1
N
AN EVALUATION OF THE PERFORMANCE
OF A BRACED EXCAVATION
-H.
BY
WALTER EDWARD JAWORSKI, JR.
.......
B.Sc., Northeastern University
(1962)
....
...
5
M.S., Worcester Polytechnic Institute
(1964)
.........
IW
.....
Submitted in partial fulfillment
of the requirements for the degree of
DOCTOR OF SCIENCE
.....
at the
........
Massachusetts Institute of Technology
(4une, 1973)
........
Signature redacted
Signature of Author:
Department of Civil Engineering (ba1/c
Signature redacted
......
I
1973)
.........
Certified by:
ia
(Thesis Supervisor
1
~......
Accepted by:
Signature red acted
(Chairman, Departmentak/Committ'ee on Graduate Studies)
Archives
JON 29 1973
'"BA R ff
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ABSTRACT
AN EVALUATION OF THE PERFORMANCE
OF A BRACED EXCAVATION
by
Walter E. Jaworski, Jr.
Submitted to the Department of Civil Engineering
in June, 1972, in partial fulfillment of the
requirements for the degree of Doctor of Science
To date, many field measurement programs have been undertaken to
monitor the performance of braced excavations constructed with steel
sheet piling or soldier beams and lagging. These programs gave insight
into the behavior of braced cuts and resulted in the development of
empirical design rules.
In recent years the use of cast-in-place concrete
walls for sheeting has gained acceptance in situations where it is
necessary to minimize movements adjacent to a braced cut. Presently,
there are only a few documented field measurement programs for excavations with concrete walls to aid in the design of this type of bracing
system.
The primary objectives of this thesis are to document and evaluate
the measured performance of a braced excavation employing a cast-inplace concrete wall, and to contribute to the development of analytical
techniques to predict the performance of these bracing systems.
The field measurements for this investigation were conducted on
the braced cut for the subway extension at South Cove in Boston.
The
excavation varies in depth from 45 ft to 50 ft, and in width from 58 ft
to 80 ft. Where the subway runs close to an existing seven-story building,
the excavation was supported by a concrete wall installed by the slurrytrench process. Elsewhere the excavation was supported by steel sheet
pile walls. The subsoil consists of 5 ft to 8 ft of fill underlain by
very stiff to stiff clays. Field measurements consisted of sheeting and
subsoil movements, pore water pressures, and strut loads. The measurements were made during the installation of the wall and excavation for the
subway.
During installation of the concrete wall in panels 10 ft to 20 ft in
length, lateral subsoil movements as much as 0.75 in.occurred immediately
adjacent the wall. The vertical ground movements were in the range of
0.1 in. to 0.3 in. The concrete wall movements while excavating for the
subway were less than 1 in. The maximum gross movement behind the wall
was 1.25 in.
In contrast, the adjacent steel sheet pile wall moved up to
- i-
Settlements behind the concrete wall were uniform and
6.5 in. laterally.
less than 1 in.
Due to seepage through and around the concrete wall, pore
water pressures dropped 8 ft to 20 ft. The measured loads in the struts
supporting the concrete wall showed a scatter of up to 50 percent from the
average strut load for a typical level.
Predictions were made of the same aspects of performance monitored
during installation of the wall and excavation for the subway. During
this phase of the investigation the capabilities of the finite element
program BRACE (Wong, 1971) were extended.
The trends of the ground movements during the installation of the
wall predicted by stress paths and finite element analysis agreed well
with the measured behavior. A study of the stability of slurry trenches
of limited length indicates the primary stabilizing forces for the
trench walls are arching and the fluid pressure of the slurry. Strut
loads predicted by BRACE and Peck's design rules gave, in general, loads
which were greater than the measured values. Due to the scatter in the
measured loads the amount by which the methods over predicted the strut
Concrete wall movements predicted by the modified
loads varied greatly.
version of BRACE agreed well with the measured values. However, the
steel sheet pile movements predicted by BRACE grossly underestimated the
measured movements. The primary reason for underestimating the steel
piling movements was the large movements which took place at a strut
Predictions of the settlement
level after the strut was installed.
behind the concrete wall due to both sheeting movements and drops in
ground water level over-estimated the measured settlements.
The effect of sheeting rigidity, sheeting penetration and strut spacing
The results show that
on sheeting movements was studied using BRACE.
sheeting rigidity and penetration have a marked influence on reducing movements below excavation level and they increase the stability of the excavation bottom. Reducing the strut spacing also reduces sheeting movements,
but the effect is not as pronounced.
Limited field
The effect of bracing details was also investigated.
measurements show that poor shimming techniques not only permit large
sheeting movements above the excavation level, but they also increase the
movements below the excavation level.
Thesis Supervisor:
Title:
T. William Lambe
Edmund K. Turner Professor of Civil Engineering
-
ii-
To:
ELAINE, C. KIM, PAMELA, JENNIFER, and DAWN
for their patience and encouragement
during the course of this investigation
-iii-
ACKNOWLEDGEMENTS
The author is thankful to many members of the faculty
of the Soil Mechanics Division for their guidance and assistance in the preparation of this thesis.
Professor T. William Lambe was my faculty advisor and
thesis supervisor.
His patience, advice and encouragement
throughout this work is gratefully acknowledged.
Professor L.G. Brownwell.provided helpful suggestions
and constructive criticism on various aspects of this work.
Professor J.T. Christian spent many hours advising and
helping with all the theoretical aspects of this thesis.
Dr. D.J. D'Appolonia spent several hours discussing the
field measurements and results of the theoretical work.
These
discussions were most helpful.
Dr. K.M. Leet reviewed and provided helpful suggestions
on the text of this thesis.
The MIT - ICEP field instrumentation staff's excellence
in installing the field instruments and retrieving data contributed greatly to the success of this thesis.
Dr. L.A.
Wolfskill headed the staff of Messrs. W.R. Beckett, J.E. Bromwell,
Mark Haley, and H.A. Russell.
Mrs. Claire Champ and Ms. Marilyn Krivitsky typed the final
draft of the text; Mrs. Louise Passos drafted the figures.
The Massachusetts Bay Transportation Authority sponsored
the research effort presented in this thesis.
-iv-
TABLE OF CONTENTS
Page
ABSTRACT
ACKNOWLEDGEMENTS
iv
TABLE OF CONTENTS
V
LIST OF TABLES
iX
LIST OF FIGURES
x
CHAPTER ONE
1
INTRODUCTION
1.1
INTRODUCTION
1
1.2
THESIS OBJECTIVES
1
1.3
THESIS SCOPE
3
CHAPTER TWO
LITERATURE SURVEY
5
2.1
INTRODUCTION
5
2.2
CONSTRUCTION OF CONCRETE WALLS
5
2.3
PROPERTIES OF BENTONITE SLURRY
7
2.4
STABILITY OF SLURRY TRENCHES
8
2.5
FAILURES OF SLURRY TRENCHES
12
2.6
SUMMARY AND CONCLUSIONS
13
16
CHAPTER THREE
3.1
BACKGROUND
16
3.2
MEASUREMENTS DURING INSTALLATION
OF THE CONCRETE WALL
17
3.2.1
Pore Pressures
17
3.2.2
Ground Movements
18
3.2.3
Ground Movements versus Fore Pressures
19
-V-
Page
3. 3
3.4
MEASUREMENTS DURING EXCAVATION
19
3.3.1
Performance versus Construction Progress
19
3.3.2
Sheeting Displacements
21
3.3.3
Ground Movements
22
3.3.4
Pore Pressures
24
3.3.5
Strut Loads and Horizontal Stresses
26
SUMMARY AND CONCLUSIONS
CHAPTER FOUR
SECTION I
27
PREDICTED PERFORMANCE
29
PREDICTED PERFORMANCE DURING
WALL INSTALLATION
29
4.1
STRESS PATHS
29
4.2
PANEL STABILITY
31
4.3
PREDICTED GROUND MOVEMENTS
33
4.4
COMPARISON OF PREDICTED AND MEASURED
PERFORMANCE DURING WALL INSTALLATION
33
SECTION II
PREDICTED PERFORMANCE DURING
EXCAVATION
34
4. 5
BACKGROUND
34
4.6
PORE PRESSURES
36
4.7
HORIZONTAL STRESSES AND STRUT LOADS
38
4.7.1
Design Methods
39
4.7.2
Analysis Methods
40
4.8
SHEETING MOVEMENTS
44
4.9
VERTICAL MOVEMENTS
45
4.9.1
Initial Movements
45
4.9.2
Consolidation Settlements
46
-vi-
Page
4.10
COMPARISON OF PREDICTED AND MEASURED
PERFORMANCE DURING EXCAVATION
47
4.10.1
Pore Pressures
47
4.10.2
Strut Loads
48
4.10. 3
Sheeting Movements
49
4.10.4
Ground Movements
50
SUMMARY AND CONCLUSIONS
51
4.11
CHAPTER FIVE
PARAMETERS AFFECTING THE PERFORMANCE
OF A BRACED EXCAVATION
53
5.1
BACKGROUND
53
5.2
STRESS PATH ANALYSIS
54
5.3
BRACE II ANALYSIS
57
5.3.1
Sheeting Stiffness
57
5.3.2
Vertical Strut Spacing
58
5.3.4
Comparison with Field Observations
61
5.4
SHEETING MOVEMENTS
62
5.5
SUMMARY AND CONCLUSIONS
65
CHAPTER SIX
SUMMARY AND CONCLUSIONS
6.1
6.2
MEASURED PERFORMANCE
66
66
6.3
PREDICTED PERFORMANCE
67
6.4
PARAMETERS EFFECTING THE BEHAVIOR
OF BRACED EXCAVATIONS
69
GENERAL
-vii-
BIBLIOGRAPHY
70
NOTATION
76
APPENDIX A
A SUMMARY OF FIELD MEASUREMENTS AT THE
SOUTH COVE TUNNEL EXTENSION
136
APPENDIX B
ACCURACY OF FIELD MEASUREMENTS
238
APPENDIX C
FINITE ELEMENT ANALYSIS OF EXCAVATIONS
276
APPENDIX D
BRACE II FINITE ELEMENT PROGRAM
291
APPENDIX E
SOIL PROPERTIES FOR ANALYSIS AND PREDICTIONS
309
APPENDIX F
BRACE II USER'S MANUAL
313
BIOGRAPHY
330
-viii-
LIST OF TABLES
Table No.
Page
Title
1.3.1
ANALYSIS OF BRACED EXCAVATIONS
81
4.7.1
PREDICTED MAXIMUM STRUT LOADS
82
4.7.2
SOIL PROPERTIES-SOUTH COVE PROFILE
83
5.1.1
SUMMARY OF MOVEMENTS ADJACENT BRACED
CUTS WITH CAST IN PLACE CONCRETE
WALLS
84
-ix-
LIST OF FIGURES
Figure No.
2.2.1
Title
Page
CAST-IN-PLACE CONCRETE WALLS BY THE
SLURRY TRENCH PROCESS
85
STABILITY OF SLURRY TRENCHES IN COHESIONLESS SOILS FOR PLANE STRAIN CONDITIONS
86
STABILITY OF SLURRY TRENCHES IN NORMALLY
CONSOLIDATED CLAY
87
3.1.1
SOUTH COVE PROJECT
88
3.1.2
SECTION AT STATION 113 + 40
89
3.1.3
SOIL PROFILE
90
3.1.4
LOCATIONS OF FIELD INSTRUMENTS
91
3.2.1
TOTAL HEAD AND SOIL MOVEMENTS DURING
CONSTRUCTION
92
TOTAL HEAD AND SOIL MOVEMENTS DURING
CONSTRUCTION
93
CHANGES IN TOTAL HEAD AND GROUND SURFACE
DURING WALL INSTALLATION
94
2.2.2
2.2.3
3.2.2
3.2.3
3.2.4
MEASURED GROUND MOVEMENT DURING CONSTRUCTION
OF CONCRETE WALL
95
3.3.1
SHEETING MOVEMENTS DURING EXCAVATION
96
3.3.2
SHEETING MOVEMENTS AT SOUTH BULKHEAD
(Sta. 107+50)
97
3.3.3
GROUND MOVEMENTS ADJACENT CONCRETE WALL
98
3.3.4
SETTLEMENT CONTOURS OF DON BOSCO SCHOOL
99
3.3.5
MOVEMENTS ADJACENT THE SLURRY WALL
100
3.3.6
TOTAL HEAD AND BOTTOM HEAVE MEASUREMENTS
101
3.3.7
PORE PRESSURES VERSUS DEPTH
102
3.3.8
TOTAL HEAD AND SETTLEMENT AT FULL EXCAVATION
103
-x-
Figure No.
Title
Page
3.3.9
STRUT LOAD VARIATION DURING CONSTRUCTION
104
3.3.10
APPARENT LATERAL STRESS
105
4.1.1
INITIAL STRESS CONDITIONS AT SOUTH COVE
106
4.1.2
STRESS PATHS
107
4.2.1
STABILITY ANALYSIS OF SOUTH COVE SLURRY
TRENCH
108
PREDICTED MOVEMENTS DURING CONCRETE WALL
INSTALLATION
109
4.6.1
FEDAR FINITE ELEMENT GRID
110
4.6.2
PREDICTED WATER PRESSURE VERSUS DEPTH
111
4.7.1
PREDICTED HORIZONTAL STRESSES
112
4.7.2
DESIGN LOAD DIAGRAMS FOR CONCRETE WALL
113
4.7.3
PREDICTED HORIZONTAL STRESSES FOR ELASTIC
BEAM ANALYSIS
114
4.7.4
STRUT LOADS PREDICTED BY ELASTIC ANALYSIS
115
4.7.5
FINITE ELEMENT GRID FOR BRACE II ANALYSIS
116
4.7.6
STRUT LOADS DURING EXCAVATION
117
4.7.7
YIELD ZONES FROM BRACE ANALYSIS
118
4.7.8
TOTAL HORIZONTAL STRESSES ON SOUTH COVE
SHEETING FROM BRACE ANALYSIS
119
4.8.1
PREDICTED SLURRY WALL MOVEMENTS
120
4.9.1
COMPARISON OF MEASURED AND PREDICTED SUBSOIL MOVEMENTS DURING EXCAVATION
121
COMPARISON OF MAXIMUM STRUT LOADS DURING
EXCAVATION
122
COMPARISON OF PREDICTED AND MEASURED STRUT
LOADS
123
MAXIMUM SHEETING MOVEMENTS
124
4.3.1
4.10.1
4.10.2
4.10.3
-xi-
Figure No.
5.2.1
5.2.2
5.3.1
5.3.2
5.3.3
Title
STRESS CONDITIONS FOR VARIOUS CONSTRUCTION
STAGES OF CONCRETE PANELS BY THE SLURRY
TRENCH PROCESS
Page
125
TYPICAL STRESS PATH FOR POINT "A" DURING
INSTALLATION OF CAST IN PLACE CONCRETE
WALL IN CLAY. OCR = 1
126
EFFECT OF STRUT SPACING AND WALL STIFFNESS
ON SHEETING MOVEMENTS FROM BRACE ANALYSIS
127
YIELD ZONES VERSUS EXCAVATION DEPTH AND
SHEETING STIFFNESS
128
FACTOR OF SAFETY REQUIRED TO PREVENT LOCAL
YIELD BELOW BOTTOM OF EXCAVATION IN CLAY
129
5.3.4
BEARING CAPACITY FACTORS FOR BOTTOM STABILITY
130
ANALYSIS
5.3.5
PREDICTED BOTTOM HEAVE BY BRACE FOR STIFF
AND FLEXIBLE SHEETING
5.3.6
131
RATIO OF Nc/NcbAT WHICH FIRST YIELD OCCURS
WITHIN AN EXCAVATION IN NORMALLY CONSOLIDATED CLAY
132
HORIZONTAL MOVEMENTS OF SHEETING DURING
EXCAVATION
133
5.4.1
SHEETING MOVEMENTS BELOW EXCAVATION LEVEL
134
5.4.2
SHEETING MOVEMENTS BELOW EXCAVATION LEVEL AT
NORTH STATION IN BOSTON
135
5.3.7
-xii-
CHAPTER ONE
INTRODUCTION
1.1
INTRODUCTION
Deep excavations within our urban areas are becoming more common
with the increasing construction of multistory buildings and subway
systems.
The excavations often are located in already congested areas
and pose an added task for the foundation engineer who now must concern
himself with the excavation's effect on the adjacent structures.
Pres-
ently there are two ways of protecting an adjacent structure, either the
building is underpinned or special construction techniques are employed
to minimize sheeting movements.
One technique gaining wide acceptance
is the use of cast in place concrete walls for the sheeting.
Support systems are highly indeterminate and impossible to design
without grossly simplifying assumptions or reliance on empirical rules.
Insufficient case records of braced excavations with concrete walls are
available to allow the derivation of empirical design rules, as has
been done for flexible walls (Terzaghi and Peck, 1967), or to conclude
whether the design rules used for the more flexible walls are applicable
to the more rigid concrete walls.
Presently the design of the rigid walls is accomplished according
to ones engineering preference.
Some engineers use modified versions
of the empirical relationships for flexible walls (Stacko,
1968); while
others rely on their engineering judgment (Thon and Harlan, 1971).
Lambe (1970) suggests two approaches to improving the state-of-the
art in the design and understanding of the behavior of Braced Excavations.
-1-
They are (1)
cases.
parametric studies;
(2) the evaluation of instrumented field
With the advent of finite element programs (Wong,
1971; Palmer
and Kenney, 1971) the parametric studies are well under way.
However,
field records, especially those of braced cuts with rigid walls, are
sorely lacking.
This thesis is directed at the second of the above stated approaches
to increase the profession's knowledge of braced cut behavior.
The
field case evaluated is the South Cove Tunnel Project in Boston, Mass.
The tunnel excavation was primarily in very stiff to stiff clays.
Where
the excavation was close to a seven-story building, a cast in place
concrete wall was employed for sheeting.
Elsewhere the excavation was
supported by steel sheet piling.
1.2
THESIS OBJECTIVES
The general objective of this thesis research is to evaluate and
contribute to the development of techniques to predict the performance of
a supported excavation.
This general objective is approached by the following specific tasks:
1.
Document the performance of an instrumented braced excavation.
2.
Identify the modes of behavior which have a bearing on
projects of this type.
3.
Improve prediction techniques employing the finite element
method.
4.
Compare observed behavior with that predicted by present
techniques. Evaluate discrepancies between the observed
and predicted performances.
5.
Define and analyze the important parameters affecting
braced excavation performance.
-2-
1.3
THESIS SCOPE
This thesis documents the performance of a braced cut employing
both a reinforced concrete wall and a similar cut supported by a steel
sheet wall.
The field measurements are summarized and evaluated.
Interpretive plots of the data are separated into two groups;
those associated with the installation of the concrete wall and those
related to the excavation of soil between the walls.
For each group
the data are reviewed for significant behavioral trends.
Emphasis is
placed on the measured data adjacent to the concrete wall and on the
relative movements between the wall and those of the steel sheet piling.
The literature on slurry trench construction has been reviewed
relative to observed ground movements and slurry trench stability
during the installation of the concrete wall.
The stability of the
slurry panel is analyzed using techniques outlined in the literature.
Predictions of ground movements and a study of panel stability are
conducted using the finite element program FEAST 3 (D'Appolonia, 1968).
Stress paths (Lambe, 1964) are used in an attempt to describe the
important aspects of observed behavior associated with the panel instal-
lation.
The finite element program BRACE (Wong, 1971) was modified. and
updated.
The new version handles bilinearly anisotropic materials and
accounts for local overstressing of the sheeting.
In addition, some
studies were conducted to evaluate the limitations of the program's predictions.
For the excavation construction phase, predictions of the (1) sheeting
movements,
(2) strut loads,
(3)
initial ground movements and consolidation
-3-
settlements, and (4) pore pressures are compared to the measured values.
The predictions are made using the methods summarized in Table 1.3.1.
Those factors which the field data indicate have a significant
bearing on observed behavior are analyzed using the new BRACE version.
The factors are: (1) occurence of first yield in the soil, (2) sheeting
stiffness, and (3)
vertical strut spacing.
-4-
CHAPTER TWO
LITERATURE SURVEY
2.1
INTRODUCTION
Slurry trenches are used for one of two purposes:
(1) to construct
impervious cut-off walls for dams, dikes, etc. or (2) to construct cast
in place concrete walls in braced excavations.
A survey on the uses
of both in the United States is given by Sherard (1969) and Kapp (1969).
This chapter summarizes the literature on use of slurry trenches to
install concrete walls.
2.2
CONSTRUCTION OF CONCRETE WALLS
Figure 2.2.1 shows a typical construction procedure for slurry-
trench reinforced concrete walls.
The walls are built in panels which
commonly vary from 6 to 30 feet in length.
As the soil is removed, a
bentonite slurry is placed in the excavation to prevent collapse of the
panel walls.
Steel reinforcement, if required, is placed in the exca-
vated panel and then concrete is placed by the tremie method.
Generally,
the reinforcement is either a steel cage or wide flange beams.
The initial phase of wall construction is the installation of guide
walls of lean concrete 3 to 10 feet deep on both sides of the trench.
The purpose of the wall is twofold;
(1) to prevent the soil at the top
from ravelling into the panels, and (2) to contain the bentonite slurry.
Installation of panels is accomplished by a variety of equipment
and techniques.
Mayer (1967)
and Saldier and Dominioni (1963)
give
detailed descriptions of the more common types of excavation equipment
used.
The equipment varies from drilling rigs (used in hard soils) to
-5-
hydraulically or mechanically operated clam buckets.
The minimum size
of panel installed is primarily dependent on the size of these units.
The plumbness of the concrete walls is important especially when
they are to be incorporated into the outer walls of a structure.
Their
vertical alignment may be maintained by using "Kelly" bars or H-piles
to guide the excavation device.
Another technique described by Kuesel
(1969) consists of installing steel soldier piles in preaugered, slurryfilled holes, which are then backfilled with gravel or lean concrete.
The slurry trenches are excavated using the piles as guides.
LaRusso
Dominioni
(1963), ,Mayer (1967), Meigh (1963), and Sadlier and
(1963)
recommend as standard construction procedure that
the level of the bentonite slurry be kept above ground water level.
Experience dictates that the minimum positive head of the slurry must
be 3 to 6 feet.
The implication is that if this head is not maintained
the panels walls fail.
In cases where the ground water table is at the
ground surface, levees are constructed along the longitudinal axis of
the wall and the guide walls are installed within them (Morgenstern,
1964).
Some authors [LaRusso (1963),
Meigh (1963),
Mayer (1967)]
describe
as standard practice the use of slurry densities varying between 65 pcf
to 75 pcf (1.04 and 1.20 gms/cm ).
construction experience.
The maximum density is based on
Densities beyond this value result in diffi-
culties in lowering excavation buckets, displacement of the slurry during
construction and pumping and recirculating the slurry.
The tremie concrete is placed at a 5 to 9 inch slump
ISadlier and
Dominioni (1963), Stacko (1968)] and often has a retarder added.
-6-
The
in place concrete has 28 day strengths of 3000-4500 psi.
reports data which suggest
Veder (1969)
that the strength of tremie concrete is
unaltered when placed in bentonite slurry.
The concrete while displacing the bentonite slurry leaves a minor
film around the reinforcing steel.
Cole (1963) quotes tests that indicate
this film results in some loss of bond strength between the concrete
and steel.
Based on his experiments, Cole recommends using bond lengths
of 1.25 to 1.50 those used in normal reinforced concrete.
Sadlier and
Dominioni (1963) report tests on the bond strength which show no appreciable reduction in bond strength.
They conclude this is a result of
cement particles replacing the bentonite clay on the reinforcing during
the normal hydration process.
2.3
PROPERTIES OF BENTONITE SLURRY
Experimental data reported by Jones (1963) on Fulbent 570 Bentonite
by Marsland and Loudan (1963) on Wyoming Bentonite, and by Piaskowski
and Kowalewski (1961) show that bentonite slurrys behave as a Bingham
material requiring a yield shear stress (T ) be exceeded before they
flow in capillary tubes.
Piaskowski and Kowalewski (1961)
is a function of the bentonite concentration.
show that
Jones (1963) reports
T
Tf
values of .01 and .2 gm/cm2 for concentrations of 2 and 10 percent,
respectively,
.7
gm/cm
2
Jones
while Marsland and Loudan (1963)
report values of .01 and
for the same respective concentrations.
(1963)
suggests the penetration of bentonite slurry in a porous
media may be expressed by the equation:
-7-
APR
2.3.1
L=
Where:
f
L
= length of penetration
R
m equivalent capillary radius for the media
AP m differential pressure
T f = Bingham yield shear stress
Marsland and Loudan (1963) investigated the factors affecting the
flow of bentonite slurrys in coarse to medium sands.
Their data show
that for uniform sands, with a coefficient of permeability for water of
.2 to .3
cm/sec, the average coefficient for slurrys are a factor of 8
greater for a 4 percent concentration and a factor of 56 greater for an
8 percent concentration.
The slurrys also exhibited threshold gradients
of 0.5 and 6 for the same concentrations,
respectively.
Mitchell (1960) showed that bentonite slurrys exhibit thixotropic
phenomena (strength gain with time).
2
of 0.4 gm/cm
2
and 2.2 gm/cm
Jones (1963)
gives shear strengths
for bentonite concentrations of 2 percent
and 7 percent, respectively, after 10 hours setting time.
al.
Piaskowski, et
(1961) show shear strenghts of 0.8 gm/cm2 and 2.5 gm/cm2 for set times
of 0.1 and 5 minutes, respectively, for an 8 percent bentonite concentration.
2.4
STABILITY OF SLURRY TRENCRES
In the literature there is no unanimity as to why slurry trenches
remain stable during excavation.
If one analyzes the stability of long
trenches as a Coulomb wedge, for slurry densities commonly employed in
construction, the factor of safety is often less than 1.
-8-
There have been
many opinions set forth as to the factors responsible for the trench
stability:
1.
Fluid pressure on the slurry against a membrane of
bentonite clay on the trench wall.
2.
Arching around short excavations.
3.
Increase of slurry density during excavation.
4.
Electro-osmotic flow of bentonite particles against
the trench wall.
5.
Strength gain of slurry due to thixotropic phenomena.
6.
Penetration of slurry into cohesionless soils.
It is generally accepted that the fluid pressure of the slurry acts
against a "cake" of bentonite clay which forms on the walls of the
trench.
Veder (1961) suggests this membrane of clay forms primarily
because of an electrical potential between the soil and slurry.
Such
a "cake" has been observed on the excavation walls both in the field
(Nash and Jones, 1963) and in laboratory experiments (Veder, 1963).
Morgenstern and Amir-Tahmassib (1965), for the assumption of a
Coulomb wedge, derived the following equation to analyze the stability
of slurry trenches in cohesionless soils:
Y
(cota (sin
N2
Y
cosa tano) +
W
=
w
where
-
Y
cosa + sina tano
i
=
Y
= unit weight of the slurry
the friction angle of the soil
Yt = total unit weight of soil
=
unit weight of water
-9-
M coseca tano)
2.4.1
t = the angle between the horizontal plane and the failure
plane
N =.(height of slurry in the trench)/(trench depth)
M = 1 -
[(depth to the ground water table)/(trench depth)]
Figure 2.2.2 shows a plot of factor
of safety versus the ratio of
yt
derived from this equation.
Nash and Jones (1963), for the assumption of a Coulomb wedge, give
the following equation for the factor of safety for trenches in cohesive
soil:
4S
2.4.2
U
H(y -Y
t f
)
FS
where
S
u
=
H, yt
the average undrained strength of the soil
Y
=
as above
The results of this equation are presented in Figure 2.2.3.
Both Figure 2.2.2 and 2.2.3 show the importance of the ground water
table and slurry unit weight in stabilizing long trenches.
Nash and Jones
(1963), Schneebli
(1964), and Tschebotarioff
suggest that for concrete panel installations,
the above two-dimensional
analysis does not apply and that a three-dimensional
to account for arching effects.
(1967)
analysis is in order
Model studies by Courteille (1969) tend
to substantiate a three-dimensional behavior.
Terzaghi (1943) has shown
that, for circular shafts in cohesionless material, the horizontal stress
required to stabilize the shaft walls decreases with depth and is less
than the initial horizontal stress because of "ring action".
Schneebli
(1964) suggests that the arching effect for a rectangular cut can be
accounted for by analyzing the problem as one similar to that of arching
-10-
Schneebli's analysis considers only the arching in a ver-
in a silo.
tical plane of the vertical stresses to the ends of the panel.
His
equation implies that for panel lengths equal to the trench depth, the
problem essentially reduces to one of plane strain (Rankine active
statel.
The implications of Schneebli's work are that the tops of the
panels are in plane strain and that below a depth equal to the excavation width arching plays an important part in stabilizing the panels.
DeNo (1969)
suggests that the factor of safety (FS)
of a circular
excavation in a perfectly elasto-plastic material is described by the
equation
[4 + (
FS
where
=
2
2
3 r+1
S
u
H
y
= the total unit weight of the soil
H
=
t
the excavation depth
Su = the undrained shear strength
r
= the ratio of the excavation radius devided by H
This equation can be used to determine a factor of safety for small panel
lengths if
the term (yt
-
Y f) is substituted for yt'
Morgenstern and Amir-Tahmassib
(1965)
state that one of the primary
stabilizing forces on slurry trenches in cohesionless soils is the
increasing slurry density due to suspended sand particles.
They demon-
strate that because of the Bingham behavior of bentonite slurrys, a
1 mm sand particle can be suspended in a clay slurry which has an initial
2
density of 1.08 gm/cm2.
Since this is the average initial density
employed in construction, it is probable that this could result in a
significant stabilizing force in sandy soils.
-11-
Veder (1961,
1963,
1969) indicates that the flow of bentonite
particles to the trench wall due to the electrical potential between
the soil and slurry environment could be a stabilizing factor.
However,
there has been no data to present to substantiate this hypothesis.
Morgenstern
(1963)
and Elson (1968)
suggest that a stabilizing
force to be considered in a stability analysis is the passive resistance from the slurry.
They recommend an extension of Prandtl's
analysis of a plastic material between two rigid plates.
For the
condition of excavation,
the slurry strength may be taken equal to its
Bingham yield strength.
However,
for such strengths this analysis
yields insignificant stabilizing forces compared to the fluid pressure
of the slurry.
Elson (1968)
proposes that the infiltration of slurry into the
soil improves the soil's strength characteristics.
The slurry pene-
tration can be described by the previously noted equation from Jones
(1963).
This equation yields a penetration
3 to 4 feet in medium to coarse sands.
length of approximately
According to Elson the thixo-
tropic strength gain and the development of.negative pore pressures in
this slurry filled zone can contribute up to an additional 10 percent
of stabilizing force in coarse sands and gravels.
2.5
FAILURES OF SLURRY TRENCHES
There are few failures of slurry trenches documented in the liter-
ature.
Schneebli (1964)
refers to the shallow sloughing of a 5-mile
long trench in granular soil when the water table rose to within 1.5
meters of the ground surface.
Morgenstern and Amir-Tahmassib
-12-
(1965)
described failures of slurry-filled
trenches 8 to 28 meters long in an
alluvium fill when the area became flooded.
The slurry in these
trenches had been undisturbed for a period of hours prior to the failure.
Both of these cases support the recommended standard practice of a 1
to 2 meter positive head of slurry to stabilize the trench walls during
excavation.
Mayer (1967) reports the failure of a trench in fine sand
due to a flocculation of bentonite slurry because of a high. lime concentration in the soil which prevented the formation of a bentonite
cake on the trench walls.
In all reported cases the failures were only
3 to 5 meters in depth.
2.6
SUMMARY AND CONCLUSIONS
The literature on slurry trench construction indicates the more
important mechanisms responsible for the stability of the slurry trenches
are:
1.
The bentonite slurry fluid pressure
2.
Increase in slurry unit weight
3.
Arching effects
4.
Slurry penetration into the excavation walls.
In most cases two or more of these factors contribute to the stability
of the trench.
Slurry trenches which have a length equal to or greater than the
depth approximate plane strain conditions and may be classified as long
trenches.
The few documented trench failures indicate that,
in granular
soils, the primary stabilizing forces for long trenches, 50 to 100 ft
deep,
are: (1) the fluid pressure against the trench walls; (2) a positive
-13-
head of bentonite above the ground water table; (3) an increase in
slurry density from suspended particles of sand.
It is felt that the
stability of these trenches should be analyzed using the coulomb wedge
method.
In the analysis a reasonable upper limit for increased slurry
density from suspended soil is 72 to 75 pcf.
Above these values
difficulties are encountered with excavation equipment and placing of
the tremie concrete.
Because of these problems,
the slurry is recycled
during trench excavation to maintain the slurry density below 72 to 75
pcf.
Slurry penetration as a stabilizing mechanism is limited to gravels
and coarse sands.
Even in these soils the penetration according to
Equation 2.3.1 would be small (1 to 3 ft).
only a minor stabilizing force.
Therefore, it would offer
For long trenches in clays it
a coulomb wedge analysis is also appropriate.
appears
However, it is unlikely
that during excavation the slurry densities will increase in cohesive
soils.
Therefore, a slurry unit weight equal to the initial
prescribed
value should be used when analyzing trenches in clay.
Where a concrete wall is installed in short panels (the length
being less than the depth) it is recommended that the stability analysis
be divided into two parts.
The first part concerns itself with the
top portion of the panel to a depth equal to the trench width.
portion of the panel should be analyzed as a long trench.
to the stabilizing forces for long trenches,
This
In addition
the guide wall offers a
stabilizing force to the tops of short panels.
At a depth below the
trench width, an additional stabilizing force is derived from the arching
of soil stresses around the trench.
in part, by using Schneebli's
This effect can be accounted for,
(1964) approach of equating the arching
-14-
action to that which occurs in silos.
ing of vertical stresses
This approach accounts for arch-
(and thus horizontal stresses) to the panel
ends.
-15-
CHAPTER THREE
NEASURED PERFORMANCE
3.1
BACKGROUND
This section. summarizes and evaluates the field measurements
for the South Cove Project.
Figure 3.1.1 gives an overall view of
Where the excavation was in close proximity to the Don
the project.
Bosco School a concrete wall was employed to help minimize ground
movements.
In this section, the excavation varied in width from 58.5
to 79.5 ft and in depth from 47 to 50 ft.
At the south bulkhead the
excavation was 40 ft wide and 45 ft deep.
Figure 3.1.2 shows a section of the excavation for station
113+40.
The soil profile is essentially the same along the excavaFigure 3.1.3 shows the soil properties determined from labor-
tion.
atory and field tests.
At all sections the bracing consisted of three levels of cross
lot struts.
During excavation for each strut level a berm
(8 to 12
ft wide at the top) was left against the sheeting until the strut was
installed.
The field measurements were concentrated primarily in the concrete
wall section.
Figure 3.1.4 shows the layout of the field instrumenta-
tion in the concrete wall section.
In addition to the instrumentation
shown, a series of settlement pins were installed on the Don Bosco
School.
At the south bulkhead slope indicators were installed on the
steel sheeting.
The field measurements are presented in two parts:
-16-
(1) measurements
made during installation of the concrete wall;
during excavation for the tunnel.
tive plots are presented for:
For the tunnel excavation interpreta-
(1) sheeting displacements;
and vertical movements of adjacent ground;
pressure; and (4) strut loads..
(3)
(2) horizontal
changes in pore water
The data from which the interpretative
plots are derived are presented in Appendix A.
construction,
(2) measurements made
The details of the
soil profile and instrumentation are also given in
Appendix A.
3.2
3.2.1
MEASUREMENTS DURING INSTALLATION
OF THE CONCRETE WALL
Pore Pressures
Figure 3.2.1 and 3.2.2 show the variation of total head during
the wall installation from December, 1968, through March, 1969.
During
this period a uniform increase of 1 to 2 ft was measured by the piezometers distant from the wall (PH-7 and PH-8).
Piezometer P-5 near the
wall showed erratic increases of total head, with maximum values
reaching 4 to 7 ft; whereas PH-4 showed a more stable behavior with
its maximum increase in total head reaching 4 ft.
This erratic fluctuation of P-5 and the 4 ft rise in head in
PH-4 is at least partially attributable to the panel construction.
A
comparison between panel installation and peak values of total head
did not show any definite correlation.
On the other hand, the steady
increase of the pore pressure in the distant piezometers is probably
more indicative of seasonal fluctuations in the water table.
The com-
plex nature of the soil and local influences, such as abandoned utilities,
-17-
coupled with construction events, make it
impossible to separate out
how much each of the above items contributed to the variation in total
head.
Figure. 3.2.3 compares the measured total heads before and after
the wall is installed with. the measured settlements.
The majority of
the piezometers behind the wall showed an increase in total head of 1
to 4 ft of water.
3.2.2
Ground Movements
Figure 3.2.4 shows the measured ground movements during install-
ation of the wall.
Horizontal movements were monitored by two slope
indicators, SI-1 and SI-3.
Vertical movements were determined by
settlement screws on the Don Bosco School.
The concrete panel lengths
in the vicinity of SI-1 were 10 ft in length, whereas those in front
of SI-3 were 20 ft long.
The maximum horizontal movements recorded were 0.7 in.at SI-1 and
0.4 in. at
SI-3.
The measurements show the maximum net horizontal
movements at the completion of the wall are less than the maximum
recorded, being 0.4 in.and 0.2 in.at SI-1 and SI-3,
respectively.
The vertical movements adjacent the wall were small.
Figure 3.2.4
shows the range of values recorded during the period of construction
from all the settlement screws.
The measurements show a net settlement
of 0.1 to 0.2 in.adjacent the excavation, whereas 40 ft from the wall a
net heave of about 0.2 in.was recorded.
-18-
3.2.3
Ground Movements versus
Pore Pressures
Figure 3.2.3 shows that at iQO ft or more from the wall the measured
heave is coincident with a pore pressure increase.
a similar trend as a function of time.
Figure 3.2.2 shows
At this distance the panel
installation will have little effect on the pore pressures; therefore,
the majority of this heave can be attributed to an increase in ground
water level.
A one-dimensional analysis predicts heaves of 0.1 to 0.2 in.
for a total head increase of 2 to 4 ft.
This is in good agreement with
the measured values.
Adjacent to the wall a net settlement is recorded at its completion,
although the measured pore pressures
increased.
In this area,
the com-
plexity of loading makes it nearly impossible to separate out elastic
movements associated with panel installation from consolidation movements
due to variations in total head.
An upper limit for the vertical move-
ment due to panel installation may be taken as the difference between the
movements adjacent to the excavation and those 100 ft or more away.
This
gives a maximum vertical movement of 0.2 to 0.4 in.
3.3
3.3.1
MEASUREMENTS DURING EXCAVATION
Performance versus Construction Progress
Figures 3.2.1 and 3.2.2 show, for the south and north ends of the
School, respectively, the settlement, total heads, and sheeting movements during construction.
In general, all the above behavioral changes
closely parallel the excavation progress.
After the excavation was
complete and the tunnel invert had been placed the measurements showed
-19-
little
further change except for the total heads measured in piezometer
PH-4.
The sheeting movements shown are those for the bottom of the
excavation.
There is a marked difference,
a factor of 8 to 10, between
the concrete wall deflections and those of the steel sheeting.
The
apparent scatter in data of SI-4 is related to the accuracy, of the slope
indicator measurements.
The average of the points is probably a good
indication of the actual movements.
The building movements are small, being less than 6 in.near the
excavation and on the order of 0.1 in.at the building line farthest
from the excavation.
The source of the settlements are the decrease
in total head and the inward displacement of the sheeting.
For the
distant points (Pt C and Pt 1) the movements are probably due entirely
to the total head changes since their distance from the excavation is
two to three times the depth.
On the other hand, the near points (7
and G) are within the influence range of the sheeting movements, since
they are within 50 feet of the excavation.
Therefore, it is most likely
the settlements and total head changes in the near points are influenced
by sheeting movements as well as drops in ground water table due to
dewatering the excavation.
The total head in PH-4 (Figure 3.2.2)
of PH-7 during the measurement period.
tively constant over the same period.
eventually increased to that
In contrast P-5 remained relaIt is felt this is related to
the construction progress from south to north.
The south end of the
excavation becoming closed up first and allowed the ground water table to
rise in this area.
-20-
3.3.2
Sheeting Displacements
The movements of the concrete wall and steel sheet piling are
shown on Figure 3.3.1.
There are two outstanding differences between
the behavior of the two bracing systems.
First the concrete wall
movements are only 10 to 15 percent those of the steel sheeting.
Second,
the concrete wall exhibits a more uniform lateral movement than does
the sheeting.
For the most part, these differences
can be attributed
to the construction technique used at each section and the relative
rigidities of the two walls.
wall vs. 3.75 x 10
(EI = 113.0 x 104 k-ft 2 for the concrete
k-ft2 for the steel sheeting.)
in construction procedure were: (1)
The major differences
the struts were preloaded in the
concrete wall section and not in the sheet piling area; and (2) dement
grout was used to shim the void between the concrete wall and wale,
whereas wood wedges were used in the steel sheeting section.
Figure 3.3.1 also shows the effect of strut removal to permit
unobstructed construction of the tunnel walls.
Both SI-4 and SI-li
(on the sheeting) exhibit significant increases over that measured prior
to the removal of strut levels C and D.
The effect of the concrete wall
having a stiffness (EI) 30 times that of the steel sheeting is illustrated,
in part, by the magnitude of additional movement of S-11 of 2.0 to 2.25 in.
versus .25 in for SI-4.
Also contributing to the higher movement of
the steel sheeting are the movements at strut B and the concrete invert,
which was constructed against the sheeting.
movement after removal of strut C and D.
SI-12 shows little additional
One possible explanation for
this is loose shimming, since struts in the steel sheeting area were not
preloaded.
Under such circumstances it
-21-
is possible for the sheeting to
deflect as if
the struts were not installed until the flexibility of
the bracing system was taken up by the inward sheeting deflection.
Figure 3.3.2 shows the deflected shape of the steel sheeting
wall at the south bulkhead area.
are shown on sketch.
The locations of the slope indicators
Comparing these movements with those on Figure
3.3.1, they are substantially smaller than those recorded by SI-l
and
SI-12, but significantly greater than those measured by SI-4 and SI-6.
Possible explanations for this decrease in steel sheeting movements
are:
(1)
corner bracing is used in this area and it
stiffer than cross lot bracing;
the bulkhead; and (3)
tighter.
However,
is inherently
(2) arching of soil pressure around
the shimming between the wale and sheeting was
the behavior at SI-9 and SI-10 is similar to that of
SI-1 in that additional movements of 1 to 1.5 in.were recorded after
strut loads C and D were removed.
3.3.3
Ground Movements
Figure 3.3.3 shows the range of concrete wall and adjacent ground
movements after the excavation was complete.
To illustrate movements
which occurred just during excavation, the position of the subsoil is
shown at the completion of the wall installation.
of the wall and
Both the movements
SI-1 show essentially constant horizontal displacements
over the depth of the excavation.
The total horizontal movement of
1.0 to 1.2 in.behind the excavation is greater than the maximum wall
movements of approximately .6 to .8 in.
SI-3 shows maximum movements
on the order of 1 in., but these movements are undoubtedly influenced by
the relatively large displacements of the steel sheeting adjacent to
this location.
-22-
The uniform displacement of the concrete wall is reflected in the
ground surface movements.
The data show the Don Bosco School exper-
ience an angular distortion (Lambe and Whitman, 1969) of less than
1/1000, which is less than the proposed acceptable limit (Bjerrum, 1964)
of 1/500 for this type of structure.
Comparison of the volumes of displacement between the end of
wall construction position and the maximum movements indicate that a
majority of the ground surface movements are initial
settlements.
The
volumetric displacement of SI-1 is approximately 4.5 ft3 versus 5.0
to 5.5 ft3 for the ground surface.
The excess surface displacement
can be attributed to consolidation settlements.
Figure 3.3.4 shows the settlement contours of the Don Bosco School.
The contours are essentially parallel to the excavation wall.
They also
indicate a relatively uniform settlement.
Figure 3.3.5 is a dimensionless plot of the vertical ground movements adjacent to the concrete wall.
This plot shows the maximum
settlements were, on the average, less than 0.1 percent of the excavation depth.
The sketch on the figure compares the data with the
empirical plot reported by Peck (1969) for excavations using soldier
piles and lagging or steel sheet piling.
The data falls at the lower
end of his suggested Zone 1 for good workmanship in soft to hard clays.
This clearly shows the concrete wall's ability to minimize initial
settlements adjacent to deep excavations.
Figure 3.3.6 shows changes in total head and bottom heave versus
excavation depth.
Unfortunately, the heave rods were destroyed before
reaching the final excavation depth; therefore, a continuous plot of
-23-
bottom movements was not acquired.
tunnel invert after its
bottom movements.
initial
Heave pins were installed in the
construction to monitor the trends of the
These readings are shown relative to their own
reading and do not show the total bottom heave.
The data show
that when the excavation depth reached 40 ft the bottom had experienced
a heave of .02
ft.
After constructing the invert, the bottom continued
to heave at a steady rate.
Figure 3.4.4 also shows that the final 24 ft of excavation depth
2
2.88 kip/ft ) resulted in 34 ft of total head drop.
(Au = 2.1
v
2
kips/ft ).
If one accepts that this excavation phase closely approximates
(Act
=
an infinite unloading, then the data suggests pore pressure dissipation
is occurring during excavation and that the measured heave is the result
of both initial
3.3.4
heave and swelling.
Pore Pressures
Figure 3.3.7 shows pore pressures versus depth adjacent to the
sheeting.
The pressure heads shown are the initial
values,
those at
maximum excavation depth, and the latest measurement recorded.
The
initial readings show the ground water was static prior to any construction events.
At maximum excavation depth the measured pore pressures
were below the static condition.
this latter
Except for PH-1 the pore pressures for
case compare favorably with the predicted pore pressures for
steady state seepage (as shown in Chapter 4) into the excavation with
the assumption that the wall permeability (i.e.,
permeability in the horizontal plane (i.e., kh).
kwall) equals the soil
In the field, seepage
was observed coming through the concrete wall panel joints directly in
-24-
front of piezometers PH-3 and P-4 to P-6.
through the steel sheeting joints.
Seepage was also observed
It is felt that this observed
in conjunction with the sand lenses in the upper portion of
seepage,
the stiff clay, is one of the main causes for the observed decrease
in pore pressure.
Some of the drop in pore pressure can be attributed
to the stress release from the excavation.
meter P-1-H on the center line.
Exemplifying this is piezo-
It shows zero pressure at full exca-
vation with dewatering only to the excavation bottom.
seepage it
should show a minimum static piezometric head of 15 ft.
the section of concrete panels in front of PH-1, little
the joints was observed.
for its
For steady-state
In
seepage through
This is probably the main factor responsible
relatively small pore pressure drop.
Figure 3.3.8 shows profiles of ground water levels and vertical
movements for full excavation at two sections.
At both sections the
phreatic surface indicates a flow of ground water towards the excavation.
SI-1 (Figure 3.3.1) recorded horizontal ground displacements over its
entire 100 ft depth.
ments, it
If one assumes constant volume elastic displace-
is possible these movements could have contributed to vertical
displacements up to 100 feet back from the excavation.
Therefore, the
source of the vertical settlements is most probably a combination of
these constant volume displacemnts
pressure.
However,
and a drop in the soil pore water
since the drop in pore pressure is also the result of
stress release resulting from sheeting displacements, the amount of
vertical displacement from each source is indeterminable.
The effect of
lowering the water table in this soil profile was probably small compared
-25-
At distances of 100 to 150 ft from
to constant volume displacements.
the excavation, where the effect of stress release is small, a total
head decrease of 8 to 10 ft over the 9 to 12 month excavation period
resulted in settlements of only 0.1 to 0.2
3.3.5
in.
Strut Loads and Horizontal Stresses
The variation of strut loads versus construction progress is
shown on Figure 3.3.9.
Strut loads were monitored in the concrete
wall section only, using vibrating wire strain gauges.
The values of
strut load shown in Figure 3.3.9 are the statistical averages and
standard deviations of the load per foot of excavation length.
Measured
values were corrected for temperature effects and strain gauge drift
as described in Appendix B.
The data show that the average B level
and C level strut loads increased 23 and 9 percent, respectively, with
increasing excavation depth.
This is contrary to what is usually found
for braced excavations using steel sheeting or soldier piles and lagging
(Peck, 1969; Lambe, et al., 1970).
The fourth construction stage, when
the two lower strut levels were removed, resulted in an average increase
of 30 percent in the B level strut load.
The scatter in the measured strut loads shown on Figure 3.3.9 is
quite large despite a rigid retaining wall, relatively uniform soil conditions, and controlled construction procedures.
were all less than the design values.
The measured loads
The largest difference occurred in
the D level struts where the maximum measured value, was about 0.5 the
design load.
The major reason for this is the design loads assume the
full initial static water pressure which acts on the concrete wall.
-26-
The Apparent Earth Pressure diagram (Terzaghi and Peck, 1967) for
the maximum average loads is given in Figure 3.3.10.
The diagram is
based on the assumption the struts carry a uniform earth pressure over
a distance half-way between strut levels.
The computed pressure diagram
falls within the range of apparent pressures (.2y tH to .4ytH) suggested
by Peck (1969)
3.4
for stiff
clays.
SUMMARY AND CONCLUSIONS
The field observations show:
(1)
Significant horizontal ground movements can occur during the
installation of concrete panels using bentonite slurry.
For panels 10
to 20 ft in length in stiff clay, measured movements were as much as
0.75 in.
On the other hand, vertical movements for the same conditions
are small, their order of magnitude being 0.1 to 0.3 in.immediately
adjacent the excavation.
(2)
one
in.
The concrete wall movements during excavation were less than
This was approximately the magnitude of movement which
occurred during the wall installation.
(3)
The maximum lateral inward ground movement behind the concrete
wall was 1.25 in.
(4)
in.
Steel sheeting experienced lateral movements of up to 6 1/2
This was 6 to 10 times the concrete wall movements.
(5)
Settlements behind the concrete wall were generally less
than one inch.
(6)
This corresponds to 0.1 percent of the excavation depth.
Settlements of the Don Bosco School were uniform.
distortion of the school was less than 1/1000.
-27-
The angular
(7)
Drops of 8 ft to 20 ft of water pressure head were measured
behind the concrete wall due to seepage through and around the wall and
total stress release during excavation.
(8)
In spite of essentially uniform soil conditions and construc-
tion procedure, measured strut loads showed a scatter of up to 50 percent
from the average.
(9)
The maximum measured strut loads were as little
as 0.5 the
design loads.
(10)
The measured strut loads give apparent earth pressures
which are within the 0.2 to 0.4 yt range recommended by Peck (1969) for
design.
-28-
CHAPTER FOUR
PREDICTED PERFORMANCE
This chapter presents predictions of the behavior of the concrete
wall section of the South Cove braced excavation.
Available methods
of analysis and design were employed for the predictions.
In some cases
where the methods did not model the problem well, they were appropriately
modified.
The predictions are divided into two sections:
1)
Predictions for the behavior during the installation of
the wall;
2)
Predictions for the behavior during excavation within the
braced cut.
Predictions for the wall installation include adjacent ground movements
and stability of the panels.
Predictions for the excavation phase include
pore pressures, sheeting movements, soil movements and strut loads.
Predictions of the measured field performance are an important
phase in the evaluation of braced excavations.
Only by comparisons of
predicted and measured performances for a number of excavations will it
be possible to establish the most appropriate analysis and design methods
to use in the design of braced excavations.
As the information of this
type is accumulated and evaluated, the design of bracing systems will
become more economical.
SECTION I
PREDICTED PERFORMANCE DURING WALL INSTALLATION
4.1
STRESS PATHS
The complexity of the stress changes and related movements during
-29-
the installation of the concrete panels by the slurry trench process is
best demonstrated with the aid of Stress Paths
(Lambe,
1967).
A stress
path describes for a soil element the continuous stress changes associated
with a particular construction activity.
peaks of the Mohr's Circles
[q =
It is generated by plotting the
(cT-a' )
2
) versus p =
(-
+
2
3]
for the
imposed stresses on the element.
Figure 4.1.1 shows the variation in horizontal stresses along the
sides of trench for each construction stage of the concrete wall at the
South Cove project.
The initial
horizontal stresses (aho' aho)
based on a correlation of overconsolidation
and Lambe, 1970).
ratio (OCR)
and K
are
(D'Appolonia
The variation of OCR was determined from the oedometer
test results in Figure A.3.1, Appendix A.
While excavating a panel, the soil experiences a decrease in horizontal stress to the fluid pressure of the slurry (a ).
The stresses
f
illustrated neglect any arching effects.
the placement of the tremie concrete.
The next construction phase is
It is assumed the full fluid
pressure of the concrete acts on the trench wall.
considered appropriate since the concrete:
This assumption is
(1) has a high slump (5 to
9 in.); (2) contains a retarder to prevent setting; and (3) is placed at
a rate of 12 to 20 ft of concrete depth per hour.
Using the fluid
pressure of the concrete the horizontal stresses after placement exceed
both the initial
total horizontal and total vertical stresses.
The stress path for a soil element at a 40 ft depth at South Cove
is shown in Figure 4.1.2.
movements.
Also shown are the trends of the related
Based on the oedometer test, results in Figure A.3.1, the
-30-
OCR of the element is approximately 3.
The pore pressure variation and
stress-strain relationships used to develop the stress path are derived
from synthesized data reported by Ladd et al (1971) for plane strain
tests on Boston Blue Clay (See Appendix E).
For the excavation stage a horizontal stress release occurs (A-B)
and results in the element undergoing an initial
the concrete
settlement.
Placing
(B-C) causes a. large increase in horizontal stress and
results in a heave of the ground surface.
Both phases are completed
within 1 to 3 days and, therefore, undrained conditions are assumed.
The remainder of the stress path is highly indeterminate.
The
excess pore pressures created by placing the concrete will dissipate
with time and result in settlement of the ground surface.
The magnitude
of the settlement and total horizontal stress on the wall at the time
of excavation will be directly related to the elapsed time since the
wall installation.
The implications of this latter behavior are dis-
cussed in Chapter 5.
4.2
PANEL STABILITY
A panel excavation of limited length compared to the depth is a
three-dimensional problem.
Theoretically correct analytical methods
for determining the factor of safety against the failure of the panel
walls in slurry trenches are not yet available.
The major problem with
present methods is they neglect the effects of arching in both the
vertical and horizontal planes.
One approach to analyzing the problem is to determine the factor
of safety for limiting conditions.
For panels of limited length the
-31-
bounds are the plane strain and axysymmetric state for minimum and
maximum factors of safety, respectively.
A coulomb wedge analysis was used to establish a lower bound of
the factor of safety for the panels at the South Cove Project.
4.2.1 shows the assumed conditions used in the analysis.
Figure
The factor
of safety was determined using both average field vane strengths and
strengths determined from Su
vo values (Ladd, et al., 1971).
The
factor of safety is taken as the ratio of the total existing slurry
pressure to the slurry pressure required for stability.
The computed
values were 0.83 using the vane shear results and 1.03 using the
ratio.
strengths derived from the S /a
u vo
DeNo's (1969) Equation
204.3.
bound for the factor of safety.
can be used to determine an upper
Using average values for the South
Cove profile, a factor of safety of 2.8 is obtained.
The actual factor
of safety probably lies closer to 2.8 because of the limited panel
length versus panel depth.
Finite element analyses were made in an attempt to determine at
approximately what length to depth ratio excavations behaved as a plane
strain problem and to verify the DeNo equation.
analysis are given in Appendix C.
The details of this
Axysymmetric and plane strain states
were selected as the bounding conditions of the problem.
The results
indicate that when the depth to width ratio approached 1, a circular
excavation began exhibiting behavior similar to one in plane strain.
This agrees
favorably with Schneebli's (1964) recommendation that plane
strain conditions prevail when this ratio approaches 1.
Comparing the
factors of safety from the axysymmetric case to that from the equation
-32-
as suggested by DeNo, the values are close.
The results show that an
)
80 ft diameter, 80 ft deep excavation fails at a slurry density (y
f
between 72 and 75 pcf. Using the average. normalized strength properties
over the depth of the excavation and a yf of 72 pcf, a factor of safety
of 1.03 is predicted by DeNo's equation.
Finite element analyses were made of a rectangular excavation
(Appendix C) in an infinite medium to examine the effect of horizontal
The initial stress conditions and soil strengths corresponded
arching.
to those on a horizontal plane through the panel excavation.
The
analyses indicate the stresses in this plane will arch around panels up
to 80 ft in length.
The only yielding of the soil occurred at the
corners.
4.3
PREDICTED GROUND MOVEMENTS
Figure 4.3.1 shows the predicted ground movements adjacent to the
panels at South Cove.
The predictions were made using the finite element
program Feast-3 (D'Appolonia, 1969) and the soil properties in Table
4.7.2.
The movements shown are for the bounding conditions of a circular
excavation, 20 ft in diameter, and an infinitely long excavation, 5 ft
in width.
The results are for the excavation stage, the effect of the
tremie concrete being ignored; therefore, they are compared to the maximum measured movements.
4.4
COMPARISON OF PREDICTED AND MEASURED PERFORMANCE DURING WALL INSTALLATION
The stress path analysis predicts that the horizontal movements
after wall installation should be away from the excavation face and
-33-
that pore pressures should increase.
However,
the measured movements
show a net displacement towards the panel excavation.
(1) the actual concrete pressure is below its
for this are
pressure;
Possible reasons
(2)
fluid
the modulus upon reloading is greater than that for the
unloading during excavation.
Nonetheless, the trends of the movements
measured in the field correspond to those predicted by the stress path
analysis.
The pore pressures followed the predicted trends, showing
an increase during the wall construction.
No major trench instability was observed in the field.
The only
instability observed was the sloughing of some clay chunks from the
panel face in the upper 30 ft where the soil was highly fissured.
This
sloughing was probably the result of the fissures opening in the very
stiff clay.
This implies that the overall factor of safety of the
panels was greater than 1.
SECTION II
PREDICTED PERFORMANCE DURING EXCAVATION
4.5
BACKGROUND
Many methods have been proposed to predict or analyze the perfor-
mance of a braced excavation.
Lambe (1970)
summarizes some of the
analysis methods and denotes what aspects of the excavation's behavior
each predicts.
This summary showed that a variety of methods are
required to cover the spectrum of braced excavation behavior.
In the past the prediction of braced excavation behavior was primarily by empirical methods or techniques with grossly simplifying
assumptions.
Peck (1943,
1967, 1969) and Tschebotarioff (1951)
-34-
developed
empirical design pressure envelopes for strut loads from field measurements.
Caspe (1966) and Armento (1970) describe methods for predicting
ground movements for known sheeting displacements.
Unfortunately their
methods require simplifying assumptions as to the pattern of ground
movements.
tion.
Prediction of sheeting movements have received little atten-
Golder, et al.
(1970) describe methods which employed elastic
analysis or engineering judgment based on field experience.
Recent developments in numerical methods and computer technology
have resulted in more rational approaches to the predictions of behavior.
Haliburton (1968), Boudier, et al. (1969), and Terahi and Balla (1968)
present numerical techniques to predict sheeting displacements and forces
as well as strut loads.
However, these solutions do not model the soil
behavior or soil-structure interaction particularly well.
The latest developments in analysis of the behavior of excavations
and retaining structures have employed finite element techniques.
Morgenstern and Eisenstien (1970) use finite elements to analyze earth
pressures against retaining structures.
Clough and Duncan (1971) inves-
tigated retaining wall behavior with a finite element analysis.
These
solutions are adequate to model the retaining wall problem but do not
represent the complex behavior of braced excavations.
Palmer and Kenney
(1971) model braced excavation behavior by combining two separate solutions.
The soil is modelled by finite elements, while the sheeting is
represented as a continuous member.
The solution is obtained by forcing
compatibility between soil and sheeting behavior using an iterative
technique.
A better representation of the bracing system was achieved
-35-
by Wong (1971)
by representing the sheeting with one dimensional bar
elements in combination with finite elements for the soil, the solution
being one step.
The approaches by Palmer and Kenney and by Wong, using
various assumptions, account for the effects of soil-structure interaction
and construction technique on the excavation behavior.
Both of these
approaches predict sheeting and subsoil movements as well as sheeting and
strut forces.
4.6
PORE PRESSURES
A knowledge of the pore water pressures in the vicinity of an
excavation is important for two reasons:
(1)
they can constitute a
major part of the stress applied to the sheeting;
(2) a decrease in
the ground water table will result in consolidation settlements of the
adjacent soil.
Both factors are of primary concern in the design of
a bracing system and require at least a prediction of the possible
limiting values.
The prediction of pore pressures outside an excavation is a complex
problem.
In cohesionless soils of high permeability the pore pressures
can range between static water pressure to pressures commensurate with
steady-state seepage.
The magnitude of the pore pressures are dependent
on the soil profile and the tightness of the sheeting.
In cohesive soils
the range of possible pore pressures is further complicated by excess
pore pressures from total stress changes.
In bracing systems employing
a cast-in-place concrete wall, the total lateral stresses undergo a
series of stress reversals during wall installation and excavation (see
Figure 4.1.2).
Superimposed on the excess pore pressures is the tendency
-36-
for a drawdown of the groundwater table due to seepage into the excavation.
Therefore, in these soils, the pore pressures are most probably
in a transient condition and are a function of the rate of construction
and the rate of pore pressure dissipation.
To predict the build up of excess pore pressures from total stress
changes, the following factors must be considered:
1.
The reversals in total stresses from the wall installation, excavation and preloading; and the corresponding
changes and accumulation of induced pore pressures;
2.
Transient condition of loadings;
3.
Two-dimensional nature and rate of excess pore pressure
dissipation;
4.
Choice of pore pressure parameters to describe pore
pressure changes in unyielded and yielded soil zones.
Since all these factors are highly indeterminate, it was felt that an
attempt to predict the excess pore pressures was not justified.
One condition of pore pressure change which was predicted was
that of steady-state seepage.
The predictions were made with the finite
element computer program FEDAR (Taylor and Brown, 1967).
This method
is selected because it readily handles multilayered anisotropic soils and
can account for seepage through the sheeting joints.
Other analytical
methods either lack the ability to model the complex boundary conditions
or are too cumbersome to use.
Figure 4.6.1 shows the finite element grid used in the FEDAR analysis.
The sheeting was modeled by a five-foot wide vertical strip of soil.
In
the analysis no absolute permeability values were assigned the soil layers.
Rather, relative horizontal and vertical permeabilities were assumed for
-37-
the various layers.
The ratio of the permeabilities selected was based
on the limited laboratory test data given in Figure A.3.1.
Figure 4.6.2 shows the relative permeabilities used in the analysis.
It is assumed that in the stiff
clay with sand lenses, the horizontal
permeability was 10 times the vertical permeability.
clay isotropic conditions were assumed.
In the underlying
The relative sheeting perme-
ability is a function of the sheeting type (concrete wall versus steel
sheeting), soil type, possible silting of sheeting joints, etc.
Deter-
ministic values of sheeting permeability are difficult to predict.
Therefore, permeabilities ranging from .01
vertical permeability were investigated.
to 10 times that of the soil
For all cases it was assumed
the ground water remained static 300 ft from the sheeting.
In addition,
the possibility of three-dimensional flow through the sand lenses was
neglected in the analysis.
Figure 4.6.2 shows the predicted pore water
pressures versus depth immediately adjacent the sheeting for the various
sheeting permeabilities.
of the sheeting.
The Unet is the net pore pressures on the back
The figure shows that the sheeting permeability has a
marked affect on pore pressure distribution even when its permeability
is one tenth that of the soil.
4.7
HORIZONTAL STRESSES AND STRUT LOADS
Available techniques for predicting strut loads are divided into
two types:
(1) empirical design methods derived from field measurements;
and (2) analytical methods based on elastic theories.
The design methods are founded on highly simplifying assumptions
for the lateral pressure on the sheeting.
-38-
These methods account for a
variation in soil type but do not separate out the effects of sheeting
rigidity, bracing type or method of construction.
They are intended to
give conservative values of strut loads.
The analysis methods, although they have a theoretical basis, often
require assumptions regarding soil and water pressures acting on the
sheeting and the deformation of the sheeting during excavation.
Both
of these factors are highly dependent on construction procedure and,
therefore, extremely difficult to foresee during the design stage.
4.7.1
Design Methods
For stiff
clays two design methods are available:
(1)
Peck's Method (Terzaghi-Peck, 1967, Peck, 1969)
(2)
Tschebotarioff's Method(Tschebotarioff,
1962)
Both methods use apparent lateral pressure diagrams determined from maximum measured strut loads on braced cuts using steel sheet piling or
soldier beams and lagging.
sis.
The diagrams are based on total stress analy-
The methods give the maximum strut load to be expected during any
stage of excavation.
Their applicability to the-more rigid concrete walls
has not as yet been verified.
Figure 4.7.1 shows the apparent pressure diagrams for the South Cove
soil profile.
The Peck diagram was modified for the bottom 12 ft from
the recommended diagram to account for the decreasing shear strength of
the clay.
This bottom section was assigned the maximum pressure instead
of the suggested linear decrease of pressure to zero at the bottom of the
cut.
No surcharge was added to the diagrams
for the adjacent building
since it is essentially a fully floating foundation.
-39-
The lateral stresses
are for average soil properties of the fill and clay.
The design strut
loads are obtained by summing the stresses acting half-way between
adjacent strut levels.
Table 4.7.1 summarizes the results of the analysis.
Figure 4.7.2 shows the pressure diagram used to design the reinforced concrete wall and to calculate the strut loads.
The diagram
assumes effective stress conditions and that stresses vary linearly with
depth.
as 0.6.
The ratio of horizontal to vertical effective stress was taken
Static water pressure and a surcharge load for the Don Bosco
School were added to the effective soil pressures.
The strut loads for
this diagram are given in Table 4. 7.1.
Analysis Methods
4.7.2
Three analysis methods were used to predict strut loads.
1.
Elastic beam analysis
2.
Elastic beam on springs
3.
Finite Element program BRACE II
They were
For the first two methods it was necessary to predict the lateral
pressure on the sheeting.
The pressures can range between the initial
lateral stresses and the active and passive states of stress depending
on the deflected shape of the sheeting.
Further, depending on the rate
of construction, the stiff clays could be in either an undrained or
drained state, or a transient condtion.
Flaate (1966) shows data which
suggests that soft to medium clays remain essentially undrained during
construction; unfortunately, there is little data on stiff clays.
For
South Cove soil profile, the sand lenses in upper stiff clays would most
probably permit a rapid dissipation of excess pore pressures from the
-40-
imposed total stress changes.
In addition, the lenses would probably
enhance the seepage into the excavation.
Therefore, effective stresses
and water pressures corresponding to the steady state seepage analysis
were used to predict the lateral pressures on the sheeting.
The lateral soil pressure (oh) was taken as:
h
h
where
v
= the effective vertical stress
a
v
K
= the ratio of the effective stresses
The appropriate value of K is purely a judgment value.
For cast in place
concrete walls, the K value is dependent on the residual lateral stress
after the wall is installed (see Section 4.1) as well as the assumed displacements and direction of sheeting movements.
Figure 4.7.3 shows the lateral stresses used in the two elastic
beam analyses for the condition of full excavation.
Because of the concrete
wall rigidity, relatively small inward displacements were expected.
Therefore, the K values were chosen more in light of the soil stiffnesses
than the sheeting displacements.
In addition, it was assumed the lateral
stresses after wall installation (see Section 4.1) were equal to their
original "in situ" values.
this section..
loose state.
it
The reasons for this are discussed later in
A K value of 0.5 was chosen for the fill
because of its
For the upper very stiff to stiff clays (K of 1.0 to 0.7),
0
was assumed that small displacements would result in large decreases
in ah; thus K was set equal to 0.5.
For the stiff to medium clays, it was
assumed little stress release would occur because of the small strains,
-41-
therefore, a K value of .4 was selected for these soils.
To the effective
soil stresses were added the pore pressures for a sheeting permeability
of 0.1 kv shown in Figure 4.6.2.
Figure 4.7.4 shows the strut loads obtained for the two elastic
methods.
level.
Table 4.7.1 summarizes the maximum loads for a given strut
In the analysis the sheeting was considered a continuous member
and given the properties of the concrete wall in Table 4.7.3.
The loads
were computed for each stage of excavation by the Computer Program STRUDL
(M.I.T., 1969).
In the elastic beam analysis it was necessary to assume support conditions below the excavation level.
A fixed support, 10 ft below the
excavation bottom was used in stage 1.
An inflection point represented by
a hinge support, was assumed in the remaining stages at 10 to 15 ft below
the excavation bottom.
The stresses from Figure 4.7.3 were applied to
the sheeting up to 3 ft below the excavation level.
Thereafter, movements
were considered to be sufficiently small that the soil stresses on each
side of the sheeting were equal; hence, only the net, water pressure was
applied below this depth.
In the elastic analysis with springs the sheeting below the excavation
level was supported by linearly-elastic springs.
The springs were assigned
a spring constant commensurate with the soil modulus.
set equal to AE/L:
The constant was
where A is the vertical distance halfway to the adja-
cent springs; L is the half width of the excavation; and E is an average
soil modulus from Table 4.7.2 which was taken as 300 a
.
The lateral
yo
stresses were applied in the same way as for the elastic beam analysis.
The last analysis method used to predict the strut loads was a finite
-42-
element computer program BRACE II.
This is an extended version of the
program BRACE described by Wong (1971).
The original program simulates the construction process for a
braced excavation in an isotropic, bilinearly-elastic material.
program performs a total stress analysis.
The
It models sequentially the
events of excavation and installation of struts.
It can also consider
such effects as prestressing of struts or additional movements at the
strut levels after installation.
Modifications to the program include
the facilities for handling anisotropic, bilinearly elastic materials and
the overstressing of the sheeting piling.
A detailed description of the
program modifications, as well as some investigations as to the program's
ability to model the problem,are given in Appendix D.
Predictions for strut loads were made for the same four stages of
excavation used in the elastic beam analysis.
finite element mesh for the analysis.
Figure 4.7.5 shows the
The material properties used for
the prediction are given in Table 4.7.2 and 4.7.3.
Since small sheeting
movements were expected the concrete walls' uncracked moment of inertia
was used in the analysis.
One particularly important aspect of the input data is defining the
value of K0 adjacent to the wall prior to excavation.
suggests K
has a significant effect on the first
0
D'Appolonia (1971)
yielding of the soil
within the excavation.
This first yielding could greatly effect the
predicted strut loads.
Figure 4.1.1 shows that the installation of a
concrete wall causes the soil to undergo complex changes in stress.
There-
fore, it is near impossible to determine what initial stresses exist after
-43-
the wall is installed.
If one accepts that during pore pressure dissipa-
tion the horizontal stress will decrease from those imposed by the concrete,
then the stresses will tend to reduce to the initial values.
Since the
clays in this analysis are overconsolidated with Ko values of about .7 to
1, it is assumed the lateral stresses on the wall after its installation
decreased to the initial "in situ" lateral stresses.
Figure 4.7.6 shows the, results of the BRACE U, Analysis and compares
them to the elastic methods.
The predicted strut loads by all methods
are within 30 percent of each other for all excavation stages, the only
exception being the D level strut load from the leastic spring analysis.
This agreement is considered good.
Probable reasons for the agreement are:
1.
The soil remained essentially elastic in the BRACE II
analysis, as shown in Figure 4.7.7.
2.
The assumed stress distribution for the elastic analysis
approximates that predicted by BRACE II, as shown in Figure
4.7.8.
The agreement between the stress distributions is somewhat fortuitous, the
assumed distribution having no formal basis.
Figure 4.7.8 shows that this
assumed distribution would be a poor approximation for the more flexible
steel sheeting.
4.8
SHEETING MOVEMENTS
Sheeting movements were predicted using the same analysis methods
and soil properties
as were employed in the strut load predictions.
all the predictions it
For
was assumed that after a strut was installed, the
sheeting was fixed at the strut level against further horizontal movements.
In addition, the effect of preloading was neglected.
Justification for these restricted movements of the concrete wall
-44-
stems from the fact that the struts were preloaded and the space between
the wales and the wall were filled with cement grout.
Elastic deformation
of the strut for the maximum design load would be, on the average,
less
than .2 in.
Figure 4.8.1 shows the predicted movements.
All the methods gave
essentially the same maximum sheeting movement of about 1 in.
the deflected shapes are quite different.
However,
This is largely the result of
the assumptions imposed in the elastic analysis methods.
4.9
VERTICAL MOVEMENTS
Vertical settlement adjacent to an excavation results from:
1.
Inward displacements of the soil adjacent to the excavation;
2.
Changes in pore pressures.
Movements from the first cause are termed loss-of-ground or initial movements; those from the second are consolidation movements.
4.9.1
Initial Movements
The complex nature of sheeting and ground displacements prevent the
use of elastic analyses to predict initial
movements.
As pointed out
in the literature review, methods have been proposed which associate
measured or predicted sheeting movements and vertical ground displacements.
However, these techniques require grossly simplifying assumptions regarding
the distribution of movements.
Presently, the only techniques which can be adapted to analyzing
initial movements are finite element compute programs.
However, even
these sophisticated analytical tools require several assumptions regarding
-45-
the bracing systems' behavior.
The finite element program BRACE II was used to predict the initial
ground movements.
For reasons explained in Section 4.8, after a strut was
installed that point of the sheeting was fixed against further lateral
movement.
The adhesion between the soil and concrete wall was considered
equal to zero.
This appears justified since the weak bentonite clay
"cake" which forms during excavation for the concrete panels is most
likely present between the wall and the soil.
BRACE II was modified to
account for this slippage by eliminating the axial stiffness of the
sheeting (see Appendix D).
The soil properties and element grid were the same as for the strut
predictions.
Figure 4.9.1 shows the predicted initial
the wall.
subsoil movements behind
In addition, the BRACE II analysis gave a bottom heave of
1.4 in.at the excavation centerline for a 40 ft depth of excavation.
4.9.2
Consolidation Settlements
Two factors causing consolidation settlements are:
1.
Dissipation of excess pore pressures from total stress
changes;
2.
Lowering of the ground water table.
As explained in section 4.6, it is virtually impossible to predict
excess pore pressures.
However, considering the general pattern of
sheeting movements, it appears reasonable to assume the total lateral
stresses decreased.
Data from plane strain tests on Boston blue clay
(Ladd, et al., 1971; Appendix E) suggest that negative pore pressures will
-46-
develop with decreases in stress.
The data further implies that the
excess pore pressures should become more negative with increasing overconsolidation ratio.
As shown in Figure 4.1.2, dissipation of these
negative pore pressures will result in a heave of the ground surface.
Based on pore pressure decreases from section 4.6 for steady seepage
into the excavation, predictions were made for one-dimensional settlements.
The compressibility characteristics used in the analysis are summarized
in Figure 3.1.3.
These consolidation settlements are added to the
initial settlements from the BRACE II analysis and are shown in Figure
4.9.1.
The total settlement shown neglects any heave which may have
occurred from the excess negative pore pressures.
COMPARISON OF PREDICTED AND MEASURED
PERFORMANCE DURING EXCAVATION
4.10
4.10.1
Pore Pressures
Figures 3.3.7 and 3,3.8 show that the pore pressures behind the
sheeting dropped below the static values.
Also, in Figure 3.3.7 are the
predicted pore pressures for steady state seepage, which are in good
agreement with the measured values at full excavation.
As explained in section 4.6, it is difficult to assess the exact
nature of pore pressure drops.
Therefore, it is not possible to state
with any degree of certainty whether or not the steady state seepage
analysis is appropriate.
However, the phreatic surface in Figure 3.3.8
strongly suggests a flow of ground water towards the excavation.
In spite of all the above uncertainties, it appears reasonable
to conclude that, in general, pore water pressures in stiff clays adjacent
to an excavation will be below static.
-47-
4.10.2
Strut Loads
Figure 4.10.1 and Table 4.7.1 compare the measured and maximum
predicted strut loads.
The maximum loads are used for the comparison
since these are the loads that the design methods are supposed to predict.
The figure shows that the range of predicted loads is as great
as the scatter in measured values.
In addition, many of the methods
predicted loads less than the maximum measured values, but greater than
the average value at a given strut level.
However, none of the predicted
values, except the conultant's design, contain a factor of safety.
It is interesting to note that the consultant's design greatly overpredicted the loads for the D level struts.
This is primarily due to the
assumption of static water conditions.
Figure 4.10.2 compares the measured and predicted strut loads
regardless of strut level.
In general, all the methods predict strut
loads greater than the measured value.
Tschebotarioff's method.
The only exception is
However, the amount by which each method over-
predicts the strut load varies greatly and, therefore, the factor of
safety against overstress would vary accordingly.
The scatter in the
data suggests that no one method is more applicable than the other.
Figure 4.7.3 compares predicted strut loads for each excavation
stage with the average measured loads.
The agreement is quite good
considering the many assumptions employed.
Probably the main reason
for the agreement is that the limited movement of the concrete wall and
the high soil strength combined to keep the soil in the elastic range.
It
is questionable whether the two elastic analysis methods would give
similar agreement in weaker soils.
-48-
Peck (1969) suggests that strut loads are directly related to the
stability number N for a vertical cut:
N
t H
S
u
Yt = total unit weight of the soil
where
= depth of the cut
H
Su = undrained strength of the clay from unconfined compression
tests
Using average values for the soil parameters, N z 1.5 for the South Cove
excavation.
Peck suggests that soils with N values less than 4 remain
essentially elastic during excavation whereas above 6 substantial yielding occurs.
The agreement between the measured strut loads and those
predicted from the elastic analysis tend to substantiate Peck's recommendation.
Also in support of this factor are the limited yield zones shown
in Figure 4.7.7
4.10.3
Sheeting Movements
Figure 4.8.1 compares the maximum concrete wall movements with
the predicted values.
The trends and magnitudes of the movements are
in good agreement, particularly with the elastic spring analysis.
Again,
this agreement is most probably a result of the limited wall movements
keeping the soil in the elastic range.
Figure 4.10.3 shows the measured and predicted maximum sheeting
movements of both the concrete wall and steel sheeting piling.
The
assumptions for the steel sheeting analysis were the same as for the
concrete wall.
The trends of the sheeting movements are in good agreement.
-49-
the predicted magnitudes for the steel sheet piling grossly
However,
underestimate the measured displacements.
This discrepancy is the
result of the large movements which took place at the strut levels after
a strut was installed.
Also, these added movements could have caused
some partial yielding of the soil within the excavation which permitted
significant movements below the excavation bottom leading to larger
sheeting deflections.
According to the stability number N of 1.5 for the cut (section
4.10.2) the excavation wall should have been able to stand without any
sheeting.
Yet, the soil behind the steel sheeting experienced inward
movements of up to 7 in.
One probable cause of these large movements
was a decrease in the strength of the over-consolidated clay from the
undrained strength to the drained value.
Therefore, in stiff clays
with either fissures or sand lenses, the use of the undrained strength
for the stability analysis is questionable.
Also suspect is Peck's general recommendation that clays with N
values less than 4 remain in the elastic state during excavation.
Whether or not these soils remain unyielded appears to be a function of:
1.
The rate at which the pore pressures dissipate;
2.
The ability of the bracing system to prevent lateral
movements and therefore keep strains below the yield
value.
4.10.4
Ground Movements
Figure 4.9.1 compares the predicted and measured ground movements
adjacent the concrete wall.
Note that the measured vertical movements
are much less than the predicted total settlement.
This difference is
due, in part, to neglecting the potential ground heave from the dissipation of negative excess pore pressures.
The predicted and measured settlements do not agree immediately
adjacent to the wall.
This discrepancy is most probably the result of
the difference in predicted and measured movements for the top portions
of the wall.
4.11
SUMMARY AND CONCLUSIONS
The measured ground movements, associated with the installation
of the concrete wall, were within the predicted limits of the ground
movements by plane strain and axysymmetric finite element analyses.
However, the predicted movements are largely dependent on the selected
soil parameters.
A significant aspect of the results is that the
trends of the predicted movements compare favorably with the measured
movements.
Comparisons between the predicted and measured behavior of a braced
cut with a reinforced concrete wall in an overconsolidated clay indicate:
(1)
Peck's (1969) recommended design pressure envelope of .4Y H
t
for stiff clays with N values less than 4 predicts strut loads on the
safe side, whereas Tschebotarioff's (1962) envelope underpredicts the
strut loads.
(2)
For rigidly constructed bracing systems one can make reason-
ably good predictions of strut loads using elastic beam theories or
finite element analyses.
technique overestimates
However, for a given strut level, each
the maximum measured strut load by varying
amounts.
-51-
(3)
Elastic beam theories and finite element analyses can give
reasonable predictions of sheeting deflections providing the bracing
system is constructed such that only minor movements occur at a strut
level after installation of the strut.
Where no special precautions
are taken to restrict sheeting movements above the excavation bottom
and sheeting deflections of an unpredictable magnitude can occur, the
elastic theories and finite element analysis can grossly underpredict
the sheeting movements.
(4)
Neglecting the effect of the negative pore pressures from
total stress release results in overestimating the settlement of the
ground surface behind the excavation wall.
-52-
CHAPTER FIVE
PARAMETERS AFFECTING THE PERFORMANCE
OF A BRACED EXCAVATION
5.1
BACKGROUND
The primary reason for employing a cast-in-place reinforced concrete
wall in a braced excavation is to minimize subsoil movements adjacent to
the excavation.
Table 5.1.1 summarizes the limited number of documented
cases for excavation with reinforced concrete walls.
They all suggest
the walls are capable of keeping lateral ground movements to less than
1.5 in.regardless of soil type or depth of excavation.
However, in exca-
vations employing concrete walls, consideration must be given to ground
movements during the installation of the wall as well as during excavation.
As explained in Chapter 2, probably the main variable affecting movements
during wall installation is the length of panel excavation.
It appears the main factors which control movements during excavation are:
1.
Rigidity of the sheeting
2.
Vertical spacing of struts
3.
Stability of the excavation bottoms
4.
Details of prestressing and wedging of bracing
Other items which must be considered are the soil type, time lag before
struts are installed, and removal of struts after full excavation.
Two approaches may be taken to assess the degree to which each of
the above-mentioned variables affects ground movements.
The first is to
employ laboratory models or analytical techniques to simulate the behavior.
The second entails collating available records -f fiild measuremeftts arid
-53-
developing interpretive plots.
Because of the sparseness of available field data the first approach
was adopted by the writer to evaluate the implications of the performance
of the South Cove excavation.
Stress paths were used to define the
fundamental behavior of an excavation.
The finite element program
BRACE II was used to investigate those factors that affect movements.
In addition, based on limited field measurements, an evaluation was made
of the effectiveness of bracing details on reducing the sheeting movements.
5.2
STRESS PATH ANALYSIS
Stress paths are an ideal vehicle by which one can gain a funda-
mental understanding of a complex problem, such as a braced excavation.
Admittedly in many cases it is not possible to obtain the exact stress
changes in the ground associated with a particular loading.
1tf is also
realized that one cannot always model the known stress changes in the
laboratory to obtain necessary soil parameters in order to predict subsoil displacements.
Nonetheless, the stress path is an effective tool
for predicting trends of movements.
Figure 5.2.1 shows the stress changes associated with the installation of a wall in a soft, normally consolidated clay.
for the simple case of a long trench.
The stresses are
This simplified condition of the
actual case was selected so as to eliminate the complex effects of
arching which exist around the more commonly used short panels.
Figure 5.2.2 shows the stress path for an Element A a few feet
from the panel face.
The effective stress changes and vertical strains
of the element for the trench excavation (A - B) are derived from the
-54-
plane strain data from Ladd et al.
(1971)
(See Appendix E).
There-
after the stress paths are the writer's judgment as to what the tendency
will be for the stress paths.
It is emphasized that the intent of using
the stress path is only to shead some light on the qualitative behavior
of excavations with concrete walls and not to give quantitative results.
The stress paths suggest some interesting behavioral trends.
During wall installation in a soft clay, the stress release (A - B)
may be sufficient to induce yielding in the element.
The horizontal
stress increase from the tremie concrete (B - C) may well bring the
element close to an extension failure.
The strains associated with
these stress paths suggest first an initial settlement and then heave.
The initial horizontal and vertical subsoil movements during
excavation could be quite large.
The data from South Cove shows hori-
zontal displacements of about .75 in.
in overconsolidated soils.
In
soft, normally consolidated clays or loose sands, it seems reasonable
that movements would be larger than those at South Cove.
Such movements
could reach magnitudes sufficient to be detrimental to the safety of
utilities and structures in close proximity to the concrete wall.
After concrete placement the excess pore pressures will dissipate
resulting in a consolidation settlement of the subsoil.
The concrete
will most likely harden before complete dissipation of the pore pressures.
Since the rigid wall will not strain laterally, further pore pressure
dissipation will result in a reduction of the lateral stress on the wall.
Even at full dissipation it seems unlikely that the horizontal stresses
will reduce to their initial
values.
-55-
The net increase in horizontal stress from placement of the concrete
is beneficial in the subsequent excavation (D - E).
stresses in the soil prior to excavation.
It lowers the shear
Thus the soil can experience a
larger stress decrease or corresponding lateral strain before yielding
occurs.
On the other hand, the soil within the excavation also is at a
lower shear stress prior to excavation.
experiences an extension loading.
During excavation this soil
Therefore, it is closer to yielding
and thus can resist less vertical stress release (or horizontal stress
increase) before becoming overstressed.
If one accepts the aforementioned effects of wall installation on
horizontal stresses, then some behavioral trends may be deduced from
the stress path (D - E).
If a rigid bracing system is constructed which
limits the soil to elastic movements, then small horizontal and vertical
displacements will most probably result.
This is born out somewhat by
Figure 4.10.3. A comparison between the predicted sheeting displacement
for an overconsolidated clay of both the steel sheeting and the concrete
wall gave essentially the same movements for the condition of no movement
at the strut levels.
As shown in Figure 4.7.7 the soil remained essentially
elastic for both sheetings.
In contrast, the measured movements showed large displacements,
especially at SI-ll (Figure 4.10.3). These movements may be attributed
to a loss of strength resulting from negative pore pressure dissipation.
But more important, large lateral movements occurred at the strut levels
after strut installation.
This large lateral strain could have caused a
yielding of the soil above excavation level.
-56-
At SI-10 where less movement
occurred at the strut levels, the sheeting displacements are significantly reduced.
5.3
BRACE II ANALYSIS
Three BRACE II computer runs were conducted to gain further insight
into the parameters that control sheeting movements.
mesh shown in Figure 4.7.5 was used for each run.
The finite element
The soil was assigned
the anisotropic properties for the normally consolidated clay from
Appendix E.
levels.
5.3.1
In each run the sheeting was assumed fixed at the strut
The ground water level was at a 5 ft depth.
Sheeting Stiffness
Figure 5.3.1 shows predicted movements for sheeting with an order
of magnitude difference in stiffness.
(The stiffness was defined as
the product of the Young's modulus, E, and the Moment of Inertia, I.)
5
The El of 4 x 10
the EI of 4 x 10
2
p-ft
4
corresponds to that for a concrete wall, while
2
p-ft
is representative of steel sheet piling.
Com-
paring the sheeting movements for the case of strut levels at 9 ft
intervals the figure shows the stiffer wall reduces the movements by a
factor 3 to 4.
Figure 5.3.2 shows the development of the yield zones during excavation.
Despite the difference in sheeting stiffness the yield zones
for both EI values merge below the sheeting at the same excavation level.
The results shown in Figure 5.3.1 and 5.3.2 suggests there is some
correlations between sheeting stiffness and yield zones.
4 x 10
4
p-ft
2
,
For an EI
=
a large increase in movement occurred at the excavation
base going from excavation stages 4 to 5.
-57-
At stage 5 the yield zones
extend almost to the bottom of the sheeting,
little resistance to sheeting movements.
Hence,
the soil offers
The main resistance is from
the stiffness of the sheeting which now acts essentially as a cantilever.
For earlier excavation stages the bottom of the sheeting was
This
embedded in unyielded soil which restrains sheeting movements.
behavior contradicts somewhat the recommendations of Peck (1969).
He
suggests that at a stability number, N m 6, the soil becomes plastic
adjacent to the sheeting and large movements will be experienced.
normally consolidated soils with a constant strength ratio of S
U
/
For
yO
=
.34, the stability number remains essentially constant and equal to 9.
Therefore, large movements should be experienced at the excavation
base for each stage of excavation.
However, during excavation no
prominent bulging was predicted until the fifth excavation stage.
This
suggests that sheeting stiffness and embedment depth may have an important influence in reducing movements.
Further examination of the yield zones in Figure 5.3.2 show they
intersect below the sheeting and eventually the base of the soil layer.
Comparing the movements shows that the stiffer sheeting did not exhibit
a significant increase in displacement at the fifth, or in fact, at
latter stages of excavation, as did the sheeting.
This is attributed to
the walls' higher stiffness and, thus, ability to resist lateral displacements.
5.3.2
Vertical Strut Spacing
Deflections of structural members with uniformly varying loads are
inversely proportional to the stiffness ratio:
-58-
4
L
where L = the support spacing
E = Young's modulus
I = Moment of Inertia
Figure 5.3.1 shows the deflections for a sheeting EI of 4 x 10
p-ft2 for 9 and 15 ft strut spacings.
The increased strut spacing
results in approximately doubling the sheeting displacements while the
stiffness decreases by a factor of 8.
A comparison of the yield zones (Figure 5.3.2) shows, as one would
expect, that the extent of the yield zones was slightly greater for
the larger strut spacing for the same excavation depths.
This increase
in yield zone area results in longer spans between the firm supports
of the struts and unyielded soil and, therefore, contributes to the
larger sheeting displacements.
5.3.3
Bottom Instability
Two stages of bottom instability are important in braced cuts.
The
first stage corresponds to first local yield adjacent to the sheeting
permitting increased sheeting movements.
The second is the gross failure
of the excavation bottom.
According to theoretical studies (Terzaghi, 1942) for infinitely
long excavations in an isotropic media, first yield occurs at the
corners of the excavation at a atability number N
c
of 3.14.
D'Appolonia
(1970) presents data on the factor of safety (Nb/Nc) required to prevent first
yield as a function of the shear stress ratio:
-59-
f
- K0
S
u
2 --
vo
is the extension strength ratio.
where S /6
U
These results are repro-
VO.-
duced in Figures 5.3.3 and 5.3.4,along with the values of Ncb recommended
by Bjerrum and Eide (1956).
The applicability of these results is sub-
ject to question since neither the theory nor D'Appolonia's data consider
the effect of sheeting stiffness.
Figure 5.3.5 shows the excavation bottom movements at various
depths for the two sheeting stiffnesses previously discussed.
Up to
the fourth excavation stage the movements were approximately uniform
and of the same magnitude.
However, at the fifth excavation stage, the
bottom bulged adjacent to the less stiff sheeting.
This heave corres-
ponds to the increase in sheeting movements (Figure 5.3.1) and approximates joining of yield zones (Figure 5.3.2).
The stiffer sheeting,
which was capable of resisting lateral displacements, did not exhibit
the same local heaving.
When the excavation reached the seventh stage
and the yield zones extended a considerable distance below the bottom
of the sheeting, significant heave occurred even with the stiffer sheeting.
These results suggest that very stiff sheeting can retard local
bottom heave when the yield zones do not extend a large distance below
the sheeting bottom.
However, in the case of deep excavations, where
the sheeting penetration is shallow and therefore large yield zones can
develop,
sheeting stiffness will have little
effect on bottom stability.
This observation is in agreement with model studies described by Peck
(1969).
Unfortunately,
there are no field observations available to
-60-
correlate with these implications
from the BRACE II predictions.
Figure 5.3.6 shows for a normally consolidated clay, the factor
of safety required to prevent local bottom heave, or significant first
yield, as a function of sheeting stiffness and strut spacing.
value of N
c
The
is computed by setting the maximum shear stress imposed on
the soil, as computed by BRACE II, equal to Su.
(The first corner
element is neglected because of the high stress concentrations.)
Where
a factor of safety of one is given, this implies that the sheeting is
stiff enough and extends deep enough to resist any instability bf the
excavation bottom.
The results suggest that strut spacing and, more
significant, sheeting stiffness are important parameters when consider-
ing first yield of an excavation bottom for H/B values less than 0.7
0.8.
For larger H/B values the sheeting stiffness is less important and
the factor of safety of 1.4 against first local yield corresponds quite
closely to that suggested by D'Appolonia's data.
5.3.4
Comparison with Field Observations.
The South Cove data gives some insight on the effectiveness of
sheeting stiffness and excavation geometry to reduce moments.
Figure
5.3.7 shows significant increase in movements occurred at SI-10 and
SI-li when the excavation depth reached 40 ft, whereas no large movements
were experienced at SI-9 and SI-6.
versus those of SI-10 and SI-l
The difference between SI-9 movements
can be related to the excavation geometry.
The soil at the 40 ft depth has an OCR of about 3 to 4.
From Figure
5.3.3 the factor of safety required to prevent local yield for B/L = 0
is about 2.6.
Using average values for the soil properties the computed
-61-
factor of safety is 2.7.
However, for SI-9, which is close to the South
Bulkhead an L/B ratio of 1 is more appropriate.
factor of safety is 3.2.
For this case the
This suggests no local yielding should have
occurred and, therefore, the movements below excavation are reduced.
SI-6 shows no marked increase in movement indicating that the
concrete walls stiffness was sufficient to prevent lateral displacements.
The importance of wall stiffness is further exemplified by the
data (Table 5.1.1) from Kuesel (1969).
The 70 ft deep excavation in
soft Bay Mud has a factor of safety of 0.8 to 1.0 against local yield
yet sheeting movements were less than 1.5 in.
Figure 5.3.6 suggests sheeting stiffness lowers the factor of
safety for first local yield yet good agreement was obtained between
field observations and D'Appolonia's results.
One possible reason for
this apparent agreement is that large sheeting movements took place
after a strut was installed (Figure 5.4.1).
This results in a tendency
for the sheeting to deflect as if the strut did not exist.
This negates
any effect of the sheeting stiffness because of an increased unsupported
span which transfers higher loads to the excavation bottom and could
induce premature yielding.
Nonetheless, this limited data suggests that
correlating sheeting stiffness and predicted first
local yield with
sheeting movements will lead to a better understanding of how to control
sheeting movements.
5.4
SHEETING MOVEMENTS
Sheeting displacements during excavation can be categorized as
follows:
-62-
1)
Movements below excavation level from elastic and
plastic deformations.
2)
Movements above excavation level which are a function of the bracing details.
For a given construction procedure and soil profile, movements
below excavation are unavoidable.
However, movements above excava-
tion level could be reduced by better construction practices which
increase the rigidity of the bracing system at a strut level.
Improve-
ment in the rigidity will no doubt reduce the sheeting movements
above the excavation depth.
However, it is probable that this could
reduce movements below the excavation level by limiting the deflection
of the sheeting and thus reducing the induced strain in the soil preventing local yielding.
D'Appolonia (1970) suggests separating the movements above and
below the excavation level by constructing diagrams as shown in Figure
5.4.1.
These diagrams, which are for the South Cove excavation, are
the total measured movements.
By plotting the sheeting displacements
for each stage of excavation, the movements above and below excavation level are readily determined.
movements below excavation level.
The shaded area represents total
The area between the maximum sheeting
movement and the line defining movements below excavation reflects the
total movement above the excavation level.
In Figure 5.4.1 SI-11 and SI-12 show a major portion of the
sheeting deflection took place above excavation level.
They also
show substantial movements of up to 4 in. below the excavation bottom.
SI-10, which exhibits maximum sheeting deflections approximately one
-63-
half those of SI-11 and SI-12, shows a majority of its
below the excavation level.
movement occurred
This latter case most probably represents
the minimum movement one could expect in this profile with steel sheet
piling and the construction techniques used.
SI-6 shows the concrete
wall experienced movements below excavation level.
The percent of the
total volumetric displacement below excavation level is approximately
the same as for SI-10.
However, the magnitude of the displacements
are much less and more uniform.
and its
In light of SI-10's nonuniform movements
maximum deflection of 2 in, it
appears that even in stiff
soils
concrete walls may be necessary to limit ground movements to tolerable
limits.
As previously stated, it is most likely that bracing details can
influence movements below excavation level.
Figure 5.4.2 shows sheeting
displacements for two test sections, A and B, at a-braced cut in Boston
reported by Lambe, et al. (1970).
the same for each section.
The soil profiles are essentially
However, because of excessive movements at
Section B the bracing details were improved at Section A.
Most notable
of these changes were the changing of wood shims from soft pine to oak,
better control of prestressing, and minimal excavation depths before
the next lower strut level was installed.
Because of the improved con-
struction, Section A experienced a lower maximum movement (4 inversus
6 in).
25
In addition, in Section A movements below excavation level were
percent lower and those above excavation level are 30 percent lower
than at Section B.
These limited data suggests that movements above excavation level
-64-
can significantly affect movements below the excavation.
However, more
field observations of this type are needed to verify this observation.
5.5
SUMMARY AND CONCLUSIONS
Stress paths for cast-in-place concrete walls suggest:
(1)
Significant movements may occur during wall installation in
soft soils.
(2)
Installation of the wall causes changes in the "in situ" soil
stresses which could result in a reduction in subsoil movements.
Analysis using the finite element program BRACE II indicate:
(1)
Sheeting stiffness is an important factor in reducing subsoil
movements and bottom heave.
(2)
The depth of penetration of sheeting below an excavation can
influence sheeting displacements, bottom heave, and subsoil movements
behind the sheeting.
(3)
Decreasing of strut spacing can result in reduced sheeting
deformations.
A review of limited field data suggests:
(1)
One important variable influencing subsoil deformations
is sheeting stiffness.
(2)
Movements at strut levels after strut installation can affect
the uncontrollable movements which occur below excavation level.
-65-
CHAPTER SIX
CONCLUSIONS
6.1
GENERAL
Field measurements at South Cove have given a considerable amount of
valuable data and insight on the performance of bracing systems with castin-place reinforced concrete walls in stiff clays.
Where possible , the
correlatioh of these field measurements with values predicted by available
analytical methods allows an assessment of the applicability of these
methods for predicting the performance of braced cuts employing concrete
walls.
Since the data in this study is from one site, the conclusions set
forth in this chapter require further verification by additional instrumentation programs and theoretical studies.
6.2
MEASURED PERFORMANCE
The field measurements program for the deep excavation in stiff clay
at South Cove show:
(1)
The installation of a cast-in-place concrete wall by the slurry
trench process, in panels 10 ft to 20 ft long and 80 ft deep,
can cause significant lateral ground movements but results in
inconsequential vertical ground movements.
(2)
Reinforced concrete walls can limit the horizontal and vertical
ground movements adjacent to a braced excavation to values less
than 1 in. if good construction methods are exercised and if the
wall penetration is deep.
(3)
Steel sheet piles can experience lateral movements of up to
twice those measured for a concrete wall for the same bracing
-66-
geometry and good construction techniques.
Where poor con-
struction methods are employed the steel sheet piles can undergo movements 6 to 7 times those of the concrete wall.
(4)
The loads in a given level of bracing supporting a concrete wall
can vary as much as 50 percent above and below the average
measured value even with good construction practice and uniform
soil profile.
(5)
The pore water pressures behind a concrete wall will drop below
the initial values during excavation.
The pattern of the decrease
in pore pressure indicates the reductions are due, in part, to
seepage through the concrete panel joints.
6. 3
PREDICITION OF MEASURED PERFORMANCE
(1)
The installation of a cast-in-place walls causes
complex stress changes in the adjacent ground which hamper
making accurate predictions of the performance of braced
excavations supported by concrete walls.
(2)
The stability of slurry trenches is not fully understood.
Presently, it appears that the method of analyzing the
stability of slurry trenches should be related to the length
to depth ratio of the trench.
(3)
The agreement between the measured sheeting and ground movements and those predicted by BRACE II is highly dependent on the
ability to estimate the sheeting deflections which will occur at
a strut level after a strut is installed.
-67-
Since this variable is a function of construction technique
and the quality of workmanship,
these predictions cannot be
made with any degree of confidence.
(4)
Predictions of ground settlements are hindered by the inability to predict:
a) the magnitude of pore pressure
change relate to the lateral stress releases; b) the rate at
which the pore pressures dissipate.
(5)
The large scatter in measured strut loads suggest Peck's
approach of using lateral pressure envelopes for determining design loads is appropriate for bracing systems with
concrete walls.
(6)
The measured strut loads gave an apparent lateral pressure
envelope in good agreement with the envelope recommended
by Peck for stiff clays.
However, it would be presumptuous
to conclude that the design envelopes recommended for the
design of bracing systems with flexible sheeting in other
soil types apply to braced excavations with concrete walls.
(7)
Predictions of pore water pressures are restricted to the
limiting condition of steady state seepage into the cofferdam.
The accuracy of these predictions are influenced by the
ability to assess the relative permeability of the sheeting
and the soil.
(8)
The indeterminate nature of the parameters governing the
behavior of braced excavations precludes obtaining accurate
predictions of any aspect of a braced excavations behavior
-68-
by analytical methods unless special construction measures
are taken to control the paramenters governing that particular
aspect of behavior.
6.4
PARAMETERS AFFECTING THE PERFORMANCE OF A BRACED EXCAVATION
(1)
Placement of tremie concrete for cast-in-place concrete walls
increases the horizontal effective stresses in the soil;
therefore, the soil can experience higher lateral strains before
yielding than it could initially.
This will result in reduced
ground movements but higher strut loads in comparison to bracing
systems with steel sheet piles.
(2)
The bending stiffness of the sheeting in conjunction with the
sheeting penetration below excavation level are important parameters in reducing the ground movements within and outside the
excavation.
(3)
Decreasing vertical strut spacing results in reduced ground
movements above and below excavation level.
However, a re-
duction in strut spacing is not as effective in reducing movements as is an increase in bending stiffness of the sheeting.
(4)
For a depth-width ratio of 1 and a sheeting penetration of .5H,
BRACE II analysis show, for excavations in soft clay, a stiff
concrete wall is no more efficient in preventing instability of
the excavation bottom than the more flexible steel sheet-piling
walls.
(5)
Parameter studies show BRACE II is an excellent method for aiding
the design engineer in making rational judgements relative to
the expected behavior of a bracing system.
-69-
BIBLIOGRAPHY
Abbreviations:
ASCE
America Society of Civil Engineers
ICSMFE
International Conference on Soil Mechanics
and Foundation Engineering
JSMFD
Journal of Soil Mechanics and Foundation Division
SGDMEP
Symposium on Grouts and Drilling Muds in
Engineering Practice
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(1963), Discussion, SGDMEP, Butterworths, London.
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Courteille, G. (1969), "The Stabilizing Action of Thixotropic
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and Foundation Engineering, pp. 189-236.
Davis, E.H., and Christian, J.T. (1971), "Bearing Capacity of
Anisotropic Cohesive Soil", ASCE, JSMFD, Vol. 97, No. SM5,
pp. 753-769.
DeNo, C.L. (1969), "Stability of Slopes with Curvature in Plane
View", 7th ICSMFE, Vol. II, pp. 635-638.
Dunlop, R., and Duncan, J.M. (1970), "Development of Failure
Around Excavated Slopes", ASCE, JSMFD, Vol. 96, No. SM2,
pp. 471-494.
Elson, W.K. (1968), "An Experimental Investigation of the Stability
of Slurry Trenches", Geotechnique, Vol. 18, pp. 37-49.
Flaate, K.S. (1966), "Stresses and Movements in Connection with
Braced Cuts in Sand and Clay", Ph.D. Thesis, University of Illinois,
Urbana, Illinois.
Golder, H.Q., Gould, J.P., Lambe, T.W., Tschebotarioff, and Wilson,
S.D., (1970), "Predicted Performance of a Braced Excavation", ASCE,
JSMFD, Vol. 96, No. SM3, pp. 801-816.
Gould, J.P. (1970), "Lateral Pressures on Rigid Permenant Structures",
Specialty Conference on Lateral Stresses and Earth Retaining
Structures, Cornell University, pp. 219-270.
Gould, J.P., and Dunnicliff, C.S., (1971), "Accuracy of Field
Deformation Measurements", 4th Panamerican Conference on Soil Mechanics
and Foundation Engineering, pp. 313-366.
Haliburton, A.T. (1968), "Numerical Analysis of Flexible Retaining
Structures", ASCE, JSMFD, Vol. 94, No. SM6, pp. 1233-1251.
Huder, J., (1969), "Deep Braced Excavation With High Ground Water
Level", 7th ICSMFE, Vol. II, pp. 443-448.
Instructions for Setting in Place the F-Series Strainmeter,
Telemac International Inc.
-71-
Jones, G.K., (1963),"Chemistry and Flow Properties of Bentonite
Grouts", SGDMEP, Butterworths, London, pp. 177-180.
Jones, J.C., (1967), "Deep Cutoffs in Pervious Alluvium Combining
Slurry Trenches and Grouting', Ninth International Congress on
Large Dams, Vol. 1, pp. 509-524.
Kallstenius, T., and Wallgren, A., (1956), "Pore Pressure
Measurements in Field Investigations", Swedish Geotechnical
Institute, Proc. No. 13.
Kuesel, T.R.,
(1970), Personal Communication.
Ladd, C.C., Bovee, R.B., Edgers, L., and Rixner, J.J., (1971),
"Consolidated - Undrained Plane Strain Shear Tests on Boston Blue
Clay", Research in Earth Physics, Phase Report No. 15, Department
of Civil Engineering, Research Report, R71-13.
Lambe, T.W., (1970), "Braced Excavations", Specialty Conference
on Lateral Stresses and Earth Retaining Structures, Cornell
University, pp. 149-218.
Lambe and Whitman, (1968), Soil Mechanics, John Wiley and Sons Inc.,
New York.
Lambe, T.W., Wolfskill, L.A., Wong, I.H., (1970), "Measured
Performance of Braced Excavation", ASCE, JSMFD, Vol. 96, No. SM3,
pp. 817-836.
LaRusso, R.S., (1963), "Wanapum Development Slurry Trench and Grouted
Cut-off", SGDMEP, Butterworths, London, pp. 196-201.
(1969-70), "Local Climatological Data", U.S. Department of
Commerce, Logan International Airport.
(1964), Manual of Steel Construction,2nd Edition, American
Institute of Steel Construction.
Marsland, A., and Loudon, A.C., (1963), "The Flow Properties and Yield
Gradients of Bentonite Grouts in Sands and Capillaries", SGDMGP,
Butterworths, London, pp. 15-21.
Massachusetts Institute of Technology, (1969), ICES-STRUDL-II,
Department of Civil Engineering, Report No. R68-91.
Mayer, A., (1967), "Underground Cast In-Situ Walls and their
Anchorage", Proceddings of the third Asian Regional Conference on
Soil Mechanics and Foundation Engineering, Haifa, Israel.
-72-
Meigh, A.C., (1963), Discussion, SGDMEP,
pp. 222-223.
Butterworths, London,
Mitchell, J.K,, (1960), "Fundamental Aspects of Thixotropy in
Soils", ASCE, JSMFD, Vol 86, Vol. 86, SM3.
Morgenstern, N.R., (1963), Discussion, Proc. Symposium on
Grouts and Drilling Muds in Engineering Practice, Butterworths
London, pp. 222-228.
Morgenstern, N., and Amir-Tahmusseb, A., (1965), "The Stability
of a Slurry Trench in Cohesionless Soils", Geotechnique, Vol. 15,
No. 4, pp. 387-395.
Morgenstern, N.R., and Eisenstien, Z., (1970), "Methods of
Estimating Lateral Load and Deformations", Specialty Conference
on Lateral Stresses and Earth Retaining Structures, Purdue
University, pp. 51-102.
Nash, J.K.T.L., and Jone, G.K., (1963), "The Support of Trenches
Using Fluid Mud", SGDMEP, Butterworths, London, England, pp. 177-180.
(1962-66), Norwegian Geotechnical Institute, "Measurements
of a Strutted Excavation", Technical Reports, Nos. 1-8.
Palmer, J.H.L., and Kenny, T.C., (1971), "Analytical Study of a
Braced Excavation in Clay", Preprint for 24th Canadien Geotechnical
Conference, Halifax, Nova Scotia.
Peck, R.B., (1969), "Deep Excavations and Tunneling in Soft Ground",
7th ICSMFE, State of the Art Volume, pp. 225-290.
Penman, A.D.M., (1961), "A Study of the Response Time of Various
Types of Piezometers", Proc. of the Conference on Pore Pressure and
Suction in Clay, Butterworths, London.
Piaskowski, A., and Kowalewski, Z., (1961), "Thixotropic Properties
of Suspensions of Soils with Different Grain Sizes and of Various
Mineralogical Types", 5th ICSMFE, Vol. 1, pp. 193-296.
Rabinowicz, E., (1970), "An Introduction to Experimentation",
Addison-Wesley, Publishing Co.
Sadlier, N.A., and Dominioni, G.G., (1963), "Underground Structural
Concrete Walls", SGDMEP, Butterworths, London.
Schneebeli, P.G., (1964), "La Stabilite Des Trenchees Profondes
Forees En Presence De Boue", LaHouille Blanche, No. 7, pp. 815-820.
-73-
Scott, J.D., and Kilgour, J., (1967), "Experience With Some
Vibrating Wire Instruments", Canadien Geotechnical Journal,
Vol. IV, No. 1.
Slope Indicator Instruction Manual, Slope Indicator Co.,
Seattle, Washington.
(1967), Soil Testing Report For South Cove Extension,
James P. Collins and Assoc., Cambridge, Mass.
Stacco, Z.A., (1968), "The South Cove Tunnel Project, Boston,
Massachusetts", Jour. Boston Society of Civil Engineers, Vol. 55,
No. 4, pp. 253-283.
Taylor, R.L., and Brown, C.B., "Darcy Flow Solutions With Free
Surface", ASCE, Journal of the Hydraulics Division, Vol. 93,
No. HY2, pp. 25-33.
Terzaghi, K.Z., (1943), "Theoretical Soil Mechanics", John Wiley
and Sons, Inc., New York.
Terzaghi and Peck, (1967), "Soil Mechanics in Engineering Practice"
John Wiley and Sons, Inc., New York.
Thon, J.G., and Harlan, R.C., (1971), "Slurry Wall Construction
for BART Civic Center Subway Station", ASCE, JSMFD, Vol. SM9,
pp. 1317-1334.
Tschebotarioff, G.P., (1951), "Soil Mechanics, Foundations,
Earth Structures", McGraw-Hill Book Co., Inc., New York.
and
Tschebotarioff, G.P., (1962), "Retaining Structures", Chapter 5,
Foundation Engineering, Edited by G.A. Leonards, McGraw-Hill,
New York.
Tschebotarioff, G.P., (1967), "Geieral Report on Earth Pressure,
Retaining Walls, Sheet-piling", Third Panamerican Conference on
Soil Mechanics and Foundation Engineering, pp. 301-310.
Turabi, D.A., and Balla, A., (1968), "Sheet-pile Analysis by
Distribution Theory", ASCE, JSMFD, Vol. 94, No. SM1, pp. 291-320.
Veder, C., (1961), "An Investigation on the Electrical Phenomena
at the Area of Contact Between Bentonite Mud and Cohesionless
Material", 5th ICSMFE, Vol. 3, pp. 146-149.
Veder, C., (1963), "Excavation of Trenches in the Presence of
Bentonite Suspensions for the Construction of Impermeable LoadBearing Diaphrams", SGDNEP, Butterworth, England, pp. 181-188.
-74-
Veder, C., (1969), "Testing Results of Bentonite Suspensions
in Trenches", Specialty Session 14, ICSMFE, pp. 21-22.
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Strutted Excavations", Technical Report No. 9, Norwegian
Geotechnical Institute.
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Thesis, Massachusetts Institute of Technology.
Zienkiewicz, O.C., and Cheung, Y.K., (1967), "The Finite
Element Method in Structural and Continuum Methods", McGrawHill Publishing Co., Ltd., London.
-75-
NOTATION
A
Area halfway to each adjacent spring in the vertical
direction in STRUDL analysis
Ap
Surface area of a hydraulic piezometer
As
Cross sectional area of a strut
B
Width of excavation
B
Half width of excavation
E
Cohesion entercept based on effective stresses
Cc
Compression index
cps
Cycles per second
Cs
Swell index
cv
Coefficient of consolidation
e
Error in a given measurement
eo
Initial void ration
E
Young's Modulus
EU0Yielded
Young's Modulus
EH
Young's Modulus from plane strain active test
EV
Young's Modulus from plane strain passive test
f
Shear stress ratio
FS
Factor of safety
ft
Feet
Gs
Specific gravity of solids
H
Depth of excavation
I
Moment of inertia
-76-
in.
Inches
K
Permeability
K
Lateral stress ratio
K
Sheeting stiffness ratio
Ka
Active stress ratio
Kd
Correction for guage drift
Kf
Principle stress ratio at failure
Kh
Permeability in the horizontal direction
Kn
Calibration constant for slope indicator torpedo
Ka
Lateral stress ratio at rest
K
Passive stress ratio
Kv
Permeability in the vertical direction
Kwall
Permeability of retaining wall
Kt
Temperature correction
L
Length of excavation
Ln
Length over which slope indicator torpedo measures tilt
Lv
Vertical spacing of bracing
Ls
Penetration of Bentonite Slurry
M
1 -
N
Ratio of height of Bentonite Slurry in a trench
trench depth
N
Standard penetration resistance
N ONVibrating
+
[depth to ground water table I Slurry trench depth]
wire frequency
N, Nc
Terzaghi - Peck dimensionless stability coefficient
Ncb
Bearing Capacity coefficient
-77-
OCR
Overconsolidation ratio
pp
9l + g3
2
73
,
2
p
Pounds
Pf
Force of bentonite slurry
PI
Plasticity Index
PSA
Plane strain active
psf
Pounds per square foot
PSP
Plane strain passive
PW
Force from water pressure
r
Ratio of excavation radius
R
Equivalent capillary radius for a porous media
S
Shear force
Su
Undrained shear strength
St
Sensitivity
To
Initial temperature of vibrating wire guage
Ti
Temperature of vibrating wire guage at time of measurement
u,U
Pore pressure
Unet
Net pore pressure
us
Static pore pressure
us
Static pore pressure
Uss
Pore pressure under steady state seepage conditions
V
Displaced volume of water in tube to cause a pressure
change of lpsf
wl
Liquid limit
wp
Plastic limit
-78-
* H
X
Calibration constant for vibrating wire guage
XpPComposite
calibration constant
Z
Depth below ground surface
[K]
Soil element stiffness matrix
[Q]
Nodal force vector
[S]
Bar element stiffness matrix
[U]
Nodal displacement vector
du
Rate of change of pore pressures
dt
E/ vo
Ratio of Young's Modulus to initial
stress
Su/covo
Ratio of undrained shear strength to initial
effective stress
vertical effective
vertical
Angle between the horizontal plane and the failure plane
in a wedge analysis
Yf
Bentonite slurry unit weight
yyt
Total unit weight
yw
Unit weight of water
Yb
Bouyant unit weight
A
Symbol indicating finite increment
AP
Differential pressure between bentonite slurry and ground
water
0
Angle of rotation
Poisson's Ratio
Tf
Bingham yield shear stress
a
Normal stress
a
Effective normal stress
-79-
C
conc
Consolidation stress
Fluid pressure of concrete
oyf
Bentonite slurry fluid pressure
aha
Active effective horizontal stress
hs h) x' x
Horizontal normal stress
Initial total horizontal stress
0
ho
avp avg azz' z
vm
Initial effective horizontal stress
Vertical normal stress
Maximum past vertical consolidation stress
a
Initial total vertical stress
a
Initial effective vertical stress
vo
vo
a 1 , 2' 3
Principal stresses
Friction angle
Friction angle based on effective stress
-80-
TABLE 1. 3. 1
Type of
Analysis
SemiEmpirical
Method
of Analysis
ANALYSIS OF BRACED EXCAVATIONS
RESULTS GIVEN BY VARIOUS METHODS OF ANALYSIS
Fore
Soil
Strut
Sheeting
Sheeting
Pressure
Stresses
Load
Movement
Stresses
TerzaghiPeck
x
x
x
x
SETTLEMENT
Initial
Cons.
Tscheboco
tarioff
Analytical
Finite
Element
1-D Consolid.
2
x
x
Beam on El.
Foundation
x
x
x
Continuous
Beam
x
x
x
x
x
x
FEDAR
x
3
x
BRACE II
1
Apparent horizontal total stress.
Variation during consolidation.
3 Steady state flow.
2
100
ON ,
1--1-
-1-----
.---.
- I
-
1.
1
x
PREDICTED MAXIMUM STRUT LOADS
TABLE 4.7.1
M
a x i mu m
S t r
u t
L
o a d s
K
i p s / f t
Elastic Beams
Strut
Level
Modified
Peck
(.4yH)
TscheboTarioff
with
Hinge
B
34.1
15.6
31.6
23.4
C
32.8
20.3
34.1
D
23.3
9.4
B
47.4
67.7
Design
Maximum
Measured
25.7
41.5
34.7
29.4
34.2
29.8
27.5
42.0
28.8
43.9
60.9
28.1
29.5
--
--
--
--
46.2
60.4
107
with
Springs
Brace
Ix
1C and D level removed.
81.60
103.8
132.2
90.3
.
a
SOIL PROPERTIES - SOUTH COVE PROFILE
TABLE 4.7.2
Soil Type
I
Depth
(ft.)
OCR
o
Unit
Weight
PCF
K
E
E
v
h
S
uv
S
uv
vo
v
vo
uh
Fill
Hard Clay
Very Stiff.
Clay
Stiff Clay
Medium Stiff
Clay
Medium Clay
0-5
5-15
5.0
.5
1.0
110
125
2504
1000
250
500
1.0
1.0
,3
,4
15-35
35-60
3.0
2.0
.87
.75
122
122
1400
340
220
200
0.77
0.57
0.69
0.65
.49
.49
60-80
80-120
1.5
1.0
.65
.52
120
120
300
250
180
150
0.46
0.34
0.63
0.60
.49
1Assumed values
for hard clay
Soil properties based on data from Ladd, et al.
TABLE 4.7.3
Sheeting Type
Concrete Wall
Arbed-Columeta
B2-350
Young's
Modulus
PSF
432,000
4,320,000
(1970)
SHEETING PROPERTIES
Moment of
Inertia
ft 4 /ft
2.25 - uncracked
0.98 - cracked
0.009
Area
ft 2 /ft
3.0
:.054
.49
30*
TABLE 5.1.1
Reference
Soil Type
SUMMARY OF MOVEMENTS ADJACENT BRACED CUTS
WITH CAST IN PLACE CONCRETE WALLS
Wall
Thickness
(in)
Excavation
Depth
(ft)
Maximum Sheeting
Movements (in)
Maximum Ground
Settlements (in)
Soft Bay Mud
Su = 300
1000 psf
58
65
1.5
1.2
Gould (1970)
Sands and Clayey
Silts ( z 330)
30
70
0.4
1.2
Thor & Harlan
(1971)
Sand (4 z 370)
and Clayey Silts
(0 z 350)
36
70
1.25
0.75
Huder (1969)
lacustrine
deposits (varied
silts, Su
350 pgf)
31.5
55
1.4
-
Kuesel (1970)
-I
(.
I
-DI
GGING CRANE
5- TREMIE-PIPES
2 -KE LLY - BAR
6-BENTONITE SWRRY
3 -CL MSHELL BUCKET
7- GUIDE TRENCH
4 -COI NCRETING CRANE
8- CONCRETE
7
-
4
00
A
4
A.
4
A
44
~
3.
8
b
44
.
6
4
fr
FIGURE 2.2.1
CAST-IN- PLACE CONCRETE
TRENCH PROCESS
WALLS
4.
4
b
4..
A.
..
BY THE
A
SLURRY
3.0
FRICTION ANGLE
SYMBOL
300
O
o
NH
35*
400
-.
Li
-
A
____
A
u-
\R
MH H
R
a
N
N.
N
N
w
N
La-
NA
0
El 'N
cl)
u-
AK
0
N
N.
N
Nm
N.
~"0 N.
N N.
N.
'N
0
-
2.0
N
- o
0
M =.50
u-
M =.75-
1.0
M= 1.0
I
R ~ADI CAcSE
N=-I
From Morgenstern and Amir-Tahmossib (1965)
'-'
I
I
1.20
1.40
1.80
L60
1
2.00
I
2.20
2.40
UNIT WEIGHT OF SOIL
UNIT WEIGHT OF SLURRY
FIGURE 2.2.2: STABILITY OF SLURRY TRENCHES IN
COHESIONLESS SOILS FOR PLANE STRAIN CONDITIONS
-86-
UNIT WEIGHT OF SOIL
110 PCF
SYMBOL
W
--
95
120 PCF
130 PCF
GROUND WATER DEPTH(D)
D
H
90
N
o
A
0
o
10g
5
-- =u-=O.321
Rvo
-
85
80
=~-~-7:jj.__
75
j
---4
--
60
0
0-
L
25
/
/
21"
i
=0.57, OCR=2
/
65
/
/
/
)
70
'-EFrom
Nash and Jones( 1963)
75
50
DEPTH OF EXCAVATION
100
125
(H)
FIGURE 2.2.3: STABILITY OF SLURRY TRENCHES IN
NORMALLY CONSOLIDATED CLAY
-87-
06
$
.-4
WAL
N
%09
DON
BOSCOE
04
STEEL
WALL
SCALE
0Of
co
FIGURE 3.1.1: SOUTH
(30.4
COVE
PROJECT
8m)
1
t
&
to
V
A
7
/
-,
DON
/
1L7
-F
5z
7'20-
7)
-5
V
//1
N
r10
C
H ARD
Y E LLOW
CLAY
Very Stiff to
H +-
to
Medium
H
-+
H
-+
H
STRUTS : 12' C /C (3.6 M)
50 FT
Bottom of Excavation
60
Stiff
I
-i
id
Blue Clay
with Sand
Lenses
H5
60.
127
Concrete
Wall
80 FT
Boston
STRUT
Blu e
B 14WF127 36WF170
C 14 WF 103 36WF 150
D 14WF 176 36WF
CLAY
100430
WALF
104
TKLL
ROCK
/
0
Ft LL
Stiff Boston
40-
I I
/
0
3'
70'
3'
/
LLI
/
0-0
BOSCO SCHOOL
I
I
FIGURE 3.1.2: SECTION
AT
STATION
-----------
113+40
U.
0
.0
2
20
SZ
-10
4.
in
40
%
F
w
2
60
0
I +e.
FILL ytal'op' t
HARD (L6Vcn
CLAY
IM C
39S3
STIFF
0
CLAY
0
30
10
x
t:4
0
c6000
-20
MEDIUM
8Oa
CLAY
0
(L l20tm
L93*m)
oo{30
44A+
u'j+ LY
S
1~
I
from Lodd,71j
(De
x
FIGURE 3.1.3: SOIL
4
10
50
10
00
H
8
rIv-
00
W.9thm)
C.
6
4
2
c
0
rtzI22psf
1
STRESS
L77
rlp
.3
.2
.1
0
*
as
Fi*Id Van*
Lob.
PROFIL E
Vans
12
60 t/nf
PH-70
LEGEND
OPH-8
*
PH
HYDRAULIC
*
P
VIB.
+SI
PIEZOMETER
WIRE
6 V
SLOPE
HEAVE
INDICATOR
ROD
o
LEVEL
POINTS
LP
ay
DON
11
BOSCO
SCHOOL
( 7 story
-brick)
0 P..7
PH-60
L Pa
PH-Se
PH-2
OLP
OPH-4
OPH-3
I6S
OPH-1
-
S I -I~IS
4
WALL
CONCRETE
STEE LSH EET
V-3 V-2
8
8+
+
V-0
P1
4
SI-7
0 02
SCALE
0
5
0 40 50
10
15
(F EET)
(METERS)
FIGU RE 3.1.4 :LOCATIONS
OF
FIELD
INSTRU MENT S
P-2
P-3
4
I
0.
0
SI- 6 (CONCRETE WALL)
120-
TUNNEL
100I-
WALL INSTALLED
WALL
ROFF
-35
*30
INVERT
SI-1l (SHEETING)
-20
+0.5-
PT
I
AT
FAR BUILDING
CORNER
-0
00
E
-0.5PT
1I-
7
L AT NEAR BUILDING
CORNER
-2
-1.0-13
-1.5--34
110-
-32
105-
z
2
-25
D 80-
PH-7
100-
30
95-
-28
2
90JI
-26
PH
85-
-24
80PT.I
SI-6 AT BOTTOM
PH-7*
OF
EXCAVATION -EL
-2
73
-0
0-
PT.
2-
7
--2
BOSCO
SCHOOL
DON
--4
L
SI-Il
AT BOTTOM OF EXCAVATION-EL. 73
--6E
PH-4
3-
-- 8
Lo
4-
-0
5-
0
'N'
D
J
'F
1968
FIGURE 3.2.1:TOTAL HEAD
M
A
'
M '
J ' A'
S
0
N
D
J
AND SOIL
F
J
MJ
1970
1969
MOVEMENTS
-92-
DURING
CONSTRUCTION
A
I
SECTION
-
I
10- - l1
LL
0
-
7,^N\77 -\Y17;N\ROOF
PORED
WALLS
-SIDE
INSTALLED
w
MUD-SLAB
80-
N
-__8O
+0.5-
+0. 5
0
w
-0
W -0.5-
Bldg.
>
0
80.-
pt. L,
NW
BId . Wt G
S 1.5- .
CORNER
--
_
100
-
-90
8 5
----
80-
0.5
-1.5
9H-D-----
85-
.-
CORNER
NE
t
100-6
-PH-8
Frozen
S
20
WALL
-
-
ePH-8
-
0
P-
5
,NWY
,NE
CORNER--8
CORNER
PH-7)
uDON
P
-
PH-4-
-. +
.
+ ..
-- 0.5
-
-0.5-
- P-5-
0 4-15- ~---
A
0
1N
1D
J
1968
FIGURE 2.2:TOTAL
F
O N
J 1A
J
1969
SOIL MOVEMENTS
HEAD AND
M
A
IM
-93-
a
Si -4
D
J
El ev. +t-73ft!
F
DURING
M
I
A
'1
1970
CONSTRUCTION
_
"
i
HORIZONTAL
-50
0
I
DISTANCE
100
50
Iq
150
-
-1
GROUND
120
WATER
TABLE
-1-
I0.
-
..........
0
o oo
--
CONCRETE
WALL
90
SY M.
o,A,E
*,A,m
o
-Is
DESCRIPTION
-800
0
F-
HEAD ALL PANELS INSTALLED
TOTAL
INITIAL TOTAL HEAD
NORTH PIEZOMETER LINE (STA. 113+20 to 113 +40)
CENTRAL PIEZOMETER GROUP (STA.112 +20 to 112-+80)
SOUTH PIEZOMETER GROUP (STA. 111-+-52 to 111-+-90)
NORTH
0J
SIDE
~
OF
0
-70
z
-.2 Z
CHOOL
.cx~.
-E
r
z
w
LUJ
0
-. 2
I.(n-
H
cr
FIGURE 3.2.3:
CHANGES IN TOTAL HEAD AND GROUND
WALL INSTALLATION
SURFACE
DURING
DISTANCE FROM PANEL FACE
i
~
-.
-.-.
w
i
SI-3
I
I
'I
101
RANGE
OF MOVEMENTS
2
HARD
CLAY
I
I
201-
f-
I
*
-INITIAL
I
I
POSITION
2
'
60-
I \
-MAXIMUM
DURING
EXCAVATION
PANEL
701
HOR.
.0J.5
.MVE
to
STIFF
I
I
I
I
I
..
50-
80-
.5
V. STIFF
40-
a
>
0
w-
30-
1-n
I 0I
FILL
-
SI- i
0
30
20
10
0
z
(FT)
NET MOVEMENT AFTER
PLACE TREMIE CONCRETE
CLAY
STIFF
to
MEDIUM
CLAY
90100-
1.0
CONCRETE
0
HORZ. MOVEMENT (IN)
SI-3
ENT(IN)
SI- I
FIGURE 3.2.4: MEASURED
0.5
GROUND
WALL
MOVEMENT
DURING
ROCK
CONSTRUCTION
OF
'1
SHEETING
2
-N
6
8
0
2
4
SI-6
6
8
SI-4
I
9-
0
(7/1/69)
--
(7/
10-
/
0
4
( IN.)
MOVEMENTS
SI-12
SI - 11
69)
(9/5/69)
(8/I /69)
-C
20--
-N
(7/22 /69)
-
8/6N
30-
I
I-D
---
-\
I
(7/17/69)
/(8/18/69)
(9/3/69)
-
N
/1/69)
(7/1/69)--
17/23/69)
(9/5/69)
I
-(9/16/69)
O/69)
(8/11/69)
L
I
40
/
(7/25/69)
+--
(9/22/69)
(0/29V 9)
-
0.
(8/I8/ 69)
8/18/69)
w 50-a
10/30/69)
1(9/12/69)
60
I--MAXIMUM
MAXIMUM MOVEMENT
(11 / 6 /69)
MOVEMENT
70-
11/26/69)
CaD
80--
STRUTS
C
REMOVED
STEEL
SHEET
FIGURE
8
D STRUTS REMOVED
PILING
3.31 : SHEETING
CONCRETE
MOVEMENTS
-4-MXIMUM
IMOVEMENi r
C 8 D STR JTS
REMOVED
MAXIMUM
MOVEMENT
C 8 D STRUTS
REMOVED
DURING
WALL
EXCAVATION
MOVEMENT
HOR IZONTAL
SI-9
120-
1
0
-0
2
FILL..
SI-10
I
0
*--B
(-B
(4/26,B
100-
-20
z
C
-- C
4/17)
p
0
(4/26/69)
(5/99)
(5/9/69)
VERY STIFF
I.-40
(5/9)
TO
-j
w
60
60-
(5/30)
(--D
(5/30/69)
w"
STIFF CLAY
-4
(5/12)
(5/2/69)
.
80-
with
fine
U(6/7)
MA
MAXIMUM
/
MOVEMENT
SAND
LENSES
I
40" -80
SI -10
HEET
PILING
1
FIGU RE 3.3.2 :
MAXIMUM
MOVEMENT
SHEETING
40
, SI-9
MOVEMENTS
AT
SOUTH
BULK HEAD (Sta. 107+50)
BOSCO
DON
SCHOOL
DISTANCE
0
.0
WALL
MOVEMENT
20
60
40
SI-I
FROM
WALL
80
100
(FT)
120
140
160
SI-3
0
II
',t
F:
I
a:
w
I
-20
.40
~0co
I
I
AT COMPLETION
I
w
I
0m
.60
N
---
MAXIMUM
OF SLURRY WALL
MOVEMENT
I
-80
1.0
0
-100
1.0
0
hI
1.0
0
HORIZONTAL MOVEMENTS(IN)
FIGURE 3.3.3.
GROUND
MOVEMENTS
ADJACENT CONCRETE
WALL
SCALE, FEET
o
.01
10 20 30
Settlement :in feet
02
.03
DON BOSCO SCHOOL
03
FIGURE 3.3.4: SET TLEMENT CONTOURS OF DON BOSCO SCHOOL
HORIZONTAL DISTANCE
EXCAVATION DEPTH
1.0
2.0
3.0
o
00
0
WALL
0A
COMPLETE (O)
__
0
0
0
EXCAVATION
_
COMPLETE (6)
SI
Il
I
A
-NAA~
U
A
--.
a
MAXIMUM SETTLEMENT (U)
0
z
w
4
U
4
w
U
w
0
0
w
t
am
0
)
1
-
O1X0
0
.02
-
0
DISTANCE FR )M EXCAVATION
DEPTH OF EXCAVATION
40
30
Q0-2
Dot 0
1.0. . I
ZONE I
( Soft
to Hard Clay
Average
V)
Workmanship)
2.0.
0
FIGURE 3.3.5: MOVEMENTS
.10
ADJACENT
From
THE
Peck
SLURRY
( 1969)
WALL
. 120X
MUD
100-
w
ROOF
SIDE WALLS
SLAB
N
INVERT
-J80-
I
I
NEW ZERO
.02
PO
0
w
0 -.02t--o
A
Fj
V-2-HEAVE
6 H P -HEAVE
T_
LO ST
ROD-EL.58.4FT i
PIN-INVERT
Fj-
2
DON
I
LA
N2
BOSCO SCHOOL
.PH-4
0,
V -2
0
100
P-I-H
INITIIAL
C
90
w
80
I
-J
70+
I.0
I-
READING
_-_III
'-1
~1
N
I-
LOST
-~-
60
O P
T
I - H -
M I A I M I J
NSOR E EV.=56.8
I
J ' A I S '0
N
J 'F
M' A
1969
FIGURE 3.3.6: TOTAL
HEAD
M 1 J'
J
A
1970
AND
BOTTOM
HEAVE
MEASUREMENTS
z
0
a
w a
0o
120A
10
PRESSURE
0
20
HEAD
60
40
I
ioo-L 2 0
80-4
-50
Stiff
Boston
___
+-- P
H-3
Blue
CLAY
PH -4\
-70
Medium
40- -80
Boston
Blue
-90
100
80
FILL
Hd. Yel.
-s-CLAY
-30
(FT)
Is
INITIAL READINGS (3/69)
A
EXCAVATION
0
FINAL
-1
0
+- P
5
\ 01 x
\
COMPLETE
(11/69)
READING (4/70)
4
47
0
0
P--H
-(ON
.
OF EXCAVATION)
\
CLAY
STATIC
Uss
K woll =Khsoil
TILL
Don Bosco High School
DDUo S
4HH
PH
n-
n 'Z 01 0
-
o nu
-
20j-h
00
rr.
P-I-H
FIGURE 3.3.7.
PORE
PRESSURES VERSUS DEPTH
-102-
DON
BOSCO
SCHOOL
.120
35
1968 (INITIAL
0NOV.
READINGS)
30-
w
-100X
0
OCT.
~
20
0C
-C
40
1969
(FULL
25- -80
LT.
DISTANCE
8.0
60
EXCAVATION)
100
I?o
30
1 0 Feet
140
50 Meters
40
10
-9-
-
z
.2
w
w
.6w
2
.8
DON
SECTION
-
(I)
BOSCO
SCHOOL
A-A
AI
1/1.
BOSCO
DON
SCHOOL
.120
-
35
1968 ( INITIAL READINGS)
NOV.
0
4
0
30
100 w
2:
-J
1969 (FULL
u0
CO
z
w
w
w
20
40
I
I
DISTANCE
80
60
I
10o
120 Feet
Meters
30
20
0
25- -80 4
I0
I-
EXCAVATION)
,
AiNr-
I.i...-"
.2.4-
.6-
C,,
.8- -2
SECTION
c
FIGURE 3.3.8:TOTAL
HEAD
AND
I-1
SETTLEMENT
-103-
AT FULL
EXCAVATION
B
STRUT
LEVEL
B
B
AVE. STRUT STANDARD
LOAD (K/FT ) DEVIATION
20.4
C
STRUT AVE. STRUT
LEVEL LOAD (K/FT)
B
C
2.6
STAGE I
2.9
3.9
STRUT
AVE. STRUT
STANDARD
LEVEL
LOAD (K/FT)
DEVIATION
B
STRUT AVE. STRUT STANDARD
LEVEL LOAD (K/FT) DEVIATION
D
Q
-
-
C
23.2
15.2
STAGE 2
-
B
STANDARD
DEVIATION
B
C
D
25.2
16.6
21.1
4.2
3.6
4.7
Invert
B
6.0
32.4
STAGE
4 ;
C
STRUT LEVELS
STAGE
3
INVERT
IN
STRUT AVE. DESIGN
LEVEL LOAD (KIPS)
B
C
D
41.5
29.8
60.9
FIGURE 3.3.9: STRUT LOAD VARIATION DURING CONSTRUCTION
AND
D
REMOVED;
PLACE
, APPARENT LATERAL STRESS
2
3
0
2
k / ft
*i
$ t /m
2
RANGE
L= 20.5
B
v
(19.3-34.7)
I
II
I
I
C
16.6
Kip
D11.4 -24
D 21.1
10L
Kips
(12.6-28.1)
*
I
//A'\7~
.2
-
0
L=13.5
-
I-ft
( STRUT LOAD
LV
-
--
i
.4(H
FIGURE 3.3.10: APPARENT
LATERAL
STRESS
STRESSES,
INITIAL
0-
0
FIL
.5
I
0
L5
EXCAVATE
STRESSES
2
(TSF)
0
4
SLURRY
TRENCH
4
2
PLACE
0
CONCRETE
2
4
6
c
Hd Yellow
I
.CLAY
P
Very
0
0*'
40.
w
0
Th
\ K\\ -\ Ii
p
to
Sti ff
Boston Blue
Clay
60
y=
n- --
vo
J
BOSTON
BLUE
LURRY
TRENCH
S-
--- 68pcf
--
7h 0
1'V\
5o
\I
--
-v
Y= 120pcf
100(1) INFERRED
FIGURE 4.1.1:
FROM OCR
INITIAL
STRESS
CONDITIONS
ON
---
\
CLAY
TILL
CONC
'I
122 pcf
MEDIUM
80-
j
V
Stiff
/
Ia.
=125 pcf
20-
AT SOUTH
COVE
1E
K
*
2~
Total Stresses
Minus Static Pore Pressures
Effect ve Stresses
B
Ttal
Stresses
2
20
CC
kips/ft 2
OA
Down<
0I
-I-+
1_aL_\__1
-
-_
P, P-Us , P
Element E
PROCESS
A
Initial
B
Slurry Wall
D
Up
(%)
CONDITION
C
-
VERTICAL STRAIN
Concrete Wall
Full Excavation
FIGURE 4.1.2: STRESS
PATH S
au
Place Slurry
Consolidate
negative
dissipates
Place Concrete
Consolidate
positive
dissipates
Excavation
negative
dissipates
Consolidate
V
+
q
A0
t
GUIDE
WALL\
Or
Su(TSF)
1.0
2.0
FILL
y=OOpd
i-
204-
ASSUMED TENSION
CRANK
Iw z
W
Hard Yellow
CLAY
Y=125 pcf
-LPW
z
7
F W
Very Stiff
40--
0Iz
60--
to Stiff
Boston
Blue
CLAY
Y=122 pcf
Yf 675 pcf
Ave.
S
Tests
Pf
onge - Field
Vane Results
Ij
rom Plane
Strain Active ( PSA)
Tests
-PSA
Medium
I ( LADD
Boston
80-
/
Field
Vane
1971
)
LL
Vane
BLUE
CLAY
S
Y = 120 pcf
Wp
100FOR
VANE - FS
FOR
PSA - FS
FIGURE 4.2.1: STABILITY
ANALYSIS
Pf
=
PPf
OF
216 =.83
2 0
216
10
2103
SOUTH
COVE
SLURRY
TRENCH
z
0
I
60
80
I
I
I
100
I
120
160
140
II
I
9
I
*
51 3
SI
I
-
I:
I
I
\1
-/
-
SYMBOL
--
"
20-
FACE (FT.)
- ------
DESCRIPTION
AXI-SYMMETRIC
ANALYSIS
PLANE STRAIN ANALYSIS
MAXIMUM MEASURED MOVEMENT
'
40-+
I. 05
-
PANEL
PANEL
40
I
FCL0
60+
I
80+M
Ioo-1
L.OQ5
i----i
0
ID
0.5
0
HORIZONTAL MOVEMENTS
FIGURE43.1: PREDICTED
-)
-
FAC E OF
20
&
FROM
DISTANCE
0
(IN)
MOVEMENTS
DURING CONCRETE
WALL INSTALLATION
-Q5~ L&J51
DISTANCE
300
200
250
L.A
2
(FT)
150
100
5,0
NITIAL GROUND WATER TAL
It
St iff
CLAY
with
0
Sand Lenses
Medium
CLAY
FIGURE 4.6.1: FEDAR
FINITE
ELEMENT
GRID
HEAD
PRESSURE
0
0-
0717
20-
LL
N.
22-
Liss (K al= 0.01 Kv)
0W60-
80
100
FILL
Hd. Yel.
_V CLAY
KH=Kv=1O
Stiff
B.B.C.
with
Sand
lenses
Kgf 10
4D-
60
40
20
(FT)
ss
=.
(K
y
STATIC
-
UNET
60-
Med.
Boston
Blue
80-
Clay
KH=KV= I
100-
sUNET=
LIS(KWalf 10 KV)
104-
TILL
KH= K;I
FIGURE 4.6.2: PREDICTED
DEPTH
-111-
WATER
PRESSURE
VERSUS
STRESS
0-
01
10
,
kips/ft
20
310
,MODIFIiD
40
m
PECK (.4YH)
.
10-
0
30-
40-
STrSCHEBOTARI0FF
50-
FIGURE
4.7.1:
PREDICTED HORIZONTAL STRESSES
-112-
IeSCHOOL
EARTH
PRESSURE
SURCHARGE
LOADS
L' 17
15'
211
-
-h
H
33
__
80 hl
ASSUME HINGE
44h 2
62.4 h2
SURCHARGE
2/3 H
LOAD SCHEDULE
q(PSF)
I 600
500
X 1 300
TYPE
NOTE: REINFORCING FOR WALL
DESIGNED AS CONTINOUS STRUCTURE
FIGURE 4.7.2:
DESIGN
LOAD
DIAGRAMS
FOR CONCRETE
STATION
111 +75 toI112 +85
112+85 to 113 15
113+ 15 to 113+65
WALL
STRESS , kips/ft 2
9
0-
I.0o
2,0
1
3p0
a.,
410
1
1
Ko=.5
K =.5
PERCHED WATER
2
TABLE
10-
..- .STAT
K 0 =41-I.3
K =.5
IC (U,)
20Z
0.
w
0 30-
c -+
____
STEADY STATE
SEEPAGE (Us)
I
___
FOR ANALYSIS
Ko =1.1-0.6
K =.4
40-
D -
p
\.
0
Is
50%
K.
FIGURE 4.7.3: PREDICTED
NII
\V
N
Inferred from OCR
HORIZONTAL STRESSES FOR ELASTIC BEAM ANALYSIS
ELASTIC
BEAM
WITH
[
ANALYSIS
ELASTIC
HINGES
BEAM
WITH
ANALYSIS
SPRINGS
0
.3 K/FT
18 K/FT
.4 K/FT
.6 K/FT
0.0 K/FT
21.7 K/FT
6 K/FT
.2 K/FT
20
U,
aw
.9
-40
27.2 K/FT
K/FT
16.8 K/FT
26,7 K/FT
a
N-EXCAVATION
LEVEL
-60
-80
1"-ASSUMED
FIGUR E 4.7.4: STRUT
LOADS
HINGE
PREDICTED
BY
ELASTIC
ANALYSIS
C0
z
L&J
DISTANCE
0
I:r
-
iJJ-
5,0
10
(FT.)
150
200
250
020.
ir_
-
40
I
ON
I
4.
I60-
a
80+ jc
loo
120 1
FIGURE
4.7.5 : FINITE
ELEMENT GRID FOR BRACE I
ANALYSIS
STRUT LOAD (KIPS / FT)
STRUT
LEVEL ELASTIC ELASTIC BRACE AVG.
HINGES
SPRING
MEAS
B
29.3
21.7
25.8
20.4
(KIPS/FT)
STRUT LOAD
STRUT
LEVEL ELASTIC ELASTIC BRACE AVG.
MEAS
SPRING
HINGES
B
18.8
16.6
20.6
234
C
31.6
27.2
34.2
15.2
STRUT
STRUT
LEVELELASTIC
HINGES
LOAD
ELASTIC
SPRING
AVG.
MEAS
25.2
16.8
14.4
16.6
26.7
43.9
21.1
20.4
17.2
C
10.0
D
38.9
-117-
BRACE
22.7
B
FIGURE 4.7.6: STRUT
(KIPS /FT)
LOADS
DURING
EXCAVATION
CONCRETE
WALL
STEEL
STAGE
///s\Y//X\Y///
SHEETING
I
\Y/A\Y//\I\\
STAGE 2
N ~
~AN Y/~ANQ'~~ANNY7~ AN\
77
~NNY77ANNY~t..(\\NY77/\N
NY 77
STAGE 3
,77777k
7 77A 7\7
7
7"\
20feet
SCALE -
p
FIGURE 4.7.7: YIELD
ZONES
-118-
FROM BRACE
ANALYSIS
TOTAL
2
0
HORIZONTAL
STRESSES
4
2
0
4
(KIPS/FT2)
6
0-r-
FILL
~L.
4.,
X
- 40.
0w
EXCAVATION
SYMBOL
DEPTH
---- _
0
----
33'
50'
K
-f+3
do'4
H d.
CLAY
VERY
STIFF
CLAY
STIFF
CLAY
0
S
60+
E
MATMED
E TIMAjTED
h-ho
C--
FOR ELASTIC ANAI LYSIS
FOR ELASTiC ANALYS,
MED.
N
STIFF
CLAY
80-L
SLURRY WALL
FIGURE 4.7.8 : TOTAL
FROM
HORIZONTAL STRESSES
BRACE ANALYSIS
STEE L -SHEETING r)
ON SOUTH COVE SHEETING
EXCAVATION
SHEETING
MOVEMENT
(IN)
II ELASTIC ANALYSIS TILELASTIC SPRING
I BRACE-Il
0
-
20
0.5
1.0
1.5
0
0.5
1.0
1.5
0
0.5
1.0
1.5
STAGES
D
r
+- C
7
4
wov
0-
-D
60-OR
SORING t---+
ANALYSIS Ml
80
ASSUMED END
CONDITION FOR
ANALYSIS 11
LEGEND
STAGE
STAGE
STAGE
STAGE
--
I
2
3
4
MAX. MEASURED
FIGURE 4.8.1: PREDICTED
.
.......
..........
SWRRY
WALL
MOVEMENTS
DISTANCE
0
20
40
60
FROM WALL
80
z
(FT)
100
120
140
160
[
MEASURED
0
.
-
.-
INITIAL SETTLEMENT
A
I
CONSOLIDATION + INITIAL SETTLEMENT (-
I
0. 4C
w
+--INITIAL
a
60
19
I
-)
POSITION
-
M EASUR ED
-
PREDICT ED
MOVEMENTS(----)
BY
BRACE (-
)
20
80
20
I
0
HORIZONTAL MOVEMENT (IN)
FIGURE 4.9.1
COMPARISON
MOVEMENTS
OF MEASURED AND PREDICTED
DURING
EXCAVATION
SUBSOIL
w
AV G.
MEASURED
C>
.6 (~3A~66
'A
B
t-,
0
I-
EMPIRICAL e
DESIGN 1
METHODS
A
.-
V
ANALYSIS
METHODS
I
L
>
DESIGN
MEASURED RANGE
PECK O.3YH
PECK O.4YH
TSCHEBOTARIOFF
ELASTIC BEAM-SP RINGS
ELASTIC BEAM-HI NGES
BRACE
C
Il
0
10
20
STRUT
FIGURE 4.10.1:
COMPARISON
EXCAVATION
30
LOAD
40
50
60
70
(KIPS/ FT)
OF MAXIMUM
STRUT
LOADS
DURING
40--
40
40 T DESIGN ENVELOPE
-
PEC K
EENVELOPE
0
Hi
30-
-
3H
30-
0
-J
0
H
.H
4.
Y
30--H
-
S
20-
HE
82H
IF~
ARIOF*F
.
-
ENVELO PE
44
E
62.4H
4
I
20-
*'
H
20
00
20
30
40
1.4H
10b
0
'o
30
20
PREDICTED
-o
LOADS
1o
00
0
0
50
40
KIPS/FT
METHODS
DESIGN
(.43
10-
-
-
1- 4
LL
40-
BRACE
30*
-
30A
00
I.
*
20
*
204
*0
10-
20
FIGURE 4.10.2:
30
40
0(D-
0
MAX IMUM
COMPARISON
OF
10
30
20
PREDICTED
0
0
1o
4b
20,o
LOAD KIPS/FT
ANALYSIS
METHODS
PREDICTED
AND
MEASURED
STRUT
LOADS
60
HORIZONTAL
0
0
1
2
a
3
i
4
i
5
A
6
a
3VEMENTS
0
7
(IN)
2
1
3
20
20
20---
p
H
40-
40
I
H
A
SLURRY
C
SI
II
0
SI
10
0
SLURRY
60
80
80
*
-
60
-
N)
LEGEND
I
LA-
PREDICTED
BRACE
MEASURED
FIGURE410.3: MAXIMUM
SHEETING
BY
MOVEMENTS
WALL
WALL
STEEL SHEETING
(KSF)
0
0
0.
2
4
6
a (KSF)
8
0'(KSF)
6
4
o7
0
8
10.
-
*.-I-20-
2
0
-
4
0
2
8
-i-
---
4
'6
S oft
X
-4CONC.
CLAY
a
40"
50+1
'---
S=120
-V
K =.5
0---
-- -
~c1v
-
a,-
U,
6
-t
--
A4
__'-
60-
INITIAL STRESSES
EXCAVATE
SLURRYPf4
WITH
68pcf)
PLACE TREMIE
CONCRETE
(6.7150 pcf)
FIGURE 5.2.1 : STRESS CONDITIONS FOR VARIOUS CONSTRUCTION STAGES OF
CONCRETE PANELS BY TH E SLURRY TRENCH PROCESS
50 FT
"A
EFFECTIVE
STRESS
TOTAL
STRESS
A-B
U-U
A-B
B-C
C- D
D- E
D-E
Excavate
trench with
slurry
Place concrete
Consolidation due to excess pore
pressures from concrete placement
Excavate within cofferdam
K
Li
K0=0.5
4
%
SE :TTLE
001
EE
E
..
Uo = initial pore pre ssure
HEAVE
E
\,A
SD
U0
IIII
',1
I,
Kf
h+V
2
2h+
C
FIGURE 5.2.2: TYPICAL STRESS PATH FOR POINT "A" DURING - INSTALLATION OF
CAST IN PLACE CONCRETE WALL IN CLAY. OCR = I
HORIZONTAL
.15
.0
.05
0
0-r
(2)
SHEN
.15
MOVEMENTS
.10
.05
(FT)
0
.15
3
.05
.10
0
3
LL
404
N4
5
-
S
-. 4
60+
t
I
80.
NON
YIELDING SHEETING
SHEETING EI=4x10 5
SHEETING EI=4x0
l
ELEMENTS
(EI
IN
P-FT 2
)
SHEETING EI=4xO 4
FIGURE 5.3.1: EFFECT OF STRUT SPACING AND WALL STIFFNESS
SHEETING MOVEMENTS FROM BRACE ANALYSIS
ON
SEI
EL=4xlo
SHEETING
SCALE
0
-
20
4
N
SHEETING
40 60 80 IFT)
1
EI=4 X10
El
IN
5
SHEETING
P-FT
EI=4xlO
4
2
FIGURE 5.3.2: YIELD ZONES VERSUS EXCAVATION DEPTH
AND SHEETING STIFFNESS
-12 8-
5
Curves
Prepared
From
Results of Finite Element
Compute Program- BRACE
-
0
4 [---
(FROM
0
DAPPOLONIA,1971)
0
.0
3
,H / 8 = 1.0-
H /B=0.25
z
LU
H/B =0.1
00
-1.0
-
0.5
0
+
0.5
+t-
1.0
SHEAR STRESS RATIO, f = 2
2
OVO
0
8
From
Plane Strain Tests
on Boston Blue Clay
LADD et al 1971)
:3
I
4
0
EXTENSION
2
w
COMPRESSION
0-
-0.5
+0.5
0
SHEAR
+1.0
+1.5
STRESS RATIO, f
FIGURE 5.3.3 : FACTOR OF SAFETY REQUIRED TO PREVENT
LOCAL YIELD BELOW BOTTOM OF EXCAVATION
IN CLAY
-12 9-
H
B
Hmax= SNcb
max.
y
I
9
-S
4
(B/L
=I)
/B/L = 0. 5
Z
8
Ncb
ur
__
)
St r ip ( B/L=O
5
0
(FROM
I
INERU
8 EIDE, 1956 )~
3
2
4
5
H /B
FIGURE 5.3.4: BEARING CAPACITY
BOTTOM STABILITY
-130-
FACTORS
ANALYSIS
FOR
EXCAVATION
El = 4 x 104
EXCAVATION
P-FT 2
El = 4 x 105
P-FT 2
EXCAVATION
STAGE
.4
4
H =36FR
.2
'2
0
10
20
10
20
Li-
5
._
.2
_2
H=45FT
z
w
.2
0
10
20
0
10
20
2
l10
7
1
20
0
i4
.4
0
H=65FT1
.A
I
10
20
-
)
.6
8
.4t
.4
-.2
.2
0
--.1
0
H =72.5FT.
I' 0
10
20
0
DISTANCE (FT)
(SEE
FIGURE 5.3.2 FOR
20
--.1
DISTANCE (FT)
EXCAVATION
STAGE
GEOMETRY)
FIGURE 5.3.5: PREDICTED
BOTTOM HEAVE BY BRACE
FOR STIFF AND FLE x IBLE SHEETING
-1l31--
SYMBOL] L(FT) f El x10 K-FT
Q
15
&
1.27
4
4
40
9
9
0.16
0.016
H
0.016
40
2.0
K
B=50FT
z
K =L
E
u
L =STRUT SPAING
-1
BEARING CAPACITY
NUMBER
N
a
1.0
0
U
N=
c
L-L
02
04
0.6
1.0
1.2
Su
f = 0.75=
I-K2
2
S
YVO
)
H / B
0.8
YH
(
z
FIGURE 5.3.6:
RATIO OF Nc / Ncb AT WHICH FIRST YIELD OCCURS WITHIN
EXCAVATION IN NORMALLY
CONSOLIDATED
CLAY
AN
0
HORIZONTAL
SI-1'
0 -r
tO2030-
0
2
MOVEMENT (IN)
3
4
0
1
SI-6
SI-9
S 1-10
2
0
I
2
0
FILL
4-
Hd.YeI.
CLAY
I
STIFF
BOSTON
BLUE
/
a
40
50..
-h
U
CLAY
60
70-80-
CLAY
-
MEDIUM
BOSTON
BLUE
- 25
S - 10
SOUTH
4 -S
BULKHEAD
I- 6
WALL
kSI-9
FIG U RE 5.3.7: HORIZONTAL MOVEMENTS OF SHEETING DURING
EXCAVATION
MOVEMENTS
S I-12
20I-
Fl LL
Hd. Yel.
sz CLAY
2
4
6
8
0
2
4
SI-6
E6
0
8
1
2
4-
7N
4--
STIFF
40+
-Is
N
SI-Il
%
0-r
( IN
)
SHEETING
BOSTON
BLUE
w
CLAY
604.
80-
MEDIUM
BOSTON
BLUE
CLAY
II
U
I.
Bo tom of
Exc va tiorr
I'
vi--
FIGUR E 5.4.1 : S HE E TIN G
[AovmenTS
0
10W
xcaytioiriLey01
M OVE ME NT S BELOW
EXCAVATION
LEVEL
PROFIL E
SECTION B
0-
WE T WALL
EAST WALL
MAX.
MISC. a
MOVEMENT
S-I
FILL
10+
PROFILE
EAST WALL SECTION A
WEST WALL
S-I
MOVEMENT BELO
EXCAVATION LEVE
.FILL
-435.*
102IO
I- 204
S-2
w
S-2
z
GANK
30+
RGANIC
SLT
a
S-3
S-3
404
re 5 TSF
=I04pcf
50-
60.
BOTTOM OF
EXCAVATION
384
L
134p
/
Ln
/o0
BEDROCK
TEST SECTION B
0
/ p
0
2
4
6
HORIZONTAL
TEST
6
4
2
0
MOVEMENT (IN)
SECTION B
FIGURE 5.4.2: SHEETING MOVEMENTS
IN BOSTON
BELOW
0
0
4
4
2
2
HORIZONTAL MOVEMENT(IN)
EXCAVATION
TEST
LEVEL
SECTION A
AT
NORTH STATION
APPENDIX A
SUMMARY OF FIELD MEASUREMENTS
AT THE SOUTH COVE TUNNEL EXTENSION
A. 1
INTRODUCTION
Under the sponsorship of the Massachusetts Bay Transportation
Authority, the Department of Civil Engineering, Soils Division, of
the Massachusetts Institute of Technology instrumented two sections of
the South Cove Tunnel Extension.
project was Dr. T. W. Lambe.
The principil researcher for the
The instrumentation was monitored from
September, 1968, through June, 1971, the construction time for the test
section, by the ICEP staff of the soils division under the direction
of Dr. L. A. Wolfskill.
Principal contractor for the project was Peter Kiewit and Sons', Co.
This appendix summarizes the results of the field measurements:
the
author's evaluation of the data and its accuracy is presented elsewhere
in this text.
A. 2
BACKGROUND
The location of the tunnel extension is shown in Figure 3.1.1.
It
is approximately 1500 ft in length and was constructed by the braced cut
and cover method.
The soil profile at the site consists of 7 ft granular
fill underlain by 100 ft of clay which varies in consistency from hard
at the top to medium at the bottom of the cut.
Two sections of the tunnel extension were instrumented as shown on
Figure 3.1.1.
The South Cove Station Area in front of the Don Bosco
School contained the major portion of the instrumentation.
-136-
In this section
U
a reinforced concrete wall, installed by the slurry trench method, was
used to minimize the building movements.
referred to as the slurry wall section.
This test section will be
The second section was the
south bulkhead area where only minimal instrumentation was installed.
The principle reason for the instrumentation in the slurry wall
section was to monitor the behavior of the Don Bosco School and obtain
data on the walls as a retaining structure in a bracing system.
The
purpose of the south bulkhead instrumentation was to give insight to
engineering problems which might occur when the tunnel is extended.
The information retrieved at the test sections were ground movements, both horizontal and vertical within and outside the cofferdam,
pore water pressures, and strut loads.
The instrumentation consisted of
settlement pins and screws, slope indicators, heave rods, hydraulic and
electric piezometers, and vibrating wire strain gauges.
A. 3
SOIL PROFILE
Borings for the tunnel and a proposed addition to the Don Bosco
School show the soil profile is very uniform along the instrumented areas
and is characteristic for this area.
It consists of a composite miscel-
laneous fill underlain by a hard yellow clay, then Boston blue clay and
glacial till overlying bedrock.
The only variations in the profile are
minor differences in strata thickness.
The fill consists of cinder ash, street sweepings and various types
of rubble fill.
It is essentially a granular soil.
from 2 to 8 ft, averaging 7 ft.
Elev. +114 ft.
Its thickness varies
It contains a perched water table to
No laboratory tests were performed on this material.
-137-
Based on visual observation of the material the assumed properties of the fill
are Y = 110 pcf
~ 30".
The yellow clay was of hard consistency, highly fissured and brittle.
The surfaces of the fissures were covered by a thin layer of iron deposits.
Torvane tests on a block sample from Elev.+104 ft gave average strengths
of 2 t/ft2
The consistency of the blue clay strata varies with depth as illustrated by the vane shear test results in Figure A.3.1.
The top of the
deposit is very stiff (Su = 1-1.3 t/ft 2 ) and fissured.
Its strength
decreases with depth to 0.5-0.9 t/ft 2 at Elevation +38 ft and slightly
increases to 0.7-1.2 t/ft 2 at the top of the till.
The till deposit overlying the bedrock consists of very dense fine
to coarse sand and gravel with a little
clay.
No tests were performed
on this material.
Figure A.3.1 summarizes the results of a series of laboratory tests
carried out on both 3-in, shelby tube samples and block samples.
Triaxial,
consolidation, permeability and atterberg limits tests were performed on
the block samples in the laboratories of the Soils Division at M.I.T.
Unconfined compression and unconsolidated undrained triaxial tests,
consolidation tests and atterberg limits were run on the 3-in, shelby
tube samples by J. P. Collins and Assoc. of Cambridge, Mass.
The results
summarized in Figure A.3.1 are those reported in their soil test report
to the M.B.T.A.
-138-
CONSTRUCTION
A. 4
A. 4.1
General
The close proximity of the South Cove Extension to the Don Bosco
School required special construction measures in that area.
To prevent
excessive settlement of the school, a reinforced concrete wall installed
by the slurry trench method was employed.
elsewhere for the excavation.
with three levels of struts.
Steel sheet piling was used
The entire excavation was cross-lot braced
At the south bulkhead area a deck was con-
structed to keep Shawmut Avenue in operation during construction.
Tunnel construction consisted of placing a mud slab; constructing
the tunnel inverts; removing C and D level struts; building the walls and
roof and then backfilling to finished grade.
This section summarizes the details of the various construction
operations and progress up to the construction of the tunnel walls.
Table
A.4.1 gives a. general summary of the construction history.
A. 4.2
Installation of Slurry Trench Wall
The slurry wall was installed by the Stang-Confor Company.
The
following procedure was used to install the concrete panels which make
up the wall.
1)
Construct 5 ft deep guide walls on both sides of the slurry wall.
2)
Excavate a 12 ft trench in between the guide walls.
3)
Prepare a batch of bentonite slurry.
The Specific Gravity of
the slurry ranged from 1.04 to 1.12 with an average value of 1.07.
4)
Excavate the panel.
The excavation procedure consisted of first
excavating 1 bucket width (6 ft) in the middle of the panel down to the
-139-
proposed tip elevation,
lowering a steel H section in the center of the
panel to guide the bucket as it excavates the side portions.
Then
excavate the side portions.
5)
Use the H section to trim the ends of the panel straight.
6)
Circular steel bulkheads are installed at the ends of the panel.
These serve as the end forms for the concrete.
For panels with installed
panels on both sides, no bulkheads are used.
7)
Set a preassembled steel reinforcing cage in panel.
8)
Place tremie concrete.
The concrete has a 5 - 9 in.slump and a
retarder to prevent setting of the concrete during placement.
9)
Remove the steel bulkheads following concrete placement by
jacking them up at about 1 in.per hour for 24 hours, then pulling them
out.
10)
At points where concrete wall joins the steel sheet piling, a
section of the piling is cast in the wall.
Figure A.4.1 shows the panel locations.
on the construction progress.
Table A.4.2 gives details
Table A.4.3 gives the as-built dimensions
of the panel top and compares the quantity of concrete the panel should
contain (based on panel dimensions: at the ground surface) with the quantity actually placed.
(In the early stage of construction panel 14, a
deep sea diver was employed by the contractor to obtain measurements of
the as-built panel.
panel.)
They agreed well with the surface dimensions of that
Table A.4.4 summarizes the measured specific gravities of the
bentonite slurry.
Figure A.4.2 shows the reinforcing details for the wall.
Some problems encountered in the installation of the panels were:
-140-
1.
Panels requiring excess amounts of concrete (see Table
A.4.3, panels 4, 5, 7, etc.) because of soil masses
spalling from the panel sides;
2.
Removal of steel bulkheads in the early stages of construction due to the adhesion of concrete and soil;
3.
Obstructions in the fill strata hampering excavation.
One problem which may have effected the measurements occurred in
panel 15.
The tremie concrete pour was stopped 15 ft below ground
surface.
To complete the concreting, a cofferdam had to be installed
around the panel and dewatered so the concrete surface could be cleaned
to create a good bond between the old and new concrete.
A. 4.3
Installation of Steel Sheet Piling
Arbed Columeta BZ-350 Belgium steel sheet piling was used at south
cove in the section south of the slurry wall.
tion is given in Table A.4.5.
The schedule of installa-
The properties of the piling are summarized
in Table A.4.7.
Installation of the sheeting was as follows:
1)
Excavate a 10-12 ft deep trench along the sheeting line.
This
allowed vertical aligning of the sheeting and driving relatively free of
obstructions.
2)
Set 10 to 20 sheets and drive them to a 40 ft depth.
3)
Start an adjacent group of piles.
4)
Sequentially drive the sheeting groups to the desired tip
elevation.
A. 4.4
Excavation Details
The excavation technique used is schematically shown in Figure A.4.3.
-141-
The numbers represent the order in which the different operations were
performed.
The excavation procedure was as follows:
1)
Excavate a 5 to 10 ft wide area adjacent the wall and install
the B level wale.
2)
Excavate a section to the bottom elevation of strut level B.
3)
Excavate an access trench in the center of the excavation
leaving 8
4)
ft wide berms against both bracing walls.
Install the B level strut and their vertical support piles as
needed.
5)
Excavate the berms and install the C level wale.
6)
Repeat items 3 and 4 for sturt level C.
7)
Excavate the entire section to a maximum of 5 ft below the D
level strut while excavating the center to the bottom of invert elevation.
8)
Install the D level wale and strut.
A summary of the excavation elevations and dates of strut installations are given in Table A.4.9 for areas immediately adjacent concentrations of instrumentation.
The purpose for the above excavation scheme was to permit excavation
by front end loaders and gradalls.
Access ramps were constructed at
various areas within the cofferdam to facilitate equipment traffic.
Figure A.4.4 shows typical contours of the excavation illustrating the
ramp areas and berms.
These particular contours represent the excavation
just prior to installation of struts B-42 and D-42.
-142-
A number of defects in the concrete retaining wall were observed
during excavation.
Some of the more significant ones are:
Large concrete protrusions were found on the face of the retain-
1)
They resulted from over excavation or spalling of the panel
ing wall.
walls during the wall installation and account for a large part of the
excess concrete required in the panels.
2)
Exposed reinforcement bars were found in the wall.
3)
Some of the vertical panel joints were observed seeping water.
At the wall face the space between these panesl was up 1 in.
4)
The steel sheeting had ripped open during driving leaving gaps
in the sheeting of as much as 3 to 4 ft in width.
Two of these openings
were found in the vicinity of SI-11 and SI-12.
A. 4.5
Wale Installation
Figure A.4.5 gives an overall view of the excavation and shows the
bracing details.
The wales were attached to the retaining walls by a
hangar assembly, which was also its lateral buckling support, and steel
plate clips.
In the slurry wall area the hangar and clips were attached
to the concrete wall by Williams No. US-8 rock bolts.
Voids between the retaining wall and the wale were shimned to reduce
movements.
In the sections with steel sheet piling, wood wedges were
In the slurry wall area, gaps were filled with cement grout.
used.
A. 4.6
Strut Installation
The bracing scheme used in the two instrumented sections are shown
in Figures A.4.6 and A.4.7.
The strut support on the wale consisted of
-143-
a section of 18[58 welded to the wale with its longitudinal axis vertical.
To the bottom of the channel a 7 in.by 19 in.x 3/4 in. steel plate was
welded to support the strut.
with steel shims.
The struts were wedged against the wale
No additional effort was made to secure the strut to
the wale.
To help minimize the retaining wall movements, the struts were preloaded when installed.
The preload was applied by two hydraulic jacks,
each with a 402 kip capacity.
The jacks were centered on the strut, one
each side of the web, and were located between the wale and a section of
36WF150 welded to the strut web.
jacked into the strut.
Eighty percent of the design load was
During the jacking operation, the strain gauges
were monitored in an attempt to calibrate the gauge and strut assembly
as a load cell.
Upon attaining 80 percent of the design load, steel
shims were wedged between the strut and the wale.
The jacks were then
relaxed and the strain gauges read to determine the residual load in the
strut.
The desired residual load was 50 percent of the design load.
To prevent buckling of the struts because of the long space across
the tunnel vertical strut supports (8BP36) were installed along the
centerline of the tunnel.
A.
4.7
Tunnel Construction
When the excavation bottom reached the invert base depth, a concrete
mud slab was poured to minimize soil disturbance.
Soon thereafter, the
tunnel invert, which was 4 ft thick, was constructed between the excavation walls.
Figure A.4.4 illustrates their relative progress during con-
struction.
-144-
After the invert concrete cured, so it could react as the lower
strut against the sheeting, the C and D level struts were removed.
The tunnel walls were then constructed using the slurry wall as outside
forms in its area and corrugated sheeting against the steel piling in
its section.
When the roof was complete the ground surface was brought
to the initial ground level.
(Details of the tunnel design are given
by Stacko, 1968.)
A.
INSTRUMENTATION
5
A. 5.1
General
The instrumentation at the South Cove extension was concentrated
in two areas, the slurry wall area (Station 111+00 to 113+80) and the
south bulkhead (Station 107+50).
The instrumentation was aimed at measuring:
a) ground movements
both vertical and horizontal within and outside the excavation; b) strut
loads; c) changes in soil and water pressure.
In the text, references to particular instrumentation and excavation
areas will be made using either the instrumentation numbers, the tunnel
centerline stationing, or both.
The strain gauges on the struts are
identified according to the struts number.
Three levels of struts were
installed in vertical sections which are sequentially numbered starting
from the south bulkhead.
The strut levels are designated as B, C, and D,
The designated number for the bottom
starting with the topmost level.
level strut in section 32 is D-32.
A summary of the instrumentation installed is given in Table A.5.1.
-145-
Their locations are shown on Figures A.4.6, A.4.7 and A.5.1.
A profile
of the instrumentation is given in Figure A.5.2.
A. 5.2
Strut Loads
Loads were measured on the struts designated by the circled section
number on Figure A.4.6.
Telemac type F-2 vibrating wire strain gauges
were mounted on all struts.
A.5.3.
Details of the gauges are shown in Figure
These gauges measure strain by electrically recording the change
in the wires vibration frequency which is a function of the strain in
the wire, which in trun is related to the struts deformation.
The theory
of the gauge operation is discussed in more detail in Bjerrum et al (1965).
On all struts, two gauges were mounted on the centerline of the web.
In
addition, on struts B-44 to B-47, two additional sets of gauges were
mounted, one 4 in.above the web centerline, the other 4 in.below.
The
purpose of these gauges was to study the adequacy of two centerline gauges
in recording'the actual strut load.
All centerline gauges were mounted
8 ft from the west excavation wall.
During the early stages of the work the gauges were found to be
temperature sensitive.
When the sun's rays were directly on the gauges
a temperature differential existed between the internal part of the gauge
and the strut.
To minimize this effect the gauges were shaded by insulat-
ing paper wrapped around the strut.
The initial gauges on struts B-31 through B-40 were incorrectly
installed.
Telemac type SB-90 strain gauges were installed on these
struts when the error was discovered.
The initial SB-90 reading was
equated to that recorded by the F-2 gauges, which, in turn, were determined
-146-
Im
from a calibration curve derived from the preloading data.
In the
remaining struts the initial "no load" frequencies were taken just before
the strut was preloaded.
The data pertinent to the strut installation
is summarized in Table A.5.2.
A. 5.3
Pore Water Pressures
Pore water pressures were monitored by both Casagrande type hydraulic
piezometers (PH) and vibrating wire piezometers (P).
sensor elevations are given in table A.5.3 and A.5.4.
piezometers used were the Geonor Model M-206.
The piezometer
The vibrating wire
Their principles of opera-
tion are the same as those of the strain gauges, the wire in this case
being strained by a diaphram deflected by the water pressure.
ing frequency value is calibrated against water pressure.
The result-
The hydraulic
piezometer sensor was an 18 in.long by 1.35 in.diameter porous plastic
tube filled with pea gravel.
of entrapped air bubbles.
tubing.
The sensor had two leads to permit flushing
The leads were 3/8 in.and 1/4 in.Polyethelene
The water level was determined by lowering an electric probe
down the 3/8 in.tube.
The installation technique for the piezometers is
shown in Figure A.5.4.
Shallow piezometers (W) were installed in the fill and just below
the hard yellow clay to determine if
the hard clay.
a perched water table existed above
The two piezometers in the fill
were 36 in.long by 1 1/4
in.diameter steel well points, two being double lead hydraulic piezometers.
Details of the installation are given in Figure A.5.5
A.
5.4
Ground Movements
Three types of ground movements were monitored at the slurry wall
-147-
section.
They were horizontal movements behind the excavation, vertical
movements both within and adjacent the excavation.
A. 5.4.1
Ground Settlements
The vertical movements outside the excavation were measured by
settlement screws on the Don Bosco School and the front of the slurry
wall and by level pins (LP) adjacent the slurry wall.
Figure A.5.6
shows details of the settlement screws and their method of installation.
A summary of the level pins is given in Table A.5.6.
Readings on the settlement devices were taken with a Zeiss level.
The recorded movements are based on a permanent bench mark installed
at the site(see Figure A.5.7).
The movements below the base of the excavation were monitored
during excavation and after placing of the tunnel invert.
While exca-
vating, the bottom movements were measured by three Borros anchor points.
Details of their installation are shown in Figure A.5.8.
Unfortunately,
these instruments were lost in the early stages of construction.
When the tunnel invert was completed, heave pins, such as those on
the Don Bosco School, were installed along its
centerline to observe
invert movements for the remainder of the construction period.
A. 5.4.2
Horizontal Movements
Horizontal movements were recorded both in the soil behind the
excavation wall and of the slurry wall itself.
made using a Wilson Slope Indicator.
The measurements were
A summary of the indicators installed
is, given in Table A.5.5.
Details of the casing installation in the ground, in slurry wall, and
-148-
on the sheet piling, are given in Figures A.5.9 and A.5.10.
The data
was recorded periodically and at times of significant construction operations in the vicinity of the slope indicator.
The slope indicator measures relative movement from its initial
starting point.
Past experience has indicated that the entire length of
casing can move laterally.
To establish a base point from which to
reference movements, an optical survey was Conducted to establish the
position of the top of the casing relative to a base line which was
parallel to the tunnel centerline.
Movements relative to the top of
the casing were determined by measuring the angle of tilt of the slope
indicator casing with a Wilson "torpedo".
The torpedo tilt angle is
calibrated against readings from a resistance pendulum housed within it.
The lateral movement at any elevation in the casing is equal to the sum
of the tilt angles (0) multiplied by the summation of a set of equallyspaced readings.
To ascertain results were reasonably correct,the inclonometer was
twice traversed down the casing at 180 degree intervals.
The sum of the
angular distortions from the two runs should equal a constant.
Readings
were taken in this manner both parallel and perpendicular to the excavation centerline.
A. 5.5
Soil Pressures
Two sets of total stress cells were installed to measure horizontal
ground pressures.
They were positioned at Station 113+40, with one set
at the tunnel centerline,
the other 5
elevations are given in Table A.5.8.
-149-
ft behind the slurry wall.
Their
The centerline stress cell was lost
during early stages of excavation in that area.
Details of the cell assembly and installation are given in Figure
A.5.1l.
The assembly has two Geanor vibrating wire total stress cells
facing in opposite directions.
in Bjerrum et al. (1965).
Their method of operation are described
The Geonor cells are housed in a steel drive
shoe, the dimensions of which are 2 in.by 5 in.by 28 in.
When installing
the cell, the base of the 6 in.steel outer casing is stopped 5 ft above
The cell is hydraulically pushed this re-
the desired sensor location.
maining 5
ft to the required elevation.
RESULTS OF.NEASUREMENTS
A. 6
A. 6.1
General
The results of the field measurements are presented in Figures
A.6.1 to A.6.26 inclusive.
In cases where the instrumentation gave
essentially the same values for a section only a representative data set
is given.
The data shown is as recorded, uncorrected for environmental
effects, etc.
Supplemental information which may have a significant
bearing on the evaluation and interpretation of the data is given in
Appendix B.
To permit correlating the influence of construction events
on the particular data presented, the figures contain a brief summary
or plot of construction progress in the vicinity of the instrument group.
A. 6.2
Strut Loads
The variation of strut loads versus time is given in Figures A.6.1
to A.6.8.
The figures presented are representative of vertical sections
of struts which have the same design load and construction history.
-150-
For
----v
example, Figure A.6.1 is representative of measurements made on struts
B-32 to B-36; these struts have the same design load and also had the
F-2 gauges incorrectly mounted.
Table A.6.1 summarizes the measured
strut loads for each excavation and bracing stage.
The plots give the values recorded from both centerline gauges as
well as their average value.
The load as measured by the gauges on a
strut can differ significantly.
For example, the SB-90 gauges on beam
B-38 and the F-2 gauges on C-38 show a load difference of 125 kips and
200 kips, respectively.
Observation of these struts indicate this
difference is often associated with large curvature or distortion of
the beam in the horizontal plane.
The gauges mount approximately 1 1/4
in from the surface of the strut web and therefore
are quite sensitive to this curvature.
their recorded strains
However, it is felt that the
lines for the average load give representation values of strut load.
Data describing the effect of temperature, the reliability of using
only two gauges, and the accuracy of the instrument itself are given and
discussed in Appendix B.
A.
6.3
Pore Water Pressure
Figures A.6.9 to A.6.14 give the variation of pore water pressure
during construction for all the piezometers.
Each figure presents a
plot of construction progress and settlement for settlement instruments
in the vicinity of a piezometer group.
Piezometer P-1 was lost before excavation in that area got underway.
Two piezometers, one hydraulic and one vibrating wire, were installed to
replace P-l.
Their data give an indication of the variation in readings
-151-
G
one can expect from the two types of piezometers.
Piezometers P-2 and
P-3 were lost when the excavation elevation reached El. 85 ft.
The shallow well observations are shown on Figures A.6.13 to A.6.14.
In addi-
They show a perched water table did exist above the hard clay.
tion, these figures present observations of water level in the slope
indicator casings, except SI-4 to SI-7,
and the open well, OW-10.
The near constant value of water pressure in the piezometers, regardless of sensor elevation, indicate the initial ground water is essentially
static (see Figure A.6.9a, P-4 to P-6).
In general, the hydraulic
piezometers show a higher water table than the vibrating wire piezometers,
in particular, PH-1.
Of significance is that those piezometers with
sensors at the same elevation record the same trends in pore pressure
fluctuation.
A. 6.4
Earth Pressure Cells
The total horizontal earth pressures measured by the total stress
cells are shown in Figures A.6.15a to A.6.15d.
In addition, the plots
contain the pore water pressures measured by piezometers P-1 and P-5,
which are at the same elevation as the cells.
The data from stress cell SC-2U is not plotted since it oscillated
over a large range.
during insta-lation.
This oscillation was probably the result of damage
The values of stress from SC-lU and SC-lL were
significantly different immediately after installation (3.75 versus
8 k/ft 2 ).
There is no obvious explanation for this deviation other than
the possibility SC-lL is in a hard zone such as a sand seam.
However,
the stress changes during construction are the same for both stress cells.
-152-
Horizontal Ground Movements
A. 6.5
Horizontal ground movements were measured behind the slurry wall,
on the wall itself, and at selected stations on the sheet piling.
The
results of these measurements are shown in Figures A.6.16 to A.6.22 for
selected construction times.
The data given is the East-West data
which corresponds to movements perpendicular to the tunnel centerline.
Observations were made on the movements of the top of each casing relative to a survey base line paralleled to the tunnel centerline.
The
recorded horizontal displacements with depth are based on the position
of casing top as established from the base line.
Thus, any movements
of the casing bottom during construction do not effect the reported
results.
Both the "North-South" and East-West" movements are given for
They are given because the grooves of the casing are parallel
SI-3.
and perpendicular to the Don Bosco School, not the tunnel centerline.
To establish the movement perpendicular to the centerline the components
of the above movements must be utilized.
When the movements between construction stages were small (less
than 1/2 in), the measured data often varied within a 1/4 to 1/2 in.
band due to the accuracy of the instrumentation.
An example of this is
SI-3 on dates 1/28/69 to 3/11/69.
A. 6.6
Vertical Ground Movements
Settlement measurements taken on the Don Bosco School, the adjacent
ground, and the slurry wall are summarized on Figures A.6.23 to A.6.31.
All the lines of settlement devices A to I on the School gave essentially
-153-
the same results.
are reported.
Therefore, only the data for lines A, C, E, and I
The slight variations in settlement for a given settle-
ment screw are a result of the survey accuracy.
The results from the heave pins on the tunnel invert are given in
Figure A.6.32.
-154-
TABLE A.4.1:SUMMARY OF CONSTRUCTION EVENTS RELATED
TO FIELD MEASUREMENT PROGRAM
Inclusive Dates
Construction Event
68
Field instrumentation installed
68
Guide walls installed for slurry trench
30 Oct
8 Oct
68 - 12 Nov
16 Nov
68 -
23 Jan
69 -
7 June 69
19 Mar
69 -
7 June 69
9 Sept 68 -
construction
24 Mar
69(l)
Concrete panels installed by slurry trench
technique
6 June 69 - 15 Aug
30
May
16 May
I-A
I-,
19 May
69
-
69 -
69
-
12
Sept
2 Oct
30
Oct
69(l)
69(2)
69(2)
69(2)
-6 Sept 69
2 Oct
69
Steel sheet piling installed between south
bulkhead and slurry wall test section
South Bulkhead test section excavated and
braced
30 Oct
70
2 Nov
70
Removed C and D level struts Station
111+80 to 112+20
5 Nov
69
Placed invert of tunnel Station 112+40 to
112+60 (SI-5 and SI-7)
21 Nov
69
28 Nov
69 -
1 ,3
27 Dec
Dec
Placed walls of tunnel at Station 111+00
69
69
26-29 Jan 70
3 Dec
69
1 Sept 70
26-29 Sept 70
Excavated and braced in front of SI-5 and
69
69
70
70
69
19 Nov
8 Aug
8 Oct
25 Oct
69
24 July 70
25-30 Sept 70
SI-7 (Station 112+40 to 112+60)
Excavated and braced in front of SI-4 and
SI-1 (Station 113+40 to 113+60)
Placed invert of tunnel at Station 111+00
to 111+50 (SI-U and SI-12)
Placed invert of tunnel at Station 111+80
to 112+00 (SI-3 and SI-6)
Removed D level struts 111+00 to 111+60
to 111+50
Removed C and D level struts Station 112+40
to 112+60
Placed invert of tunnel at Station 113+40
to 113+60 (51-1 and SI-4)
Placed walls of tunnel at Station 111+80
to 112+00
Removed C and D level struts Station 113+40
to 113+60
Placed wall of tunnel at Station 112+50
Placed walls of
113+60
tunnel of Station 113+40 to
RELATED
(Continued)
Inclusive Dates
Excavated and braced in front of SI-lland
SI-12 (Station 111+00 to 111+50)
Excavated and braced in front of SI-3 and
SI-6 (Station 111+80 to 112+00)
3-4 Nov
6 Nov
TABLEA.4.1:SUMMARY TO CONSTRUCTION EVENTS
TO FIELD MEASUREMENT PROGRAM
Construction Event
Placed roof of tunnel at Station 111+00
to 111+50
Placed roof of tunnel at Station 111+80
to 112+00
Placed roof of tunnel at Station 112+50
Placed roof of tunnel at Station 113+40
to 113+60
Removed B level struts Station 110+50
to 112+20
Removed B level struts Station 113+50
to 114+00
Removed B level struts Station 112+25
to 112+60
Removed B level struts Station 112+60
to 113+50
(1) Table
(2) Table
sequences.
gives detailed information on panel installation
gives detailed information on excavation and struting
TABLE
A.4.2:
CONSTRUCTION PROGRESS OV CONCRETE
SLURRY TRENCH PANELS
Concrete
Placed
Time to Place
Concrete
1
2
3
4
5
6
7
8
9
10
11
12
13
14
12/18/68
1/ 4/69
1/ 9/69
1/ 9/69
1/30/69
2/13/69
2/21/69
2/19/69
1/21/69
3/ 4/69
3/ 1/69
2/ 5/69
1/18/69
11/30/68
1/ 3/69
1/ 8/69
1/14/69
1/20/69
2/ 3/69
2/20/69
2/27/69
2/22/69
1/23/69
3/ 8/69
3/ 6/69
2/13/69
1/21/69
12/18/69
3.5 hrs.
4.5 hrs.
5.25 hrs.
15
16
11/26/68
2/26/69
1/ 3/69
12/31/68
1/ 8/69
1/11/69
1/18/69
2/ 1/69
2/19/69
2/26/69
2/22/69
1/23/69
3/ 6/69
3/ 5/69
2/12/69
1/20/69
12/ 3/68
12/ 6/68
11/30/68
2/27/69
1/ 4/69
2/ 3/69
3/ 8/69
2/15/69
3/17/69
3/14/69
1/28/69
1/29/69
3/24/69
3/ 1/69
3/11/69
2/18/69
3/17/69
3/14/69
1/29/69
1/30/69
3/24/69
3/ 4/69
1803)
-
17
1/20/69
-
Excavation
Complete
-
Panel Excavation
Started
Panel No.
11/30/68
2/28/69
1/ 4/69
2/ 4/69
4.5 hrs.
5.75 hrs.
4.25 hrs.
6
hrs.
4.85 hrs.
4.25 hrs.
6
hrs.
4
hrs.
3.5
4.5
hrs.
hrs.
2/ 1/69
19
20(1, 2)
21(1)
22
23(1)
24(1)
25
26
2/ 4/69
3/10/69
3/12/69
1/23/69
1/27/69
3/15/69
2/26/69
5.25 hrs.
5
3.3
hrs.
hrs.
(1 )Encountered obstructions during excavation.
(2)Problems with bulkhead removal.
(3 )Started excavating north end 1/20/69 to free steel bulkhead on panel 15: 2/1/69
began excavation of south end.
(4 )Delay of pour. Used deep sea diver to make measurements to check dimensions of
panel excavation.
-156-
TABLE A.4.3:
AS BUILT DETAILS OF CONCRETE PANELS
Tm
7
Panel Dimensions at
F
Ground Surface (ft.)
I
Volume of Concrete
41
Computed
L
yd
11.9
14.9
84
126
11.85
11.15
11.6
22.25
18.9
14.85
14.15
14.6
83
85
82
83
80
212
202
123
117
117
12.05
13.5
82.5
118
11.43
4.1
14.4
7.1
21.0
19.5
13.0
19.1
82
84
86
82
84
84
84
84
83.5
84
82
119
48.
214
211
142
213
20.4
16.0
22.1
13.15
16.2
12.8
14.32
15.8
11.32
.
I
TABLIA.4.4: SI'ECIFI&
yd
OF BENTONITE SLURRY
Speci fic
Gravity
3
(g/cm
20
40
60
73
1.04
1.09
1.108
1.108
18
20
1.097
14
25
30
1.05
1.10
3
20
32
1.07
1.087
6
20
1.125
5
10
1.08
18
20
1.097
1
s
-157-
c
11.5
3.3
4.5
35.8
15.4
4.1
11.9
23.6
20.0
0
22.8
14.8
14.6
11.7
-12.7
42.5
40.0
I ______
Ii
GRAV.TY
M
125
142
113
218
219
211
167
135
122
234
234
192
132
71.5
180
147
60
214
259
163
244
259
172
130
212
164
154
149
149
117
Depth
(ft)
Panel
No.
c
I
Percent di".
4
44
4
-1
T
I
1
2
3
4
5
6
7
8
9
10
11
12
13
14
15
16
17
18
19
20
21
22
23
24
25
26
Volume 3 (V")
Volume 3 (V
Depth
)
Panel
No.
Measured
_ __
TABLE A.4.5:SCHEDULE OF STEEL SHEET PILING INSTALLATION
t-n
00
I
Station
14-18 Jan. 69
19-22 Jan.
28 Jan. 69
29 Jan. -21 Feb.
21 Feb. 69
21 Feb. 69
4 Mar. 69
4-14 Mar. 69
20 Mar 69
107+94 - 108+34
108+34 - 108+99
107+68 - 107+84
108+20 - 108+75
107+50 - 107+68
South Bulkhead Wall
South Bulkhead Comp
107+50 - 108+20
5-12 May 69
14-16 May 69
16-21 May 69
21 May 69
3-30 June 69
7 June 69
100+00
109+00
110+00
109+30
111+00
111+30
- 111+00
- 110+00
- 111+00
Excavation Wall Driven
West
East
Remarks
x
x
x
x
x
SI-9 installed
SI-8 installed
.
Date
x
SI-10 installed
Deck for Shawmut Ave.
was complete
X
x
x
x
- 111+30
- 111+50
Install SI-12
Install SI-ll; sheeting
installation complete
- - -
II
TABLE A.4.6:
SUMMARY OF BRACING DETAILS
FOR SOUTH BULKHEAD
Strut
Level
Strut
Size
Corner Strut
Wall
Size
Size
B
C
D
12WF45
14WF103
14WFI19
8WF35
12WF58
12WF72
36WF150
36WF 300
36WF 300
These are corner strut and first
struts at station 107+62.
TABLE
cross lot
A.4.7: PROPERTIES OF STEEL SHEET PILING
Manufacturer
Arbed-Columeta
Size
BZ 350
Moment of inertia
180 in /ft
Section modulus
31.1 in 3/ft
Area
7.86 in 2 /ft
Width (L)
19.68 in
Height (h)
11.62 in
(Belgium Steel)
Thickness
.375 in
(t)
43.88 lbs/ft 2
,
Weight
I
h
L
TABLE A.4.8:
WALE SIZES FOR
SLURRY WALL AREA
Wale Size
Strut Level
Centerline Station
T
B
111+00
111+80
-
-
111+80
113+85
24WF176
36WF170
-159-
C
36WF150
36WF150
D
36WF170
36WF300
TABLE A.49 :
-
Station 11112
Slo-pe Indicators S111
111+50
-
and 5112
StntiOn 112+40
-
SUI*IARY OF EXCAVATION ELEVAT10NS
AND STRUT INSTAI.JAPTT0N
Slope
112+69
SI5
Indicators
6 June 69
16-24 June 69*
25 June 69
1--15 July 69
10 July 69
19-24 July 69
22 July 69
25 July-12 Aug.
14 Aug. 69
15 Aug. 69
I-
b
*
a
(ft)
Elev.
Berm El.
Ext. El.
112
112
110
97
112
112
107
97
95
81
Installed
C
80
80
71
71
D
-
1ll+SO
16 P;ay 69
31 June 69
14 Aug. 69
20 Aug. 69
25 Aug. 69
2 Sept. 69
3 Sept. 69
23 Sept. 69
1 Oct. 69
1 Oct. 69
2 Oct. 69
5
69
El.
Excavation
Berm
30 May 69
July 69
IS July 59
1 Aug. 69
1
10 Aug. 69
21 Aug. 69
25 Aug. 69
3 Sept. 69
5 Sept. 69
12 Sept. 69
22 Sept. 69
Elev.
(ft)
Strut Level
Installed
95
92
92
19 March 69
1 April 69
9 April 69
17 April 69
26 April 69
29 April 69
9 May 69
30 May 69
7 June 69
95
85
80
C
82
80
78
73
D
73
73
6
-
.1
JI
Slope Indicators SIl and
113+60
Excavation Elev.
Berm
El.
S4
11.6
11
106
8
95
110
95
88
88
88
91
85
C
82
73
El.
Installed
(ft)
Exc.
El.
Exc. El.
ill
73
73
D
19 May 69
1 Aug. 69
5 Sept. 69
5 Sept. 69
23 Sept. 69
2 Oct. 69
10 Oct. 69
16 Oct. 69
29 Oct. 69
29 Oct. 69
30 Oct. 69
Slope Tndicatorj
SIR, S19, SIll
Exr
vatio
El..
El
EV.
StruL Level
Installcd
(ft)
Exc.
113
113
106
106
106
96
95
95
80
72
95
85
80
72
B
1
Date
Date
111
95
Strut Level
Date
ion 207,50 (Scth Suliheid)
-Berm
Slope Indicators S13 and S16
112+00
1S
Exc. El.
111
106
16
Station 113+40
-
Excavation Elev. (ft)
Berm
in front of 5I-11 Leading to top of Excavation (EL 113)
Station
Date
I
Strut Level
Installed
Strut Level
Excavation
Datc
unl S17
108
108
108
112
108
95
92
92
95
80
92
80
76
72
72
72
B
C
D
EF.
Start to install
steel deck for
Shamut Ave.
Deck cozplete.
C
D
___________
TABLE
Type Instrument
I0a
A.5.1:
SUMMARY OF INSTRUMENTATION AT SOUTH COVE
Number in
Slurry Wall Area
Number in
South Bulkhead
11
12
8
9
3
3
P
8
8
W
4
4
SYM
Slope Indicators
Hyd. Piezometers
Vibrating Wire
Piezometers
Shallow Observation Wells
Vib. Wire Total
Stress Cells
Settlement Screws
Surface Level
Points
Vertical Movement
SI
PH
Total Num.
Installed
SC
4
52
LP
9
Rod
V
3
3
Heave Pins
Permenant Bench
Mark
HP
6
6
BM
1
4
(45 on Don Bosco School, 7 on Slurry Wall)
Strut
No.
Strut
Size
(2)
B-31(2)
B-32(2)
B-34 (2)
B-36(
)
B-38 ( 2
2
)
B-40
B-41
B-42
B-43
)
)
B-44( 3
B-45(
B-46 33)
B-4 7
B-48
B-49
C-31
C-32
C-34
C- 36
C-38
C-40
C-41
C-42
C-43
C-44
C-45
C-46
C-47
C-48
C-49
D-31
D-32
D-34
D-36
D-38
D-40
D-41
D-42
D-43
D-44
D-45
D-46
D-47
D-48
D-49
14WF184
14WF184
14WF184
14WF184
14WF184
14WF184
14WF11l
14WFll
14WFll
14WFll
14WF127
14WF136
14WF136
14WF103
14WF1O 3
14WF78
14WF78
14WF78
14WF 78
14WF78
14WF84
14WF87
14WF87
14WF87
14WF103
14WF103
14WF 111
14WF95
14WF87
14WF87
14WF142
14WF142
14WFI42
14WF142
14WF142
14WF150
14WF158
14WF158
14WF167
14WF176
14WF176
14WF176
14WF158
14WF142
14WF142
Strut
Area
2
(in )
54.07
54.07
54.07
54.07
54.07
54.07
32.65
32.65
32.65
32.65
37.33
Design
Load
(kips)
Jacked
Preload
(kips)
370
482
482
482
402
414
497
497
497
536
536
39 98
39.98
30.26
30.26
22.94
22.94
22.94
22.4
22.94
24.71
25.56
25.56
25.56
30.26
30.26
32.65
27.94
25.56
25.56
41.85
41.85
41.85
41.85
41.85
44 .08
46.47
46.47
49.09
51.73
51.73
51.73
46.47
41.85
41.85
CONSTRUCTION DETAILS OF INSTRUMENTED STRUTS
458
379
278
249
323
323
323
269
303
363
363
363
426
426
426
363
302
218
543
707
707
707
589
633
760
760
760
769
769
769
657
1
Residual
Load
(kips)
420
320
410
490
495
400
470
465
260
160
261
330
300
345
351
195
450
450
380
320
370
290
175
270
204
204
253
249
158
182
64
125
230
200
340
280
310
330
330
270
225
160
590
550
88
138
232
119
136
138
169
132
133
92
324
293
287
209
205
354
294
297
360
325
232
365
254
136
122
540
430
440
530
660
550
535
570
615
510
380
235
(1)
Date
Installed
7/31/69
7/31/69
8/ 1/69
8/14/69
8/14/69
8/14/69
8/14/69
8/14/69
9/ 4/69
9/ 4/69
9/ 5/69
9/11/69
1/11/69
9/11/69
9/11/69
9/ 2/69
9/ 3/69
9/ 3/69
9/ 3/69
9/ 9/69
9/23/69
9/23/69
9/23/69
10/10/69
10/10/69
10/10/69
10/15/69
10/15/69
10/15/69
10/15/69
9/22/69
9/22/69
9/22/69
10/ 1/69
10/ 1/69
10/24/69
10/25/69
10/27/69
10/28/69
10/29/69
10/29/69
10/30/69
10/30/69
10/30/69
10/31/69
(1)
Residual strut load after shimming the strut and removing the preload jacks
(2)
The F-2 strain gauges were improperly installed on these struts.
gauges were installed from 17 Oct 69 to 27 Oct 69.
(3)
Six F-2 gauges were installed on these struts; the others had two.
-162-
Date
Removed
9/29/70
9/29/70
9/30/70
10/30/70
10/30/70
11/ 2/70
11/ 2/70
11/ 2/70
11/ 2/70
11/ 2/70
10/20/70
10/ 8/70
10/ 8/70
10/ 8/70
10/ 8/70
11/ 3/69
11/ 4/69
11/ 6/69
12/ 2/69
12/ 2/69
12/10/69
12/12/59
12/15/69
12/15/69
12/16/69
12/16/69
1/ 2/70
1/ 2/70
1/ 2/70
1/ 2/70
11/ 4/69
11/ 4/69
11/ 6/69
11/19/69
11/19/69
12/11/69
12/11/69
12/14/69
12/16/69
12/16/69
12/16/69
1/ 2/70
1/ 2/70
1/ 2/70
1/ 2/70
Two additional Tel ,
SB-90
Strain
Gauge
Type
SB -90
SB-90 & F-'
SB-90
SB-90 & F-2
SB-90 & F-2
SB-90 & F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2 & SB-90
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
F-2
&
TABLE A.5.2:
strain
TABLE
Instrument
No.
No.
Date
Installed
Installed
1
PH-1
PH-2
PH-3
PH-4
PH-5
PH-6
P1-7
PH-8
PH-9
PH-10
Elev. of
TABLE
4
Sensor
+75.17
+74.93
+74.94
+74.96
+75.20
+75.04
+75.01
+75.0
+75.0
+72.45
+75.0
1
South Bulkhead
South Bulkhead
South Bulkhead
A.5.4: LIST OF VIBRATING WIRE FIEZOMETERS
Date
Installed
Elev. of
Sensor
P-1
9/25/68
+62.70
P-2
P-3
P-4
P-5
P-6
P-7
P-8
9/19/68
9/17/68
9/25/68
9/17/68
9/17/68
9/26/68
9/26/68
+39.0
+13.32
+93.0
+65.5
+40.5
+93.0
+92.1
5/ 3/69
5/ 3/69
+58.2
+56.8
P-1 Replacement
P-lH
Remarks
Sensor
11/ 6/68
10/30/68
11/ 5/68
11/ 7/68
11/ 1/68
11/ 1/68
11/15/68
11/ 4/68
11/21/68
11/22/68
11/15/68
PH-11
Instrument
No.
A.5.3: LIST OF HYDRAULIC PIE ZOMETERS
-163-
Remarks
Lost P-1 (shorted out
4 April 1969)
Both installed in old
P-1 Hole.
TABLE
Instrument
No.
A.5.5: SUMMARY OF SLOPE INDICATORS
Date
Installed
Station
Location
SI 1
113+42
20 Sept. 68
In soil 5 ft. back
of the slurry wall
SI 3
111+83.5
25 Sept. 68
In soil 29 ft. back
of the slurry wall
SI 4
113+42
28 Feb. 69
In slurry wall
SI 5
112+44
5 Feb. 69
In slurry wall
SI 6
111+83.5
28 Feb. 69
In slurry wall
SI 7
112+44
18 March 69
In slurry wall
SI 8
Center of south
bulkhead wall
28 Jan. 69
On steel sheet
piling
SI 9
107+38
21 Feb. 69
On steel sheet
piling
SI 10
107+25
4 March 69
On steel sheet
piling
SI 11
111+39
3 June 69
On steel sheet
piling
SI 12
111+14
7 June 69
On steel sheet
piling
-164-
TABLE
A.5.6: LIST OF VERTICAL MOVEMENT RODS,
LEVEL POINTS AND BENCH MARK
Instrument
No.
Elev. of
Sensor
Date
Installed
V-1
V-2
V-3
9/24/68
9/20/68
9/23/68
+71.3
+58.4
+39.2
LP-1
+118.6
+118.2
+118.4
+118.0
+118.8
+118.1
+116.9
+114. 3
+114.9
LP-2
LP-3
LP-4
LP-5
LP-6
LP-7
LP-8
LP-9
BM
10/8/68
Date
Installed
Elev. of
Sensor
W-1
W-2
W-3
W-4
7/14/69
7/14/69
7/12/69
7/12/69
+92.6
+115.6
+96.3
+114.4
SC-1,
SC-1,
SC-2,
SC-2,
Cell
Cell
Cell
Cell
#51-67
#46-67
#31-67
#33-67
month
LIST OF SHALLOW WELLS
Instrument
No.
Instrument
No.
Contractor installed the level
points over a period of one
+7. 8 T i'p Elev.
TABLEA.5.7:
TABLE A. '.8:
Remarks
Remarks
Pourous plastic piezometer
36" Long Wellpoint
Geomeasurements, Hyd, Piezometer
36" Long Wellpoint
LIST OF VIBRATING WIRE STRESS CELLS
Date
Installed
9/24/68
9/24/68
10/ 1/68
10/ 1/68
-165-
Elev. of
Sensor
Remarks
+65.5
+64.7
+64.1
+63.3
Not operating
TABLE
A.6.1:
STRUT LOADS VERSUS CONSTRUCTION STAGE
Preload "B"
Level
Excavate for
"C" Level
Preload "C"
Level
Excavate for
"D" Level
Preload "D"
Level
Load
Temp
Load
Load
Load
kips
Cc
kips
260
160
261
330
300
345
351
195
27
17
31
22
22
23
21
22
22.4
21.9
20.3
18
18
18
18
352
283
340
390
338
384
250
185
221
214
254
294
208
194
185
Excavate Final
Stage
remove
"C" and
"D" Level
Strut
No.
B-31
C
kips
Temp
C
Temp
C
Load
Temp
Load
kips
kips
'C
kips
Temp
0
Load
Tenp
C
kips
Cc
490
380
470
430
480
480
360
330
320
350
430
390
410
390
300
13
8
8
23
23
14
12
19
27
27
27
29
21
27
0
349
250
330
408
370
366
269
220
235
211
276
268
203
179
187
22
23
17
22
17
15
15
17
17
22
22
20
20
21
21
371
302
370
467
380
349
292
223
245
227
285
267
215
196
199
32
15
15
17
23
19
14
14
19
17
16
11
13
14
15
394
285
362
475
383
328
290
232
237
214
247
283
215
187
193
15
16
16
22
19
19
19
15
17
12
10
23
11
16
15
380
316
362
475
383
382
295
237
237
213
261
268
194
192
221
15
31
16
22
19
9
11
9
9
10
9
11
3
4
2
64
125
21
21
180
182
184
232
31
31
15
22
88
138
232
119
136
138
169
132
133
92
16
16
16
15
15
15
12
12
12
12
104
166
269
144
247
230
152
138
155
106
23
14
14
16
17
15
14
15
14
14
122
122
179
216
192
103
176
265
105
231
205
161
158
143
202
15
15
15
24
22
18
18
15
14
14
19
11
15
14
9
182
182
179
216
182
117
162
272
126
203
209
155
147
109
149
31
31
15
24
16
9
11
6
11
10
D-31
324
17
D-32
293
17
D-34
287
17
D-36
D-38
D-40
209
205
354
18
18
0
270
170
290
19
15
11
D-41
D-42
D-43
D-44
D-45
D-46
D-47
D-48
D-49
294
297
300
325
232
365
254
136
122
10
16
8
11
10
9
9
9
9
241
209
210
297
178
322
281
190
122
12
12
12
4
10
8
5
2
0
B-32
B-34
B-36
B-38
B-40
B-41
B-42
B-43
B-44
B-45
B-46
B-47
B-48
B-49
'-C
Temp
C-31
C-32
C-34
C-36
C-38
C-40
C-41
C-42
C-43
C-44
C-45
C-46
C-47
C-48
C-49
204
204
253
249
158
182
,
28
19
21
20
18
15
15
17
27
19
18
23
19
22
22
9
8
2
1
3
x
I>
C
PERCHED
0
20
I
40 600
3
.
4.0
3.0
2.0
1.0
0
50
6.0
TEST SAMPLE STRESS HISTORYI TYPE
No.
E, E
t.c
TEST
loom)
I
40
4
w-6
-
-60
.
BOSTON
7-122
-- -N(OIUM
V-.-
BLUE
60
CLAY
'00
i
cANGLE
3ITWIST
\
TILL
W~p
114
OF
AT
FALURE
0%
SHEAR TESTS BY MIT
SHEAR TESTS
0 TORVANE
* FIELD VANE SHEAR TESTS FROM
J.R COLLINS SOILS TEST REPORT
-
TEST RESULTS FROM BLOCK SAMPLES
TEST RESULTS FROM 3"SELBY TUBE SAMLES
PERMEABUTY TESTS
ST SAMPLE
No Elex.(ft) :PE lAAT)
m
C-I
84
4.0
0.8
C-2 -
74
7.0
2.5
1
04
4.0
2
84
6 0
60.0
3
84
5.0
3.0
C-3
74
6.5
1.0
4
84
420.0
60
C -4
74
20.0
0.8
2.0
6.0
C-5
104
70.0
20.0
5
74
TEST
x (/) x max REMARKS
Py
14 2043640580521000 -
1.4
2.04364 34' HLO YELLOW CLAY
380
LOE
130 52
30
UU
241.14
3
94.0
1.30
52
3.2
UU
2.6
4
84.0
1.61
4.7
1.18 3S8798
CJU
5
64.0
1.61 4.7 .78 6.0 10.0 T U
16 2.0 56570 .43
300271 5.1
2.55565 268
6
84.0
1.61 4.7 78 6.0 10.0' CIU
&8 1.883441.18 .31 940O3.11.5
1.85 345 34
1.05 4.0
1.23
2.8 L28 2,86098 32 430 26 104 120 343 205
8.0, C I U 46
1.26
251
1.15
315 340 302 13 1.15 2.05 342
7
74.0
1.92 4.3
90-92
1.40 5.2
U
7.0 122
3z
(9-2)-2
70-72
2.0 42
U
32 .55
41.RO
19-31-3
51-53
3.60
33
U
75 .2
0
33
27
62-63
2.25 3.8
U
184 .82
85-87
1.55 4.9
UU
7.3 L03
(8-2)-1
70-72
2.0 4.2
UU
8.1 .99
28
10-2)-I
98-100 1.35 5.6
UU
i8 .55
20
(10-31-
8-I -I
I ALL
3q
RESULT
STRESSES IN
SITU
2.IN
.
CLY
680
19-11-I
RESULTS FROM TESTS ON BLOCK SAMPLES
FIGURE A.3.1: LABORATORY
1.42 4.0 80 C I U
-
94.0
FELD VANE
COEFFICIENT OF CONSOLIAfON
TESTSAMPLREc MPRESSION VMIN
xxr3
3910-3
N. Elv.(d
/
A 0
.88 5.7
AT (
2
I
\
)'" lpcf
-20
104.0
RESULTS
TEST
TRIAXIAL
FILL
HD YEL
T
G
CLAY108
V STIFF
TU STIFF
0a
STRESS (T/FT2)
W(.
WATER
2
STRESSES
3" SHELBY TUBE
SAMPLE RESULTS 47
REPORTEG IN J P
SO
2
TONS/FT
5. E-SECANT MODULLS a
6. Ou z SIN-' -2
P
f
DON
BOSCO
I
SCHOOL
uSI -6
11
-
+SI-3
T-7' i
0
0
0
0
00
I
12
1
I')
-4----
-- i----
-H-
I
9
1 -4
.............
I
I
8
i-4 o
I
7-
-_-_
19
214
FIGURE A.4.1:CONCRETE
WALL
PANEL
LOCATIONS
Concrete Strength
B
Vertical Reinforcement
(All bars # 11)
IL
SYM
C
L(ft)
Bars/f t
Spacing
10'
7'
0.27
1.34
1.07
44 1/2 in.
9
I1 1/4
D
1I'
1.25
9 5/8
E
F
G
H
I
J
K
L
M
18'
20'
12'
8'
22'
9'
10'
6'
5'
2.88
0.90
0.27
1.97
2.88
1.25
1.02
0.71
0.27
4 1/4
13 3/8
44 1/2
6
4 1/4
9 5/8
12
17
4 4 -1/ 2
A
B
C
D
L
ottom of
Excavation
E
= 3000 psi
Yield Strength of Steel
11 Bars - 60 OOOpsi
Others - 36,060 psi
HorizontalI bars
H-or*z7nCa) 1'- b6 /c#
/c
H1-6
X
Vertical
Bracing -
#
I
T
5 Bar
K
Reinforcement
stirrups - ;
5 Bars
L
1<--3 ft -+
FIGURE A.4.2:TYPICAL
CAGE
REINFORCEMENT FOR STEEL
IN CONCRETE PANELS
-169-
+120
I
FILL
Hard
4
Yellow
Clay
L
7{= 125 pcf
Very Stiff
+100
(SZDesign)
-E
to Stiff
9
//
+80
D9
(
Boston Blue
CLAY
Max.
t =122pcf
+60Med. Bostcn
Blue CLAY
+401--
-
Concrete Slurry
t = 120 pcf
Wall
TILL
0
10
0
30
Scale: Icm=lOft
0
SLURRY
FIGURE A.4.3: EXCAVATrON PROCEDURE
WALL SECTIONSTATION
113+40
-170-
\
N
Bosco High School
-m..Colled
Nort h
/
(
+95 = ELEVATION OF EXCAVATION
BOTTOM (B.C.B. -4- 100.0)
Don
MUD
S L A B
r
+9
+7 +79
-4-97
RM
R
+08
B- LEVEL
INSTALLED
9
+i-9*
+:71
-4
0
8
+79
++1
0
8
4.
LO
4
N
N
I
I
0
I
I NVERT
MUD
SLAB
POURED
+76
+7
+83
o
1o
20 30 40
50
D-LEVEL
INSTALLED
Scale: 1cm = 10 ft
FIGURE A.4.4:CONTOURS OF EXCAVATION
BOTTOM WHEN
INSTALLING
STRUTS 41-+42
I
~
-*
C
~t1~
~~Ac
41
I
~\
FIGURE A.4.5: VIEW OF THE SITE
-172-
SP-7
7W-6
LEGEND
7
DON Bosco
80C
SCcOOL.
la"
22
LP - Level Pin
P - Vibrating Wire Piezometer
PH - Hydraulic Piezometer
SC - Total Stress Cell
SI
- Slope indicator
W- Shallow Well
Struts with Srain Guages
"'"-C
*Lp.8'A
k
-n\CALL E
46
NORT
+
LP-r
*3r-3
L10fr 0
+
I-+
- sr - u
R
4
a
33
su
s-1u
23
2
23
26
rvE-
cs
2
27
i~
f 120
-71AWSu~~~~cN~~~~mnemr zA mw oopf - ryi 00
t
7
412
*
90
+
SUE
ii~i..IIIz
Song
L enses
+60
Rled.
1,7
-i -7
J
Blu
~ s~-~'
glue
50
b'-
-
/
-Sbf
4 lA.6
'e .
Eor
SCALE
- ,V
0
c0
IC.,-
FIGURE A.4.6: SOUTH
--
0
-- ..
40 50
CLvuR
40
--
WAu
0
SSECT/ON1\
g
_
__.
10 Oct
COVE SLURRY WALL TEST SECTION
INSTRUMENTATION
,,,
AND
BRACING
SCHEME
.
22
2
s (
%
09
7
-
--
/
-- ---s-
owwwom-
-
0
0
00.
0
N
/
-9.
z
0
IU,
PH-l1
7o
36WF
FOR DECK
MSI-1O
,
Z9t-
14WF103
Corner Bracin
-if(
SI-80-
Deck IArea
I0
ON
14 WF 10
0
TOP OF
SHEETING
PLAN
PH-lO
+9
SL-9
VIEW
Ground Surface
120-
0
0
STRUTS
FILL
0
co
L) 100-
Hd. Yel.
CLAY
z
Stiff
0
-
/
1
12WF65
1
1B
24WFIIO
STRUTS-14WIlI
z
IC
36 WF300
STRUTS- 14 WF127
Blue
~80-
w
-J
w
II
1
II
1I
CLAY
60-I
YZX\\'>WA\
with fine
Sand
Lenses
Bottom of Excavation
SHEET
C ROSS
0
10
Z)
30
40
PILING TIPS
(
EL 40
SECTION
50
Scale (ft.)
BL ULKHEAD
SOUTH
COVE
FIGURE A4.7: SOUTH
INSTRUMENTATION AND BRACING SCHEME
-174-
I
11
15
23
L P-9
3
3
3
3
0 0
20
Scale -cm =10 f t
41
5
DON
BOSCO
HIGH SCHOOL
LP-2
10
X
19Q
ELP-3
w ,9
7
7L
LP-8
,"
vp37
L
22 L
3
X
30L
LP-7
46
38X
38 L
F
L- 5
L-
LP-6
si-6
si-5
S -6
s-5
FIGURE A.5.1:
LOCATION
POINTS
OF
LP4
LUWRRY
WALLs-4
SETTLEMENT
E
G
SCREWS AND
LEVEL
I
Don
Bosco School
.~p -4
of
o-i-P -
Fill
-
+/0
70. 5
58.5 to 79.5)
SI
-
(Varies
Hc1rd ye//ow
Clay
+/00
1ery S/iff
to S/if
Lhue
..lo y With
-
.*-.51-
Sand
/
+80
H -a
Lenses
8C-9
ON
sc-I
II
0*~~
II
I,
I'
II
II
.f.
-f
II
Med Blue
Clay
4*0
Cu
3
0.
-0
-
I
7Wl
*
8 enchmrr
erA
0
77r7r
, ,
,,,,7
0
/0
cO
Sca/e: /cm
FIGURE A.5.2. PROFILE
OF
INSTRUMENTATION
AT
STATION
113 + 40
30
= /D
ft
I
WATERPROOF
CABLE ELEC TRICAL
ELECTRICAL
CAB LE
DIAM.= 6.5 M m
PLUG
FLANGE
GROUNDING
RING
SETTING
0"
SCRE W
RING
CONE
MUSIC
WIRE
BRASS
CASING
RODS
0
LI
MAGNETS
I ~
HOLES
I/4" DIA
FIGURE A.5.3: SCHEMATIC
VIBRATING
-177-
OF TELEMAC TYPE F-2
WIRE STRAIN GUAGE
I
PIPE CAP
TANDARD
Existing
Ground
10' *i
+-5 FT tLength 5" or 6"
, Capped Protective Pipe
Bockfill w/
* I
Blasting Sand to
within 10ft.t! of
Ground Surface
( UNCOMPACTED)
2 1/2" Standard
Pipe
( Bottom length>--10 ft.
)
with no Coupling at
Bottom
5
Nol Blasting Sand Backfilled
Lifts
Compact in IFT.
5*
-.
SENSOR
5FT. Compacted Bentonite
or Chemical Grout Seal
#-IFT.
-
4 1/2'
I 1/2'
;2
FIGURE A.5.4:HYDRAULIC
-178--
B o
PIEZOMETER
1 Blast ing
Sand
INSTALLATION
VENTED CAP
.1-Li
.
GROUND
SURFACE
I I
In'
fill
-
3 LENGTH
OF 2/2 " PIPE
Li
1
.
b~.
-
I" RISER PIPE
-1
I:.
SAND FILTER
4"IWELLPOINT
-
00
I.
~
I
II
I.v,
FIGURE A.5.5:SCHEMATIC DIAGRAM OF
TYPICAL OBSERVATION WELL
-179-
I
I
"Co/
/o07_7
' A2D 5n/s
s
sAee/ /-n'c hoo-- screw,
/he ornchor
Sef to />ro/ect
«|.
I, r<'>O
-
anchor ,Vo
.922O
or
egwivo/ent
0
N
e-
9
1J/.
I
p
______________________
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I
I
I
I
I~ii-~j
r<&
I
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/
67>
/ 4/4" ,n$o
fate :
dr-,ed
the coArrnn
W/hen
the
aere
/%e
-- 0447,2' /XeOd
to r-eCe r
on, edge of heod
37/,/" screw
a
J+'''~scre'
2ree/
pa
0
/c/e
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elnelt observorlor,
/,n..//,- t/)epoe
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0
1
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rr
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fI
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f
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S.S 2'"DIAM.
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TO PIPE
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SCREWED PIPE
FLANG ES
2"EXTRA HEAVY
PIPE ASTM
DESIGNATION A120-63aT
GASKETS
BRASS
S S 2"EX TRA
2
0-RING-
HEAVY PIPE
WATER STOP 7
3 2"EXTRA HEAVY PIPE
3 2"EXTRA HEAVY
A.ST M DESIGNATION
PIPE
A120-63oT
2"EXTRA HEAVY
PIPE
2
I
CONCRETE
SLAB
3'-0"
GLACIAL TILL OR
BEDROCK
WASHED
- 2'- 0".
TOP ASSEMBLY
DETAILS
HOLE
STEEL POINT WELDED
TO PIPE
1*
FIGURE A.5.7: DETAIL OF DEEP BENCHMARK INSTALLATION
-181-
STANDARD PIPE CAP
2 2
DIAMETER PIPE
CONCRETE
I"DIAMETER PIPE
1
4" ROD TO
DRIVE TIP
BORRO SETTLEMENT
POINT
EXTENDING ANCHOR
PRONGS
6
FIGURE A.5.8:DETAIL OF BORROS HEAVE ROD
-182-
STANDARD
PIPE
CAP
1/2
3' of 8" dia.--
P"-TOP
3'x3' concrete slab-5"thick
Y//A\\\ GROUND LEVEL
//A\\v//\
STEEL
JERSEY
"Steel
bore
S lope
CAP
Li-NEW
1 1/2'
-6"
OF WALL
hole
-Slope
indicator
-
6"
#s_2 blasting sand
compacted
6"
[1
Blasting Sand
No 1/2
Pipe
Indicator Casing
3# 7 Reinforced Bars
At
Equally Spaced
Continuous Weld
Each Coupling
ack Weld All Around Couplings
Make Up Couplings Tight m/ Pipe Dope
(Joint to be Water Tight)
co)
<
6
bore hole
4
Slope indicator
casing
( NOT
GROUND
INSTALLATION
FIGURE A.5.9: SLOPE
TO
SCA LE)
INDICATOR
-L1
WELD ALL
AROUND
SLURRY WALL
I NSTALLATION
Dia
AOut side face
SLOPE INDICATOR
DETAIL
PIPE TIE
WELL INSTALLATION IN SOIL AND
SLURRY
WALL
P/. A
'e
-r
I-
o p~p / from /op of
$h? et p/e. we/& ,o o-
--.
bo/
fop
s'?s
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p//rOJ/
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70d
A
-
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p-re
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w/ee/ p/o.
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to/fom coup/eng on
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ill
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HEAVY PIPE
5
ga
PUSHEI
U-1
ABOUT 6'[i
2"1
STRESS
DETAILS OF STEEL
DRIVE SHOE
INSTALLATION OF CELL
FIGURE A.5.11: VIBRATING
WIRE
TOTAL
ST RE SS
UNIT S
CELL ASSEMLY
0
qak~
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0
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2/21/69
A
3/11/69
PANEL CONSTRUCTION
PANEL DATE LOC.FROM
SI-I
NO.
14
17
20
1/28/69
12/20/S9
1/28/69
2/21/69
North
South
Front
FIGURE A.6.16: HORIZONTAL GROUND MOVEMENTS
DURING INSTALLATION OF THE SLURRY WALL
-222-
SI-I
(-)
3
2
East
-H--
_+)
MOVEMENT IN INCHES
I
I
0
2
x
+*
\
x
106-
Berm elev.
t, elev.
I-
/
/
x +-
-
40
0/2/69
z
z
10/29/69
0
-J
8A4169
w
z 86
w
w
I
20
w
uw
3
2
+
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125
I
West
66
60 a-
10/30/69
46 -80
26 - 100
SYM.
+
x
DATE
8/13/69
10/2 /69
10/27/69
10/31/69
2/4/69
10/10/701
Base line
)=&east, in.
0.38
0.16
0.25
0.16
0.35
FIGURE A.6.17: HORIZONTAL GROUND MOVEMENTS
DURING EXCAVATION
-223-
SI- 3
)
(-
)
North
IN INCHES
MOVEMENT
122.5
3
2
I
West
'
0
1I
2
2
-I
East
MOVEMENT IN INCHES
I
0
1
[I
102.5
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LL
I-
z
LU
w
3
5
6
60 z
I-
0
SYM
42.5--80
0
---
22.5
Berm elev
(I e ley.
PANEL CONSTRUCTION
PANEL DATE LOC. FROM
NO
DAT
51-3
w 82.5 -40
-I
3
I
20
I-
z 62.5
0
2
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/
South
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Y
x
2.5 L 120
0
1/14/69
2/3/69
2/2/69
DATE
North
North
East
W Base line
[)M=Aeasi,in.
1/28/69
3/11/69
5/14/69
7/31/69
8/26/69
10/2/69
2/4/70
10/10/70
(~
0,8
0,4
07
FIGURE A.6.18: HORIZONTAL GROUND MOVEMENTS DURING EXCAVATION
I
I
/
7/18/69
r
8/25/,69-
9/12r'69
-4
SI-4
East
West
IN INCHES
I
0
MOVEMENT
124-
2
3
1
I
2
4/2/6
I
4/29/69
-
I
114
3
-10
5/16/169
104- -20
8/13/69
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w
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z 84-
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10/2/69
50
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10/29/69
elev.
10/30/69
S
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5/23/69
8/13/69
0.0
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10/2/69
0.3
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0.4
v
10/27/69
A0
10/30/69
2/4/70
+
+
54- -70
I
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AT Base line &(S14 S 1)
DATE
10/10/70
0.3
0.1
0.5
44- -80
FIGURE A.6.19:HORIZONTAL MOVEMENTS
SLURRY WALL
-225-
OF THE
SI-7
SI-5
West
2 35
3
2
West
East
East
MOVEMENT
MOVEMENT IN INCHES
1
0
I
2
2
I
I
IN
0
,I
113.5 -- 10
5/r649
103.5 - 20
7/31A9
I-
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LL
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w.
w
z
z 83.5- -40 z
0
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w
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0
1o
69
63.5
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DATE (4i.Aeastin. -1decr,in
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5/23r9 0.0 0.0
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7/31/69 0.1 -0.5
8/2069 04 04 -03
10/2/69 01 -0.3 - 03
2/4/70 0 -03
02
10/0/701 0.1 -031 - 00
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535-70
-
a.'
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8/25/69
0
43.5--80
FIGURE A.6.20:HORIZONTAL MOVEMENTS
OF THE SLURRY WALL
INCHES
I
I
2
I
3
A
SI-10
East
West
118
I
3
2
I
MOVEMENT IN INCHES
I
0
1
I
i
2
3
I
I
Berm elev
108- -10
4/17/69
t elev
98 + 20
C
I-
w
1-
z 88 + 30
z
-
0
5/9/69
z
78- -40 a
5/30/69
Id
6/7/69
68 t 50
58- -60
481 -70
SYM.
DA TE
o
5/1 2/69
A5/3
0/69
6/1 2/69
8/1 3/69
*
0
a Base line
(+)=& east,in.
0.0
-0.7
-
Q6
-0.6
FIGURE A.6.21: HORIZONTAL MOVEMENTS OF SHEET
PILE WALL AT THE SOUTH BULKHEAD
-227-
I
SI- II
East
West
6
114.8 6
4
MOVEMENT IN INCHES
22
0
2
4
4
6
I
x
104.8 + 10
I
94.8+20
Iw-
H
z 84.8- -30
LL.
w
I
7/10/69
x
w
z
m
Fa5:
-40 w
74.8
w
a
IJ
w
7/24A9
8 elev.
erm el.
8/25/69
64.8 -50
SYM.
54.8 -- 60
x
448- 70
+
v
0
DAT E
7/9/69
S 7/24/69
S8/20/69
8/26/69
11/26/69
2/4/70
10/10/70
Base line
east,in.
0.1
-0.6
-0.6
0.3
1.2
0.8
FIGURE A.6.22:HORIZONTAL MOVEMENTS OF SHEET
PILE WALL ADJACENT SLURRY WALL
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to
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F-I
APPENDIX B
ACCURACY OF INSTRUMENTATION
B.1
INTRODUCTION
Proper evaluation of a field measurements program requires
an assessment of the accuracy of the instrumentation readings.
To
assess this accuracy, factors such as the effects of instrument environment, sensitivity, repeatability, and stability require study.
Both field and laboratory investigations were conducted to
establish the degree of accuracy of the South Cove instrumentation.
The investigations consisted of:
instruments; 2)
3)
1)
field calibrations of the
field and laboratory studies of temperature effects;
monitoring of pertinent field data which would assist in estab-
lishing the measurements accuracy.
This appendix summarizes the
results of these studies and describes the accuracy limits of the
instrumentation.
Where applicable, correction constants are estab-
lished for the various field measurements presented in Appendix A.
B.2
STRAIN GAUGES
B.2.1
General
Telemac F-2 and SB-90 vibrating wire strain gauges were
used to monitor the strut loads.
The basic operational principles
of the gauges consists of relating axial strains to the changes in
frequency of a vibrating wire which is a function of the wire tension.
Strut loads are computed by the equation:
Ps = ASEX(No 2 - N1 2 )
-234-
B.2.1
where:
=
As =
strut load (kips; positive value is compression)
cross sectional area of the strut (in 2
)
P
E
=
Young's Modulus of the strut (ksi)
X
=
calibration constant for the gauge
No=
the no-load vibrating frequency of the wire
Ni
the vibrating frequency of the wire at the
=
measured load
For the F-2 gauge the theoretical constant, X, is
3.125 x 10~.
9
The accuracy of the measured strut loads as determined from
the above equation are dependent on the following factors:
B.2.2
1.
Temperature inertia of the gauge,
2.
Accuracy of the calibration factor,
3.
Drift of the no-load frequency.
Temperature Effects
There are two ways by which temperature can cause errors in
measured strut loads.
The first is by environmental conditions causing
the gauge and strut to be at different temperatures.
The second is
caused by the gauge having a different thermal inertia than the strut.
Both of these conditions can cause a shift in the no-load frequency of
the gauge and result in a substantial error in the computed strut load.
One cause of temperature gradients between a gauge and strut is direct
sun rays.
housing,
These cause the temperature within the protective guage
(see Figure A.5.3) to increase above that of the strut.
-235-
Figure B.2.1 shows the fluctuation of frequency readings on unloaded
struts for gAuges which are both exposed and shaded from the sunlight.
Over a 24 hour period the exposed gauges showed a frequency variation
of about 35 cps whereas the shaded gauges showed a variation of
approximately 10 cps.
Figures B.2.2 and B.2.3 show additional data
on the thermal inertia of both shaded and unshaded gauges.
This data
indicates that shading the strain gauges reduces greatly the effect
of the sun rays.
However, it does not completely eliminate the effects
of temperature on the no-load gauge frequency.
Evidence of differences in thermal inertia between the strut and
strain gauges is given by the 10 cps frequency change over a temperature change of 10*C
for the shaded gauges in Figure B.2.1.
The
effect of this frequency change is shown by Figure B.2.5 which compares strain gauge sensitivity verses gauge frequency.
The data on
Figure B.2.2 indicates a 10*C temperature change results in a 10 cps
change in gauge output frequency.
For gauges set at approximately 600
cps at zero load this would result in approximately a 55 kip error in
a measured strut load for a strut which has cross-sectional area of
.
50 in 2
Figure B.2.4 shows the temperature variation for the construction
period at South Cove was on the order of 50*C.
The temperature fluc-
tuation shown was recorded by the United States Weather Bureau in Boston
which is approximately 2 miles from the South Cove project.
For com-
parison, the figure shows periodic air temperatures at the project site
as well as temperatures recorded by sensing devices within the strain
-236-
-1
gauges.
The data shows the air temperatures at the site and at
the Weather Bureau are in close agreement, hence it is assumed the
recorded temperatures by the Weather Bureau are representative of
the average temperature variation experienced by the shaded gauges
during construction.
Table B.2.1 also shows the temperature of
struts in the instrumented section of the excavation varied both
longitudinally along the axis of the excavation and vertically at
a strut section by as much as 6C during the time it took to record the strain gauge readings.
Based on the measured no-load
frequency changes for the shaded gauges in Figure B.2.1, it is obvious the aforementioned temperature changes could induce substantial
error in the measured strut load if the loads are not corrected for
the differential temperature inertia between the gauge and the strut.
B.2.3
Temperature Corrections
The differential temperature inertia between the gauge and
the strut is primarily the result of some gauge components being made
of materials possessing thermal coefficients unlike that of the steel
strut.
A composite correction constant for the thermal inertia can
be obtained for the gauge by attaching the gauge to a steel member
and under zero load conditions vary the temperature of both the gauge
and steel member in a constant temperature environment.
Both the
gauge and the strut plate must be allowed to come into equilibrium
at a given constant temperature.
If changes in the no-load frequency
are measured they can be used to develop a composite correction
factor for the gauges to be applied to the field data.
-237-
Figures B.2.6 and B.2.7 show the results of tests of the aforementioned type, which were made on Telemac F-2 and Telemac SB-90 gauges.
The gauges were attached to a 1/4 inch steel plate and the temperature
of the composite unit varied over the range of temperatures experienced
in the field.
Two sets of tests were undertaken,
International Inc.,
the other by the author.
show a wide scatter, however,
one set by Telemac
Both sets of test results
the data shows definite trends.
The F-2
gauge frequency decreased with increasing temperature, whereas the
SB-90 gauge frequency increased with increasing temperature.
A temperature correction factor is obtained from this data by recognizing that, under no-load conditions, a
1*C change in the tempera-
ture of the gauge and steel plate will result in a given difference in
the term (No -
N1 ) as defined in Equation B.2.1.
in the (No -N1
) term will be constant,
This difference
irregardless of the No value
since this difference is dependent only on the tension in the wire which
will vary linearly with the temperature change.
The equation for strut loads corrected for temperature effects can
be expressed as:
Ps = As [EX (No2 - N1 2
Kt (TO-Tl)]
B.2.2
where
T0 =
Gauge temperature at time of reading N0
Ti =
Gauge temperature at time of reading N1
Kt =
(N02 - N 1 2 ) /
(T0 - T1 ) at no-load condition
from thermal inertia tests.
Table B.2.2 gives values of Kt for the F-2 and SB-90 gauges which
"-238-
were determined from Figures B.2.6 and B.2.7 and theoretical computations.
The theoretical values are those reported by Telemac International Inc.
The strut loads reported in this text were corrected using a Kt =
87.5
psi/*C for the F-2 gauges, and Kt = +46.0 psi/*C for the SB-90 gauges.
B.2.4
Gauge Calibration
The strain gauges and struts were calibrated as a load cell
during the initial preloading of the struts.
The purpose was to define
a composite factor so that strut loads could be evaluated from the
equation:
Ps = Xp (N2 - N)
where:
Xp = AEX as defined in Eq. B.2.1
The term XP would have accounted for the effect of manufacturing tolerances resulting in some variation in the actual strut area
and Young's Modulus from those values published by the manufacturers.
(For example, according to the American Institute of Steel Construction,
Inc.
[AISC] specification the cross sectional area of a strut is rolled
to within
2.5% of the published area for a given strut size).
Figure B.2.2 and Table B.2.3 compare the strut loads computed
from the strain gauge measurements versus the strut load determined
from the calibrated jack pressure gauges for various steps during the
preload operation.
The scatter is quite large and the data shows a
definite tendency for the jack loads to be higher than the theoretical
loads.
-239-
Figure B.2.9 shows the percent difference between straingauge-strut load and the measured jack loads at the maximum preloads
placed on the struts.
than 10%.
The percent differences are, in general, greater
This is much higher than the expected composite error of 3%
as determined from manufacturer tolerances for struts published by AISC,
and the accuracy of the strain gauge calibration (X) (which is published
as
.l%) reported by Telemac International Inc.
This large difference between the computed strut loads and
the loads determined from the jack is attributed to friction in the
jacks.
A review of Figure B.2.8 supports this hypothesis since the
error tends to increase with jack load.
If the error was the result
of variations in strut size and gauge calibration it would be essentially constant with strut load.
Figure B,1.10 compares the computed
strut load from both SB-90 and F-2 strain gauges which were mounted on
the same strut and were simultaneously monitored during preloading.
This
data is actually a representation of the accuracy of the gauge constant
(X) since the strut area and Young's Molulus are common factors.
agreement suggests the manufacturers value for X is within
This
0.1%.
The strut loads reported for South Cove were computed using
the manufacturers recommended gauge constant (X), the published strut
areas, and a Young's Modulus of 29,000 ksi.
This approach was chosen
because of the excellent agreement between the two strain gauges as
shown on Figure B.2.10 and the uncertainty in using the preloading data
to determine a composite calibration constant for the struts.
-240-
B.2.5
Gauge Drift
The primary causes of drift in the no-load frequency (NO) in
vibrating strain gauges are:
1) loosening of the lock nuts connecting
the gauge to the support posts; 2) bending of the gauge support posts;
3) shock of the gauges by impact loads resulting in a relaxation of the
tension in the vibrating wire; 4) Creep in the tensioned vibrating wire
with time and fluctuation in temperature.
a decrease in No.
Each of these tend to cause
In the event No does decrease during the measurement
period, theoretical loads figured on the basis of the initial (NO)
value will be too high.
Figure B.2.11 and Tables B.2.4 to B.2.6 show the apparent
residual loads (or drift) as recorded by the strain gauges after the
struts were removed from the bracing system.
The results given in
Figure B.2.11 are the average loads obtained from readings of the
gauges on a strut and are corrected for temperature effects.
Wire creep appears not be be a cause of the residual strut
loads for the following reasons:
1.
Figures A.6.1 through A.6.8 show no definite trend in
variation of the B level strut loads with time at full excavation
depth; (some struts loads decrease with time others increased).
2.
steel.
The vibrating wire in the Telemac gauges in of tempered
Tests by Browne and McCurrich, 1967, have shown creep in this
type of tensioned wire is negligible for the temperature range experienced during the measurements period.
-241-
Io regards to other factors affecting drift, Figures
A.6.1 through A.6.8 show no definite large change in strut load
during construction which could be attributed to gauge shock or
damage to the gauge supports.
Further, upon completion of the
strut load measurements the stain gauges were checked and found
to be functioning properly.
Therefore,
the possible reasons for
the residual loads appear limited to loosening of the support
nuts, bending of the gauge supports, and gauge shock or racking
of the strut when removed from the bracing system.
Some struts
were observed to be distorted after removal from the excavation
prior to the recording of the final no-load readings.
However,
there was no correlation between this factor and the amount of
gauge drift.
If is not possible to determine which of the above (or
combinations thereof) resulted in a shift in No or when or how an
estimate of a correction for drift should be applied.
Because of
the uncertainty asto the cause and time of the shift in No,
the
reported strut loads in this text were not corrected for drift.
The
error associated with this procedure is discussed in Section B.2.7.
B.2.6
Accuracy Limits of Strut Loads
The overall accuracy of the measured strut loads can be
evaluated by performing a theoretical error analysis (Rabinowicz,
1970).
One can express the total error in the strut load by the
following general equation:
-242-
2
(an)
2
B. 2. 4
n
where:
F
=
equation describing a corrected strut load
n
=
a term in the equation (F)
en
=
the error in the term n
e
=
total error in the strut load
The general equation for the corrected strut load is:
Ps = AE X [(No2 - N 1 2 )- Kt (To - Tl)] - Kd
B.2.5
where Kd is the correction for gauge drift.
Undoubtly drift has caused some error in the measured
strut loads since essentially all the struts showed the expected
trend of a residual compressive load when removed at the completion
of the measurements program.
However, as explained in Section B.2.6,
it is not possible to assign an exact value of Kd to each strut's
monitored load.
The limiting error in each strut load is its measured
residual strut load.
reasonable
However, some struts indicate this is not a
correction.
For example strut D-48 had a residual load
of 122 kips and a maximum measured strut load during construction of
190 kips.
In addition, there is no definite trend between measured
strut load and the amount of residual load in the strut.
For pur-
poses of assigning some error to the strut loads, Kd is assumed equal
to 20
20 kips.
These values correspond to one-half the average of
the residual strut loads, and one-half their standard deviations.
-243-
It
-I
is emphasized that this is an intuitive assumption based on a
review of the data and has no formal origin.
e2
=2
A
E2X2 (N 2 - N 2)2 + e 2 A 2 (N 2- N 2)2
0
1
Ex
s 0
1
+
The equation describing the error in a strut load for
factors other than drift is:
2
2
2
2
2
e fA
2As2E2X2
2
EX + ett As
s (T 0o1-T 1 )2 + ettKT A SB26
B.2.6
where
eA = error in the cross sectional area of the strut
eEX= error in the product EX
ef = error in the monitor value of No and Ni
et = error in temperature correction
ett= error in recorded temperature
the following values may be assigned to the error terms:
eA = .025 As
eEX= .01 EX
e
= 1 cps
et = .02 ksi/oc
ett= 20 C
The range of error in the strut loads can be estimated using
the following range of measured values and correction factors for the
measured strut loads which ranged from 110 kips to 430 kips.
(N0 2 - N 2)
N
= 45.8 x 10+3 to 122.5 x 103
= 490 cps to 560 cps
2
in 2
As = 25.56 in to 51.73
-244-
(T
-
T0 ) = 0 - 12*C
KT= .088 ksi/oc
EX = 94.2 x 10- 6 ksi
The estimated error by equation B.2.6 was
kips for the small struts and
20 to
8
25 kips for the large struts.
This results in an error approximately
strut
5 to
8 percent in measured
loads.
The overall accuracy of the measured strut loads can be
estimated by the following equation:
Ps
m
where
+
0
)
( .08)
P
m is the measured strut load corrected for tempera-
ture inertia effects.
B.2.7
Method of Gauge Installation
On struts B-44 to B-47 (see Figure A.4.6) six gauges on
the neutral axis of a strut gave reliable results for strut loads.
The gauges, in pairs, were mounted at the neutral axis and 4 inches
above and below it.
Figures B.2.12.
A typical set of the measurements is given in
The comparison of loads are shown at the bottom of
the plot for several arbitrary dates.
The gauges, if everything is
properly mounted and correctly functioning, should show a straight
line variation of load from the top to bottom gauge.
The error inherent in using just two gauges on a strut can
be estimated by comparing the average strut load from the two neutral
axis gauges with the average of the six gauges.
-245-
Another check would
be to compare results of the neutral axis gauges with the results of these comparisons which show that the use of two gauges
at the strut web centerline leads to only a minor error in the
measured strut loads.
This error is, in general, less than 10
percent, which is within the measurement accuracy that can be
expected from the F-2 gauges used at South Cove.
B.3
SETTLEMENT AND HEAVE MEASUREMENTS
The individual readings were obtained by a skilled
survey crew using a Ziess level and a Philadelphia Rod.
The
vertical movements recorded were referenced to a permanent bench
mark on the site.
in
.003 ft.
Closures of a survey level loop were kept with-
In addition to the closure error, the accuracy of
individual reading have inherent error resulting from estimating
a 0.001' on the level rod.
However, if one accepts the estimate
is accurate to
.002', then the accuracy of an individual reading
is better than
.005 with the probable average being
.003' be-
cause of compensating errors within the level loop.
B.4
PIEZOMETERS
B.4.1
Hydraulic Piezometer Accuracy
The maximum permissable rate of change in measured
pore pressures can be estimated by the following equation (Penman,
1961; Kallestenios, T. and Wallgren, A., 1956):
du = Aut
dt
2k
B.4.1
-246-
-.4
where:
the maximum permissable rate of change
dt
of measured pore pressure (psf)
Au
=
the permissable error for the calculated
)
piezometer level (ft 2
k
A
=
permeability of the soil (ft/day)
=
the surface area of the piezometers (ft 2
=
the volume of water required to move in
F
V
or out of the piezometer top in order to
cause a change in water pressure equal to
1 psf (ft 5/lb)
The detailed dimensions for the hydraulic piezometers are
given Section A.5.3.
From these dimensions the following variables
are appropriate:
A
= 0.53 ft 2
V
= 1.65 x 10
p
-5
ft
3
The range of k values were assumed equal to 2 x 10-8 cm/sec
to 6 x 10-8 cm/sec.; the permissable error (Aut) was taken as 0.5 ft.
Based on the above, the allowable rate of change of pressure
head ranged from 0.03 ft/day to 0.1 ft/day.
B.4.2
Electrical Piezometers
The pore pressure from the Geonor vibrating wire piezometers
(Model M-206) are determined from the following equation:
U = M (No -_NJ 2)
7-247-
A
Where:
u
=
pore pressure
M
=
calibration constant.
N
=
zero pore pressure frequency (cps).
N
=
frequency at the measured pore pressure
(psf).
(cps).
The piezometers were calibrated in a pressure chamber
to evaluate the constant, M, which varied considerably between
individual piezometers.
The individual constants for each piezo-
meter checked within 0.5% of that specified by the manufacturer.
Values of M from the pressure chamber tests ranged from 0.00489
to 0.00614.
The piezometers were also calibrated in the field by
lowering them in a 4 inch casing which was driven in the ground
to a depth of 100 ft.
This field calibration showed the piezo-
meters to, be. temperature sensitive.
the N
A change of up to 40 cps in
value occurred when the piezometer was first lowered into
a 4 inch casing.
To account for this temperature effect, the tem-
perature of the piezometers was allowed to equalize with the water
temperature within the casing, at increments of 10 feet of depth
for the 100 ft. of casing to obtain a correlation between frequency
and pressure head of waters.
This date was extrapolated to zero
pressure head to obtain a new N
for each piezometers.
The field
data indicate that this extrapolation could result in an error of
2 ft. to 7 ft. in the determination of initial pressure head at the
-248-
-.
4
piezometer tip elevation.
With regards to the changes in pore pressure occurring
during the construction periodthe temperature sensitivity of the
gauges has no effect on the results because of the essentially
constant temperature of the groundwater.
B.5
SLOPE INDICATOR ACCURACY
Horizontal ground movements were monitored with a Wilson
"Torpedo" which measures the tilt of the Slope Indicator Casing
(see Figure A.5.9) by an encased accelerometer.
The relative
horizontal movement (A) between two points along the casing is
determined by:
n
2
2
E K (N 2- N 2) L
1 n n
n=1 n1 o
A=
Where:
Ln = incremental length over which the torpedo
measures tilt.
K
= calibration constant for the accelerometer
N
= initial frequency of the accelerometer at
given point when the casing was installed
N
= measured frequency at the same point as N0
The factors affecting a given set of readings along a
casing are:
1.
Physical characteristics of the device.
2.
Environmental influence on the casing grooves which guide
the torpedo (roughness of grooves).
-249-
3.
Installation of the casing (denting of casing
during installation).
4.
Observer error in positioning the device.
5.
Data reduction procedures.
A comprehensive discussion of these effects on measured
deformations is given by Gould and Dunnicliff (1971).
They.!
summerize the accuracies of Slope Indicator measurements reported
in several references.
The reported accuracies ranged from 0.001
to 0.003 radians per 100 ft. of casing length.
Bromwell et al
(1971) reports an accuracy of 0.1% in measured deformations in
controlled laboratory conditions.
A review of the first three items effecting accuracy
suggests the most desirable approach to determining the overall
accuracy of Slope Indicator measurements is to perform in place
field tests.
for SI-1
Prior to any construction, Slope Indicator casing
(See Figure A.4.6) was installed and repeated measure-
ments of the casing slope were made.
The measured maximum difference
in movements between any two sets of readings is considered the
composite error in the measurements.
Figure B.6.1 shows the re-
sults of these measurements in addition to some measured movements
of the Slope Indicator casings which were installed in the concrete
wall.
These
of this wall.
measurements are the net movements parallel to length
It is assumed the wall did not move in this direction
and that these results are also an indication of the overall accuracy
-250-
I4
of the Slope Indicator measurements.
The data in Figure B.6.1 suggests that the accuracy of
the Slope Indicator measurements are better than
.05% of the
length of casing over which the measurements were taken.
The accuracy of the total movement measured by the Slope
Indicators is dependent on knowing
the exact location of some
point on the casing since the Torpedo gives only relative movements
between two points of the casing.
was the top of the casing.
The point on the casing chosen
A survey line was set up which was
essentially parallel to the wall and swing offsets were measured
to the casing top for each set of readings taken.
that these readings were accurate to within
-251-
.02'.
It is estimated
TABLE B.2.1:
VARIATION IN STRUT TEMPERATURE ALONG INSTRUMENTED
SECTION AS MEASURED BY STRAIN GAUGE THERMOCOUPLE
Time of Day
Readings
Recorded
Average Temperature *C
of Strut Section
Date
Weather
Start
End
Fair
20
21
-
0830
-
Sections
48 & 49
9-15-69
10-10-69
Cloudy
Fair
30
19
31
21
31
25
1300
1140
10-29-69
11- 6-69
1400
1505
Fair
Rain
17
19
9
11-24-69
12-10-69
12-30-69
14
9
1000
0715
1400
0900
Fair
Cloudy
Fair
15
7.5
6
13
9
5
1100
1400
1500
1225
1600
1- 8-69
Fair
-2
-1
2- 2-69
2-20-69
3- 5-69
-2
Rain
Fair
Rain
1200
1300
.6
8
1
6
12
2
6
12
2
100
1300
0820
1115
-
6
7
4
3-20-69
4-10-69
Fair
Fair
17
17
20
14
20
16
1100
1100
1200
1200
4-24-69
Rain
9
10
10
0945
1050
5- 8-69
6-30-69
Cloudy
Rain
13
22
13
23
15
23
0815
1100
0910
1220
1020
7-24-69
Fair
36
37
40
1130
1230
8-17-69
Cloudy
28
9-14-69
29
30
Rain
1105
-
tx,)
I,
Sections
40 & 41
-
8-27-69
Sections
38 & 39
15
18
17
1000
1100
10- 2-69
Fair
19
19
20
Notes:
1) Maximum temperature variation across a vertical section is
being 1"C.
5*C, the average
2) Maximum temperature variation between north and south side of strut
average being <1*C.
30 C, the
-1
TABLE B.2.2:
TEMPERATURE INERTIA OF STRAIN GAUGES
Telemac F-2
2
Af /oc
Mi
Max
690
1165
K
T
Avg
Avg
935
-87.5
1100
-103.5
(Psi)
0c
Remarks
Range
22.5
From laboratory tests
Theoretical value
recommended by
Telemac Ltd.
Telemac SB-90
740
1714
735
46.0
18
Based on laboratory
tests
Value selected represents author's interpretation of data.
Algerbraic sign corresponds to increasing temperature.
Load
corr
= LoadI
meas
[K AAT]
T
-253-
I
TABLE B.2.3.: COMPARISON OF MEASURED STRUT LOAD FROM
JACK PRELOAD AND STRAIN GAUGES
C)
-)
--
-
I,
U,
-Is
-.
C-3
-3
-
4
4
B-41
B-42
B-43
B-44
B-45
B-46
470
485
B-47
B-48
B-49
320
370
290
-
-
C-31
C-0
C-32
-
0-4
175
230
150
-
4d H9I
D-31
D-32
D-34
D-36
D-38
D-40
590
551
503
D-41
D-42
D-43
D-44
D-45
D-46
530
660
550
535
570
615
386
457
617
360
447
441
541
D-47
D-48
D-49
510
380
235
481
188
135
D-2 505
-4
-oC-8
-
B-31
B-40
B-32
B-34
B-36
B-38
B-40
540
-D-
4
4
4
2
3
-S D-3
43398
C-36
C -38
C-40
-
3
44 38
-
2
-
-
0-44
c-45
C-46
230
200
340
280
310
330
3.30
186
275
370
248
270
310
267 (F-2)
c-47
C-48
C-49
330
270
225
160
276 (SB-90)
250
178
102
494
448
C-41.
C-42
-
-
C-43
450
380
380
372
321
290
278
329
267
Avg. Diff. = 11% 6%
138
186p-4p0
139
Avg. Diff. =13% 5%
Overall Avg. 19% 6%
550
540
430
440
-
-n-
B-3
-3
-d
cd- - 41
.4 C-32
4LZ
co0
B-32o"- Z- (1z1o)
39c
B-3
0
0
0.4
Br-
41A
-
a)
409
398
Avg. Diff. =27% 5%
I
TABLE
B.2.4: B-LEVEL INITIAL AND FINAL ZERO LOAD STRAIN GAUGE FREQUENCIES
Average
Strut
No.
B-31
Strut
Size
14WF184
Strain
Gauge No.
B-31-Nl
B-31-SI
B-32-NI
B-32-S1
B-32
14WF184
B-34
14WF184
B-34-NI
B-34-SI
B-36
14WF184
B-36-NI
B-36-Sl
B-36-N 2
B-38
B-36-S2
B-38-NI
B-38-SI
B-40
B-38-S2
B-40-NI
B-41
14WF111
B-40-N 2
B-40-S
B-41-N
B-42
14WFll
B-43
14WF111
B-38-N2
B-40-SI2
B-41-S
B-44
14WF111
B-42-N
B-42-S
B-43-N
B-43-S
B-44-NT
B-44-N
B-44-NB
B-44-ST
8-44-S
B-45
14WF127
B-44-SB
B-45-NT
B-45-N
B-46
14WF136
B-45-NB
B-45-ST
B-45-S
B-45-SB
B-46-NT
B-46-N
B-46-NB
B-46-ST
B-46-S
B-46-SB
B-47
14WF136
B-47-NT
B-47-N
B-47-NB
B-47-ST
B-47-S
B-47-SB
B-48
14WF103
B-48-N
B-48-S
B-49
14WF103
B-49-N
B-49-S
Initial
Freq. f
(cps) "
Final
Freq.
Te, C
f (
882
707
838
775
850
848
754
835
805
808
882
781
823
834
842
858
816
835
587
558
612
609
594
544
574
626
590
603
596
585
614
610
595
596
595
592
609
602
591
625
618
599
606
615
588
620
628
619
594
589
592
599
Gauge
Air
17.0
16.9
16.8
17.0
18.5
17.7
17.8
17.5
16.0
17.8
20.0
19.0
20.0
19.0
20.0
20.0
41.1
41.1
22.1
19.9
39.1
39.0
21.9
22.8
23.2
22.9
21.0
21.8
21.6
21.1
20.2
22.9
20.0
19.6
19.6
20.0
f
(cps) Freq.
3C
3
768
884
816
15.3
15.0
15.9
15.0
16.0
8.3
27.7
27.7
874
811
8.3
15.0
15.0
226
329
234
285
15.0
879
827
745
798
791
927
875
812
825
807
8.3
5
Residual
Load
Initial
SB-90
Load
GaugeAr
____
16.0
Tem
at
A (kips) (kips) PeiCq)
Residual4
Load
18.5
24.2
17.5
23.5
20.0
17.5
15.2
18.0
13.5
13.4
15.0
15.2
9.9
9.9
10.5
299
-196
741
120
200
235
135
136
113
519
-640
310
262
381
401
537
360
377
404
413
461
160
187
397
531
123
402
549
54
10.5
548
14.0
27.7
27.7
20.0
20.0
22.1
18
18
18
34
10.5
10.5
588
632
595
613
601
574
618
593
583
622
616
595
624
616
591
602
591
578
19.5
18.0
14.0
11.1
13.1
15.1
17.0
14.3
31.3
24.0
30.1
29.5
36.1
26.9
609
611
583
625
630
617
587
601
28.5
27.0
27.3
28.4
27.5
27.6
28.0
29.0
18
19.2
19.2
17.5
10.5
10.5
19.9
19.9
19.9
-48
-22
-19
-35
-17
39
-15
71
49
-101
-81
-10
-67
-60
0
105
123
92
-13
19
22
-21
-11
10
23
-38
+27
+7
-5
-36
-34
19.9
1Telemac
type SB-90 strain gauges installed after strut in place to replace incorrectly installed Type F-2
strain gauges.
2
Incorrectly
installed Type F-2 strain gauges.
3
Temperature from U. S. Weather Bureau climatology data.
.
4Residual load in strut determined from f
5
Average residual strut load corrected for temperature effects.
6Temperature corrected load as recorded by F-2 gauges when SB-90 gauges installed.
-255-
6
A
TABLE
B.
. : C-LEVEL INITIAL AND FINAL ZERO LOAD STRAIN GAUGE FREQUENCIES
4
InitialTemp.
Strut
No.
.n'
Strut
Size
C-31
14WF78
C-32
14WF78
C-34
14WF78
C-36
14WF78
C-38
14WF78
C-40
14WF84
C-41
14WF87
C-42
14WF87
C-43
14WF87
C-44
14WF103
C-45
14WF103
C-46
14WFll
C-47
14WF95
C-48
14WF87
C-49
14WF87
lAll
2
Strain
Guage No.
Freq. f
(cps)
C-31-N
C-31-S
C-32-N
C-32-S
C-34-N
C-34-S
C-36-N
C-36-S
C-38-N
C-38-S
C-40-N
C-40-S
C-41-N
C-41-S
C-42-N
C-42-S
C-43-N
C-43-S
C-44-N
C-44-S
C-45-N
C-45-S
C-46-NI
C-46-Sl
C-46-N
C-46-S
C-47-N
C-47-S
C-48-N
C-48-S
C-49-N
611
675
567
592
550
587
598
624
582
608
548
619
587
582
576
586
544
627
712
671
654
653
662
629
727
742
C-49-S
670
f pC
0
Gauge Air2
21.0
21.0
21.0
21.0
23.9
23.9
23.0
20.5
Freq. ff
(cps)
24.2
19.8
19.8
19.8
15
16
16
16
15.5
15.5
15.5
15.5
15.5
15.5
17
12
650
684
630
673
12
12
___
Residual
Load
_00(kips)
Gauge
Air
612
575
572
592
547
588
538
650
609
549
545
620
587
582
575
582
545
627
701
636
641
639
660
617
742
15.1
15.4
14.3
14.4
11.3
10.3
0
-0.9
0
0
6.5
6.4
7.1
5.6
3.2
3.9
2.4
2.3
0
0
0
0
-5.4
-5.0
-6.0
13.3
647
670
692
637
626
-7.0
-6.3
-5.0
-5.3
-5.0
-6.1
648
-4.9
13.3
9.4
-2
-2
7.2
7.8
0.5
0.5
-1.6
-1.6
-6.1
-6.1
-6.1
5
0
-13
0
6
-2
147
-65
-67
147
8
-3
0
0
0
8
10
0
43
131
49
50
9
47
-111
-65
-24
-19
147
3
Average
Residual
Load
(kips)
14
6
24
87
72
22
20
26
27
115
82
73
19
128
67
strain gauges are Telemac type F-2 except C-46-N (SB) and C-46-N (SB), which are Telemac Type SB-90.
Temperatures are from U. S. Weather Bureau climatology data.
3
Residual Load in strut determined
from ff.
4
at
f mC)
Average residual strut load corrected for temperature effects.
A
I -TABLE
Strut
No.
Strut
Size
B.2.6:
Strain
Gauge No.
D-LEVEL INITIAL AND FINAL ZERO LOAD STRAIN GAUGE FREQUENCIES
Initial
Freq. f0
(cps)
Temp. at
f (0C
.
I...,
U,
14WF142
D-32
14WF142
D-34
14WF142
D-36
14WF142
D-38
14WF142
D-40
14WF150
D-41
14WF158
D-42
14WF158
D-43
14WF167
D-44
14WF176
D-45
14WF176
D-46
14WF176
D-31-N
D-31-S
D-32-N
D-32-S
D-34-N
D-34-S
D-36-N
D-36-S
D-38-N
D-38-S
D-40-N
D-40-S
D-41-N
D-41-S
D-42-N
D-42-S
D-43-N
D-43-S
D-44-N
D-44-S
D-45-N
D-45-S
D-46-N
D-46-S
D-47
14WF158
D-48
14WF142
D-49
14WF142
D-47-N
D-47-S
D-48-N
D-48-S
D-49-N
D-49-S
580
602
617
589
606
613
571
606
620
602
771
795
746
811
784
775
565
667
615
628
677
657
659
612
618
649
642
650
b41
605
Air2
16.5
16.5
16.5
18.0
18.0
0
0
10.0
15.6
8.4
11.0
10.0
9.0
9.0
9.0
9.5
1
Freq.
(cps)
Gauge
D-31
Final
1
577
602
621
587
603
592
584
594
632
616
754
789
740
800
787
768
566
656
611
620
676
658
651
610
624
651
626
648
628
615
1
All strain gauges Telemac type F-2.
2
Temperature from U. S. Weather Bureau climatology data.
3
4
Residual load in struts determined from ff.
Average residual strut
load
corrected for temperature effects.
f
Temp.
T (R
f
at
Gauge
Air
17.2
16.5
16.1
11.1
11.0
11.0
1.5
4.1
1.1
1.6
12.8
12.8
9.4
-2.2
-2.2
7.7
7.7
2.7
-0.7
-0.6
0
0.9
0
0
-6.8
-6.9
-5.1
-6.0
-5.0
-4.9
-9.0
-9.1
-1.7
-1.7
-1.7
-6.1
-6.1
-6.1
-6.1
_
Residua13
Load
(kips)
14
0
-20
9
14
99
-37
+58
-59
-68
107
38
37
75
-18
48
-6
68
24
49
6
-6
51
13
-32
-9
79
88
65
-44
Average 4
Residual
Load
(kips)
5
7
68
64
8
35
35
72
65
80
45
99
42
122
75
_
_
_
-1-1 - -. , -
I
"I
j
TABLE B. 2.7:
STRUT
DATE
Preload
2 Oct 69
1 Nov 69
1 Dec 69
5 Jan 70
2 Feb 70
2 Mar 70
AVERAGE STRUT LOADS
44
1/2 EP
1/6 EP
190
199
215
187
250
320
250
45
+ P
B
T
T
192
199
213
196
263
270
260
% Diff.
1.1
0.0
1.0
4.6
4.9
18.5
3.8
1/2 EP
1/6 EP
220
255
260
245
340
295
290
+ P+
B
P
T
217
235
237
238
318
297
273
% Diff.
_
_
1.4
8.5
9.7
2.9
6.9
0.7
6.2
tU,
DATE
48
46
STRUT
1/2 EP
1/6 EP
+ PB +
T
% Diff
1/2 EP
1/6 EP
+ PB + PT
% Diff.
Preload
253
261
3.1
250
227
10.1
2
1
1
5
69
69
69
69
220
282
217
279
1.4
1.0
315
305
3.3
176
212
193
252
160
207
183
241
10.0
2.4
5.4
4.5
2 Feb 70
360
350
2.8
290
290
0.0
2 Mar 70
3 April 70
320
355
312
345
2.6
2.9
252
288
246
281
2.4
2.5
Oct
Nov
Dec
Jan
50
I
1-I
7
Time of I Air Temp. 0 C
18/23 1 8/27
day
SYM.
08:40
5
Ad
5
2
3
40
4
6
5
6
7
8
9
10
II
-3
6
2
cr
30
(LJ
6
10: 30
12:00
13: 45
15:00
16:00
18:00
20:00
23-00
03:00
07:00
24.4
20.0
26.0
18.8
25.5
17.2
17.2
17.2
18.9
16.1
18.2
.2
-
5
w
I
o GAUGES UNCOVERED
-
Il4~ Icr%.II
ao
wo
2__
.
8
20
o
(8/23/69)
0~
I I0
A 0 GAUGES SHADED
FROM SUN(8/27/69)
__8
-
1-
STRUT UNLOADED AND
IN BRACING POSITION
620
630
590
600
FREQUENCY
FIGURE B.2.1: FREQUENCY
B-44-N
a
I
585
64C
VARIATION
610
(CPS)
VERSUS TEMPERATURE
615
6B-44-S
m
SYM.
A
U
0
-,
SCALE
I
I
2
2
2
GAUGE
B-41-N
B-41-S
B- 42-N
B-44-N
B-44 - S
SYM.
+
SCALE.
2
2
GAUGE
B-45-N
B-45-S
READINGS
FOR A 24 HOUR
PERIOD 8:00,8/27/69 to8:00 8/28/69
~40
I
STRUTS
INSTALLED
BUT
UNLOADED
wL
2
"r~
Lu
I(10:00 hrs.)
z
20.
.
'II
zi
U,
oi0
560
I
I
I
N
I
600
620
500
520
540
I4u
a
I
580
'
0
6 40
I
560
FREQUENCY (CPS)
FIGURE B.2.2: STRAIN GAUGE FREQUENCY
SHADED STRAIN GAUGES
VERSUS
TEMPERATURE
FOR
I
I
40
w
w
l'-
30
20
I
I
SY m.
GAUGE
0
K>
B -31-N
B -31- S
B-32-N
A
A
I
B- 32-S
B- 34-N
B- 34- S
B-36- N
2
t
4-
0
B-36-N
B-36-S
B- 36- S
0
4 'o
000--
S+
A
A
1-
0
ov--
A
0
cy%
0
G
MEASUREMENTS
ON
STRUTS IN BRACING
GAUGES
101
8 00
I.
UNLOADED
POSITION
NOT COVERED
I
____________________________
820
810
FREQUENCY
FIGURE B.2.3: RANDOM
A
READINGS
OF
830
840
(CPS)
GAUGE
FREQUENCY AND
TEMPERATURE
R C F -
0C
ATHER
F CC F C
RAIN
CLOUDY
FAIR ,SUNNY
C F F
FF
F
RR
F FF C F C
F R F
F
F
FRR
F
FFF
R F F
F
F F
R F C F F F C
C F
F
C F
C F F F F
R
C F F
R
F
3530.
25w
20.
0
404
15.
0oI
r-j
(ON
5.
I-
10
0
-5.
20-
-10
-
0EMERA
I5
0
Scale lDops
-20
AUGUST
SEPTEMBER
FIGURE
16FEAUS
OCTOBER
NOVEMBER
B.24:TEMPERATURE
2E0N 3D
TEMPERATURE AT SITE
AILYAIR
DECEMBER
RECORDED
JANUARY
BY U.S.
WEATHER
EBRUARY
WEATHER
MAR04
APRIL
BUREAU,
LOGAN
MAY
BUREAI
JUNE
INTERNATIONAL
TEMPEMATURE
JULY
AIRPORT,
AT
TIME
9F
AUGUST
BOSTON
THERMOCOJPE
SEPTER
MASS.
READING
OCTOBER
k
8001
a-
z
w
7001
D
7,O
w
1
i:
Ooo
r
LL
w
6001
z
5001
W:
4001
.0 r
.08
.09
.10
.11
SENSITIVITY
FIGURE B.2.5: TYPE
F-2
STRAIN GAUGE
.12
(KIP/
.13
IN 2
.14
.15
/CYCLE
)
I'
SENSITIVITY (KIPS/IN 2/CYCL.E)
'I
I
-
60l
o M.I.T. DATA
II
A TELEMAC
I
'
Ia- //
-A-
40
-t
I
/
/
50
DATA
I
4.
-
4-
'I
-AGAUGES
I
OVEN
(fo = 6 22 @ 22.80 C)
30
a
a
.. 20
101x
07
CAUJES
IN
REFRIGERATOR
Oa
(f o =659 @ o't
-I0
0
GAUGES
IN
-
FREEZER
20
30
610
___1~~
620
630
640
650
FREQUENCY
FIGURE B.2.6: TEMPERATURE
F-2
STRAIN
660
(CPS)
INERTIA
GAUGES
-264-
670
680
OF TELEMAC
-_
70
----
OM.I.T.
60
OF
L!TELEMAC DATA
RANGE
PROBABLE
/I
DATA
READINGS
/ AF
I
(fo=722@22.8*)
40
GAGES
3EN
I
ii
/
LL
IN
Ii
I20
RANGE
OF
READINGS
10
-
GAUGES
IN
REFRIGERATOR
0.
0
G
( fo =832 @ 265 *C
)
0
-7
GAUGES
IN
F R E E ZER
7)
7.O
/
6 (fo
'd
-20
NOTE:
-
30L
700
--
INITIAL FREQUENCY AT
ROOM
TEMPERATURE WAS CHANGED FOR
BOTH
___________
_ I
III
/0
M.I.T. TESTS
L
h
720
740
760
780
FREQUENCY
FI GURE 8.2.7: TEMPERATURE
TYPE
SB-90
800
CPS
820
INERTIA OF TELEMAC
STRAIN GAUGES
-265-
840
J
r
r
________________
I
I
S
)a
t
JI
-
I.
500
-
CL)
y~
4 I.
*
0
0
0
I
0
S
0
0
400
________
4.
4
4
I
V.
I
300
I
________I
__________
*S.
0
w
..4
0
0
0.
t
0
0
I
.*
I
III.
200
0
10DID
w
e
-
loo
I
I.
+
I
Ip
001
C
.
0
100
_
200
3200
THEORETICAL
400
JACK
50
9-0
LOAD (KIPS)
7 00
z
FIGURE B.2.8: COMPARISON OF JACK LOAD AND STRAIN GAUGE LOADS
0
A (1
+80
,6(102
+-70
0
D
+6C
I
*~
0
-J
Ns
w
w
2
)
0
0
b
D
I -
+50
z
+
3
o0~-0--
-12'
I"
N
-'4'0
0
%mo
I
00
GOa
-.01
0
0
0
11
0
-14
z
w
u
w
0.
_______0
0
-2
-30D'
0
0
I
100
&
I
200
STRAIN
300
GAUGE
400
K
J
500
A
600
700
LOAD (KIPS)
FIGURE B.2.9: PERCENT DIFFERENCE BETWEEN
AND STRAIN GAUGE LOAD
JACK
PRELOAD
-1
I
300
0
020C
a)
-0
ID
C')
w
00i
0
(U)
w
0
-J
0
100
LOAD MEASURED
FIGURE B.2.1O:COMPARISON
TYPE
200
300
BY F-2 GAUGE (Kips)
0 F STRUT LOADS MEASURED BY
SB-90 AND F-2 STRAIN GAUGES
i
OVER ALL AVG.= lO
KIPS
B40)
AVG= 244 FOR SB-90(B31-B40)
o 285
o->~262
AVG = I12 FOR F-2 GAUGES
I1
AVG=-8 FOR F-2 (B41- B49)
R
LEVEL
(B31 -
z
0
AVG=55 KIPS
C
LEVEL
AVG= 54 KIPS
cr.
ILOAD
D LEVEL
I
-100
-50
0
RESIDUAL
50
LOAD
IN
ARE CORRECTED
FOR TEMPERATURE EFFECTS
SB-90 GAJGES STRUTS 83-840
0
F- 2 GAUGES STRUTS 831-840
LLECORRECTLY
F-2
I
I
100
150
STRUT
FIGURE B.2.11: RESIDUAL LOADS IN STRUTS
STRAIN GAUGE READINGS
E
-
w
-j
w
200
(KIPS)
AFTER
REMOVAL
BASED
ON
I5TRUT
B
Top
+
3wo
LEVEL
"
0
100
Q
.L.
44
S
A
-t
0
+
i
3
0
AIi/.iI~T
I
~ qrDTPt.1nv~
~ ~ ...
.... or aF
.
LOAD 0
/00
9
I
rr M~V~P4t~4fR
Ara Af .-...- j...*
.
t2Fe~t~1RFR
ilr
JANLIAR Y'
ma-
-
ZARLiARf
h
-
KIPS
3. '0 500
STRUT
six
ll~
DISTRIBUTION
TOP GAUG E
D vsiQ-N
W.~~~~ F'- 11
LOAD
.#
,
AL
+So.+
lW )
44
-o
OK
....-.
.- ..- --.....
-
+
---.
-. 1
0
-GAUCFE
;AUrb
F 4' Fill
FIGURE: 8.2.32 TYPICAL MULTIPLE GUAGE STRUT LOAD MEASUREMENTS
C
4a
randi'ng
9E4,
radin
MARCi.
..
A
PA
SLOPE
SLOPE
INDICATOR MOVEMENT x 100
LENGTH
INDICATOR
(+ NORTH)
(- SOUTH)
.2
.1
.I
0
.2
-
0-I-
* II
0
SI
/
z
i
*I'
* II
a-
~
-. NORTH
:11
z .5-4w
a-
''I
I
1.0 .L
SYM.
REMARKS
Slope Indicator
ST- I
DIFFERENCE OF TWO SETS
READING AT SAME TIME
Sr -
RANGE
TO THE
4
Sr-5
SL-7
FIGURE B.6.1: ACCURACY
OF MOVEMENTS
SLURRY WALL
a
OF SLOPE
MEASUREMENTS
-2 71-
OF
PARRALEL
"
0
INDICATOR
-*1
APPENDIX C
ANALYSIS OF SLURRY FILLED EXCAVATIONS
C.1
INTRODUCTION
The behavior of vertical excavations which are supported by
a bentonite slurry is, at present, a poorly understood subject.
This
Appendix presents the results of some analytical studies which were
conducted to gain insight into the behavior of slurry filled excavations of limited length in cohesive soil.
Basically, the studies were
aimed at obtaining some understanding of the phenomena of stresses
arching around the excavations and the stability of the trench.
C.2
METHOD OF ANALYSIS
Currently there are no theoretically exact methods available to
analyze the behavior of a rectangular trench of limited depth in bilinearly elastic-plastic material. Meyerhof (1972) proposed analyzing
this case using existing theories for linearly elastic materials to
determine deformations and modified stability analysis techniques for
cylindrical cuts to determine the factor of safety against a trench
failure.
Analytical studies, along the line suggested by Meyerhof, were
made using a finite element program Feast-3 (D'Appolonia, 1968).
Briefly, the program analyses plane strain or axisymmetric problems
and considers materials with bilinearly elastic stress-strain relationships and anisotropic strength properties.
geometry is applied in increments.
The loading on any problem
Within each increment, the stress-
strain relation is linearly elastic, but o'ver the whole range of loading
-272-
it
C.3
is not.
ANALYTICAL STUDIES
C.3.1
General
One
The analytical studies were divided into two parts.
set of analyses examined the movements and stability of a rectangular
slot in a bilinearly elastic material.
Plane strain conditions
were assumed parallel to the axis of the slot.
The intent of these
studies was to gain insight into the soil behavior adjacent to a rectangular slurry trench at a given depth in the soil.
The second set
of analyses studied the stability of circular and infinitely long
slurry filled excavations in bilinearly elastic material.
The pur-
pose of these studies was to shed some light on when the length to
depth ratio of a rectangular trench was sufficiently large so that
the trench may be analyzed as a plane strain problem.
It is recognized that neither of the aforemention analyses
exactly simulates the field conditions.
For example, the first set
of analyses assumes the intermediate principle stress is parallel to
the axis of the excavation and is constant.
In the field condition,
this stress, which is the vertical stress, varies in magnitude and
direction with depth.
Nonetheless, it is felt that these studies will
give some insight into the behavior of slurry filled excavations.
C.3.2
Analysis of a Rectangular Slot
Figure C.3.1 shows the finite element grid used to analyze
the case of the rectangular slot.
Only one quarter of the overall
geometry was analyzed since the problem is symetrical around the
Y axis.
X
and
The width of the slot was varied to model excavations of varying
-273-
-1,
lengths (Y) to width (X) ratios.
The sides A-B and D-E are re-
flective boundries and were restrained from moving in the Y and X
The sides B-C and C-D were unrestrained
directions respectively.
against movements in either the X and Y direction.
The stress re-
lief (AaL) at face of the slot was taken as the net difference
between the total horizontal pressure in the ground before excavation
of a slurry trench and the fluid pressure of the slurry at a given
depth.
The net stresses in the surrounding soil were set equal to
zero.
Only the case of trenches in normally consolidated clays
were analyzed.
The soil parameters employed in the analyses were
taken from the plane-strain-passive triaxial test results on resedimented Boston Blue Clay reported by Ladd et al, (1971).
The
parameters chosen were as follows:
E
=
150 ovo
Suh
=
0.19 jvo
After yield EH
=
0.15
vo
Figure C.3.2 shows the results of the finite element analyses.
Three cases were analyzed, each with the same soil conditions and slurry
unit weight but with varying dimensions of L.
The cases represent pro-
bably the worst conditions for a slurry trench excavation in clay.
Slurry unit weights will be in general
higher than 62 pcf (68-72 pcf)
and the shear strength of the clay used in analyses was
minimal.
The results show that as the length of the excavation in-274-
creased the movements were a smaller percentage of the slot length.
The dimensionless area of the yielded zone also decreased with increasing slot length.
None of the cases show any evidence of a
gross inward failure of the slot walls.
The limited size of the
yield zones also suggests substantial arching of the stresses
around the slot.
C.3.3
Analysis of Slurry Filled Excavations
Figure C.3.3 shows the finite element grid use for the
analysis of circular and long slurry filled excavations.
The dimen-
sion of the trench was varied to study the effect of the depth to
The stress relief
diameter ratio on the stability of the trench.
along the face of the excavation was set equal to the net difference
between the total horizontal pressure in the ground before excavation
of the trench and the fluid pressure of the slurry at a given depth.
Excavations in both normally consolidated (OCR = 1) and
overconsolidated clays (OCR = 2) were analyzed.
Anisotropic strength
parameters for the clay were taken from Ladd et al, (1971).
Figure C.3.4 shows the predicted
movements adjacent to
slurry filled circular excavations in normally consolidated clay for
several different diameters.
The slurry unit weight for the analysis
was 72 pcf which corresponds to the unit weight for a factor of
safety of 0.9 for an infinitely long excavation in normally consolidated clay.
The excavations 10 ft. and 20 ft. in diameter exhibited
small movements, whereas the 80 ft. diameter excavation had proportionately larger movements below a 35 ft. depth.
-275-
Figure C.3.5 compares predicted movements adjacent to
an 80 ft. diameter and a long excavation in normally consolidated
clay for a slurry unit weight of 72 pcf.
Adjacent to the long
trench, the soil yielded within a zone delineated by a 1 on 1
slope from the bottom of the excavation to the ground surface indicated by the magnitude of the movements.
This is as expected
since a wedge analysis of this condition yields a factor of
safety of 0.9 against failure.
For the 80 ft. diameter excavation,
only small movements occur 30 ft. back from the excavation face.
This suggests arching of stresses around the excavation and the
yielding of the soil is confined to a small area adjacent the
excavation.
Figure C.3.6 presents the movements adjacent to slurry
excavations of various dimensions versus the unit weight of the
slurry.
The figure shows that a circular excavation with a di-
ameter equal to the excavations depth and an infinitely long excavation, experienced failure at the face of the excavation at
approximately the same slurry unit weight.
Figures C.3.7 and C.3.8 give the results of predictions
for excavations in clays with an OCR = 2.
The data shows essentially
linear elastic behavior for both circular and infinitely long excavations.
Therefore, stability of slurry excavations in this clay
appears not to be a problem.
C.4
CONCLUSIONS AND RECOMMENDATIONS
Because of the limitation of this study it is not possible
-276-
to determine at exactly what length to depth ratio a slurry filled
trench should be considered a plane strain problem.
However, re-
cognizing that the behavior of a rectangular excavation will be
between that of the circular and infinitely long excavations, the
data suggests excavations with a length to depth ratio of 1 or
greater should be analyzed for stability as a plane strain problem.
-277-
DIMENSIONLESS
)
VARIED
B /2
DISTANCE
(
A
3.0
2.0
0
REFLECTIVE BOUNDRY
i
B
4----
w
0
z
N
-J
A 6 L~
S 6 L c=(
U)
U)
w
0-
6
-6f)
1
I
Ef #+4
-i
z
0
00
H
cn
z
w
L0
x
0
-1
REFLECTIVEBOUNDRY
U
C
N
FIGURE C.3.1: FINITE ELEMENT GRID FOR SLOT ANALYSIS
-A
0
.5
0
.5
1.0
.0
LENGTH
.2
0
.5
ID
HORZ. MOVEMENT x 10
4
i
I
'-EXCAVATION
L/2
FACE
Y
L .5-
.7,
X
L =20'
IELD
ZONE
B = 4L/2
AOL= 1.05KIP/FT2 L2
1.01
HORZ. MOVEMENT KOID
LENGTH
.4 .2
0
.5
0
L --
AECTION A -A
X /L
E XCAVATION
1.9
FACE
50'
Y
L .5
S=62PCF
YIELD ZONE
oV
L = 40'
B =4'
]*
AOL= 1.25 KIP/ FT
2
A
A
1.0-
0-
HORZ. MOVEMENT
LENGTH
4
.2
0
/L
1.0
.5
AL
Iq
=r - Yf Z
EH = 525 KSF
L12.
5L
EXCAVATION
0l-=
FACE
0.62 KSF
i
YIELD
ZONE
L= 80'
B= 41
Acj= 1.25KIP/FT
2
1.0-
FIGURE C.3.2 : MOVEMENTS AND YIELD ZONES FOR A RECTANGULAR
MEDIUM
ELASTIC
OPENING IN A BILINEAR
-279-
I
a
Iii
HORIZONTAL
+80-
~>29
I
50
100
DISTANCE
150
( FT)
200
250
F
+60+40I-
z
0
i
I~
w
F--.
0
-201-40
-60-80-100'
-
0
+20-
D
C
FIGURE C.3.3: FINITE ELEMENT GRID FOR STABILITY ANALYSIS
'
DISTANCE WX)
HOR IZONTAL
f'
20
.
40
60
EXCAVATION
FROM
80
100
120
FACE
( FT)
160
110
180
-
t-.005
z
0WO
X
.01B
EXCAVATI
(FT)
20
10
C
30
40
ON
SYMBOLDIAMETER (D)
I
0
- 20
20
ILi~
1
NJ
00
F-A
I
Soil yields
4
below 35'depth
I.-
0 I
'40
L
a-
L&J
0
0
Local
yield
-6
-
0
.20
.10
60
10000)~72PC
-8 0
.30
&
0
IOFT.
2OFT.
8OFT.
0
HORIZONTAL
.06
.04
.02
- -80
0
MOVEMENT (FT.)
FIGURE C.3.4: PREDICTED MOVEMENTS OF A CIRCULAR BENTONITE SLURRY
EXCAVATION IN A NORMALLY CONSOLIDATED CLAY
A
HORIZONTAL
0
20
FROM
DISTANCE
60
40
80
EXCAVATION
FACE
120
100
(FT)
140
180
160
.
U-
0.2.
A
.
~
-
A
0.1
-0
X (FT)
-0
0.34
20
10
30
Ua
g0.40
-20
I
NJ
00
r-j
I
SYMBOL
--20
IIL.
EXCAVATiON
TYPE
Circular (Dio.8Oft
0
-40
.40
Infinite
w
.60
60
80ft
f =72 PCF looft
I OI~
1.0
.8
.6
.4
.2
0
HORIZONTAL
.6
.4
.2
-80
0
MOVEMENTS (FT)
FIGURE C.3.5: MOVEMENTS OF INFINITELY LONG AND CIRCULAR BENTONITE
SLURRY EXCAVATION IN A NORMALLY CONSOLIDATED CLAY
0
I-
1.2
l.l
2
85
I
0.9
0
1.0
i
80
75
0.8
70
F. S.
65
.10
x
PC F
lz
L =20
L
'
-
Y
.
o
30.
.0.O
A
L= 80'
PLANE
STRAIN
45
H
w
Yf-
.60z
'1
35
C.
7.5
1
PCF
>-
W LL
-101
L 82 0
Iod
0
LJ
Z
"L
.30-4
"A
=O'-)fZ
S
N-#'/' VYWNA A
N
.50- L
PLANE,-I
STRAIN
L= 80'
F.= -
4Su
H(s- f)
(
Io
5
L20.
-
B
.50-
FIGURE
FOR CIRCULAR BENTONITE SLURRY
EXCAVATION IN NORMALLY CONSOLIDATED CLAY
.3.6: MOVEMENTS
L
HORIZONTAL
S0
4.0
I
DISTANCE (X) FROM
190
80
60
I
FACE (FT)
EXCAVATION
li0
120
160
180
-
6.005
.oo5;
(D%
-
;.o
'005
w
in0
0
SYMBOL
EXCAVATION
DIAMETER (D)
20 FT.
x
(i-f)
40
23n
0
0
-20
-20
8OFT.
LA.
La..
00s
.. 40
z
40 I-
0.
0.
LU
D
w
a
80.
60
-60
=72PCF OOF
.08
.04
- 80
0
80
.08 .04
0
HORIZONTAL MOVEMENT (FT.)
FIGURE C.3.7: PREDICTED MOVEMENTS OF A CIRCULAR BENTONITE SLURRY
E XC AVATION IN AN OVERCONSOLIDATED CL AY (OCR =2)
190
I
85
0
L80
I
80
1.70
[.o
I
75
1.50
70
I
1.30F
2.
F. S.
1.
65
Y? PCF
0..
idL
=20 FT
.02.
2
L=80 FT
.04-
z-.064.
-j
4 0.
C.)
Lii
35'
cloy
Surry
B
H
45
.08
PLANE STRAIN
.lo
9 1 I5
c,
80
75
70
I
65
PCF
I
I=2n1
'L
= 0H -f Z
FT
z
L=80 FT
.
0
N
0
PLANE
4
FS=
Su
STRAIN
.20-
FIGURE C.3.8: PREDICTED MOVEMENTS FOR CIRCULAR BENTONITE
SLURRY EXCAVATION IN AN OVERCONSOLIDATED CLAY;OCR=2
0* 5
HORIZONTAL
40
DISTANCE (X)
60
80
FROM EXCAVATION
100
120
FACE (FT)
160
140
*0
-
LL
20
Z .02W.04
0
a:8
0
0C'
20
A
30
0
08
SYMBOL
(FT.)
X
.06.
EXCAVATION
TYPE
Circular
Di.=80FT.
Infinite
0
20
20
40
-.. 40
w
60
a.
w
a
-60
80.
Y= 72 PCF looFT
.15
.10
.05
,80
0
HORIZON TAL
-80
.15
.10
.05
MOVEMENT
(FT)
0
FIGURE C.3.9: MOVEMENTS OF INFINITELY LONG AND CIRCULAR
EXCAVATION IN AN OVERCONSOLIDATED CL AY (OCR=2)
IQ0
APPENDIX D
BRACE II FINITE ELEMENT PROGRAM
D.1
BACKGROUND
The use of the finite element technique in analyzing complex
engineering problems is well described in the literature (Zienkiwiez,
1967).
Briefly, this technique models a problem by an assemblage of
discrete triangular or quadralateral elements.
The forces
(Q) at the
nodes of these elements is related to the node displacements (U) and
the global stiffness (K) of the element assemblage.
A system of
linear equations results which can be described by the equation:
[4] lo
(D.1)
= [Z)
This system of equations is solved to obtain nodal displacements.
These
displacements are then used to determine element strains and stresses.
A recent application of this technique was in the development of
of a computer program BRACE (Wong, 1971) for analyzing the behavior
of braced excavations.
The program models soil by discrete elements with
bilinearly-elastic stress strain properties.
The retaining membrane
for the excavation is simulated by one-dimensional linearly elastic
bar elements.
The program simulates a specified excavation and
bracing-construction sequence.
The loads from a particular construction
operation are applied incrementally.
For each load increment a speci-
fied modulus is used until the yield strength of the soil is attained.
Thereafter, a reduced modulus is used for each additional load in-287-
crement.
This appendix describes modifications made to BRACE resulting
in a new version BRACE II.
In addition, results are presented of
studies to evaluate the programs capabilities to model the soilretaining wall interaction.
D.2
PROGRAM MODIFICATIONS
Modifications to BRACE II consisted of improving the programs
ability to model the behavior of the retaining wall represented by
bar elements and to simulate the anisotropic stress-strain properties of soil.
In the presentation of these modifications an under-
standing of the finite element technique is assumed.
D.2.1
Retaining Wall
Two capabilities introduced in BRACE II were the ability
of the retaining wall to develop a plastic hinge when a specified
yield moment was exceeded and the ability of the soil behind the wall
to slip unrestrained relative to the sheeting.
important in the analysis of braced excavations.
Both capabilities are
In braced excavations
in soft clay with large strut spacings, the sheeting may become overstressed and large movements will result.
In cases where a concrete
slurry wall is installed, a bentonite clay cake remains between the
concrete and the soil.
strength, restraint
Since this clay has essentially no shear
of slippage between the soil and concrete wall may
be considered non-existant.
The element stiffness matrix [S] for an arbitrarily oriented
one-dimensional bar element is as follows:
-288-
r
u
V
1
u
1
121
2(
121+~
A)X
L2
L2
(1
12+y2
V
2
2
p 2+AA 2 )
2 (L_A)y
L2
L
X
-
-A)Ay
.. (_12Ix2+A12
L2
L
(12 1 '2+A A2
L
61
61
L
61
-LI
=
cos
0,
-L
61
L
L
41
where y
61
61
-
2
)
02
E
L
Symmetrical
I
61
L
21J
X = sin 0, 0 is measured from the vertical axis
A = area of cross section of element
E = Young's Modulus of element
L = Axial length of element
I = Moment of Inertia of element
u and v are the translations in the horizontal and vertical
directions respectively, and
6 is the rotation of the bar
element at the nodes.
Subscripts 1 and 2 refer to the two end nodes.
If, during the application of incremental loads, the yield moment
of the sheeting is exceeded at a bar element node, that node is made a
plastic hinge for all additional incremental loads.
This is accomplish-
ed by statically condensing out the rotational resistance of the
-289-
-1
yielded element node.
This approach was suggested by Prof. J.T.
Christian (1971).
The equilibrium equations for a bar element can be written
as follows:
Sli
S12
S13
S14
S15
S16
S22
923
824
325
026
V
S3 3
534
535
S 36
u2
S4 4
S45
S46
V2
2
555
S56
1
M
S66
02
M2
Symmetrical
1
V1
1
=
2
D.2
where:
V ,V2
denotes the shear at nodes 1 and 2
P1 9 P2 denotes the axial load at nodes 1 and 2
M ,M2 denotes the moment at nodes 1 and 2
Equation D.2 can be written in following partitioned matrix:
Si
S2
U
P
S2
S
0
M
k
D.3
Since no external moments are applied to the elements M can
be set equal to zero to obtain the following equations:
[S ]
(U)
+
[S 2
(0)
=
P
D.4
[S2]
(U)
+
[S 3 ]
(6)
=
0
D.5
Solving for 0 and substituting in D.4 yields
-290-
[S
(U)
+
[S21
3
[S2 ]T
(U)
=
P
D.6
Therefore the effect of the initial rotational stiffness at
a yielded element node has a direct influence on the unyielded,
adjacent element nodes.
Yielding at one end of a bar element is
accounted for by setting the resisting internal moment, M2 , at that
and equal to zero in equation D.2.
02 ~
~
Solving for 02:
K61+ K62 + K 6 3 + K 6 4 + K 6 5
2V
K66
v1
D
u
D.7
v2f
u2
M
This equation for
62 is substituted through the rest of the
elements stiffness matrix to condense, out the effect of the yielded
node.
Unrestraint slippage between the soil and sheeting at the interface was simulated by assigning zero axial stiffness to the sheeting.
This is accomplished by setting the v1 and v2 terms in the element
stiffness matrix (Equation D.2) to zero.
Figure D.2.1, through D.2.3 compare sheeting movements and adjacent ground movements predicted using non-yielding and yielding bar
elements.
The importance of considering yielding of the bar elements
is evident from these figures when making predictions of the behavior
of a braced excavation.
-291-
-4
D.2.2
Anisotropic Stress-Strain Relationships
Ladd et al (1971) shows that in plane strain tests
both the undrained modulus and the undrained shear strength
of clay are dependent on the orientation of the principle stresses
applied to the soil.
In the construction of a braced excavation
the adjacent soil experiences a reorientation of principal stresses
(Wong 1971).
Therefore, the accuracy of an excavations predicted
behavior will be dependent on how well the soils anisotropic
properties are modelled.
The anisotropic variation of modulus was simulated
based on proposed methods by Christian (1971).
He recommends the
following approximate relation to account for the effect of stress
reorientation on the undrained modulus:
EH -
E=
(EH-EV) Cos 4
where EH and EV are the undrained modulus for tests with
the major principle stress applied to the soil in the vertical and
horizontal directions respectively, 0 is the orientation of the major
principle stress from the vertical plane and E0 is the modulus used
to compute the deformations of the soil.
To account for anisotropic shear strength properties the
following yield criterion recommended by Davis and Christian (1971)
was used:
(xay
2
-
Suv~ Suh
2
2 + -x
2
x
-292-
a2 = a 2
-*1
where:
b
a
a ,a
S45
--S S
uv uh
,T
= conventional total stress components
in the x,y plane
= shear strength of a soil sample
Su45
oriented at 45 degrees from the
vertical.
S UV'
uh
= shear strength for compression in the
vertical and horizontal directions.
D.3
ACCURACY OF BRACE II RESULTS
There is inherent incompatibility in the BRACE program between
the deformations of the bar elements and the soil elements.
The
formulation of the soil element stiffness matrix assumes each soil
element has a constant strain in any direction and, therefore displacements along the side of the soil elements vary linearly.
On the
otherhand, the bar elements are not restricted in the elastic line
they can assume.
For the general loadings from the soil the bar element
line will be at least a 4th order curve.
Hence, horizontal
Novements
between the soil elements and the bar elements will be compatible only
at the connected nodes.
This incompatibility could result in gross
errors in the predictions of sheeting forces if sufficient bar elements
are not used to model the sheeting between two support points so the
soil deformation and sheeting deformations at the nodes can lie essentially on the correct elastic line for the imposed loads.
A series of computor analysis were made to determine the number
-293-
of bar elements necessary to model a cantilever sheeting.
This
problem will give, theoretically, a 5th order elastic line for
sheeting deflection and represents the highest degree of incompatibility expected in a braced excavation analysis.
Figures D.3.1 through D.3.4 show the results of these
analyses for sheeting elements of varying length and varying
stiffness.
Normalized soil properties for normally consolidated
clay listed in Appendix E were used for the analyses except for
test runs 7 and 8.
Except for test 10 the vertical stiffness of
the sheeting was neglected to simulate slippage between the soil
and sheeting.
Table D.3.1 compares moments and shear forces predicted by
BRACE II with those obtained by static analysis using the horizontal stresses (ax) on the sheeting predicted by BRACE II.
The
stresses for the static analysis were extrapolated to the face of
the sheeting using the stresses at the center of the first three
elements adjacent the sheeting.
(The variation of these stresses
are shown on the figures.)
The results of this study show that the number of bar elements
required to get reasonably accurate answers for sheeting moments and
shear forces is primarily dependent on the stiffness of the bar
elements and the homogenuity of the soil profile.
It appears that for
most problems the use of 5 bar elements in between bracing supports
6
is sufficient to give accurate results with stiff sheeting (EI = 10 kif)
4
and 4 bar elements is sufficient for the flexible sheeting (EI = 10 ksf).
-294-
The incremental load technique in BRACE II poses the problem
of determining the number of load increments required to simulate
load applications.
If, for example, the load relief due to an
excavation stage is applied in one increment, those soil elements
in elastic range before the excavation stage behaves elastically
for this construction stage.
This can result in under-predicting
the deformations of the retaining wall below excavations level
if the elements would have yielded under one-half the load release if two load increments were used.
Figure D.3.5 shows the effect of using 1 versus 2 load
increments per construction stage.
For the case of 2 load in-
crements, the sheeting movements were approximately twice that
for a single level increment at the lower depths.
Clearly this is
a significant factor in the predictions of the behavior of braced
excavations.
It is not possible to assess how many increments should
be applied for each load application since this will be a function
of the problem geometry, soil parameters, and sheeting stiffness.
Hence, it is recommended that test runs be made to determine the
optimum number of load increments to be applied to each problem.
-295-
Table D.3.1 Effect of Bar Element Size on Predictions of Shear and Moment Values
Moments
Sheeting Properties
Test
No.
Shear
(K-Ft)
(Yips)
Excavating
EI
10 (K-Ft2)
Bar El.
Length (Ft)
Depth
(Ft)
Brace
1
Computed
Brace
Computed
1
18.5
5
15'
+5.8
+10.8
-1.2
+3.9
2
18.5
2.5
15'
21.2
20.8
4.4
4.0
3
0.58
2.5
15'
18.9
20.4
3.5
5.0
4
0.41
3.0
9'
6.5
7.8
1.4
2.3
5
4.05
3.0
9'
3.2
4.4
.8
1.5
6
0.41
3.0
15'
15.6
22.0
3.2
4.5
3.0
12'
3.8
10.0
1.1
3.6
7
18.5
8
0.058
3.0
12'
5.6
10.6
1.7
3.8
9
4.0
3.0
12'
19.5
20.6
3.5
5.15
10
4.0
3.0
121
17.4
16.8
1.4
4.2
1.
Moments computed using stresses on sheeting predicted by brace.
HORIZONTAL
j syE04Exvation
.2
.1
0
DISPLACEMENT (FT)
.5
.4
0
I
.3
.2
-
r/77777
9u
9
18
2
0,~.
-20
I--
3
27
4
36 0
a.
4
w
45
0
6
.40
a
55
&
I4
0
SHEETING
I:t 4x (k -f?-
N ORMALLY CONS
CLAY
I
8
0
NON YIELDING
BAR ELEMENTS
FIGURE D.2.1 . COMPARISON OF SHEETING MOVEMENTS
NON YIELDING SHEETING ELEMENTS
80
YIELDING
BAR ELEMENTS
FOR YIELDING AND
(FT-KIPS)
MOMENTS
2
0
+20
-20
60
40
20
0
-20
-40
60
40
20
0
-20
YIELD
MOMENT
8
)
20-
-
0-
EXCAVATON
STAGE
3
4
0. 40+
wJ
a
I
27 U
36
5
45
6-
55
k
w
4
60+
I
I
00
I
-
:3
I
8
4th STAGE
SHEETING E= 4x 104K-FT 2
SOIL- CLAY
6th STAGE
5th STAGE
E
OCR =L
-
NOTE
: SHEETING
A
TICt
I
YELD MOMENT=25 FT kips
MOMENTS IN ELASTIC RANGE THROUGH 4th EXCAVATION
FIGURE D.2.2: COMPARISON OF SHEETING MOMENTS
YIELDING SHEETING ELEMENTS
FOR YIELDING AND
STAGE
NON-
-40
EXCAVATION
0
20
HORIZONTAL
40
60
DISTANCE
100
80
(FT)
120
140
z
160
-
01
0 w
.2
20-
I40-
w
a
z
w. 2 -.I..
T10
-.
.82 FT.
+TO 0.83 FT.
-
- - --
60-1.
YIELDING
ELEMENTS
NON YIELDING ELEMENTS
EXCAVATION
DEPTH = 55 FT.
80.L
FIGURE D.2.3 : COMPARISON OF GROUND SURFACE MOVEMENTS
AND NON - YIELDING SHEETING ELEMENTS
FOR YIELDING
PREDICTED
MOMENTS
(K-FT)
O
2.0
HORIZONTAL DISTANCE
)
6 x kksf)
0.5
71-
0
I
FROM
SHEETING
I.0
7.5
2.5
1.0
0
0.5
1.0
I
I
0
0.5
C)
10 15
74
1.0
I1
TEST
5
H
H
SHEET
PROPERTIES
6 x(ksf)
67x(ksf)
E =
T
EXCA.
EPTH
20
W
I
4 8 2 00 0
k/ft
= 2.25 ft#
No. of el.
i
I
= 3
TEST 2
5
E = 482000 k/ft2
I = 2.25 f t4
No. of el. = 6
10
15
a.
20
O
I
~~1~
I
-I
TEST
3
5
E = 4320000 k/fta
I=.0135 f t 4
10
-
0
---
No. of el. = 6
15
20
FIGURE D.3.1: HORIZONTAL STRESS VARIATION BEHIND SHEETING WALL FROM BRACE
HORIZONTAL
2.5 FT.
MOMENT
K -FT.
00
DIS TANCE FROM SHEETING
7.5FT.
O-x
o-x (KSF)
10
0
.2
A
.6
.8
0
.2
(KSF)
.4
.6
.8
TEST 4
4
E =4.3 x 10
k sf
1 =.0093t
H
No of EL=3
121
H = 9 ft
0
0
10
0
.4
.6
.?
.8
.6
4
TEST 5
4
E=4.3x10
8..
I
-
2
.
ksf
w
a
.2
.
161
-
.093
No of EIr-3
12--
H = 9 ft
0
K)
0 r----
C
.2
.4
.6
.8
(
16JL
)
.4
.g
.8
TEST 6
E= 4.3xlcf
ksf
I =.0093
4.
No of EI=5
H = 15ft
121
16
.2
77M
FIGURE D.3.2- HORIZONTAL STRESSES AND SHEETING
MOMENTS FOR CANTILEVER SHEET PILING
FROM BRACE
-301-
MOMENT
K - ft
5
1
0 .2
K0
.6
0.
DISTANCE
.8
1.0
FROM
0
.2
SHEETING
75 FT
.4
6
.8
1.0
TEST 7
-
0
0
.
HORIZONTAL
2.5 FT
-=250
4
E=4.8x io
IH
4E
8
ksf
1000
-
12-
1=2.25 ft4
No d E=4
H= 12ft
:.
a
0
)
5
10
0
.2
.4
.6
.8
1.o
0
.2
.4
.6
.8
01
E=250
44
8.
=-000
TEST 8
E= 4.3 x10 6
ksf
I=0. 087
Nod EI=4
H =12 ft
16-
FIGURE D.3.3 : HORIZONTAL
CANTILEVER
STRESSES AND SHEETING MOMENTS
SHEET PILING FROM BRACE
FOR
I-
HORIZONTAL
0
MOMENT
2.5 FT.
K-FT.
10 20
O'x (KSF)
0
30
I
i
p
0
.2
.4
.6
DISTANCE
7.5 FT.
ox (KSF)
.8
1.0
0
.2
.4
.6
.8
1.0
TEST 9
4
E =4.3x 105
7
i
I
I = .093FT
No of EI=4
12
-3
-~
0
0.
10
-
2)
30
No VERTICA
STIFFNESS
0
.?
.4
.6
.8
1P0
S.4
.6
.8
1.0
TEST 10
I
4.
E = 4.3 x105
81
I =.093FT4
12
With VERTICAL
STIFFNESS
16
FIGURE D.3.4 :
HORIZONTAL
STRESSES
CANTILEVER
SHEETING
AND SHEETING MOMENTS FOR
S HEETING
.10
.05
0
MOVEMENTS (FT)
.15
.10
.05
0
0
I
-*Am
204
I
I.-
a.
w
2
2-+
3
40. 3 -- m
3-
C
4
U.)
604
80 1
I LOAD INCREMENT
PER STAGE
2 LOAD INCREMENTS
PER STAGE
SHEETING EI=4 x10 4 P-FT 2
SOIL- NORMALLY CONSOLIDATED CLAY
FIGURE D.3.5 : EFFECT OF NUMBER OF LOAD INCREMENTS
EXCAVATION STAGE ON SHEETING MOVEMENTS
PER
APPENDIX E
SOIL PARAMETERS FOR ANALYSIS AND PREDICTIONS
E.1
PARAMETRIC STUDIES
Soil parameters for analytical studies were based on synthesized
stress-strain data as reported by Ladd et al (1971) for plane strain
triaxial tests on resedimented Boston Blue clay specimens.
The
selected values were as follows:
1
2
4
0.5
0.72
0.95
E /oYO
250
340
450
EH /avo
150
200
230
S /a
uv vo
0.34
0.57
0.95
Suhlavo
0.19
0.37
0.67
0.49
0.49
0.49
OCR
K
Modulus values correspond to the secant modulus determined at a
principal stress difference equal to one-half the soils shear strength.
After yielding the modulus was set equal to 0.1% of the unyielded mod-
ulus and Poissons ratio was set equal to 0.4000.
The total unit
weight of the soil was taken as 120 pcf.
E.2
SOIL PARAMETERS FOR SOUTH COVE PREDICTIONS
This section outlines the origin of the selected soil parameters
shown in Table 4.7.2 which were used to predict the behavior of the
-305-
South Cove braced excavation.
For the topmost fill layer the strength parameters were
estimated from correlations of the blow count from standard penetration tests and frictionlangle (Terzaghi and Peck 1967).
the blow counts and a visual examination of the material a
Ko
=
4
of 30'
The Ko value was determined from the
was selected for the soil.
empirical formula
Based on
1-Sin *.
Values surmized for the soils unit
weight, elastic modulus and Poissons ratio were 100 pcf, 250 Jvo,
and 0.3 respectively.
The parameters for the hard clay were determined from test
results on block samples.
Strength tests consisted of one con-
solidated isotropic under-drained triaxial test (CIU) and Tor-vane
shear tests.
One consolidation test was performed on the block sample.
The results of these tests are given in Figure A.3.1.
results, it was estimated the soil had an OCR of 5.
Based on these
The value of
Ko= 1.0 was extrapolated from the data given in Section E.l.
shear strength value of 1.0 aoy
The
was based on the torvane strength tests.
The value of Ev = 10000vo was estimated from the CIU test on the block
sample.
A value of EH = 500
was judged reasonable.
0vo,
which is one-half the vertical modulus,
Poissons ratio was set equal to 0.4 because of
the fissured nature of the hard clay.
A substantial number of 3 inch shelby tube samples were obtained
from the medium-stiff and medium clay strata.
In addition, field Vane
tests were made through the entire depth of these soil strata.
Block
samples also were obtained from the strata during the construction phase.
-306-
Consolidation and strength tests were performed on many of the
samples, the results of which are summarized in Figure A.3.1.
These
results were synthesized to obtain a plot of OCR versus depth as
shown in Figure E.2.1.
Ko, Suv9
Based on the OCR at a given depth, values of
Suh, Ev and EH were determined from the test data on re-
sedimented Boston Blue clay summarized in Section.E.l.
Figure E.2.1
shows the variation of K0 , Suv and Suh with depth.
The presents of the till was ignored in those analyses predicting
the stability and deformations adjacent the excavation.
It was
assumed that this strata was dense and stiff enough, when compared
to the medium clay, to be considered essentially a material equivalent
to the underlying bedrock.
In all finite element analysis, after yield of a soil element its
modulus was set equal to 0.1% of the unyielded elastic modulus and
Poissons ratio was increased to 0.4999.
-307-
a
LL
LL
-j
OCR
LU
0
1
122
4
2
P-7
0
6
N
FILL
115-
100-
-100
Hard
Yel.
CLAY
Very
Stiff to
Bot.
OCR
PLAN E
I
Boston
72E
Blue
CLAY
62,
Medium
oI
Bt. of
SlurryW
80tm
-40
DATA F RO M
( K.
Boston
DAPPOLONIA et c 1 1970)
(PLANE STRAIN ACTIVE)
Blue
(STRENGTH
CLAY
100-
S
PA,SSIVE
Stiff
of
40- -80
60L60
I.10
L ADD
20
et ol 1971
18
BED ROCK
O
FIGURE E.2.1: SOIL
MMM
.2
.4
.6
.8
1.0
1.2
Ko
CHAAACTERISTICS
USED
FOR
FROM
RATIO
PREDICTIONS
)
w
Massachusetts Institute of Technology
Department of Civil Engineering
Soil Mechanics Division
AITENDIX
F
USER'S MANUAL
PROGRAM BRACE II
Date:
December 1970
Modified:
February 1972
Language:
FORTRAN IV (G Level)
Programmer:
J. T. Christian and I. H. Wong
Modified by:
W. E. Jaworski and J. T. Christian
I.
DESCRIPI'ION
BRACE II is a finite element program for analyzing braced
excavations.
which include
It models excavation and bracing construction stages
prestressing of struts and sheeting displacements due
to the deformation of shims.
Problems are restricted to plane strain conditions.
Con-
stant strain triangles or quadrilaterals represent the soil mass. The
sheeting is represented by one-dimensional bar elements.
It is
assumed that struts are installed horizontally and their deformations
are neglected.
Excavation stages are simulated by specifying soil
elements and associated nodes to be removed.
Material properties for the soil may be considered as linearly
or bilinearly elastic and isotropic or anisotropic.
be specified for the sheeting.
Yield moments can
The program performs a linear elastic
incremental load analysis for each construction stage.
-309-
Total stress
analysis is used and pore pressures are evaluated on this basis.
compressible materials cannot be specified.
In-
The program computes
for each construction stage the strut loads, sheeting forces and soil
stresses and deformations.
II.
PROGRAM CAPABILITIES
The following restrictions are placed on the size of problem
which can be input.
Nodal Points
- 290
Elements
- 260
Soil Types
-
20
Strut Sizes
-
2
Sheeting Elements -
25
Soils must be input as layered systems.
Initial stresses can
be input for each element or generated by the program.
Plotted output can be obtained from a modified version of
CNTRPLOT available at the Soil Mechanics Division, M. I. T.
Computer running time depends on the band width of the
global stiffness matrix.
For typical problems the running time varys
between 1.0 and 1.75 minutes per load increment at each excavation
stage and .25 to 1.0 minutes per bracing stage, the greater time
resulting from specifying a prestress.
III. INPUT DATA FORMAT
A.
Title Card - Format (18A4)
Any title or comment in card columns I through 72 will
be reprinted at the top of the output.
be provided.
-310-
This card must
W."
B.
Control Card - Format (2015).
This card contains control information for the program
as follows:
Card Column
Information
1 - 5
Number of Nodal Points in original
configuration of problem before
excavation
(NUMNP) Maximum = 290
6 - 10
Number of Elements in original configuration
(NUMEL) Maximum = 260
11 - 15
Number of soil materials
(NUMMAT) Maximum = 20
16 - 20
Number of strutting materials
(NMSMAT) Maximum = 2
21 - 25
Number of sheeting element
(NSHEET) Maximum = 25
If NSHEET = 0 Omit Piling Material Card
26 - 30
Gravity Stress Indicator (IGRAV)
If IGRAV = 0 initial stresses are set
equal to 0. If IGRAV = 1 initial
stresses are calculated from Y and
K from Card D-1
0
31 - 35
Excess Pore Pressure Indicator (NPOREC)
Put NPOREC = 1 if excess pore pressures
due to total stress changes are to be computed for any one or all layers
36 - 40
Initial Pore Pressure Indicator (NPORE)
-1 hydrostatic initial pore pressure
0 no initial pore pressure
1 nonstatic initial pore pressure
-311-
41 - 45
Plotting Indicator (IP LOT)
IPLOT should be * 0 even if no plot is
to be generated before excavation is
initiated (see Instruction Card).
46 - 50
Surface loading card (ILOAD)
ILOAD = 1 surface load present before
excavation and before sheeting is driven.
Deformations due to these surface loads
will not be superposed to deformations
after sheeting is driven.
ILOAD = 2 surface load present before
excavation and after sheeting is driven.
51
-
55
Settlement Indicator (ISETLE)
If ISETLE * 0 Parameters needed to
compute I D settlement or heave due
to dissipation of excess pore pressure
will be read in.
56 - 60
Anisotropic Strength Indicator
Set NDELSU = I for card E I to be read in.
61 - 65
Soil sheeting Interface Card. If NADD > 0
slipping between soil and sheeting is
allowed.
66 - 70
Number of load increments (NLDINC)
NLDINC should be - 1.
At each increment the load applied will
be total load/NLDINC.
71 - 75
Capillarity Indicator (ICAPIL)
If ICAPIL * 0, capillary pore pressures
will be considered in calculating effective
vertical stresses.
75 - 80
Nodal Point Update Indicator (IUPDAT)
If IUPDAT > 0, nodal coordinates are
updated.
-312-
C.
Piling Material Card - Format (3Fl0.0)
Omit if NSHEET = 0
Information
Card Column
1 - 10
Young's Modulus, E
11 - 20
Moment of Inertia
21 - 30
Cross Section Area
31 - 35
(IVSC) Axial Stiffness = 0 if IVSC = 1
36 - 45
Sheeting Yield Moment (YMOM)
If YMOM = 0 the sheeting has no
rotational stiffness.
D.
Soil Material Cards - Format (15, F5. 0, 3F10. 0, 2F5. 0,
2F10.0, 2F5.0)
One card per soil type from number (one) to a maximum
number of 20. Cards to be input sequentially.
Information
Card Column
Soil Number
1 -5
6
-
10
11
-
20
21
-
30
K
0
Unit Weight,T
Young's modulus in the Vertical
Direction, E
V
31 - 40
Young's modulus in the Horizontal
Direction, Eh
41 - 45
Poisson's ratio from Vertical to
Horizontal, v VH
46 - 50
Poisson's Ratio from Horizontal to
Horizontal, v
HH
-313-
51 - 60
Blank
61 - 70
Cohesion, C or undrained shear strength
)
when the major principal stress is vertical
(S uv
71 - 75
Friction angle,
76 - 80
Yield Factor
Notes: If EH is input as zero, an isotrc pic material is assumed
with E=EV, andv =v H. For anisotropic soils the
modulus for an element is a fun ction of the angle (6)
between the major principle str ess and the vertical
plane and is taken as
-
Material Properties - Second Set
os
0.
Card (15,5X5F10.0, 512)
If (NPOREC * 0) this set of cards w ill be read in. One card
per soil type from number 1 to a m Lximum number of 20.
Input cards sequentially.
Card Column
Information
Information
1 -5
-
10
Blank
11
-
20
Shear strength when major principal stress
is rotated 900 (S
30
Shear strength when major principal stress
is rotated 450 (S u45). If left blank:
)
5
-
E.
= EH-(EH - E
)
E
21
Su45
=
2 ( 5 uv
uh
The shear strength Suo of an element is
dependent on the angle 0 between the major
principle stress and the vertical plane. An
elliptical strength variation is used (Davis
and Christian, 1970) in the program. The
equation describing the yield criterion is:
(a
.2
- a X -S uv
-314-
2
2
Suh) + o 2 a
xy 7=
b
2
a
.4.
uv uh
where b/a =
a=
S
uv
+5
uh
2
31
-
40
Henkel pore pressure parameter a.
a is a constant
AU = A oct + a'roct
41
-
50
Blank
51
-
60
Yielded Poisson's ratio
61
-
62
Modulus Indicator
1 Modulus normalized with respect to av
0 Constant modulus with depth
1 Modulus normalized with respect to Oct
63 - 64
Strength indicator
1 Strength normalized with respect to av
0 Constant with depth
-1 Strength normalized with respect to at
65
-
66
Pore Pressure Curve Indicator
1 Henkel parameter a will be a continuous
function with maximum shear strains
67
-
68
Yield Poisson's Ratio Indicator
1 Bulk modulus computed using yielded
Poisson's ratio
0 Bulk modulus constant during shear
69
-
70
Anisotropic Strength Indicator
If zero, isotropic shear strength is used
for that material
-315-
E. 1 Nonuniform Henkel Parameter Card - (215, 2FlO. 0)
If column 65 - 66 in Card E is nonzero, the following set of
cards will immediately follow that card for that material.
Information
Card Column
F.
1.- 5
Number of data points of maximum shear
strains to be read in.
6 - 10
Number of data points of Henkel's pore
pressure coefficients a to be read in.
11 - 20
Constant Henkel's pore pressure coefficient a.
21 - 30
Maximum shear strain beyond which
Henkel's coefficient a is constant.
One-Dimensional Settlement Card (15, 5X, 5F10.0)
If Column 51 - 55 (ISETLE) # 0, the following cards must be
supplied, one for each material.
Card Column
G.
Information
1 - 5
Soil Number
6 - 10
Blank
11 - 20
Void Ratio at L of Layer
21 - 30
Virgin Compression Index
31 - 40
Compression Index in 0 - C range
41 - 60
Effective stress at (L of Layer
Strutting Material Card - Format (15, 2F10. 0)
If no strutting material is desired, NMSMAT should be set
to zero and this card may be omitted. No more than two
types of strutting can be used.
Card Column
1 -5
6
16
-
15
25
Information
Material Number
Young's modulus
Cross-sectional area
-316-
HI.
Nodal Point Cards - Format (15, F5.0, 4F10.O, 15)
Cards should be input in increasing order of number of
nodal points. If cards are omitted, the nodal points will
be generated along a straight line between the two points
before and after the omitted ones. All such generated
points will be unrestrained and will have no load on them
except as caused by gravity stress and excavation or
bracing.
Information
Card Column
Nodal Point Number (N)
1 -5
6
-
10
Loading Code:
Code
UX(N) is a force
UZ(N) is a force
UX(N) is a displacement
1
UZ(N) is a force
UX(N) is a force
2
U(uN) is a displacement
3
tUX(N) is a displacement
UZ(N) is a displacement
X coordinate (X (N)
21 - 30
Z coordinate (X (N))
31 - 40
UX(N) if necessary
41 - 50
UX(N) if necessary
51 - 55
KOD If # 0, all generated succeeding
)
11 - 20
points will have the same code.
I.
Soil Element Cards - Format (615)
Cards should be input in order of increasing element number.
If cards are omitted elements will be generated by adding one
to each of the nodes of the preceeding element.
-317-
Material
numbers are kept constant in the generation.
element must be input.
Card Column
The last
Information
1 - 5
Element Number (M)
6 - 10
Node Number I
11 - 15
Node Number J
16 - 20
Node Number K
21 - 25
Node Number L
26 - 30
Material Number
Note: -Nodes must be numbered in a rotational order from the
positive X to positive Z (axes), i.e., counter clockwise
for the usual convention of Z positive upwards and N
positive to the right. Triangular elements are described by making K = L. The maximum difference
between nodes for any element must not exceed 26.
J.
Sheeting Element Cards - (615)
If NSHEET = 0, omit this card.
Card Column
Information
1 - 5
Element Number
6 - 10
Nodal Number I
11 - 15
Nodal Number J
Intermediate Element Cards will be generated. Only end
Cards need be input. Maximum allowable difference between first and last sheeting node number is 26.
K.
Plot Control Card - Format (12, F9. 4, F5.1, 6F8. 3)
This card is used only if plots are requested.
Card Column
1 - 2
Information
Integers '01'
-318-
-'I'
3 - 11
Blank
12 - 16
Distortion Factor for displaced mesh,
DMESH. Displacements are multiplied
by this factor to obtain an exaggerated
plot.
Vector scale factor for principal stress
plots, SPLOT. Value is length of
largest vector in grid units.
17 - 22
23 - 28
Delta Sigma X contour plot code
29 - 34
Delta Sigma Z contour plot code
35 - 40
Tau XZ contour code
41 - 46
Tau maximum contour code
47 - 52
Maximum shear strain contour code
53 - 58
Excess Pore Pressure contour code
The contour codes are interpreted thus:
59 - 64
Total Sigma X contour plot code
65 - 70
Total Sigma Z contour plot code
0. or blank - no contour plot desired
Positive
-
value is the interval between contours
value is the number of desired contours
The plotting program will find a suitable, even interval.
After these cards are read the computer sets up the problem and solves for any initial loads, displacements or
stresses. The following input can follow:
Negative
L.
-
Grid Reduction Card - (2F10.0)
To restrict the contours to region of interest.
plot is wanted.
Card Column
1 - 10
11 - 20
Omit if no
Information
Extreme Left X-coordinate of grid
to be shown
Extreme Right X-coordinate of grid to be
shown.
-319-
M.
Pore Fluid Card (2F10.0)
If NPORE (Column 36 - 40, Card B) is -1,
be supplied.
Information
Card Column
1 - 10
11 - 20
this card must
Depth to water table below ground surface as a positive number
Unit weight of water
Or if NPORE is 1, the following card must be supplied.
Card M-1 (15, IPE12.4)
Card Column
Information
1 - 5
Element number
6 - 17
Pore Pressures
Card M-1 must be repeated for every element for which
there is non-zero pore pressures.
N.
Layer Thickness Cards - Format (8F10. 0)
These cards are used only if IGRAV is not zero, that is, if
initial stresses are to be calculated. Each card contains
up to eight numbers, each of which describes the thickness
of one layer of soil. The layer numbers correspond to the
soil material numbers on cards D-1 and E-1 and must
increase with decreasing depth. Thus, soil 1 is the top
layer, soil 2 the next and so on. A maximum of three
cards may be needed to describe all twenty permitted
layers. The first would read:
Card Column
Information
1 - 10
Thickness of soil 1
11 - 20
Thickness of soil 2
21 - 30
Thickness of soil 3
and so on.
-320-
0.
Instruction Card - Format (18A4)
This card must have one of the following five sets of characters in card columns 1 through 4:
a.
This signals end of problem, if the next four
**
columns also contain '****' execution ends.
Otherwise a new problem is read starting with
card A.
b.
'EXCA' This means excavation will occur as described
under cards P through T below.
c.
'BRAC' This means bracing will occur as described
under cards U and V below.
d.
'LOAD'
This means new surface loads are input.
e.
'STLE'
This means l-D settlement or heave will be
calculated due to dissipation of excess pore
pressure.
P.
Excavation Control Card (815, 2F10. 0, 15)
Card Column
Information
1 - 5
Number of elements to be removed
6 - 10
Number of nodes to be removed
11 - 15
Number of new surfaces exposed by
excavation (one surface per element
exposed)
16 - 20
Plotting indicator for new configuration, IPLOT
21 - 25
Highest degree of polynomial used for
extrapolating stresses at sheeting surfaces. The maximum permissible
value is the number of elements in a
horizontal row to be removed minus 2
(set = 1 for first 'EXCA')
26 - 30
Number of Load Increments (2: 1)
31 - 35
Element number for the corner element
between old surface and sheeting
-321-
36 - 40
Highest degree of polynomial used for
corner element. (Cannot exceed the
value in cc 21-25 above).
41 - 50
Z-coordinate above which slipping between sheeting and soil occurs. If
NADD = 0 leave blank
Q.
51 - 60
Normalizing factor for plots.
61 - 65
1 results for
Output control. If
each load increment output. If blank
only results from final load increment
output.
Elements Removed - Format (815)
Numbers of the elements to be removed are listed, up to
8 per card. As many cards as are necessary are used.
R.
Nodal Points Removed - Format (815)
Numbers of the nodal points to be removed are listed, up
to 8 per card. As many cards are used as are necessary.
If no nodal points are removed the card should be omitted.
S.
New Surfaces Exposed - Format (215)
Card Column
Information
1 - 5
First node
5 - 10
Second node
Repeat S for each new surface.
T.
Boundary Cards - Format (1615)
These are needed only if plots are requested.
The number of
nodes at the end of straight lines on the boundaries are listed,
16 per card, on as many cards as needed, up to IPLAT nodes.
-322-
_11
Cards P through T should follow the 'EXCA' card. The program will solve for the effects of excavation, print output,
and read the next card 0.
U.
BracingCard - Format (215, F10.0, 15, Fl0,
25)
Information
Card Column
1 - 5
Node at which strut is installed
6 - 10
Strutting material number
11 - 20
Prestress in bracing program com-
putes preload for strut. Input stress
in positive direction as negative value
+z
Input stresses as a
Input stresses as - a
+
x
21 - 25
Plotting indicator, as in 16 - 20 for card P
26 - 35
Crushing of timber wedges at % of movement already occurred at strut level
36 - 40
Load Increment (2 1)
41 - 45
Output control.
If set 21 results for
each load increment output. If blank
only results for final load increment
output.
V.
Boundary Card - Format (1615)
These are needed only if plots are requested and are identical
to card T.
Cards U and V should follow the 'BRAC' card. This program
will solve for the effects of the bracing, print output and read
the next card Q
If "load" in card 0, cards W and X will be read in.
W.
Loading Control Card - (15)
Card Column
1 - 5
Information
Total number of nodes subject to
external loads
-323-
6 - 10
11 - 15
Plotting Indicator as in CC-16-20 for
card P
Number of Load Increments (NLDINC)
If NLDINC = 0, only the nodal codes are
changed.
X.
Loading Cards
Information
Card Column
1 - 5
Node number
6 - 10
Nodal Code
31 - 40
UX(N)
41 - 50
UX(N)
See Card H
As many cards are needed as number of loading points.
If 'STLE' in card 0, the following cards will be read in.
Cards Y-2 - (1615)
Element numbers in a string which contribute to heave of
settlement. As many cards as needed.
Cards Y-3 - (15)
Card Column
1
Total number of nodes in a string
5
-
Information
which heave or settle.
Cards Y-4
-
(1615)
Node numbers in a string which heave or settle.
cards as needed.
As many
Hardware Requirements
The program required that scratch discs or tapes be set up
on logical units 8, 11, 12 and 13. If plots are desired the
required data can be written on data set reference number 7,
as described by Job Control Language (JCL) cards or the
-324-
required cards punched.
bytes of computer core.
The program requires 450K
Recommended JCL cards for the 370-MI55 IBM computer at the MIT Information Processing Center are
as follows:
//
'Name', CLASS=B, REGION=450K
/*MITID
USER =((MMMMM, NNNN)
/ *SRI
TIME=TT, LINES= LL, CARDS=CC
/*MAIN
DD
//C.SYSIN
SOURCE DECK
/*
//G. FT08F001 DD UNIT=SYSDA, SPACE=(816, (1200, 200)),
//
DCB=BLKSIZE=816
//G. FT11F001 DD UNIT =SYSDA, SPACE=(1280, (600, 100)),
//
DCB=BLKSIZE=1280
/1G. FT12F001
//
DD UNIT =SYSDA, SPACE=(0312, (034, 010)),
DCB=BLKSIZE=0312
//G. FT13F001 DD UNIT =SYSDA, SPACE=(0560, (110, 030)),
// DCB=BLKSIZE=0560
DATA
Where
MMMMM = Problem Number, NNNN = Programmer Number
T T= Maximum time to run problem
LL = Maximum lines of output, thousands
CC = Maximum cards output, hundreds.
-325-
BIOGRAPHY
Walter E. Jaworski was born May 31, 1939 in Woonsocket,
Rhode Island.
He matriculated from St. Mary's High School,
Milford, Massachusetts in 1956.
In June, 1958, he received an Associate of Science in
Mechanical Engineering from Worcester Junior High School, Worcester,
Massachusetts.
In June, 1962, he graduated from Northeastern
University with a B.Sc.
in Civil Engineering.
In September, 1962, he entered Worcester Polytechnic Institute, Worcester, Massachusetts.
While at Worcester, he worked
as a teaching assistant in the Department of Civil Engineering.
In June, 1964, he graduated with a M.Sc. in Civil Engineering.
He joined the faculty at Northeastern University in September,
1964, where he has been a full-time member to present.
this time he also engaged in
During
consulting work in the field of Soils
and Foundational Engineering.
He entered the Massachusetts Institute of Technology in
September,1965~, as a special student.
In 1968, he was awarded a
N.S.F. Faculty Fellowship and assumed full-time study at M.I.T.
for that academic year.
He is an associate member of the American Society of Civil
Engineers.
He is also a member of Xigma Xi and Chi Epsilon, the
Civil Engineering National Honorary Fraternity.
-326-