Stress Paths---Summary
The use of stress paths to illustrate the various
possibilities assists the engineer in his judgement.
Qualitatively the pore pressure-strength-strain relation is
clarified in one's mind. Similarily Henkel (1970) uses
stress paths to illustrate active and passive earth
pressures for excavation in normally and highly
overconsolidated clays. For comparison, several of Henkel's
plots are reproduced in Figure 2.1.5 showing a slightly
different graphical portrayal of stress paths. Henkel
recognizes the utility in applying this type of approach as
an aid to the engineer in his understanding of how a given
excavation may behave.
2.2 THE PERFORMANCE OF BRACED EXCAVATIONS VS. TIME
Cohesive soils are assigned low permeabilities compared
to cohesionless soils which have high permeabilities. In
general for the time associated with most construction
projects involving loading or unloading of the soil or
changes in flow conditions, some degree of drainage occurs
in cohesive soils. For sands the condition of full drainage
is achieved rapidly.
2.2.1 Braced Excavations In Cohesionless Sand
For excavations in sand deposits steady-state seepage
conditions are achieved rapidly. Figure 2.2.1 shows stress
paths illustrating this behavior. Since sand has a
relatively high permeability pore pressure dissipation is
virtually instantaneous. The total stress path equals the
26
effective stress path and the steady state pore pressure can
be obtained from a flow net. Therefore time does not play a
significant role. Although the following two sections are
not specifically bound to the thesis objective, they serve
to illustrate the different type of role time plays in the
excavation in cohesionless soil (vs. cohesive soil).
2.2.2 Piping, Suffosion
Where silts and sands or gravels, are present, time
plays a role with respect to migration or erosion of soil
particles via high seepage gradients into the excavation.
This migration is referred to as piping or suffosion.
Figure 2.2.2 graphically illustrates the mechanisms
responsible for suffosion and/or piping. Peck (1969) points
out that inadequate groundwater control may lead to damaging
settlements resulting from migration of sand into the cut.
Over a period of time migration of silt size particles can
clog construction dewatering wells and render them
ineffective. Particles can also be piped through cracks or
holes in a sheet pile wall. Serious conditions can be
anticipated by field montoring of flow into and out (pump
effluent) of the excavation for the presence of suspended
soil particles. Also at the bottom of the excavation piping
failure by heave may occur. The mechanics of this phenomena
were elucidated by Terzaghi (1942) and discussed by Terzaghi
& Peck (1967) and D'Appolonia (1971).
2.2.3 Effects of Frost Action
Field observations have shown the effect of frost action
27
to be substantial. Several cases show the loads in the
struts may increase several fold. Pappas & Sevsmith 1
monitored performance of a deep braced cofferdam in
sensitive Leda clay. Abnormally large increases in strut
load were detected during December excavation. Since
heating of the adjacent soil resulted in rapid load
reduction, the problem was concluded to be one of frost
act ion.
DiBiagio & Bjerrum (1956) observed after ground freezing
within and at the surface of an excavation trench, decreases
in load in the upermost struts, and severe increases in the
lower bracing levels buckling some of the walers and
struts. Although reliable deformation measurements were not
available, visual observations led to the conclusion that an
upward expansion of the soil during freezing caused the
unusual changes in distribution of pressure. Figure 2.2.3
shows the dramatic time vs. strut load behavior.
Terzaghi & Peck (1967) report that braced cuts in Oslo &
Chicago were subjected to below freezing weather. They
observed earth pressures climb to magnitudes several times
greater than those prior to freezing.
Ground freezing as these cases illustrate can result in
unpredictable and radical behavior. Since the formation of
frost lenses is time dependent, the potential for damage can
be anticipated by monitoring the wall movements and strut
loads.
-------------------------------------------------------------IRefer to Clough and Davidson (1977)
282.2.4 Braced Excavations In Normally Consolidated Clay
For clays one expects a soft normally consolidated clay
to have a lower coefficient of consolidation than an
overconsolidated clay since this value is generally found to
be higher in recompression than in primary consolidation as
shown in Figure 2.2.4 for a varved clay. We would therefore
expect soft clay to dissipate excess pore pressures at a
slower rate.
At least one braced excavation case illustrates that for
normally consolidated clay the ideal undrained condition is
not sustained over the course of time for a temporary
excavation. DiBiagio and Roti (1972) measured the
magnitude of total earth pressure, pore pressure and
deformations of a braced slurry-trench wall in a soft clay
in Oslo. Below the clay the wall was keyed into a pervious
bedrock. Measurements were made before,during and after
excavation. Figure 2.2.5 shows the soil profile and
geometry of the excavation as well as the instrumentation
layout. Instrumentation included fifteen earth pressure
cells and two piezometers.
The measurements during the excavation stages shown in
Figure 2.2.6 indicate only minor decreases in pore pressure
occured until the excavation approached bedrock and
groundwater control was facilitated through this layer.
Consequently, the pore pressure at the bottom of the clay
was significantly reduced, accompanied by a marked reduction
in total earth pressure on the wall. The excavation
remained open to this depth for approximately 200 days. A
29
downward gradient in the clay contributed to consolidation
settlements during this period before the "sealing" of the
excavation bottom with the construction of a base slab. The
pore pressure built up steadily from this point on
decreasing the rate of settlement and increasing the total
earth pressure back towards its initial value.
Over the course of the construction of this excavation
the total earth pressure varied dramatically due primarily
to changes in pore pressures. For the center of the clay
layer, piezometer PZ-9 shows a small decrease in pore
pressure over the time of construction. Coupled with the
fact that the location of the resultant earth force moved
upward towards the center of the layer with time, this
writer concludes that this central point of the soft clay
behaved partially drained. As a practical matter, in this
specific zone the undrained assumption may be reasonable.
However, looking at the clay mass as a whole the assumption
is not correct as piezometer PZ-8 clearly shows. After
careful study, the writer concludes that because the
decrease in pore pressure in PZ-9 occurred before dewatering
and remained relatively constant, dewatering did not
significantly affect the center of the clay and the negative
increment of pore pressure was due solely to changes in
total stress. Conversely the lower portion of the clay
layer shows the effect of both changes in total stress and
changes in flow conditions.
30
Although admittedly this is only one case, it
illustrates that pore pressure dissipation can be a factor
even for impervious walls in soft clay. Dissipation of
positive excess pore pressures resulting from dewatering can
lead to large pressure decreases against the wall during
construction. More rapid dissipation of excess pore
pressures can be expected for sheet pile and other
non-slurry-trench walls which are characteristically more
permeable.
2.2.5 Braced Excavation In Overconsolidated Clay
Some interesting conclusions from field data suggest the
undrained assumption to be seriously in error for the
overconsolidated clay.
Referring now to Figure 2.2.7, Lambe (1968) observed
piezometer readings from the medium clay layer at the bottom
of a temporary braced excavation for the CAES building
foundation at MIT. The excavation was open for 24 days.
Very litle dewatering was attempted in this excavation. The
stress paths for a point at piezometer P-1 are reproduced in
the figure. Note that u. , is the pore pressure immediately
after excavation. Thereafter the pore pressure increased to
a value greater than u. but still less than u. (the
pre-excavation static porewater pressure). The dissipation
of negative excess pore pressure moves the effective stress
path towards failure and results in a consolidation heave.
It follows as Lambe concludes that there is a progressive
decrease in strength leading to a decreasing factor of
31
safety with time. Therefore the end of the unloading period
is the most critical time for this excavation. The end of
the unloading period is defined as just before the load is
replaced in the excavation. Hence, the most critical
condition occurs some time after completion of the
excavation.
Observations and measurements were made for braced
slurry-trench and sheet pile walls in stiff clay for a
subway excavation in Boston (see Figure 2.2.8). Pore
pressures, deformations and strut loads were measured.
Figure 2.2.9 shows a redrawn compilation of strut load and
pore pressure versus time for a test cross section at the
slurry wall. Observations of piezometers indicate
excavation was accompanied by steady decreases in water
pressure behind the wall from an initial total head of 106
feet decreasing to 82 feet measured in piezometer P-4 behind
the wall after the final cut had been made. Following mud
slab placement, axial strut loads increased steadily as did
pore water pressure measured in the piezometer P-4 as shown
in the figure. Since measurements showed very little
outward wall movement (less than 1 inch) the change in pore
water pressure can be associated with altered flow
conditions resulting from construction. Seepage through
wall panel joints was observed. Figure 2.2.10 shows
measured pore pressure vs. depth. Steady state seepage
predictions compared well with end of construction pore
pressure measurements. Hence, it can be argued that the
32
drained parameters,C and are more applicable for this "end
of construction" case.
In constrast to the slurry wall case in the soft clay in
Oslo, pore pressures throughout the stiff clay drop
significantly. Imposition of new boundary water pressure
conditions are met with a relatively quick drainage. Also
the strut loads at the end of excavation do not accurately
reflect the strut loads which the bracing will experience
over the life of the structure. The trend in the figure
shows an increase in strut load by a factor of approximately
1.5. Similar trends are seen in adjacent struts.
2.2.6 Braced Excavation In Stiff Fissured Clay
One interesting case for excavation in this type of soil
is the experimental trench in stiff marine clay Bjerrum &
DiBiagio (1956) studied. Figure 2.2.11 shows a section of
the excavation, corresponding soil profile and test data.
Strut loads during the Autumn (Sept. to mid Nov.) show a
gradual and marked increase followed by enormous winter load
increases attributed to frost action and previously
discussed in section 2.2.3 (see also Figure 2.2.3). Autumn
rainfall data compared to the total earth pressure show
sharp load increases following periods of rainfall. The
most intense period of rainfall is followed by the maximum
value of pore pressure observed for Autumn in mid November.
Pore pressures plotted in Figure 2.2.11 show the measured
pore pressure versus time over this period. The fissures in
the clay facilitate access of water to the clay mass. Total
33
stress release from excavation in fissured and weathered
clays can lead to further opening of the fissures and
increases in permeability. Prolonged construction under
such conditions leads to progressive weakening or softening,
as well as increases in the succeptibility of the soil to
rapid changes in pore pressure.
The effect of fissuring on stability is dramatically
illustrated by the short term slip failure of an unsupported
excavation in London Clay reported by Skempton and
LaRochelle (1965). The results of this study lend
understanding to braced excavation in stiff fissured clay as
well. The slip occurred shortly after excavation had been
completed. The estimated mobilized shear strength in the
clay mass was approximately 55% of the average undrained
strengths measured from triaxial tests. The authors
attributed the difference between the measured strengths in
the triaxial tests and the actual failure to (1) pore
pressure migration within the intact clay, and (2)
fissuring. It was shown that by increasing the time to
failure for the triaxial tests, strength decreased.
Conventional tests run at a time to failure of 15 minutes
correspond to strain rates much higher than and
unrepresentative of the strain rate due to excavation. This
influence of strain rate on strength is attributed to pore
pressure migration. Figure 2.2.12 shows the results of
strength versus time to failure for a series of undrained
triaxial tests performed to study this behavior. Secondly,
34
the authors conclude that at least five factors related to
the fissuring can weaken the clay during excavation.
Several of the important factors are, absorption of water
along fissures leading to increases in water content
(softening), reduction of strength to zero as fissures open,
and progressive failure.
Quantifying the effect of fissures and the rate of
softening on stability and deformation to predict
performance is difficult.
Corbet et al (1974), and Cole and Burland both reported
observations made for a raker braced diaphragm wall in a 20
meter deep excavation for a basement in stiff fissured
London Clay. In view of the uncertainties cited in the
previous paragraph, detailed construction performance
monitoring was undertaken. Wall movement and verticality
were monitored. Earth pressure however, was not monitored.
Several of the conclusions drawn by the authors after
observing the construction performance are reiterated below.
(1) Wall movements showed a clear time dependence.
(2) Early internal support placement is essential
to minimize movements.
(3) Performance monitoring is important for construction
control and is of research value for future exca-
vat ions.
(4) Inward wall movement is associated with a
progressive softening and reduction in stiffness
of the clay. The greatest reduction in stiff-
35
ness was found to be near the ground surface behind
the wall. This is attributed to the opening of
fissures due to lateral stress release.
In summary, one can clearly see that time is a very
significant factor for load and deformation behavior of
excavation in stiff fissured clay. Although additional
observations and measurements of field situations are needed
to advance our understanding of this case we can make
several conclusions of practical value at this time. The
undrained assumption is not applicable to excavations in
stiff fissured clay which remain open for several months,
construction performance monitoring is important and is
recommended, and minimizing construction time is desireable.
2.2.7 Braced Excavation in Silt
Very little well documented information on excavation
performance in silt has been reported in the literature.
Lambe (1969), et al, and MIT (1972) reported on a study of
two instrumented Boston subway cuts in fill overlying organ-
ic silt, underlain by glacial till as seen in Figure
2.2.13. The test sections, Section A and Section B were
monitored for stresses and deformation and compared against
predicted behavior. No significant strut load variation
with time was measured. However, one conclusion reached in
the studies was that for braced excavation in sands and
silts the pore pressure behind the excavation can be ex-
pected to be less than static due to construction. Some
agreement between the final measured pore pressures and
36
steady state pore pressures predicted by finite element
analysis (FEDAR, Taylor and Brown 1967) indicates the pore
pressure in the silt was probably in a transient state near
steady state. See Figure 2.2.14.
Although these two test sections penetrate silt, soil
conditions are not truly representative of an excavation in
silt due to the upper fill and stiff lower strata.
Therefore general conclusions by the writer regarding
excavation in silt cannot be made. However the cases are
representative of the fact that many excavations do
penetrate several soil types and cannot be easily
categorized.
These studies are valuable case histories. They
illustrate the type of construction monitoring data which
predictions of construction performance can be evaluated from.
As a result, our understanding of braced excavations can be
extended. The reader is referred to the references for
additional details. Use of finite element computer programs
like BRACE and FEDAR to model braced excavation and flow
respectively showed encouraging predictive capabilities.
2.3 METHODS OF ANALYSIS AND DESIGN
The following currently used methods of analysis and
design will be discussed as they relate to the incorporation
of time effects.
(1) Semi-empirical Methods
(2) Numerical Analysis Methods
(3) The Stress Path Method
37
2.3.1 Semi-empirical Methods
Several well known semi-empirical design methods are the:
(1) Terzaghi and Peck Method (1948, 1967)
(2) Tschebotarioff Method (1973)
(3) DM-7 (Navdocks) Method (1971)
These methods are similar and are commonly used in
practice. They are intended to estimate maximum lateral
strut loads. Only the Terzaghi and Peck Method (1967) will
be discussed to illustrate this design approach.
Terzaghi and Peck presented the original apparent
pressure envelopes in 1948 for braced excavations in sand or
in clay. Peck, Hansen & Thornburn (1974) is the most
recent published version. The design envelopes were
derived from classical earth pressure theory and actual
measurements of strut loads in excavations. Hence the
method is referred to as semiempirical. Difficulty in
reliably measuring stresses against the sheeting required
the apparent stress distribution to be inferred or
backfigured from the strut loads. The envelopes are
intended to give maximum values of strut load to be
experienced and are not truly representative of the actual
stress distribution. Over the years modifications and
adjustments have been made to accomodate newer observations
and experience. Figures 2.3.la and 2.3.1b show the design
envelopes of the Terzaghi and Peck Method for sand and clay.
The pressure envelope for sand is based in part on strut
load measurements for braced excavations in New York,
38
Munich, and Berlin. An equivalent uniform earth pressure
was fit through the apparent pressures in the excavation.
Several qualifications attend the application of this method
for sand.
(1) The authors emphasize that the diagram has been
developed from a limited number of excavations
varying in depth from 28' to 40'.
(2) The diagram is a pressure distribution for
estimating the maximum values of strut loads to be
expected for similar cuts.
(3) The groundwater level is assumed to be below the
bottom of the cut. Consequently, (in the writer's
opinion), seepage or static water pressures should
be added to the pressure distribution. Ries (1973)
points out that for impervious diaphram type walls
which typically retard drainage of the soil behind
the wall, this assumption (water level below cut)
is likely to be seriously in error.
For sands detrimental time related effects have been
associated with piping and piping heave at the bottom of the
excavation. This problem cannot simply be incorporated into
any pressure envelope. Application of the Terzaghi and Peck
method for sands must accompany an awareness of such
conditions. Also, in the writer's opinion the groundwater
pressures which may be experienced over the life of the
structure should be added especially since these pressures
can be substantial.
39
Several relevant facts one should be aware of when using
the clay envelopes of the Terzaghi and Peck method are
summarized below.
(1) Cuts are assumed to be temporary. The)=0 condition
or undrained condition is assumed to prevail.
Average undrained strength of the clay along the
side the cut is used.
(2) Although the data from the test cuts (Oslo and
Chicago) from which the envelopes evolved included
measurements showing large increases in strut load
due to freezing during cold weather, these
measurements were not used for formulation of the
envelopes. Therefore the effects of freezing
should be considered separately.
(3) The authors (1967) admit that due to little
available data on cuts in stiff intact clays, sandy
clays, and clayey sands etc., design envelopes have
yet to be worked out. However more recently, Peck
(1969) reports that the trapezoidal clay envelope
fit the measured values of five cuts in stiff clay
conservatively and therefore could be used as a
guideline. Peck suggests that limited data for at
least one deep cut in dense clayey sand and stiff
sandy clay shows drained behavior and the
appropriate drained strength parameters can be
input into an equation for Ka. Presumably, the
equation for earth pressure for excavation in sand
40
is then modified to include this expression.
(4) Use of the trapezoidal envelope for stiff fissured
clay is discussed by the authors (Terzaghi and Peck
1967, Peck 1969) in light of the measurements made
by DiBiagio and Bjerrum (1957) which showed time
dependent behavior. This case has been discussed
by the writer in earlier sections of this paper.
The authors suggested that the lower value for
maximum design pressure, .2WH be used if movements
of the sheeting are minimal and construction time
is short. If these conditions do not apply, use
of the higher pressure value of .4VH is recommended.
Hence Terzaghi and Peck, in an indirect way have
assigned some importance to time effects for stiff
fissured clay.
In summary, it is the writer's opinion that these
methods provide a useful guideline for estimating maximum
strut loads. The method has been tempered with experience
although to extend the scope of the Terzaghi and Peck method
more experience for a large variety of cuts is required.
The limitations of the method should be realized and may be
understood by studying the evolution of its formulation.
Major braced cuts open for extended periods should not be
designed on the basis of this method alone. The engineer's
understanding of the excavation problem may benefit by using
other more fundamental techniques.
41
2.3.2 Finite Element Analysis
Over recent years the advent of the high speed digital
computer has enabled numerical analysis to become a popular
and viable technique for solving complex engineering
problems
sets of
mathemat
sets of
complica
Appl
now well
Zienkiew
analys is
. The use of the electronic c
equations by matrix algebra al
icaly model problems in detail
equations could be set up but
ted to achieve practically by
ication and theory of the Fini
documented in the literature
icz 1967, Bathe and Wilson 197
. Subsequently, over the last
omputer to solve 1arge
lows the engineer to
* Previously such
the solution was too
hand.
te Element Method is
(Desai and Abel 1972,
6) for engineering
decade several finite
ent computer programs to model construction of braced
vations have been developed and used by several
stigators including Wong (1971), Jaworski (1973) and
gh and Duncan (1971).
Advantages of using this technique include:
(1) Complex soil and structure geometries and
properties can be modeled.
(2) Construction sequence and details can be simulat
(3) Detailed information on stress changes and
deformations can be obtained.
(4) Parameter studies can reveal the relative import
of input data (strengths, moduli, sheeting stiff
and penetration etc.) for a given excavation.
(5) Critical stages of excavation can be predicted.
elem
e xca
inve
Clou
ed.
ance
ness
42
Some limitations have yet to be overcome. One such
limitation lies in the difficulty in modeling the soil
structure interaction (i.e., soil sheeting interface).
Nevertheless, even in the relatively young stage of
development for finite element modeling of excavation,
encouraging results and insights have been gained. Indeed,
in practice, MIT (1972), Cole and Burland (1972), Palmer
and Kenney (1972), Clough and Tsui (1974), Duncan and Chang
(1970), and others have demonstrated the powerful
applicability of the method.
The finite element program developed at MIT for analysis
of excavation is called BRACE. The reader is referred to
the theses by Wong (1971) and Jaworski (1973) for details.
The program in its present form will model the stress strain
behavior of the soil via a linear or bi-linearly elastic
modulus. The program does not model the effects of time on
strength, modulus, loads or deformations. If necessary,
these effects must be indirectly considered by inputing soil
parameters which include the effect of time (e.g., pore
pressure dissipation). Conventional soil parameter input
for such a program is the undrained soil strength and
undrained modulus to model the undrained case (total stress
analysis) or drained parameters, C and for the drained
case (effective stress analysis).
Pore Pressure Dissipation
Recently, new insight into the evaluation of the
validity of the undrained assumption and total stress
43
analysis for excavation has been presented by Clough and
Osaimi (1979). The authors used the finite element
technique to model excavation sequence for given rates of
excavation (meters per day). The purpose of the study was
to show how pore pressure dissipation occurs during and
after excavation. The name of this computer program is
TIME-CONSTRUCT and at this writing is not available for
distribution (Clough 1980) since its documentation is not
complete. In any event, the published results by the
authors are of interest to the engineer and will be
discussed in the following paragraphs.
Using TIME-CONSTRUCT Clough and Osaimi analyzed several
cases. Results of the analysis of a one dimensional
excavation are shown in Figure 2.3.2. For wide excavation
the one dimensional unloading and drainage represent the
central protion of the bottom of the cut. The figure shows
that decreasing the excavation rate corresponds to an
increase in the percentage of consolidation at the end of
construction for the wide range of permeabilities analysed.
Figures 2.3.3a and b summarize the analysis results for
two dimensional excavation for cases of both linear and
nonlinear soil behavior. Contours of negative excess pore
pressure shown indicate that the pore pressure response at
the bottom of the excavation is essentially one
dimensional. Dissipation of the excess pore pressures will
occur vertically. Of general interest, in Figure 2.3.3a
contours of negative excess pore pressure for no wall in
44
place are shown for comparison with the supported excavation
in the linearly behaving soil. As would be expected the
presence of an impervious wall markedly inhibits
consolidation. Therefore we can conclude that for the
realistic case of a more pervious wall (sheetpile), the
magnitudes of negative excess pore pressure would be at
values somewhere between the two cases. In any event more
consolidation would occur for the sheet pile wall than for
the impermeable type.
Note that this analysis by Clough and Osaimi does not
consider the role of drainage due to dewatering which
accompanies excavation construction. As a result, the
dissipation of excess pore pressures due to shear alone will
produce an upward movement of the soil mass behind the
wall. Since measurements and observations for many case
studies consistently show settlement occurs behind the wall
with time, we can conclude for such excavations one or more
of the following:
(1) The dissipation of negative excess pore pressure
due to shear behind the wall is negated by equal
or larger positive excess pore pressures resulting
from dewatering.
(2) The heave movements are non-existent behind the
wall due to greater settlement resulting from
inward wall movement and/or loss of ground towards
the bottom of the excavation with time.
45
(3) Other factors including surcharging by traffic or
equipment as Clough & Osaimi suggest.
One important conclusion was drawn by Clough and Osaimi
as a result of their study. The undrained, "end of
construction" case is unlikely to occur even in thick
deposits of low permeability clay. They further note that
consolidation in the field is likely to occur more rapidly.
This latter conclusion is probably based on the commonly
accepted fact that consolidation in the field generally
occurs several times more rapidly than predicted by
available methods. Extrapolating test results using small
specimens in the lab in analyses of the field situation where
undetected drainage layers may exist may explain some of
the discrepancy.
2.3.3 The Stress Path Method
Lambe's Stress Path Method is a useful method for both
understanding and analyzing a given geotechnical engineering
problem. Lambe (1967) and Lambe and Marr (1979) describe
the fundamentals and applications of this method. By
selecting and analyzing representative or "average" soil
elements, one can attempt to estimate the stresses and
strains that will occur in the field for each phase of
construction.
For a supported excavation one can subject
representative soil specimens to the in-situ initial
stresses, apply the total stress changes associated with
excavation and measure the strains, pore pressures (and
46
hence, effective stresses), and coefficients of
consolidation and lateral earth pressure, K..
Of course the method has some limitations one should be
aware of. Sample disturbance affects measured values of
coefficient of consolidation, strength, and modulus. K,
determined from reloading to the initial stresses is not
truly representative of the initial K. in the field. Stress
path tests can be both expensive and difficult to run. Many
of these limitations however, are present to varying degrees
in the other methods available to the geotechnical
engineer. The advantages of the method outweigh the
limitations, particularly if the engineer is aware of the
limitations. The value and the applicability of this method
lie in the evaluation of the predicted versus measured
performance of actual geotechnical projects.
2.4 CONCLUSIONS
The measurements from the case histories cited from the
literature show that time effects (pore pressure
dissipation) are of significant importance in braced
excavation. For cohesive soils partially drained conditions
exist during construction and the undrained assumption is
only a design simplification which is in error and may be
seriously in error in some excavations.
Recent numerical techniques allow analysis which may
include the effects of time, but most use linear
stress-strain parameters. Other numerical techniques which
have non-linear parameter models, are expensive and
47
difficult to use and don't consider time effects. Although
Clough and Osaimi have developed a computer program
(TIME-CONSTRUCT), which considers non-linear and time
dependent stress strain behavior, their studies to date have
only treated a simple, hypothetical excavation.
Nevertheless their work is a significant advance in this
area of study. The writer looks forward to an evaluation of
predicted versus actual performance of an actual excavation
using TIME-CONSTRUCT.
The Stress Path Method provides a rational way to select
soil parameters which include time effects. One may combine
the Finite Element Method with the Stress Path Method. The
powerful numerical technique can be used to model sequential
excavation and obtain the complex stress distribution in
conjunction with the Stress Path Method used to determine
the appropriate soil parameter input. The deformations of
an excavation could then be predicted using this fundamental
approach.
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a
+
a-ESP
k-k 0
54
DEWATER
he-
ss5/7
//%\ ow w
o= hea.d drop
SZD(s ss)
EXCAVATE
time
06
h
aL=head~ drop
A h5 -h-LL( 6 -)
tofca/ /ead >
DEWATER + EXCAVATE t o-ti heoxd = h - G -as cdewoter- e xcvate
DETERMINATION OF TOTAL HEAD SUBJECT
- TO EXCAVATION AND/OR DEWATERING
a +er- La-mbe.(19618)
K0
r0
FIG. 2.14
iota/ /tead -lo
E X C AVAT
0 1000 2000
STRESS
3000 4000 5000 6000 7000
Oh; pSi
PATH FOR SPECIMEN 0* ACTIVE CASE
/ p
th P
250
20C
t5o
3.0 6m2.0
1.0-- 0
7 7' %
---
2500
2000
1500
Ov P-s I-
1000
50C
0
14000
12-000
8000
0 500 1000 5000
t1 pst
STRESS PATH FOR SPECIMEN N ACT!VE CASE
N P -
faces%- 7to .7p
.6 p
0
0 500 1000 1500 2000 2500 3000
0r, ps.
STRESS PATH FOR SPECIMEN N. PASSIVE CASE
AND PASSIVE CONDITIONS FOR NORMALLY ANDOVERCONSOLIDATED SOIL
0 2000 4000 e 6000 8000
Stress History For Specimens N Gnd 0
FROM HENIKEL (131O)
FIG. 2.1.5
L
5000
4000
3000
G00ps0
2000
A N0 -
T
I 30/
0.
0
soSC
5000I
40001
3000
O5 p s.I
2000
1000
0
b
N-I-fl
0 1000 2000 3000 4000 5000 6000 7000
hl1 P I.
STRESS PATH FOR SPECIMEN 0: PASSIVE CASE
STRESS PATHS FOR ACTIVE
4 000 -
2000-
~1-a
0
N.
TV ru
T> P o It-4I
+
l RVo,
10 0
FLOW NE T
ressire ot4 - e\eJQ oVI-h0ea.d heod Vecxd
STRESS PATHS FOR EXCAVATION IN COHESIONLESS SAND FIG. 2.2.1
4-EXCAV'ATtoM
Nun
U. (StATIC)
EtP
'V
1/
;u 171
4-5
-4
2a
hp,, . Y
pressure
Flow
-A] Movementof Fines into
Coarse Structure
Flow
Piping ofFines from
Coarse Structure
FRoM CLOuG6 a' DAvI1OSoij ('977)
PIPING/SUFFOSION MECHANISMS
F I G. 2.2.2
//IA\\WI/
U,
I.
FILL
SILT
-SAND.-. --
GRAVELLY'. B ,'SAND o - - -
. 5 6 - -.
58
06&CJ, Air teripero;breC C -- - - ~ p%
/ 02 0 2
EF-2-0- c
- E
c
'D level struts
- ---- --~ --- ---- ~ ---
C levei stru s
Q3
LA!~IA.erg -tr. lod .lt -e ~ cpr
Avercge strut lood, ion vrsus cepto0202 0C 202 0 2 202 0 2 4 6 2
LWIJ~H
'A' level struts
C D level struts
0-
I - A-mx vOd tes0-- Min. voues
0- E'level struts-
OF -- EXcavotion cod braccing cornpiefed0[ L.--Al wedges removed
-
.0
0--
* F level srt
.0 A
O__.. % '2 20 26 1 6 14 22 30 1 7 15 23 1 9 17 2 5 1 10 1 26 13 11
-19.15 seot.+ - Oct -Nov - -DeC 47956 Jan. - wi Feb.-
D/o/AG/a f V.TrAPUA (1957)
INCREASE IN STRUT LOAD VS. TIME FOR ABRACED CUT IN STIFF FISSURED CLAY F I G. 22.3
2-0-.0 __
7 -06-05-0
2. 0110
0L..
9-
7,6-5.4-
3.
2-I.
109.6-7-654.3.3-2
76
3.
2'I.
I -a ert -i , L- r.I I,1j0 .6I .aa
59
C
/2
A/o&7-HAMP7Zw VAKYEZ) dAY
(DD V /sSA97) I Ilt
'2- 2. - /0 Z 0 5-0 /00 20 00
0.8 __ ____ _ ___ _ ___ __ _____ CYtCLE_ _ _
-ruTA LOADING
I . II
2 5 /0 20 50 /00 CoO 4o
CA S c 7.EST -rs '\ AFTEZ LADO 973)
VARIATION IN COEFFICIENT OF CONSOLIDATION
FIG. 2.24
t t 1VK~..
0
I-LL
(-)
IV -IT
rS-61 ITS-62 S-631 T S-190
-- L- -L - - --------------------- -- ---- --- ------ -
2
EARTH PRESSURECELLS
E Water con!ent Unit UndrainedSSoil description weight shear strength Stt /m?
20 30 40 50 J /m3 1 2 3 . 5
FILL
CLAY CRUST 87
186 50 ---- 8- w 18-
185
1 89 5
SILTY CLAY-5 187 5
t 868
I 192 + V+i
187 7
194 11
-1 2 01 -
-112 - 20 ___2_184
208 li 2
w natural water content S = sensitivityw : liquid limit V z conewp= plastic limit z vane
-5'
6
71
81
101
11 1121
131
141
15 I
S - Settlement boltsP - Piezometer
-02
-3.6
ON'0
-14.1
-15.3--E
0 5 10
Scale, m
Typical soitl data Instrumentation details - Tdlefonhuset.
FROM DtAw'io ; RoTi (97Z.)
SOIL PROFILE, EXCAVATION, AND INSTRUMENTATION DETAILS FOR A BRACED SLURRYWALL IN SOFT CLAY, OSLO
F-IG. 2.2.5
Key
0-
5
E
0
w
-10
-15 _
P p_9z
61
0 100 200 300 /00 50c 600 700
nistollation cf Indr ottm 1
is B EARTH PRESSURE
16 *
.12 -
-- T0 I10 -~ , - . ---- d
--P ,-- 3
8--
5. C TOTAL FORCE
-200
041
T-. LOCATON
o~~~N a>r~ * i
0
038
0 130 rC C 0 0 0 0
ESUMMARYOF .OBSER.E PORE PRBECSSUR E
--.. . - . 2---------PP 9
2
E -1 F SETTLEMENT-
2 - -
Po'i 5-63 -'-" ...-
0 100 20C 30C 400 Soo 60070DAY NUMBER
Summary of observations from Dec 1969 to Sept 1971
FROM 1D'tB,,Aci0 Ren' (1577.)
SUMMARY OF OBSERVATIONS FOR BRACED CUT IN SOFT CLAY, OSLO
F IG. 2.2.6
62
SI-2 T
S4 is-
+20 -
0 - IS-2 s___oSand /WAAA7AaywA
JS-3' WIS-1
Medium clay P5-25 o o-25 P1layer I
S4-39 -39 S2
layerclay P6-52ijo o-52%, P2
Soft clay-sand P3 -77;,layer III
Tif N -92BlBMf 0 -991,
2 -
B -
p, > (tons/m2) 10 20
2DL
STRESS PATH FOR PIEZOMETER P-1
BELOW THE CAES FOUNDATION EXCAVATION F G.2.7
-20
-40
-60-
-80
+
3 ,
12
N
N
Fl LL
HARDY E LLOW
CLAY
Very Stiff to
Stiff Boston
Blue Claywith Sand
Len sesIv
Stiff to
Medium
BostonBlue
CLAY
TILL
ROCK
70
122
ld
H *-
H ~-
-1*---
-- + H
_+ H
-+ H
Bottom of Excavation
ConcreteWall
STRUTS Z 12' C /C (3.6 M)
50FT
P-_
l80FT
STRUT WALF
B 14WF127 36WF170'C 14 WF 103 36WF 150D 14WF 176 36WF30
SOUTH COVE SUBWAY EXCAVATION TEST SECTION PIEZOMETER P-4
FIG. 2.2.8
0
7
22
0-0
-520-
1040-
15
/
/ /
V. /*/
DON BOSCO SCHOOL
///////
I-a.Wa
60-
100-
-20
30
(.,3
10FRom 3AWORSKi (1973)
, , ,3
b
Ut~~~ 44oSDE WALLSn Mo ao-
/ 00-
-44L&PPR L04". (a
-0
- -- -Zoo
-4 /00- - - ThTS R4jM0Y*D BEAM
48-
//0 - -//0I//0
/Ole 4- iA/Jr/AL.ocr /969
949
86
78 -78AFTEP, DATA FRoM .3AvOtKi (1973) -
AuG SEp OC-r NOV DOC JAN FES JAA. AP1 . AY E :3-"L.Y
9 1970
STRUT LOAD AND PORE WATER PRESSURE VS. TIME-SOUTH COVE SECTION
FIG. 2.29
65
PRESSURE HEAD (FT)
z
I-i
_j ww -120f 0
-10
00- -20
Medium
Boston
Blue
CLAY
-100
I TILL
/777
0 20 40A
0
< - P 4
PH-3 p_--
PH-4 N
-__ _____
60 80 100
INITIAL READINGS (3/69)
EXCAVATION COMPLETE (11/69)
FINAL READING (4/70)
P 5
P -- H(ON . OF EXCAVATION)
n--P6
STATiC
UssK walI =Kh
soil
P-I-H
FROM ~SAWORSKI (11-7)
MEASURED AND PREDICTED STEADY STATE
PORE PRESSURES-SOUTH COVE TEST SECTION F 1G. 2.2.10
FILL
Hd. Yet.
-CLAY
StiffBoston
Blue
CLAY
-30
80_h40
-50
-70
-80
-90
40-
20
Don Bosco High School
nu I0 ~ ~ P "7 I
.-PH -4 H -
I-
66
EA
E .
Open tiroe,' panel (2-0x 1-33m)E V, [p, Steei waling(.70NP)
E . Ock beading plre
E ~ - Screw type steel strutsEE
E Stee' re'erence s/QkC
0 2m0-9 M
E
4
MaxImLim winter strut lood(ton/m rur)0 ; 7 6 5 4 3 2 1 0
A -+
Max'mum frost penetration
4 Feb. C-.
27.ar..956
F-
0 1 2Distance scc'e rn
Autumn stru load (t m run)0 7 2 4 5
.. 15Nov
-25 Nov.
, Isept(Before excavation)
20 Sept.
Lines of zero pore wcter pressure
(Ju20 Sept? 21 Oct.T 5 Nov. FROM WERLLM
25 NO V. AN[) DSAGio (1950
TIME DEPENDENT BEHAVIOR OF TEST CUT IN STIFF FISSURED CLAY
FIG. 22.11
Water conten (%) B.lk density(ton/-' Shear stre-grh (ton/m2) Sens;-10 20 30 40 50 60 70 1-7 1- !-9 2-0 2-1 24 6 81012141618 tivity
PL LL
Uncor fined1 3compressior tests
I fIn1 /one boring 27
-_ _ -_--_-_-_-_-- - 30120
C, ~ 5i5
g
67
100%
80
"rheolog;cal" decrease in strength
ecrease due to 6")
60- -Block 6
2 31
20-
Water content41 increase in
0 A w
15 mns. 6 hrs. l aCy 2 5Time to failure t (fi scale)
c= rzor sirength, time to fcilure =15 mins.c1zsheor strength, tirr.e to failure = t
ROm sKEMP-roej Aus LAROCHELL1E J% >
EFFECT OF VARIATION OF TIME TO FAILURE
ON STRENGTH IN UNDRAINED TRIAXIAL TESTS
10 15 20 days
FG. 2.2.12
2c Z 29 psi. 2ct= 23 psi.
w %Le31-0 17 p.s. 30-9 - 14-5 psi.
31-0 -- 9-. p.si. 320 - 14 p-s-i.
- Shear zone
15 min. test p =30 ps 7 day test
c = 2100 ps.f. Ct= 170 0 ps.f.
Aj 4 9.3 51 .20-8 A = +0-6C
c
cl
68
At-/h SAM^' 5'(o p~s ) |
10
-50
/cc/e/03eH
71s T S-c1roN A
4 *5)',-i,
Se - -
4- -
O#9d7J / /
r /0 * 4/s/ o cr o pio
A- .e4 '/Oa0*~ cm/s, -
ro-.sc /7"58ro -, //7~ s
.$a*! /L .
/0,4 -'/Xie oGP-ay .5o7 C/ey A/5O7O'
7 5sr SEcf/oA B FRMo\ LAMBE 0570)
EXCAVATION PENETRATING SILT: MBTA TEST SECTIONS A & B
-50
L60
AoArxdrw*N
IL4
oc/47
T74 1
0.
/0.
&0.
#0.
GO -
K
NK
N
28I7m7 y
S.'leebn; SO./ A 1-4 cl"eceSee/c-/./
FIG. 2.2.173
... ,awn any,,-,,,,,
i- - lr7, 0 "*C
r
0
/0
20-
30
40
60 -
x M
0 StOfc, LIS
SteadySeepcpe , U4.5
Teas4jred 20 ft from wa//easured a/ong wo// //IG4/68
0 I 2 3 4
PPc'SLPE /N K/PS'/T 2
POPE PPESSUPE SEC 7/ON B
MEASURED AND PREDICTED PORE PRESSUREMBTA TEST SECTIONS A AND B FIG. 2.2.14
69
Steoy S..pgeBest Estrnoe-
Stoic , /4s0 x
K
Kj
AMLasured Pote Pressuresx 40 feet from wo//* /8 feet from wo/Io 9 feet from woll on 6/1/69
F/LL
S1L T
7LL
0 / 2 3PPESSLE' IN K/PS/FT 2
POPE PPESSUPE .ECT/ON A
0
/0 C
K
t~J
K
k
Q
FIL L
SILT
TILL
30-
40-
60- I I I I
GANDY CLAY5,CLAYEY SAN D5
STRUT --
H
O.65kH
K, = TAN2(i5-/)
TiER-AGHi AtN4DkCK (Ie7)
0.65KVH
K=TANz(45-/) I HTAN(+S-
PE.CK (tc3(3)
TERZAGHI & PECK DESIGN ENVELOPES
FIG. 23Ia
0
71=1 ":--
P0EC K, "ANSON A wTHORNBURN (1'314)
YT-4UT --- 4
H-4CFOPt N z--mg
,'~ '7,,,.'
o.751
CLAYSY H
N=c0.25H N H
0.75H
6.
fP~ f(~~/(
0.s H
0.15 H
L!
YH-c N *~oiZ' 47
YH -2j ozHN > 6 " 8, LARGE DEPTH 0 -HOF CLAY BELOW CUT
TERZAGHI & PECK DESIGN ENVELOPES
os H
0.25 H
FIG. 2.3.lb
'PECK (1%69)
TIME TO COMPLETE EXCAVATION0 5 daysA 50 dayso 500 days
Ift = 0.305m
80-k = C04 Cm/seC
60-
40- 10-6 cm/sec 1.5 x 10-5 cm/sec
20 - YO~8 cm/Sec-
0 2 40 2 4 6 8 10
RATE OF EXCAVATION
EXCAVATION
7VT50'
-I
. in ft/day
FROM CI0ot-H A, OSAIMI (979)
CONSOLIDATION AT THE END OF ONE-DIMENSIONAL EXCAVATION
FIG. 2.3.2
0
0Cl)
z
0
0
a0
CE
0LuLX00
k)
73
NS. '.(&T7
-500Imeavation Rate - 0.5 ft/day
K K - 10-8 o/cI y
3o crr_^ ll
wall
30
Vro- O"SPm (i9T7) ONTOURS OF EXCESS PORE PRESSURE (pof) AT THE END OF EXCAVATION.
CONTOURS OF NEGATIVE EXCESS PORE PRESSURE
FOR TWO-DIMENSIONAL EXCAVATION
FIG. 23.30
-500
-500
60 rr
Strut
K . 100
-2500
-ow
3o FT 40 $r
KI . 300
Cnod"ks ,,wmvbe-)
-2500
Ffw OSAIMI (s9TT) CODNiOURS OF EXCESS PORE PESURE (psf) AT THE ED OF EICAVATIO.
CONTOURS OF NEGATIVE EXCESS PORE PRESSURE
FOR TWO-DIMENSIONAL EXCAVATION
FIG. 2.3.3b
74
W- all
Strut Wall
'iT
75
CHAPTER THREE
PROPOSED PROCEDURE FOR INCORPORATING
TIME IN THE PREDICTION OF THE PERFORMANCE OF A
BRACED EXCAVATION
3.0 INTRODUCTION
In this chapter the writer outlines a proposed procedure
to incorporate a major time effect (pore pressure
dissipation) in the prediction of construction performance.
Acknowledging advantages of certain aspects of available
Finite Element techniques and the Stress Path Method, lead
the writer to use both in a composite procedure. Hence this
procedure integrates the use of a sophisticated computer
model as a tool for a Stress Path Method approach.
The fundamental basis for this method is the Stress Path
Method. The most recent edition is presented by Lambe and
Marr (1979). The basic principle behind the Stress Path
Method is the effective stress principle, which states that
the effective stress equals the total stress minus the pore
pressure.
TI
Soil behavior is dependent on the effective stress history
and future changes in effective stress.
Table 3.1 illustrates the types of geotechnical
problems; force, deformation, stability, and flow, which the
geotechnical engineer must face in practice. Generally, he
76
must predict the performance of constructed facility. Table
3.2 summarizes the steps used in the Stress Path Method to
predict performance for the classes of problems presented in
Table 3.1. The method developed in this chapter by the
writer follows the steps in Table 3.2 for force and
deformation.
3.1 OUTLINE OF THE PROPOSED METHOD
This procedure is an iteration whose solution can be
improved with each pass through the steps outlined. Actual
field performance measurements during construction can be
input to update or adjust the solution. The proposed
procedure is as follows:
Step Description
(1) Using a finite element computer program which
models braced excavation (i.e. BRACE), input
an initial geometry and initial soil and structure
properties (i.e. strength, moduli, pore pressure
parameters).
(2) From Step (1) obtain total stresses (approximate).
Ac~
(3) From Step (2) obtain usE(the excess pore pressure
generated by the change in total stress due to
excavation).
(4) Estimate u (the excess pore pressure resulting
from the altered flow conditions associated with
77
the excavation). A hand drawn flow net or other
method (FEDAR) may be necessary if steady state
conditions are anticipated.
(5) Run dissipation analysis for actual drainage
condition (i.e. one, two, or three way drainage)
with estimated coefficient of consolidation.
Estimate the percentage of dissipation, D.
(6) Select a representative element(s) or prediction
point and an undisturbed soil sample(s).
(7) Consolidate the soil specimeNs) to the estimated
in-situ stresses. Measure the coefficient of
consolidation.
(8) Shear the sample by applying a total stress pat
obtained for the representative element. In
other words apply ACrF. Measure the actual uE a
the shear strains 2S . The first set of tests
be sheared under conditions of no drainage. If
second set are executed after the first pass th
the procedure, partially drained tests may be
required.
(9) Using the results of (5) and of (7) and the
estimated time t for the excavation constructio
determine the percentage of dissipation,D.
(10) Then,
h
nd
might
a
rough
n,
AUJT~ D +_ 4UE)
1. One might use superposition as shown previously inFigure 2.1.4 to obtain the initial net excess pore press.
78
(11) Apply to the soil specimen. Measure the
consolidation strain, £c-
(12) Find;
E = ES+Ee
This is a convenient simplification of the actual,
partially drained behavior. Here it is assumed
that the summation of the strains for undrained
shear followed by consolidation are equal to the
strain for a partially drained stress path
terminating at the same final effective stress.
Later, one might use stress path tests with partial
drainage to refine the solution.
(13) Compute a soil modulusE and Poisson's ratio ,v;
V Eva ~ fL ~ 2EV
Here we assume that the actual stress strain curve
can be replaced with a single modulus passing
passing through the actual stress level.
(14) Re-input parameters established from the test data
into the computer program (BRACE). For a series of
tests, contours of equal strain can be plotted as
seen in Figure 3.1. From such plots moduli can be
determined for any element by plotting the total
79
stress path on this plot, estimating the
"undrained" strains and adding them to the
estimated consolidation strains.
(15) Check against original total stress paths and
reiterate procedure if necessary to refine solution.
Below are summarized the main simplifications and
assumptions made in the presentation of the method;
(1) Deformations for an undrained unloading followed
by consolidation to a final effective stress point
equal deformations for a partially drained stress
path terminating at the same final effective stress.
(2) The actual stress strain curve can be replaced with
single modulus passing through the actual stress
level.
(3) The dissipation analysis is not too affected by
details of initial pore pressure distribution.
(4) Contour plots of equal vertical strain can be
constructed similar to the plot for the normally
consolidated soil in Figure 3.1.
Obviously this method relies on a laboratory testing
technique to obtain a soil modulus. Sample disturbance and
other factors influence this value. Another limitation lies
in the difficulty in correctly assessing the pore pressure
dissipation. However the writer believes the application
of this Stress Path Method type technique is a rational
approach for obtaining input for the computer program. As
Finite Element parameter studies have shown for supported
80
excavation, notably the Palmer and Kenney (1971) study, the
soil deformation or stress strain modulus is found to have
the greatest influence on the solution. One of the proposed
method's main objectives is to obtain and use this modulus
in a way which considers the effect of pore pressure
dissipation.
81
Para -TYPE EXAMPLE Predicted meters
Needed
F
Forces,ET U, ()h
Wall
H s,EDeforma- H c w
v Au withOff-Shore Structure cycling
FSStability Dam F
FS
Natural Slope
q, ht
Flow Stability k-- - of soil
Flow through Dam particlesunderprtce
from Lambe and Marr (1979)
GEOTECHNICAL PERFORMANCE
TABLE 3.1
82
Step Force and deformation Stability Flow
(1) (2) (3) (4)
1 Estimate stress- Estimate location of Estimate flowstrain pattern surface of minimum pattern
stability2 Select average element or select several elements3 For selected element(s), determine stress paths for:
(a) Geological history, i.e., past; and (b) design lifeof structure, i.e., future
4 On soil existing at or near selected elements in field situationRun tests along stress paths determined in step 3(Orient samples for tests same as in field)(Use field permeant in flow tests)
5 Obtain stress strain Obtain strength or Obtain permeability;or modulus strength parameters Measure any
movement ofsoil fines
6 Force = f(stress) or Factor of safety = Determine flow netf(modulus) average strength/
average shear stressDeformation = Average shear stress Evaluate stability of
f(strain or) from: Force polygon; soil to flowf(modulus) or elastic analysis;
or frnite elementanalysis; orstress paths
from Lambe and Marr (1979)
STRESS PATH METHOD
TABLE
0
--I
H -t 2
Il
I I --
3
ba
b
1
0.
V
3
~1*~ -~
------2--
62
0.70
=. 2 %
0.071%
5
2= (kg/cm2)
LAN|LLMNILLAS CLAY FROM LAMSE 1 WHrTMAW (1963)
CONTOURS OF EQUAL VERTICAL STRAIN
FIG. 3.1
--7-
1 2
00~w~
I I
a , , 3 -,
S 0 w
-EN-
84
CHAPTER FOUR
ILLUSTRATION AND APPLICATION
OF THE PROPOSED PROCEDURE
TO A CASE STUDY
4.0 INTRODUCTION
To serve as a vehicle to develop and illustrate the
application of the method to handle partially drained
conditions outlined in chapter 3, the writer will make a
prediction of the performance of a braced cut in Tokyo,
Japan. The excavation is currently underway and is
scheduled to take over two years to reach a design depth of
29.5 meters. Specifically, the writer will attempt to
predict the loads and deformations for a critical excavation
stage. The excavation depth at this stage is approximately
20 meters.
It is emphazized that the available soil test data is
insufficient. Particularly lacking is adequate information
on the stress history of the deposit, as well as a lack of
good information pertaining to the character of the soil
below 45 meters. Initial groundwater conditions (i.e.
static or non-static) have not been measured. Summarizing
the missing information:
(1) stress history (oedometer tests)
(2) u.
(3) limited soil samples and tests
85
Nevertheless, the main purpose of this chapter is to
illustrate the proposed method and demonstrate its
applicability. In the process a prediction is made using
the available test and boring data.
4.1 TYPE OF PREDICTION
In Lambe's (1973) Rankine lecture "Predictions in Soil
Engineering", Lambe classified predictions as follows:
PREDICTION WHEN PREDICTION RESULTS AT TIME
TYPE MADE OF PREDICTION
A Before event ---------
B During event Not known
Bl During event Known
C After event Not known
C1 After event Known
The writer's prediction will be a classification A
prediction. A prediction of early stages of the excavation
currently underway would fall into the class B category.
The writer did not have measurements of performance of these
stages prior to prediction.
4.2 PROJECT DESCRIPTION
4.2.1 General
The excavation is for construction of a pumping station
facility for the Nakagawa Sewage Treatment Plant in Tokyo
86
and is designated site B. The excavation support system is
designed as a cross-lot braced, concrete filled pipe pile
wall. Plan dimensions of the excavation are approximately
79 meters long and 52.5 meters wide from wall to wall,
internal dimension. Cross lot bracing will have inter-
mediate vertical support derived from columns supported by
deep concrete piers. Within the excavation, lime
stabilization of the upper 25 meters of soil has been
executed. Figure 4.2.1 shows a plan view and cross secton
showing these construction details. Also shown are recent
boring locations.
4. 2.2 Cons truct ion Schedule
Figure 4.2.2 depicts the schedule for construction and
pertinent events during this period.
4.3 SUBSURFACE CONDITIONS
4.3.1 Soil Conditions
Limited preliminary soil data indicated that the soil
profile consists of intensely stratified deposits of
relatively soft sandy silt, fine sands, clayey silt and
silty clay extending to a depth of approximately 45 meters.
Below these lies a relatively dense stratum of sand and
sandy clay layers.
The cohesive soil deposits are of moderate to high
plasticity. Blow counts (blows/30cm) in the top 30 meters
of soil are typically less than 10, gradually increasing
below this depth to approximately 50 at a depth of 50 meters.
874.3.2 Groundwater
Groundwater level is encountered in borings at
approximately 1 to 1.5 meters below ground surface.
Elevation on the boring logs coincides with depth (i.e.
depth = 5 meters, elevation = -5 meters).Unfortunately it is
not known to which datum (i,e., sea level) these elevations
are referenced to.
Unfortunately initial pore pressure data are not
available to verify static or nonstatic groundwater
conditions at this site. Static groundwater conditions are
assumed in this study. It should be noted however, that
previous work in the Tokyo area by Lambe (1968) at a
Kawasaki City Site revealed that pore pressures were
non-static due to long term pumping from the lower acquifer
55 meters below the ground surface. The total head in this
layer was ultimately reduced to -5 meters. Therefore the
static initial condition assumption may be in error.
4.4 NAKAGAWA PREDICTION
4.4.1 Initial Model of Field Situation
Following the steps of the outlined procedure in Chapter
3 the writer selected an initial soil profile based on the
limited preliminary soil data. A finite element mesh was
developed to model this excavation using the BRACE III
computer program. The mesh geometry was organized to
closely model excavation and bracing levels as well as
simplified soil stratification. Areas of anticipated high
88
stress concentrations (i.e. e
modeled with smaller elements
lements adjacent to wall) are
The initial profile, and
finite element
is a detail of
excavation and
Linearly e
mesh are shown
the excavation
bracing levels
lastic stress s
in
and
to
tra i
the initial BRACE III analysis.
to henceforth as Run A. Values
E ,/W equal to 75 which corres
/St= 250. These values were us
deposit. Input values of Poiss
were selected for cohesive and
w
po
ed
on
co
Figure 4.4.1.
wall zone sh
model the fie
n moduli were
This run will
ere obtained
nds rough
to model
's Ratio
hes ion les
ly
th
of
s s
t
e
0
0
Figure 4.4.
owing the
ld situation.
employed for
be referred
using an
o a value of
ent ire
.45 and 0.40
ils,
respectively.
4.4.2 -Stress Paths
To check the re
(Run A) total stres
selected "average"
excavation. Severa
in Figure 4.4.3 for
stress paths illust
unloading one might
of the excavation.
near the wall, (e.g
paths become less e
path extended be twe
s
s
e
1
from Initial BRACE III Analysis
ults of the initial BRACE III analysis,
paths were plotted on p-q diagrams for
lements both behind and within the
of these total stress paths are shown
this linearly elastic analysis. The
rate the compre
expect behind
Within the exc
., element No.
asily categoriz
en the initial
ssion and extension type
the wall and at the bottom
avation for soil elements
345), the total stress
ed, although a s traight
and final points indicates
the simpler extension mode of unloading.
2
Etx
4.4.3 Laboratory Test
General
Five "undisturbed"
Shelby tubes and 7 jar
89
s And Revised Soil Profile
1 meter
samples
long, 8 cm diameter
from recent borings
brass
adjacent
to the excavation were made available to the writer for
stress path testing at MIT. In all, 4 anisotropically
consolidated undrained triaxial stress path tests, 10
Atterberg limit index tests, and 6 combined sieve analyses
(mechanical and hydrometer) were executed by the writer to
supplement the available information on soil conditions.
Visual classification and torvane tests were also made.
Results of one unconsolidated undrained triaxial test
performed at MIT were supplied by Mr. Matthew Southworth.
In addition to the above, results of six isotropically
consolidated compression tests performed in a Japanese lab
were also obtained.
Soil Classification
Table 4.1 summarizes the index
classification of the tube and jar
through A.6 in the appendix give c
for the combined sieve analyses.
plasticity chart . Atterberg Limit
properties and visual
samples. Figures A.1
omplete grain size plots
Figure 4.4.4 gives the
s plot near the A - line.
90
Stress Path Tests
Figures 4.4.4a through 4.4.4d illustrate results of the
stress path tests executed by the writer for four
"undisturbed" samples of sandy and clayey silts.
In order to simulate the stress changes that the writer
predicted would occur in the field the procedure in Chapter
3 was employed. Based on the length of construction time
for this project, it was assumed that consolidation due to
altered flow conditions would occur, particularly within the
excavation where construction dewatering is inevitable.
Consequently the testing procedure consisted of the
f o 1o w in g 3 steps.
(1) Consolidating the sampl
stress state.
(2) Shearing undrained via
unload ing
envelope w
sample).
Stress Pat
subs tantia
The tes
pressures
decrement
(3) Applying a
to simulat
to the failure
hile maintaini
This procedure
hs from BRACE
1 zones of loc
ts were stress
were allowed t
of
c
e
the coefficient
Unfortunately no tes
es to the estimated in situ
ompression
envelope (
ng integrit
reflects t
III Run A w
al yielding
controlled
o stabilize
applied stress.
hange in pore pre
post shear consol
of con
ts were
ssure (
ida t ion
solidation.
performed whe
or extension
define failure
y of the
he trends of
hich indicated
Pore
for each
back
and
pressure)
measure
re negative
91
excess pore pressures were allowed to dissipate although
this situation is likely to occur in the field as well.
This is mainly because the writer originally assumed deep
dewatering was to be implemented, later to learn that this
was not the plan. Not having a sufficient knowledge of the
initial pore pressure distribution in the field, initial
pre-shear effective stresses for the triaxial tests were
computed assuming static pore pressure conditions.
Admittedly, this assumption may be in error for the actual
case.
Coefficient of Consolidation
Table 4.2 summarizes values of coefficient of
consolidation computed from both pre-shear and post-shear
consolidation. Both square root of time and log time
methods were used. The small load increment ratios often
yielded type III curves (Ladd, 1973) for log time and the
time for 100% consolidation was difficult to obtain from
this type of plot. Thus, this writer has more confidence in
the reliability of square root of time CV data. Figures A.7
through A.14 show typical square root of time vs. volume
change plots for the triaxial tests.
Revised Soil Profile
Based on the recent soil borings (October 1979) and the
results of the tests run at MIT on the samples taken from
these borings a revised soil profile was constructed for the
prediction. The profile is shown in Figure 4.4.6. Strata
92
below 45 meters was inferred from older borings in the same
general area.
The strength profile shown in the figure was selected
for extension and compression unloading. The selection of
this simple profile was influenced by:
(1) available undrained strength data -- although
high Su/ T, values determined from triaxial tests
(to be discussed later)indicated some over-
consolidation , useful stress history data is
missing and would help put such data in
perspective. Other factors such as high strain
rate used in CIU compression tests and the
unquantified effects of the lime stablization of
soil may have increased the test strengths.
Because of the lack of better information on this
deep deposit, the writer used strength data from
the tests run at MIT, selecting the lower values of
Su/T-Y to represent the deposit.
(2) Limitations of the BRACE III computer program --
BRACE III is limited in the number of soil
materials that may be input. The writer elected to
use the maximum number (20) of materials to model
different moduli at the "cost" of assuming a simple
strength profile described by two values of the
ratio Su/0 0o.
The CIU compression tests shown in Figure 4.4.7 are
plotted with the triaxial tests performed at MIT. The wide
scatter of
attributed
of samples
sheared at.
was applied
time to fai
2.2.12), th
CIU tests,
likely give
because por
subs tantiat
performed b
significant
increments
93
strength from the Ja
to the wide range of
from a given depth w
The high strain ra
is also a source of
lure as Skempton and
e measured strength
tested at a strain r
unrepresentative an
e pressures are not
ed by the controlled
y the writer which i
ly longer time (15 +
as shown on Figure
panese CIU
confining
ere consoli
te at wh
error.
Laroche
will dec
ate of 0
d higher
allowed
stress,
ndicated
ich
By
1 le
rea
.5%
un
to
st
th
tests can be
pressures a series
dated to and
the shear stress
increasing the
showed (Figure
se. Similarly the
per minute most
drained strengths
equalize. This is
ress path tests
at it took a
5 minutes per increment,
4.4.5) for pore pressures to
equalize during undrained shear.
Consequently the writer applied a correction based on
the Skempton chart. First, strengths for the sample depths
extrapolated from each series of tests were obtained by
estimating the strength associated with the insitu
octahedral effective stress.
ocT 3
In other words for
samples retrieved from
example, if two undrained
a depth of 10 meters were
tests of
sheared at
94
consolidated pressures of 5 TSM and 10 TSM, and FOc-r in the
field was 6.3 TSM the approximated strength was extrapolated
in the following fashion:
This strength was then reduced by 20% to correct for the
excessively high strain rate. Table 4.3 summarizes the
corrected CIU strengths as well as the other triaxial test
data from the MIT tests. The agreement between the two is
good for the compression tests. The extension tests have
strengths equalling approximately 50% of the strength
measured in the compression tests.
In comparison, undrained triaxial test data on normally
consolidated Kawasaki Clay (P.I. = 40 + 10%) with a slightly
higher P.I. from Ladd et al (1965) show values of s
equalling .445 and .225 for compression and extension
respectively.
Although the strength data showing large Su/:,0 ratios
indicates that there is probably some overconsolidation from
18 to 28 meters, stress history data is not available to
define the degree of preconsolidation. As previously stated
because of the lack of sufficient data, for the purpose of
this study the strength for the undrained deposit will be
assumed to increase constantly with depth equal to 0.43 V,
for compression and 0.22&,o for extension. These values
agree well with Ladd et al (1965) although they may be
95
slightly conservative values in the light of test data
available from 18 to 28 meters.
Contours of Equal Vertical Strain
Based on the available CIU and CAU tests, plots of equal
vertical strain were constructed. Figures 4.4.8a and 4.4.8b
show contours of vertical strain on a p-q plot drawn through
the effective stress paths for the CIU tests. The two
stress path compression tests TC-1 and TC-2 are also plotted
and show reasonably good agreement. For extension unloading
the two stress path tests TE-1 and TE-2 are plotted on
Figure 4.4.8c and estimated contours of equal vertical
strain are averaged through these effective stress paths.
These plots will be used to extrapolate strains for various
loadings at any depth to obtain soil moduli.
4.4.4 Revised BRACE III Analysis for Partially Drained
Conditions
A. General
As previously stated, this prediction will address the
excavation when it has attained a 20 meter depth. One
reason for selecting this depth is illustrated with the help
of Figure 4.4.9. This figure shows a simplified soil
profile and the Nakagawa excavation geometry for the sixth
excavation stage (depth = 19.5 meters). At this depth,
bottom stability due to uplift becomes critical. A simple
computation for the geometry in the figure determines the
critical depth, ZCR for equilibrium of the soil mass against
uplift;
96
-Z br - (2- xu)Y'
- 45x I. - C4-5-1-S9 lO
~/cm 2.1MIETERS
At this excavation depth, the effective stress is zero
at 45 meters and the upward force equals the downward force
(i.e., the factor of safety is one).
Consequently this prediction will treat the loads and
deformations to be anticipated for a depth just prior to the
depth where bottom instability is imminent. Two new BRACE
III runs B and C are made for this prediction. For the
final Run C updated stresses from the previous run are used
to obtain moduli.
B.-Computation of Partially Drained Moduli
The partially drained moduli incorporate both undrained
and consolidation strains in the relation:
w h ere:
w h e r e:
ET = -crA. vCn TiA^L STRAW E o= - AL. VO&..u.MET IC SnTrA W
1) Undrained Shear Strains
Using the linearly elastic BRACE III RUN A, the writer
selected approximately 50 elements inside and outside of the
excavation and tabulated the values of total stress change
in terms of p-u0 and q for the initial and sixth excavation
97
stage. (Table 4.4 shows such tabulated values of stress for
the final BRACE III RUN C). Total stress paths are then
constructed from these values. Using the appropriate plot
of strain contours,(see Figure 4.4.10)the total stress path
(minus the static pore pressure) can be superimposed and the
undrained strains can be scaled off between contours along
the estimated effective stress path.
2) Consolidation Strains Due to Dissipation of Excess Pore
Pressure within the Excavation
Given the final effective stress point established by
the undrained shear, the excess pore pressure ~u. due to
shear (typically negative in this case) can be obtained by
subtracting the value of (p6 - u) for the sixth excavation
stage from the value of 'p. Refer to Figure 4.4.10.
The average degree of consolidation for this geometry is
estimated using one dimensional consolidation theory. Within
the excavation the impermeable wall is likely to inhibit
lateral drainage. The Terzaghi equation for the
dimensionless time factor Tv is;
CVtV H
For a depth of excavation equal to 20 meters the time
factor within the excavation is computed using:
-3 '(a) C 3 x 10 c-wn/sec-nd
selected as an average value from the stress path
test data
(b) t :.St 5 8 inoriths computed from
98
Figure 4.4.11 to represent an average time for
unloading similar to the procedure used to compute
settlement time for embankment loading (see Lambe &
Whitman p. 414 1969)
( c ) HA = (45- L(.s) - 2 = ..LS vneters
The drainage path length for the consolidating
cohesive layer. This assumes double drainage.
The time factor then equals 0.05 for this stage of
construction for the soil at the bottom of the excavation.
For a linear distribution of excess pore pressure this
will yield an average degree of consolidation of
approximately 2 % obtained from Figure 4.4.12a. Having
observed the dist
within the excava
considering the s
dimensional verti
pressure the writ
degree of dissipa
any depth or disc
excavation. (Ide
and constructed t
of excess pore pr
ribution of
tion to be
negative
reasonably
excess pore
linear and
pressure
also
tudy by Osaimi (1977) which indicates
cal
er
tio
ret
all
o r
ess
u
n
e
y
e
U
degree of dissipation Uz
pressure one estimates t
dissipated. This value
to be added to the posit
due to surficial or shal
dissipation o
ses Figures 4
for negative
element belo
a plot of Z
present the a
re.) By mult
t i
he
is
ive
low
me
ac
th
s t
tua
en
he
1
ta
excess
dewate
f nega
.4. 12b
exc es
w the
vs. Uz
c tua 1
iplyin
in it i
nega t i
bu late
por
ring
a one
tive excess pore
to approximate the
s pore pressure for
bottom of the
would be derived
initial distribution
g the specific
al excess pore
ve excess pressure
d (Table 4.4), later
e pressure dissipated,
within the excavation.
99
The excess pore pressure due to shallow dewatering can
be analyzed using Figure 4.4.13 for a change in piezometric
level at one drainage boundary only. This case also
experiences double drainage and therefore the time factor
will be the same as previously computed. The dissipated
excess pore pressures are obtained using this figure and are
computed for various depths and elements (AuD).
The summation of Au, + 6u = Au, gives a net value of
dissipated pore pressure in TSM. Using Figure 4.4.14a
and/or 4.4.14b based on the post-shear consolidation from
the series of stress path tests, volumetric and vertical
strain can be extrapolated for a given depth. Since stress
path tests with dissipation of negative excess pore pressure
were not run the plot in Figure 4.4.14a shows inferred lines
for this case. Since this is an unloading problem, the
strains due to dissipation of negative excess pore pressure
are not likely to be in the zone of primary consolidation
and therefore are likely to have a similar relationship to
changes in pore pressure (but opposite in sign) as the early
recompression stage of the stress path tests.
(3) Consolidation Due to Pore Pressure Dissipation Behind
the Excavation
The steel pile wall is assumed to be impermeable. The
steel interlocks between the piles have been grouted to
prevent leakage.
100
Below the lower silt layer the total head is assumed to
remain constant and this stratum is assumed to have free
access to a distant source of groundwater.
Given these assumptions, dewatering due to shallow
pumping within the excavation will not affect pore pressures
behind the wall.
The stress path test TC-1 indicates that the upper
organic silt has a coefficient of consolidation of
approximately 7 x 10 cm /sec. Computing a time factor for
this doubly drained stratum for the time associated with the
20 meter deep excavation we get:
T .0 1
and from Figure 4.4.12
U(k 12 */A'VG
similarly for the lower clayey silt we compute
.03
A'/G
Continuing with the use of one dimensional theory as was
used for within the cut, strains are estimated for consolida-
tion and added to the undrained strains obtained from the
contours of equal strain plots. See Table 4.4.
(4) Contours of Moduli
Contours of moduli can then be constructed across the
finite element grid using the tabulated values for the
101
discrete elements analyzed. Figure 4.4.15 shows the
contours of moduli in tons per square meter. Stress strain
moduli for the granular layers above and below the cohesive
soils were estimated using empirical formulas based on blow
count (Mitchell and Gardner, 1975) and compared with ranges
of modulus for granular soils in the literature. The
modulus in the layered lower sand and sandy clay stratum was
reduced from 1650 outside of the excavation to 1000 within
the excavation to recognize the fact that the shear stress
level in this interior zone is higher and therefore the
secant modulus will be lower.
The values of modulus from this plot were substituted
into the BRACE III computer program by initially assuming 14
soil materials (moduli) and assigning the appropriate values
of moduli to the finite elements. Total stresses from this
computer Run B were used to improve the solution by
determining new or updated moduli as shown in Figure
4.4.16 The final Run C was made using Fig. 4.4.16 and 20
materials. The current version of BRACE III will handle a
maximum of 20 soil materials. Ideally zones of soil with
similar modulus and Poisson's ratio would be contoured,
since these values can also be estimated as shown in Table
4.4 for the final run. Unfortunately, the current version
of BRACE III is as previously stated, limited to how many
materials may be input.
Ultimately the writer observed that in general
Poisson's ratio will follow the trends outlined below:
102
a) Zones Inside the Excavation
1) where dewatering prevails and the net excess
pore pressure is positive, dissipation will
tend to increase Poisson's ratio
2) where the net excess pore pressure is negative,
dissipation will tend to decrease Poisson's
ratio
b) Zones Outside of the Excavation
1) where the net excess pore pressure is positive
dissipation will tend to reduce Poisson's ratio
2) where the net excess pore pressure is negative,
dissipation will tend to increase Poisson's
ratio
Ultimately an average value of 0.49 was selected for the
cohesive soils, this value at the upper range of typical
values (0.2 - 0.5) usually used in finite element analysis
modeling of soil. In general the undrained strains in this
case tended to dominate the deformations.
(5) Ko
Ko = 0.5 for all soils was selected based on review of
stress path Ko consolidation data on these soils, the 1-sin
i equation and empirical correlations of Ko versus P.I.
(6) Anisotropy
To account for anisotropic shear strength, The Davis and
Christian (1971) yield criterion was used in BRACE III.
Also, any yielded Elements are assigned a modulus equalling
103
1 to 10% of the unyielded modulus, depending on the initial
modulus value.
C. Predicted ,Loads Deformations and Stability
(1) General
From the final BRACE III run, contours of
deformation and stress are shown in Figures 4.4.17
through 4.4.23. Figure 4.4.23 is a contour plot of
undissipated excess pore pressure and is intended
to show the approximate trends prior to
consolidation.
Related aspects of techniques used in the BRACE
III computer program to model the soil structure
interaction are discussed in this section.
(2) Deformations
Figure 4.4.17 illustrates the deformed geometry
after the sixth excavation stage. Inward movement
of the steel pile wall towards the excavation range
from 12 centimeters at the top of the wall to a
maximum of nearly 44 centimeters at a 27 meter
depth. Bottom heave varies from 40 centimeters
near the wall to 76 centimeters at the center
line. Figures 4.4.18 and 4.4.19 show selected
lines of lateral displacement and bottom heave,
respectively.
Maximum ground surface settlement behind the
excavation is 20 centimeters occurring 25 meters
behind the excavation. Figure 4.4.24 compares
104
ground surface settlement with Peck's chart. Note
the uncharacteristic heaving of the excavation
face. Ground surface settlement begins to occur at
a distance of 3 meters behind the excavation. In
the 3 meter wide zone between the wall and this
point a very small net upward movement occurs. In
the writer's opinion, since this is contrary to the
downward movements we would normally associate with
past experience, this lifting of the excavation
face is due in part to the way which the BRACE III
computer program models the soil-structure
interaction. When excavation release is simulated
the upward forces applied to the nodes on the newly
exposed surfaces of the excavation include an
upward force applied to the nodes shared by
sheeting elements. Consequently, as the bottom of
the excavation heaves, the sheeting may also heave
slightly.
This behavior is enforced by the technique
intended to allow the soil to slip behind the
wall. This technique involves setting the axial
stiffness of the sheeting equal to either the soil
stiffness, or zero above a selected level (usually
above the depth of excavation). BRACE III runs in
this study set the axial stiffness of the sheeting
equal to the soil stiffness.
105
Simulation of strut installation
fixing the node at which the strut is
the horizontal direction. A BRACE II
study by the writer comparing this me
fixing the node in both vertical
is made by
applied in
I computer
thod with
and horizontal
directions
directly b
method. A
mented in
the node i
approach.
modeling t
BRACE
install at i
a) no
b) ze
as a p
s trut
indicated that more settl
ehind the excavation face
ithough this technique was
the current BRACE III anal
n both directions may be a
Therein is evidence of th
he soil structure interact
III Input options used for
on in the final Run C
prestressing of stru
ro crushing of shims
ercentage of movement
level
inc
ts
(or
alr
ement
w i th
not
yses
more
e dif
ion.
s tru
occurs
the latter
imple-
fixing
realistic
ficulty in
t
luded:
timber wedges)
eady occurred at
Al th
employed
Run B, s
revised
ough
for
h im
Run
c
B
no pr
the i
rush i
( the
e
n
n
stres
itia 1
g was
first
s ing of the struts was
linearly elastic Run
set to 50%. The firs
of the two runs using
y drained modulus) showed
movements (up to 1 meter
agnified by the crushing o
Acknowledging this error,
C assumed no crushing of s
large inward
at a 20 meter
f the
the final BRACE
hims. Unfortu-
1
g
m
A and
t
the
partial
shee t in
depth)
shims.
III Run
106
nately, the writer does not have knowledge of the
actual prestressing and shimming techniques and
details for this excavation.
Movements above the excavation level are also
as Jaworski (1973) states, a function of bracing
details. The selection of the related input
parameters are usually drawn from the engineer's
experience or knowledge of actual performance in
the field.
(3) Lo ad s
Figure 4.4.25 shows the strut loads and total
lateral stresses behind the wall. The negative
values in several struts can be attributed at least
in part to the lack of prestressing in the struts.
Furthermore, since for early stages of excavation
the sheeting typically behaves as a cantilever
followed by a more convex distribution of
displacement, the tendency for the top of the
sheeting is to bend back away from the excavation
during later excavation stages. This would give
negative load to a non-prestressed upper strut, as
indeed the upper two strut loads illustrate.
Consequently the load is shifted to the lower
struts. The unusual negative load in one of the
lower level struts is probably a result of several
factors. While lack of prestressing may contribute
to this result, Jaworski notes that four to five
107
sheeting elements between strut levels may be
required to accurately model the sheeting
behavior. Otherwise,uncharacteristic sheeting
forces may be transferred to the strut. Only two
to three sheeting elements were used between strut
levels in this study.
4., Stability
Stab il ity
is typically
analysis for
for an excava
analyzed using
shallow excava
tion of these d
Terzaghi's bot
tion (H/B< 1) a
imensions
tom heave
s shown
in Figure 4.4.26. Bjerrum and Eide would be used
for cases where H/B> 1.
basal sta
indicat in
meters.
beneficia
support p
of these
view area
For th
bility ana
g instabil
Also note
1, stabili
iers have
columns is
of the bo
e Nakagawa
lys is
ity a
that
z ing
been
a ppr
t tom
t
f
e
n
0
0
For comparison recall the
made in section 4.4.4A
an excavation level of 2
or both analyses any
ffects of the concrete
eglected. The plan area
ximately 13% of the plan
f the excavation.
excavation, the average
strength along the assumed failure surface can be
estimate
(1)
d by:
selecting several average elements along
the assumed failure surface
1
108
(2) constructing stress paths including the
effects of pore pressure dissipation
(3) selecting a strength for each element b
on an assumed failure loading path star
from the final effective stress point
determined in (2)
(4) average the strengths and input the Ter
equation for Factor of Safety shown in
Figure 4.4.26
The extended time for which this excavation
remains open has resulted in a dissipation of n
negative excess pore pressure for elements outs
of the excavation and for the majority of the
elements within the cut. Consequently, the soi
strength decreases with time as the soil swells
the water content increases. Hence in the
stability analysis the actual or "partially
drained" value of c is input into the equation
below. This method is used instead of the comp
used " 5= 0 concept where Su, the undrained
strength is substituted for c;
Factor of 1
Safety H -c/B
ased
t ing
zaghi
et
ide
a
and
monl y
Although the
simplification,
"I5= 0" approach is a convenient
if it is used for this case the
109
Factor of Safety may be unconservative.
Figure 4.4.27 shows the excavation geometry at the 20
meter stage and several "average" elements. Several of the
stress paths for these elements have been plotted in Figure
4.4.28 and an average strength selected from strengths
estimated from the stress paths is 8 TSM (refer to Table
4.5).
The r
Using the
29.5 mete
some erro
the s tres
different
strength
assumes)
depth is
esulting factor of safety using this method i
same strength for the final excavation depth
rs, the computed factor of safety is 1.1. Al
r is involved in making this extrapolation be
s paths for the "average" elements will be
as well as the strengths. In other words,
does not equal a constant (as the Terzaghi so
and the factor of safety for the full 29.5 me
probably overestimated.
s 1.6.
of
though
cause
lut ion
ter
110
NORTH
80-' M
APPROX. 4CROSS LO
3.250
2-1 - iiIK?~ Jz54 W"
LOFT BRACING
- 6~ 16 I 6 * 6 6 6~6 I 9 U ~TYTI
o1
3.25
80 7.013.2 5
1.52 M DIA.CONCRETEFILLED STEEL PIPE PILES
9.0 1 9.0 1 9.0 17.5 5.5 7-5 6- 15 7 t3.25
2-5 M DIA.VERTICALSUPPORT PILES TODEPTH OF 55 M. (TYP.)
PLAN VIEW
50-5 M±ZONE A
15MtZONE B VZONE A
(0.7M C/c -ALT 14MAND 25M DEEP HOLES
(.4 M C/C)
0.4M AUGER HOLESSTABILIZED WITH LLIME -25
,- *
CAST-IN- PLACE CONCRETE PIERSFOR VERTICAL SUPPORT COLUMNS
CROSS SECTION X-X
PLAN VIEW AND PROFILE OF NAKAGAWA SEWAGETREATMENT PLANT EXCAVATION - SITE B
FIG. 42.1
'-RIN
1-2.N
-i-I19.2SM
K
34-75M
15M±I
if "*/'-*"jjll1l1j.14-
-25-
I W f
w
wa
-0
-- 10
-20
-30
-40
50
L60
2
I
) & cnii
T I
, , , , , ,
SOIL5TAB i LlZ.)RIVE STEEL PtLE WALL
- 1 L'C6NCRETEF PIER 5
- 3.0
I . , , i I I I I I II I I I
1978 1979 1990
EXCAVATION STAGES - (ELEVATIONS M ETERS)
6. o -03.5 -14.0 -rr.0 -20.0 -e3.0 - 0- 29.5
1980 1991 19 3
SITE B CONSTRUCTION SCHEDULE
F IG. 42.2
ift,
38 ~s Wo 2ZZ
i-i~ I-r '9
"7
23g-3
'7
,7
G07 4--7q
5-04
5,01so
30
4ii49749'495444
4149489 -
4821 -4U.
44 9
4-9
s'l
4-1
4-0
38
37
'4
'3
2
Ci
"3
97
t4-
9+'
" -
9e'
95.
--32
30
3o
4
3
G.
(-7-
go'I
SAND350
$1~ L
.=W 0
SILT
aSAND
4) -35'SANDflCp =L35'
sf .3Z
0-
V20
30-
4o_
50 -
50
60
3-
33 44 0)
- ^
40
,A4-
N
'9
'4
-J
0
8
'9
8
I.,
'3I.N
'4
'I
1
ALL SOILS: Y-, =- %15-fCM E/,=75) K,=0.5
54
5-4
4-3
28
25
11)
/
2
4I
I"
-T73
.7'
70
INITIAL PROFILE AND FINITE ELEMENT MESH FIG.. 4.I
7
4-7
4-
45
44
f.
I
493
*
C'4-
08*
-V t ; t2,3
2 24
* e 9O
EXCAVfli BACM0 srAGE
/0-
20 -
30.
40
60
I II I I -I I I0 /0 20 zs 3o 40 So o 7o
DETAIL OF FINITE ELEMENT MESH IN ZONE OF EXCAVATION F IG.4.4.2
8
6
If
z
0
TYPICAL TOTAL STRESS PATHS INITIAL BRACE III
ANALYSIS LINEARLY ELASTIC MODULUS FIGAA.3
/06/ *
tJ
i~/ ~
- I
4
/0
I
4~1
a
115
ST = SHELBY TkAEE. MOTE: SANOy SPAfl>JF - No roiz'RAlZ
SOIL INDEX PROPERTIES, AND CLASSIFICATION
TABLE 4.1
-rMf A M P E uo S W
MIETEKS T/ ( 0/ .s,.e F
- ST .4 (.3 33 3o 2.7
-. 9 2-2 ST r.8.Z 59 'A0 O - 3- OZ-MH
1c30 3 51 44.Z 53 3/ 22 3.6' 22. "
21.4 2-2 ST 55.4 57 3/ Z( 3.2. - MW
2.o 0s.5- 5) 29 22 4. 3 CH -MR2-1 ST CTE-2)
-35 93 6 -- . ML
31.3 Z-Z 7AW - 38 2- (3
3+.o Z-1 Z- 40 5 41 3z 20 S 2 ML(TE-0
34-'> Z-Z .15' 37 2IZ 25 - 'z cL
4-2.15 z-Z _T Z_ 'sZ Z-zs*.5 - *3( H-
r I I
00
0
CH
7oeeV
_ _ MH _ - _ _
m a I I
50 go 70 80 CI0 /00
IJQLUD 41/-I7
PLASTICITY CHART
FIG. 4.4.4
K 56
z
U
/0
0
CL
x'7
4-
/0 .o 30 4o
i if
z
401 i --
oo Z- ---
o< MLG
'ESS-
~5?-
5- /0/5
P P T5M
-~ T/0 ' Zo
/5
/0
5
E
0/
(3 4. 5*
C0MWR~E~S seN')
/v7zs :
j. STRESS Cc1NT ot.LEID C K,.C(u)
2. BORt4G Z-2. DEPT" 10 MEnERS
STRESS PATH TEST TC-1 FI G..4.5
/
zg
FTL
u
3-
' ~7' 8
,~ -'a-
eZs'
/0
/ 0 /5 6Z-30 3z..
E
______S ___ __ ~_
5" /0
/
E Z
5-
A/oTES :
1. SrRsS COOT I.nLET- CiUC (u.)
2. BORIWG 3 DEPTH 19 METERS
30 >-5
STRESS PATH TEST TC-2
1.'
CT
/0
5-.
ESP
0
S
"-34 '4J~
~
~J)1.A) 0U
0 V-"
c
-1-sm
FIG.445b
~20
I I -
z5- -3--i
-/0C -A TMIJSIOhJ)
EvJEr CAIW .
3 I0
STRESS CONTR-kEk> Ciz. E (u.)
2. BoRiNG 2-I DEPTH 34 METEAS
P -F StSTRESS PATH TEST TE-1 FIG. 4.4.5c
zI-
0~
-
o-(
/~ 2. 3 35
[ _ _ _ _j _ _ _ _
U
t~7I
4
0I
I 2
/V
i
_
11 ~ j
s-+
i
zG Zs5 "N30
/0 ______________ /0_
6Z5-
0 __00 . pp SA 25
-/0 __ _ _ _ _ _ _ _
7
o-
in
5 3o
p
/0 -
~10
-.5-.
I
84- ' 7
(ExTEInsIL)
E V~P.TCA
k)0
NoTECS.*
I. STRLss CoNTRO LLED C<LE C.)
2. SoBC r-IG 2-) IDEPTH 28 tAETERS
FIG.445dSTRESS PATH TEST TE-2
U)I-
tvc
15'
Triaxial TC-1 TC-2 TE-1 TE-2Test
Depth m 10 19 74 28
FittingMethod log t V log t logt log t
0.27 0.30 7.1 1.1 4.5 0.74 0.26 0.60Pre-Shear 0.03 0.12 3.5 2.0 2.1 0.65 0.22 0.33
1.4 5.0 0.26 0.85 0.26 0.721.15
Average 0.15 0.21 3.0 2.7 2.3 1.3 0.25 0.55
1.3 3.1Post-Shear 0.08 1.8 4.6
0.06 5.0 2.0 0.602.8 0-30
Average 0.07 3.4 2.7 1.3
COEFFICIENT OF CONSOLIDATION CM/sEC x 10 -3
SUMMARY OF COEFFICIENT OF CONSOLIDATION
DATA FROM STRESS PATH TESTS TABLE 42
0 TV 4S MIEIA$ iMASO 4V
0
'I)
441~-114 So -
4-
4.0 -
60
B840VAIS 7-1, zz,-
- U..Aw. s6 P2AY
F/A/E 5/9A/D
a. etA
whelt& .9C YEYS
Y*/. 75c g
-rPUS 7. - , - .A1DI M . SA All)
LT. GAYF/*-. SAM01
Y eI.S 7C'f
SILT
3/uTY CL4
S-54AJD
.- ~AJO CY
r32*
soIL COAJnoafs FA.&tj 4Cl
040o A &)* A'A 04ri $SPM.G l5CAfE499L Aite4
8 oftims 3aopteW4 1-1
I >
01 I
0
bp'
o lo z6 3a 40
DG ~ L~
Wp WJI'
LK
0 20 40 (0W..ATAX COW-rsr *.
REVISED SOIL PROFILE
:1 U
K
19i.-- --
+ IT- AP
=.43I ASSU
TEMTc-TO
MED
3g.Z2
so s - ,e '- to cs~
F IG. 44.6
rI'Q
123
00
20su t/ma
30 40
= 5,10,20,&.46 t/mL
UNDRAINED STRENGTH:
TRIAXIAL TEST RESULTS VERSUS SAMPLE DEPTH
FHG.4.47
boring2-2-2
0-
10-
20-
.30]
40-
50-
60-
0-0
-- 0
Lu
Ca-EUU
cmuc-
CAUE-1u
I-10
I0
LUU
S /ow 46 /oZO46-
C7
124
DEP77 7uST - - a CtA7sm As rsmE
/o.zt C/aCU) 7.1 5.7 (AZTER.5) (75A4)
/0.31 C/uc-) 5.? -9 4 4 7
/0.3+ Lt1L() 3.9 43
o.3± CALUCc') 3.9 -3 - 43
/9.4± cia.c-G2 /*. -/45 7 93 . Z. >
/f.3 CA UaC(t) c. 0 / 7.e(, /0./ .51 TC-Z.
Zo.± c I tcL) /4. f / 8 /6.7 i/., .77-
26.4± cC C J-) 16.9 20 /0.Z ~ /3.7 /3.7 .66 @
29.3± CAWE (k) 4.6 22.0 // /4.7 .-27 -
34-3! coquE(u) 6.2 27.5 /3.7S( /8.3 - . 22
/3.0 Z8.3 /4./ /8. ? /0.1
I I ________ I _________ t ________ ________
.37
C/l" srREcA-N3. EXTRAPoA.ACO FROMl 4 SERI6s OF TESTS 4T D/FFEREA)-CoNF/A// AJG PRE.5.5$ ES (SEJE TEXT)
0 7,EsTs RUA/ AT o ST RAI /A) R 4TJE OF o.S; /, P.Q. MmotarE
SUMMARY OF SHEAR STRENGTH DATA
TABLE4.3
35-.3 t
3C
2E
2C
5 10 15 20 25 30 35 40 45
dep+h = 10.00- 1tO1 'm
0
2C
\u i
I r-I
.- 17- -_
)-
5 10 15 20 25 30p t /tIr
L -rc - I at ., %Lx I MT stcain
35 40 45 50
cdept1a. S I..9--.o '
ESTIMATED CONTOURS OF EQUAL STRAIN - COMPRESSION
FIG. 4418a
125
%
-(4
.) *Ir
q -t/I
5
C0
25
IE
i0
E
O
12630
5 10 15 20
-re- t bo - % 3 0 1 p t / dr
L i, 4*. STAa
3C
2C
2C
25 30 35 40 45bor qK 2.-i
deptK -z.00 - It.som
5 10 15 20 25 30 35 40 45
Ds
borvn ?.-I
de ptK = 35.o o -3S.To -m
ESTIMATED CONTOURS OF EQUAL STRAIN - COMPRESSION
FIG. 4.4.8S
1 1
-1
15
10
5
050
9 15
IC
C
p t /'Mz
i i
47.
-%
25-
20[ ll.
----
5
10 ---- --
TE-2 TE-
100
q t/m2 o -__
0 5 10 1S 20 25 30 35t/mO
ESTIMATED CONTOURS OF EQUAL STRAIN - EXTENSION FI .48
TY ,,,
CLAYEY SILT ~wfs 1 ~Fis5arny c,7 x 10 cm/s
CLAYEYSILT with
f sand lenssand 4hl Is
-3c = 3 x 10 cm /s
rN)
-GANDdSANDYCLAY L AYERS _
SIXTH EXCAVATION STAGE SIMPLIFIED PROFILE
FIGA4.9
0-
10-
20-LI)w -
F-
30-
0 40-
50-
6 0 i
-
-
~IS/ODE £XCAVATI O4
Xf Ja ag *Acff ulsusrsot aLi lt u -a 4LLO V CcE Ej
227T 31.3 Z5.( 7.9 135' Iz.5 io. 5 I.ss 4.o - 915 o.94 - 194 1.4 -oi -0.o4 -oo 1.4 o.-1 %.3 ~~.14 .63 382.
231 30.9 18.6 f..- 10.6 3.0 9.Z o.9 lc-. - 1(..o 69 - 6.1c -0.45 ~o.o -o.o -o.a. 1.S -0.9 ::8 ~0.32. .52. 333
2357 24.8 14.9 5.0 8.6 01 9.0 o4- 14.3 .24 ~-7.o 0.17 *3.5 2.9 *0 *.5 o-1 +o.I1 s.5 ~os 'I-C.1 '0.S .35 3S3
Z94 4,.,' U. B 81 5.0 13.) 9.4 I.S9 1.1 - -13.o o.-7Z - 9.4 -. 4 -.2-1 -. 0 -. 09 ~ -0.z t. 'sl- Z5 -93 45?.
290 29.'5 1. 5.9 10.1 -92 7.8 8.1 11.1 - -1.0 0.ot - 0.4 -0.4 -0-1 -o-oc-o.of 'I.I ~SC *l.o4 -. C4 .59 3.5'
29S Z1.8 13.1 4.4- 3.3 ~4.7 -5.S 0.19 IG.4 63 -130 o40- *.3 -7.3 +3.o + 4+ 4 o.4-7o +35 -. (, 3.3 .5 3o0
342 43 Z.I C1 7.7 9.8 8.4 9.8B 1.1 - -'1i. o.- - -11.Z -o.3 -o. -6. 4'o- lo-o a o -15 - C5
3+4 3(.3 2S . Z -7.3 & .+ -3.:3 7 r .3Z -- 1.0 0.6-4 -0-77 -0.7 ~0.c, -b-ov-o-oa 70.~ i - 6is 3/7 ./39 0.1:3 (S348 Z3.s 1.4 5.9 S'.2- - 3. -5.G 011 11- - -14bo o.z - -os -'s 3 -o- -oS -O..s -0 .o *-S . Z241 0.43 4SI
352 23.3 14.b 4.' 4- -11-0 -4.2 0.30 1S.3 .44 ~18.0 0.33 . -&.1 *a.9 Io 4-s +.2 . LX -4,5~ *Z.7S 4.A Z.47 0.G8 416
398 43S 24C.1 87 T.I f7.1 C.5- 1,8a 9.1 - 17. 0 72 -- -'1.2 -1?.z Z.3C /Z 7-t -.o5 tor -17 09S' ~.5. 2588
3-1 313 7.3. 7.9 r_9 5. 4-o i.ss- +.a - 181 o8 - _? - Z '_ Z6 07 0-c l os L4 -o(s ~.1-4 S405- l7. 94.1 S.(& 49 -7.3 -3:3 o .(.5 12.1 .03 -15 6.04f '04 0.7 0.3 -. 17 -oj -o4 -Z.3 4.15- -2-34 f.i1 -. 47 ~70
401 3o. 20.4 1. '.O -1.2 4.1 1.14 7.8 - o -01,5 - -0.14 -6-14 ~o4! -015 0 19 ~3.0 1-9 -3oIS' 1-495 .45 7Z3
427 39.3 Z3.4 7.9 6.9 3.2 -4.o i.5 4.0 -- 0 0.0 I - ~014 -0.96 ~.06 7~.oz oz -Z.S '1.-Z -Z.9Z 1.Z3 o.49 952
410 20.3 12.1 4.1 3.( -Kg. 4 -Z.z 0O(. 11-4 -7 Z9 o.8g 'IS- -15:1 14-1 ~-~1 ~.7 :51 -1.! 0o-~5 ~2.o7 f./8 0.09 409
454 435 12. I S-1 7.7 (-t 3.4 1.,R 1. - -7.v o:72. - ~12. -9I. IL 7. 1Z. ~%-z ~o.t o-67 ~.Z7 4 ' .17 39U.
4S5 31.3 Z3.( 1.9 r.. 2.4 -2- 1,S 4.0 - iZ.o .9 - -1.9 -1.9 -K . S -0-5 ~-S10 05 -/s- 1.45 .43 938145'1 34.6 20.4 .8 .o ~3.z -2.1 1.14 7.8 - 13.5 0.09s - o.z --. 2 O3 -'os ~ 03 -/.o *0.5- -/-1.3 ".47 -4. -7 8
4(. 74.% 18 6 .3 4.G -/o-o -25 0-57 /?.9 .08 -/1 0 . /.o -1.4 84 IL 7071-.5- -/ to.- -/.57 *o- .4-3 1/63z
483 35.3 73.6 7.9 4.I */.3 -/. /.SS i. o - 3-. o./( --. -Z./ /2 .64 .04. -0-s .ozr -. 54 #.j. 4o 3333
488 2M.3 17.4 5.9 - 7.Z "43 .?7 .I - o.o oo -0.4 oO -. 2 -o4 .4 -o 9- o .- f -1 . 8
414I Zb.3j /2.2 I±f.1 3.C -s.i-z.-ooC 174-
. COA4PESSE ISTR.AIS 455uMEo -PosTwE
23.o O.84 T;7
SUMMARY OF DETERMINATION OF PARTIALLYDRAINED STRESS STRAIN MODUI i TA BLE 4.4
~ 57 -/S I.07n -7.07 -. /g x0(607f- f -/9.81-0.7 1~/.7
BZH/MD EXCA9 V47TOA/
Eess ExCE0 QissfPorfb
M* fp.,.9. /aU AtY *-C 6 u Avet avc E,, E., IE44 E %f iEN E,/5 288 17.3 5.8 '7. /C S (-.7 88 o - -13 C1 013 -013 C> o a o a .46 - z o.us45 0-So 4-00
/9 34'i Zu.9 7o /2. o /9.3 9.3 /-3 0 - -2. o o08 08 0 0 0 +O-7 ~0.15 0. -~3 6Sb 388
ZI i / 0c /0.8 3.( .Z r.+ 08 0.4 o c.o 0 0 0. -zd,4 o4 .iS 0553:s
2 .S o z2.8 7.(p 13.11 Z1.1 9.Z 1.52 Z C 0.13 0 Z4 7U4 :4 7 0? 702 #61 -4 t0-78 7. I o.94 4-lb
7 Z5 5 /s3 5.1 8.6 /3.9 C.7 0.41 C - c> .o' o Cd c1 0 0 0 'o SS ~o.W "o-56 -obx o.Sb 5
30 13 ./ 2.7 4., 7-/ 4.6 C- - - ) 0 0 0 C 4 'i.2Y ~4 k.Z5~ 0.6-0 ZC
3( 3.8 l .4 /1-0 A9 e8 /.o 0 -- 3. .01 0 0.04- o 0 0 0 1. o.. -. 0., 0.5- 400
38 a. -7 3.7 7./ lo.1,'-S o-1 3. o 42. 0 /44 -/44 -/.35 -0.4& -4 Y/.I o5' *U.6 ~/.o 1.54 5-9~39 /3s 8. Z.I 4. 4.7 4. - - 0 0 C o 0 a 4 .4 -70 +14 -0.7 0.5- 2C
44 386 Z-8 ~7.( 3.1 /?./ /0.7 /.Z o - -4. 3 o..3 o ~oSZ -. 52 :3 - / -o 1/.8 -o. .79 -1/ OSI 3441
44 ZGS 17.1 .7 y 8 13-1 6.4' .8 0 - -3-1 0.01! 0 - 05 o s a /.4 ~. -. ~ 0.s-o 344
41 Z03 /Z.z +.b 70 -7& 4.3 0.3o 0 - -3.o 033 -1.0 -. 0 // -. 36 f73 .4 t.sz -13 /66 5,5-
5(. 38.0 Z7g 7.4 /5.1- /4-9 /1.3 /.Sz 0 - -(..o oa-3 c -0.78 -o.78 ,63 .o- 6f - Z, S /.Z 2.45 -/. z. os/ 3oo
59 IT-o 1. Z S4 q-3 /.9 . 0.7( a - -6-0 O.Oz 0 -a./ -0.10 0 0 0 /4 o.7 1.4 .7 o-So 38.
4S '.o r.o 2.0 3.4 5. / .7- 0.47 o - -2.0 C 0 0 0 0 0 (0 O5 zS SOS -. z5 OS-O 30o
97 414 24.8 e.3 14- %cb.5 - /.7f o - -i.o o-z c, -7.14 -7-14- -o.19 ~o.o6 0.0(. -Z.S '/.zf 'Z.4- 7.3/ zS 331
91 33.. Zo.Z 4.7 /0.1 //Z/ a - -- 5 -o4 0 -0.38 -08 4, 06 0 4O/4 .9.5 1.9 0-9S 0. -0 3S9q 7- 27.4 /6.4q 5 .4- 5.9 . o 0.79 0_ - c1. & _ -0. Z3 -0.23 -. / a .32 -3 z7 /-2__ -72 ss- S .5 4 43
75 i-4. /zs #8 z 73 8.i (-.2 0.37 a - -3.0 o.23 o ~o.7 -o.-7 0.4 - z - z /.0 -0.5 "0.8 -. 70 0.83 489
/3? 43.5 Z./ 8.7 /5-0 iz-.3 /5.9 /.g o _ - -1-o 6-76 0 - 12.5 -0.3C 0/Z -0.1Z '35 779 3.39 -/97 0.55 308
14- 139 3 ?_ /373'.1( 721 135 ._ - -15.U 0. . 0 .0 ~3jo -0/s -o.65 -O.- "3. Z -/. "3.15- -/(5' 0-5-2 .317
/45' Z13 /7.4 5.1 /0./ 5.0 9.4 0.9. 0 - -10. 0.61 0 -. / 1 0 0 OV5~ -6.7S /.57 ~o.7f aSo 333/5b 2/8 /3.1 4.4 75 /0.o 3.-7 0 0 __- -4.o 0.19 0 .7(. b7 '.7S- ~ 25 0+ #0.2 -4 45 C - /~
5-s 11 4.1 Z./ 3.( 5.0 2.0 0.9_ 0 ~0 -1.0 -o 0 6 o o L O o .o33 -0. -033 -o0S* o-D 300
£MitoR b'E 7- CapWUT-1- ( .RO5
TABLE 44( Cc. ork"LI)
0
MerE: o ~e tkEs ti' A/Or
'r"SP FPoA ORACE 22
RSP £S'4 TEO FRop-7Ms P*r T E-2 ~--- TEI
S. *'
\0OI1-0I
15 eI
- 5 10 15 20D t/ma
-0
_ - - -
27.
25 30_ __ 35
SUPERPOSITION OF TOTAL STRESS PATH
ON CONTOUR PLOT OF EQUAL STRAIN
F IG. 44.10
5t-
\0-1
5-1
t/mri
w
-O
TIME in MONTHS
-2
24
DEPTH OF EXCAVATION VS. TIME
FIG. 44.11
)
Lii
25-
30--
I I I I I I I I ~
4 4 4
___I >~~~~___I. I - A
1 20
0
20
4040
0
0
80 -
1000
4.4. 12 a
0.1 0.2 0.3 0.4 0.5Time factor T
0.6 0.7 0.8 0.9 1.0
TIME FACTOR VS. AVERAGE DEGREE OF CONSOLIDATION
AeY~Iu0 Ue/uo
4.LLA .LC4 7 77E1.
< T = 0
po 10 -0.15 00 0- 0.40 0P fh
0.60 - --- 0.848--
070.90
Y17
0.4 0.5 0.6Consolidation ratio U,
0 0.1 0.2 0.3 0.7 0.8 0.9 1.0
NORMALIZED DEPTH VS. CONSOLIDATION RATIO
FIG. 44.12
133
- --- --- -
-7
L _ __ - _ _- -_ _ __--_ _
0
42
N
1
2
4.4. 12b
P- -. XA,72-&
", '7! f'%
/E
~W~qVR/N6AT SUPPICF-11V714L k zx LRs FG/9.SA4Em1sU____
0.2 04 0.6 0.8
uz
NORMALIZED DEPTH VS. CONSOLIDATION RATIO FOR
TRIANGULAR INITIAL EXCESS PORE PRESSURE
FIG. 44.13
0
=4Z
0.5
1.0
1.5
2.00
0.1
0.2 0..4(.4 8
EXCESS4 3 . m rro A . L- E XcESS P. fEL. Z35
,4 U , -7.2 19--2F= o.% Tsm\ = LL eF// ~ A Lk~~ -~C 10 aoe =L Ss5 TSE~
_____~~~~~~ -I__ ____ =\L L 14 .3-i .-.
1.0
SLT
SAYtErs
I'2
d)
0
0i
.- 30
0.
x'-i,
5
4.
3
-2
______ ______ ______ ______ -I
S -20 -'
15 3-
;- -/0
I i -
I, -
E/IoA
RA-SEn)ER % SA5 Tie O56
/ s rcPAXh'
//,- P
3
-4--
5-
-i
5
slr~vss461
LLi
, I~ 10 t6
Nvo-r.E
DiSSIPATE -&L.(EXCU5S) DiSsiPAiE +AU, (ExCESS).
VOLUMETRIC STRAIN VS DISSIPATION OF PORE PRESSURE
F IG. 44.140
V
'I.)
0.
wI1oT"
U.)
IVeRrIcA S.T:WAIj ___
A A.S7PA 67ED
I
ov
z
I')10.0 Zo.0 30-0 40.05 0~
0.0 4 J4
.o
2.0
"3.
0zoo /.07
VOLUMETRIC STRAIN VS. LOGARITHMIC DISSIPATION OF PORE PRESSURE
FIG.4.4-14b
I
TS^
Z; v I r~tZ - "50 r '05- s -
4-a
4-00
ooors I-
CONTOURS OF MODULI - BRACE III ANALYSIS - REVISED RUN NO. B
FIG. 44.15
E CAA t,
400
oo
40-
TSAI
450
w
35o7-,400 Tsm
E 50 TTSA
E/000I,~ -s Z /=1650 TSM'
CONTOURS OF MODULI - BRACE III ANALYSIS - FINAL RUN C
FIG.4416
20
40
60
,\\\\\CA
'TF-
I 11
,E T!5^1
n
SCALE METERS
0 0 0 020 00
DEFORMED FINITE ELEMENT MESH
FIG.4.417
w v 0 0
/2 C0'4
4.-
SCALE AlErCRSp S I I
/0 /6 2o
-I
zs*
SELECTED LINES OF LATERAL DISPLACEMENT
0
F IG. 4.18
0 IV
0 9 S S0
0 0
.Is
SCA4E ^fCTERS
0 5 ,o /5~ Zo
SELECTED LINES OF VERTICAL HEAVE
FIG.4419
76I~
0 0 0 0
0
t+IXAr UM =0 . 3 /7 TsmA
MINI MUM = -12. 3835 Tsm
CONJTOUR INTERVAL - 0.900
( 1I Kil 2( 11r
.317
7
j _I ---
P7
OVN
SCAA.E METEkSI I p I I
00 is /0 2:5Z
CONTOURS OF SHEAR STRESS FIG. 44.20
7
/
/
0 0 0
C-)
MAXITMUM - 0 . 235 TSM TOTPL SiGMr7 LI+ MTNTMUM - -96. 860 TSMUONBJR I.NTLRVRL 7. 000 TSM
5'- o /0 20 is
CONTOURS OF TOTAL SIGMA Z
F !G.4-21
MRXIMUM - 0. 0 00 -rSme MINTMUM - 74. 760 TSM
CONTOUR INTERVAL 5. 000 T-Sm
TOTRL SIGMAi X -
O.0o
o.co
Go ---
-/-04-
SCALE AfE7ERS
fr.J
.Ls
05 7 /0 /5. ZO 25.CONTOURS OF TOTAL SIGMA X
FIGA4.22
Ii
S C S4D
MH XlJ MUM 21.500 TSMMINIMUM +0. 683 TsM
CONTOUR INTERVAL = 4. 000 rsM
S/0 ! zo rs
SCA.E -AIrERS
PORE PHESUPE (LuNjDs-sipATED)
V CU
.3
)
CONTOURS SHOWING TRENDS OF NEGATIVE EXCESS PORE PRESSURE
FIG.4423
0
9
I-'
*
PREDICTE D// BA CE iTI
=?,>FGUIe 4 f.0
146
3 Z aM-l(
Zone ISand and Soft to Hard ClayAverage Workmanship
Zone IIa) Very Soft to Soft Clay
1) Limited depth of clay below bottom ofexcavation
2) Significant depth of clay below bot-tom of excavation but Nb<Ncb
b) Settlements affected by construction dif-ficulties
Zone IIIVery Soft to Soft Clay to a significantdepth below bottom of excavation and withNb5 Ncb
COMPARISON OF PECK'S CHART WITH SETTLEMENTBEHIND THE NAKAGAWA EXCAVATION
FIG 4424
5E/7-4kE/xNT ADJACEArr Tco EYcAVAIO/AISPEC/< (1369)
V/S~TAA/CE FROMA EAcAVrrI/0
AMI/Uf DEPin OF rXC4vq7oN/
r IA0
V..
I~11z14)
LQ
7z
73RAr'E 77T STRU7T LOA)S AND 9H DISTR/iBtIOAw
Z= 317..S. L .
- 57.7 -- ,
- 40.3 -'
+2m. / -)
-8.03 -0
I 4L
ZO /0 0
'lN
/0
~Thuo
~0
r0 30 40
5T7xE5S -7r-S-l
ESTIMATED TOTAL STRESS DISTRIBUTION BEHIND
THE STEEL PILE WALL FIG. 4425
(w.I'' 57 T.,F ( Av'A
DEP VAe 4".'
rioA
4 --1
60 S'o 40 o0
SWAL -
148
L= Length
H
k--B-.
10
8
6
4
Factor of Safety = -;
0 1 2 -3
CNcy
H/B
Bottom Heave AnalysisB JERRUM
for Deep Excavations (H/B >1)-& EIDE (1956)
B
L/ / W
D
450
Factor of 1Safety H
C N cy- C/0.78
Factor of 1 CNcSafety H - C/0.7D
Bottom Heave Analysis for ShallowTERZAGHI (1943)
Excavations (H /B<1)-Fr-GuGE FQO M CLou44 CI SC.WA1Tr C19-1-7)
BASAL HEAVE ANALYSIS
F IG. 44.26
4 5
H0.7B7H + -
0.7 B
S Scuorcuor (BAL = 1)
B/L= 0. 5
0000,00 4-e-000", 00011-
Strip (B/IL=O)
B-50.- mrmes
Z r 80 sErRs
MlER
T
( 114 EXCA-ATlJM VTAGE
/7=fe r"+ -= a
I-
4,L- - - - - J
IA r 2 iS
represe t ... ....- - - - -
surfoTce 4 36
-I - .. .. - --
' 4
f,9,
/
E S. = - c /'..- c17,8
I ex .3
/.5 . 7 g
SIXTH EXCAVATION STAGE - STABILITY ANALYSIS
1.63
F IG.4.427
iN
-4-4
3
N
-45445"
1;;
1-r 0 r
FtGiTE EMEM4ESH .
G = T s.*I
A= '37 / .4' N% x C.3
-- -1
-A a
DR AlEb 8.EMAVI O v
AOWiER $ANJDY LA.CY~e 0
4-0.
/O 2
'ESP f 20T' SO
.-' AStA.Akrr.. FA ItI UPko.0 AfC7
XCA'4A1rof STAGE
A.OADiAIC7
SpoE PReITAL Ld,5,#/04P
ESITI AL LIoAJ~IuG
2o
STRESS PATHS FOR STABILITY ANALYSIS FIG.4+28
/0 -
%,TSM 0
/0-
s--
/0 -
10 30_
40
au,
o -. 04 -.04 /
0 -/1.4- -/.4 *
0 -. /7 -. 7 3
o e s /30
0
Z.0o
0.
-3/
0
-1.0
-.47
0
*.3/
0
0
/3-
4
7.5
4A4k TZ.A1S - ^AlrE. , 4,JO MTr C TOAJS S=STREA/7)Y
08 JSE0 ,,F aR &JE ,,V6,-O H E 0- 4o, - 33
AVG. 4-078.3
TABLE 4.5
D[P77/| ?.-u. 4J/q z4g Aub, siFL .
34
38
4o
43
293
376
397
4Z9
482Z
253
6.4
3.9
.76
9. z-
//./
5.0
5. 1
97
(.8
U111
0
0
0
0
0
OS
^-0
0
P - u.
-/2./
-' .
/70
-4.3
*0..6
31.75
44.o
24's
295'
4"-
34-. C
45t.
-510
2.6
3.6
/-3.
'3.0
"'/.?
-7-0
/.'
/9./
/2.7
5.3
3.4
/7.(.
Z9. 1
/7.,f
24.-
-0.B /04
0
0
a
0
0
f/.,/Y19
718
0
6.IT
o.57
0 77
/. Ao
/'.9 g
.0t
.42
0-'7
0
0t
.laz
au,,
-3. 5
-3-1
-/o
0
0
- 5-0
0
-208
-27.9-U.
0
41,9
0
0
0
0
(-4,
-- '- / . I
0
/o.'
zt .f
e
g.
152
CHAPTER 5
SUMMARY AND CONCLUSIONS
5.0 GENERAL
Review of
role of time i
case history
s important.
braced excavat
Dissipation o
ions show
f excess
s that the
pore
pressures,
response,
with time.
the best t
behavior,
environmen
Similarly,
d iss ipat io
which can
in braced
thes is
intende
effects
p
d
the major time effect influencing the
will
The
0 ften lead to a
writer conclude
echnology av
for the safe
t (general p
deformation
n. The magn
be tolerated
excavation,
resents
to bet
of pore
the fo
ter pre
pressu
decreas
s that
ailable because o
ty of the workers
ublic, existing s
s are also a func
itude of movement
and anticipated
particularly in u
rmulation and app
dict such movemen
re dissipation.
soil
e in factor of
an excavation
f thi
and
truc t
t ion
s and
are a
rban
1 icat
ts by
The a
safety
warrants
s time dependent
the surrounding
ures, etc.).
of pore pressure
deformations
major concern
areas. This
ion of a method
including the
pplication of
the method shows that one can combine the Stress Path Method
with the F
inc lud ing
consolidat
The as
braced exc
inite Element Method to obtain a powerful app
the effects of non-linear soil behavior and
ion.
sumption that undrained conditions prevail in
avation in cohesive soil need not be merely
roach
153
accepted for lack of another method. The proposed method to
model partially drained conditions can be applied. Although
it is illustrated for a major excavation, the method can
also be applied to less complicated problems on a simpler
scale to estimate stability and deformation. One does not
need a large finite element program to apply The Stress Path
Method. However for complex geometries such as the Nakagawa
Treatment Plant Excavation, the finite element program BRACE
III is a powerful tool, and the use of it is justified.
5.1 LIMITATIONS
One limitation of the proposed method as applied in
chapter four is that there is
work involved. However, compa
additional tests and testing t
not unreasonable. In this cas
construction performance was p
illustrate that for one to pre
excavation stages the entire p
each stage. A better approach
to execute the hand work for o
and then fit a polynomial curv
through the stress strain data
a significant
red
ime
e s
red
d ic
roc
to
t
i
t
e
to the al
the amou
udy only
cted. Fi
performa
dure must
model all
amount of hand
ternat
nt of
one s t
gure 5
nce of
be it
stage
iv
ha
ag
.1
0
er
s
ne or two excavation
e or hyperbolic
An
e of
ndwork
e of
helps
ther
ated f
would
stages
is
to
or
be
equation
initial modulus for the
r stages can be obtained from the undrained stress
test data. Kondner (1963), and Duncan and Chang
present and discuss utilization of the hyperbolic
(refer to Figure 5.2) where;
ear lie
strain
(1970)
model
154
The current version of BRACE III however, does not have
this capability to model such nonlinear stress strain
behavior, although the program could easily be modified.
As a practical matter one should be aware of several
limitations when using the proposed method.
(1) Large finite element computer programs such as
BRACE III require experience for proper useage and
effectiveness. It can also be expensi
program and analyze the results.
(2) Stress path tests are both expensive a
to run, although several carefully exe
path tests may be of significantly mor
a larger number of more common laborat
(e.g. UU Tests). Well known factors su
disturbance can significantly influenc
test results.
(3) Program modifications to make the prop
and BRACE III more compatible are need
modify and dimension the program to ha
many varied soil properties)
5.3 SUGGESTIONS FOR FURTHER RESEARCH
Some suggestions by the writer for further
work in this area of interest include:
(1) Additional studies of the type made by
Osaimi, applied to actual excavations,
ve to run the
nd difficult
cuted stress
e value than
ory tests
ch as sample
e laboratory
osed
ed (
nd 1 e
method
i.e.,
input of
research and
Cl
i.
ough and
e.,
predicted vs
deformations
actual performance (e.g.
). Such study for a variety of soil
155
conditions would help evaluate the appl icability of
the sophistica
hopefully broa
time in braced
(2) Evaluation and
paper, and oth
method by comp
For the Nakaga
oedometer test
test data avai
data and field
input can be r
(3) Improvement of
modeling in th
is an inherent
and sheeting e
Displacement c
inter-element
dimensional sh
dimensional el
wall elements
per node assum
elements have
maintain
position.
ted numerical technique,
den our understanding of
excavation.
updating of the predict
er predictions using the
arison with actual field
wa site, additional test
s) are needed because of
lable to date. From add
measur
ev ised
the so
e BRACE
s train
lements
ompatab
boundar
e e tin g
ements
wh ich
e cur
only
h
ve
tw
and
the role of
ion in this
proposed
results.
ing (i.e.
the limited
itional test
ements the BRACE III program
and the prediction updated.
il structure interaction
III computer program. There
incompatability between soil
(Jaworski 1973).
ility is not maintained along
ies between the one
elements and the two
of the soil continuum. The
ave three degrees of freedom
d deformed positions. Soil
o degrees of freedom and
straight boundaries in the deformed
156
(4) Develop Brace III program capability to model
nonlinear stress strain behavior, perhaps by
employing a subroutine which uses a hyperbolic
relationship as proposed by Kondner (1963).
5.4 CONCLUSIONS
The objective of this thesis was to formulate and
apply a method which includes the effects of pore pressure
change with time to predict deformation in an excavation.
This thesis shows that one can combine the Stress Path
Method to obtain soil parameters with the Finite Element
Method to perform analysis in a way that includes the
effect of time on deformation of a braced excavation. The
approach is outlined in Chapter 3 and applied in Chapter 4.
Advantages of the approach outlined in Chapter 3
are:
(1) Complex nonlinear time analysis is avoided.
(2) The fundamentals of the field problem are examined
by the engineer as he selects his input so that he
retains a feel for the problem.
(3) A few tests can be used and extrapolated to cover
an entire geometry.
The proposed approach is an iterative procedure. The
application to the example in Chapter 4 suggests that two
to three iterations give a reasonable solution, especially
for the overall deformations.
For the example case consisting of an excavation in a
157
clayey silt, the effect of time on deformations within
the excavation is significant although not major. Heave
and inward wall movements within the excavation appear to
increase by approximately 10 to 20 percent with consolida-
tion.
Within the top 8 meters (depth of 20 to 28 meters) of
the bottom of the excavation at the 20 meter excavation
depth, the negative excess pore pressures due to unloading
and the positive excess pore pressures due to shallow or
surficial dewatering, substantially offset each other.
This results in a relatively small net excess pore pressure
for this case. In other words, because of the dewatering
the dissipation of large negative excess pore pressure
and resulting heave is minimized.
For this particular soil the negative excess pore
pressures due to excavation are offset by the excess pore
pressures generated by surficial dewatering. Consequently,
the effects of time on predicted movement are relatively
small in this case. With other soils and other conditions
of flow, the effect of time on deformations could be con-
siderable.
In addition, for this particular case the example
shows that the undrained heave is large when compared to
the consolidation movements. For other situations invol-
ving higher factors of safety against bottom heave the
contribution of time to total deformation will be more
significant.
158
Again, a practical advantage of the proposed method
is that such fundamental behavior can be observed and
quantified by the engineer during his analysis of the
field situation.
DRAIMAGpE EC, V T- o ^> Po-AG- - - - -io A' - A-
I ro
*4
ClIRVE ) - ACTLA L
Ci'J\ E (3) -
BEHANJIOR. - FrTS ALL 5TAGES
TDE~L-ZEEh ~SEWWIFO APPLLCABLE Mr PREDICMlON
OF er EXCAVATicb sTAGE - LSIN9,T+E $AMEMotu.e.LU.S FoR- PREi'ou.5 EXCAVA~TON STAGES WLL
R ESLLLT 1 MJ AN O\JER-PREtkCTlON OF STR AiN.SIW C, $A tE S CANIT MoO .US FoR. >EEPER
5TAGES vJiLL RESLUT It-)AM UJDER-PREIt 'C-nONOF 5TRAIN MA-GllTUDE.
LINEAR VS. NON LINEAR STRESS/STRAIN FIG 51
159
A:G7G
8EQTCAL.
160
Asymptote = (a 1 - u3)lt 1/b
1
E' = 1/a Ct
Hyperbolic Stress-Strain Curve
C4
-4
U.)
4
U
-4
formed Hyperbolics-Strain Drve
b
Axial Strain, E
STRESS STRAIN CURVES
FRoim OSA/4/ /977
FI G. 5.2
Trans
Stres
HYPERBOL IC
161
LIST OF SYMBOLS
a Intercept of K line
C Cohesion intercept based on effective stress
CAU Anisotropically consolidated triaxial testsheared under undrained conditions
CRSC Constant rate of strain consolidation test
CAES Center for Advanced Engineering Study
CIU Isotropically consolidated triaxial testsheared under conditions of no drainage
c Coefficient of consolidationv
E. Initial tangent modulus
ESP Effective stress path
H Height of cut
h Elevation heade
h Pressure headp
h Total headt
Hd Length of drainage path
K Active earth pressure coefficient
K At rest earth pressure coefficient, horizontalground surface
OCR Overconsolidation ratio
P.I. Plasticity index
P 9 3X P T 4 ____
S Undrained shear strength of soil
TSP Total stress path
TSP-u0 Total stress path minus initial pore pressure
162
u Pore pressure
u Steady state pore pressure
ue Excess pore pressure
u D Excess pore pressure due to excavation unloading
u E Excess pore pressure due to dewatering
UU Unconsolidated undrained triaxial test
wn Water content, natural
wy Liquid limit
w Plastic limitP
z Depth
z Depth to water table
U Consolidation ratioz
U Average degree of consolidation
Slope of the K line
Unit weight
YT Total unit weight of soil
Unit weight of water
Strain
E Young's modulus
(Ty Total vertical stress
(Y Total horizontal stress
'3 o r Octahedral stress
Vertical effective stress
a Horizontal effective stress
y, Initial total and effective stresses, vertical
So, Initial total and effective horizontal stresses
Friction angle, friction angle based on effectivestress
163
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Mitchell, J.K., and W.S. Gardner(1975), "In-Situ Measure-ment of Volume Change Characteristics,6th PSC,ASCE,pp. 379-391.
Peck, Hansen, and Thornburn, Foundation Engineering , Wiley,1974.
Osaimi, A.E. and G.W. Clough (1979) "Pore-Pressure
Dissipation During Excavation, Proc. of ASCE, JGED
Vol. 105, No. GT4, pp. 481-499.
Osaimi, A.E., "Finite Element Analysis of Time Dependent
Deformations and Pore Pressures in Excavations and
Embankments" Phd Thesis, Stanford University Dept.
of Civil Engineering 1977.
Palmer, J.H. Laverne, and Kenney, T.C. "Analytical Study of
a Braced Excavation in Weak Clay" Can. Geotechnical
Journal Vol. 9, No. 2 pp. 145-164 May 1972.
Peck, Ralph B., (1969). "Deep Excavations and Tunnelling in
Soft Ground," State-of-the-Art Volume, Seventh
International Conference on Soil Mechanics and
Foundation Engineering, Mexico City, pp. 259-281.
166
Ries, Carol, Predicting Lateral Earth Pressures for the
Design of Tie-Back Retaining Walls',' Master of
Science Thesis, M.T.T. 1974.
Skempton, A.W. and P. LaRochelle, (1965). "The Bradwell
Slip: A Short Term Failure in London Clay"
Geotechnique, No. 15 pp. 222-242.
Taylor, R.L. and Brown, C.B., "Darcy Flow Solutions
Free Surface, "Journal of the Hydraulics D
ASCE, Vol. 93, No. HYZ, pp. 25-33, 1967.
Withiv is ion
Terzaghi, K., Theoretical Soil Mechanics, John Wiley and
Sons, New York 1942.
Terzaghi, K. and R.B. Peck (1967). Soil Mechanics in
Engineering Practice, Wiley and Sons, New York.
Tschebotarioff, Gregory P., (1973). Foundations,Walls and Earth Structures, McGraw Hill,
RetainingNew York.
Wong, I.H., Analysis of Braced Excavations,M.I.T. 1971
Phd. Thesis,
Zienkiewicz, O.C., Cheung, Y.K., The Finite Element Method
in Structural and ContinuTm~Hicnics,~Ion~on,McGraw-Hill Publishing Co., Ltd 1967.
_
167
APPENDIX A
GRAINSIZE PLOTS AND SQUARE ROOT OF TIME
CONSOLIDATION PLOTS - NAKAGAWA SOIL SPECIMENS
U.S. STANDARD SIEVE SIZE
100
90
so
-U
z-I
mmm>0
j E 4 --
G)C
m m{ >
T0
60
50
40
30
20
t0
0100
2 IN. IIN.3/41N.1/2 N. NO.4
10
NQIO NQ20 NQ 40 NQ.6 NO.OO NO. 200
1.0 0.1 0.01
GRAIN SIZE IN MILLIMETERS
GRAVELCDBLESC 0
COARSE IE
SAND
ICOARSE1 MEDIUM FINESILT OR CLAY
UNIFIED SOIL CLASSIFICATION SYSTEM
TEST sYM MATERIAL SOURCENO, SOURC
S-I 0 - Tokyo, JapanL-1REMARKS
lakaiaga. Treatment PlantI,te 13
F6. A-
- - - - - - - - -~- - _ __
---...-.- ----.-.-.-.-.- I
. - . _ ___ -.-- - - -........._ -- _ -_
-_ I -__TI - -I- _
.-- - - - - - - -- - -
- - - - --
- -I -- -. _ _
-I - - - --- - - - - -. -
0.001
i
. - .
I
I
U.S. STANDARD SIEVE SIZE
0
r q
x
-4
< 0 -- m
O I
m z
XI -3 _j(flW
90
z
>0
IN. 3/41N.1/2 IN. NO.4 N .UO NQ.20 RQ.4OfQ6ON.IOO NO.20100
90
so
-0.01 0.001
GRAIN SIZE IN MILLIMETERS
GRAVEL SANDCDBLES SILT OR CLAY
COARSE FINE COARSE MEDIUM FINE
UNIFIED SOIL CLASSIFICATION SYSTEM
0 Tokyo, JapanA
Hakagawa Troatment Plant2ite B
FIG. A.Z
IN.
0.1
TO
60
50
40
30
20
10
0
I - i l l-- - -- I
- - t~-- - I 1 -
I--- - - - _
-- T T - r -I- T --I - - -_ _- -
I -T~ 1 I I
~ I - - ~ -- ~ ---- -I
--- -I F~ ---------_ _ _ I-I - -I-
100 10
TEST
N-2
REMARKSSYM. I MATERIAL SOURCE
I
2
1.0
I
-
rf
mt
-X (n CD
m z ->
:4 >
F'
z
ci
CO01X
100
90
so
Io 0
60
so
40
3D
20
10
0100
2 IN.U.S. STANDARD SIEVE SIZE
I IN. 34IN.1/2 IN. NO.4 NAI0 k.20 NO.40 NO.60 NO.100 O.200
I0 1.0 0.1GRAIN SIZE IN MILLIMETERS
OBBLES GRAVEL SAND SILT OR CLAYCOARSE COARS EA MEDIUM FINE
UNIFIED SOIL CLASSIFICATION SYSTEM
TEST SYM. MATERIAL SOURCE REMARKSNO. ___ ____________ ________ _
0-3 0 Tokyo, Japan Hakagawa Treatment PlantH-3 A Site B
- _ F/6. A.3
----- I - - T - - _ _
--- - -- - a-a
T - - T - - -- T -
- - a- - - - - - - - - -
- - - -- I- -- _-
- - -
- - -- - --
' 1 1
- -- -a g---
I j I I I I ______-
C
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100
90
90
V 70f"uAl
60so
50
40
30
rrI -IOV)>W
- -
y z
0 0C C
00
z
20
to
0100
U.S. STANDARD SIEVE SIZE
. . .N ./ .N12I 0 O 0 04 00N.0 O 0
---- -T-
- .1
- -T -
- -I I -- i 1
-- -F-
-- ~ -- --- - - -I ~ A
-- - -- - - - -
- - -F - - ~
GRAIN SIZE IN MILLIMETERS
09LE - GRAVEL SAND SILT OR CLAY00LSCOARSE FINE COARSEI MEDU FINE
UNIFIED SOIL CLASSIFICATION SYSTEM
TEST SYM. MATERIAL SOURCE REMARKSNO.
- Tokyo, Japan 1Jakarawa Treatment Plant11-4 A : itr B
6/G. A.4
1N IN 3'4!Nf/2 f Ni 4 N O 40 NO63NO.lOO m. O
0.01 0.001.0 0.110
t00
90
90
-4
TI
z
-4
;u -- acn w
r Gm z
x11
-3 v"(i
TO
60
so
40
30
20
to
0too
U.S. STANDARD SIEVE SIZEIN. IN.3/4 IN.1/2 IN. NO.4
10
N .10
1.0
NQ.20 NQ40NO.60NO.100 ON.200
0.1 0.01
GRAIN SIZE IN MILLIMETERS
GRAVEL SAND
UNIFIED SOIL CLASSIFICATION SYSTEM
TEST SYM MATERIAL SOURCE REMARKS
S-5 0 Tokyo, Japan-5 A
lakagrawa Tratment Plant:"I t f B
I - --- - - - - -- i --- .- -
- - - - - - - - - - - - -
___ T V1 T I --
- - --- - - - - - --- - - - -
I - I - - - - - - I
I-I - -- - - - -- - -. -
-- - - -l - - - -
-I -~-~ -
- I I - -'- ~ - ___- I I I - -- .-- -
0.001
1'0
f16.A.s
I2f
U.S. STANDARD SIEVE SIZE
1
2z
< G)Z
22
n >
c2
0 0,
>
uT -1 in~I
r -icn P
2IN. IIN.3'4IN./2 N. NO.4 NPO Np20 NQ40NQ6ONO.IOONO.2O
1.0 0.1 0.01
GRAIN SIZE IN MILLIMETERS
GRAVEL SAND
COARSE FINE COARSE MEDIUM FINE
UNIFIED SOIL CLASSIFICATION SYSTEM
TEST SYM. MATERIAL SOURCE REMARKSNO. ______________
s- 6 0 Tokyo, Japan Hakagawa Treatment PlantL-6 A Site B
90
0
TO
60
50
40
30
20
10
0
I I ITII
- -....... I __ _ _ - -- I . . -- - --- _
T I I
i ii I--I i 1T I -_ ~ . ~ . . . .
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-. I - - - - - i- - - - - - - --
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0
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1.40
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182
APPENDIX B
VERTICAL EFFECTIVE STRESS VERSUS DEPTH - NAKAGAWA
183
0 10 20 30 40
TSM
NAKAGAWA
VERTICAL EFFECTIVE STRESS VS. DEPTH
FIG.B.I
uhrLU
LU
LU
0-
I0-L
20-
307
40+
50-
N
i