THEE/“"3

This is to certify that the

dissertation entitled

BOND AND SLIP OF STEEL BARS IN FROZEN SAND

presented by

R IYADH M . ALWAHHAB

has been accepted towards fulfillment
of the requirements for

Ph.D. degreein C'iV'll 8: Sanitary
Engineering

cw R,MM

Major professor

Date November l0, 1983

MSU is an Affirmative Action/Equal Opportunity Institution 0-12771

 

 

 

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BOND AND SLIP OF STEEL BARS
IN FROZEN SAND

by

Riyadh M. Alwahhab

A DISSERTATION

Submitted to
Michigan State University
in partial fulfillment of the requirements
for the degree of

DOCTOR OF PHILOSOPHY

Department of Civil and Sanitary Engineering

1983

 

 

 

 

 

ABSTRACT

BOND AND SLIP OF STEEL BARS
IN FROZEN SAND

By
Riyadh M. Alwahhab

Structural members (reinforcement, piles, ground anchors) embedded
in frozen ground for the transfer and distribution of loads, tensile or
compressive, involve load transfer at their common interface. Engineer—
ing design requires prediction of both the short— and long—term adfreeze
bond strength for a range of freezing temperatures and sand-volume fract—
ions. The design method must account for surface type (steel, wood, con—
crete), roughness, loading geometry, and the presence of any impurities
in the sand—ice material. Adfreeze bond strength at the interface can be
described as the shearing stress between the structural member and surround-
ing frozen soil at failure. For steel bars embedded in frozen sand the
adfreeze bond includes three components: ice adhesion to the bar, friction
between the bar surface and sand particles, and mechanical interaction
between frozen sand and the bar surface roughness.

This experimental study involved measurement of the adfreeze bond
components using model structural members (steel bars) embedded in frozen
sand samples. Pullout loads, bar displacements, and temperatures were
monitored for both constant displacement rate tests and constant load (cre-
eP) tests. The observed behavior showed that initially ice adhesion combi-
ned with friction prevents slip. Very small bar displacements (about 0.002
in.) characterize the start of tertiary creep for shearing stresses below

those required for rupture of ice adhesion. After rupture and partial slip

 

Riyadh M. Alwahhab

a much smaller residual adfreeze bond controlled the pullout load. When

a single lug was added to the bar, bearing action of frozen sand on the
lug greatly increased the pullout load. Larger displacements with higher
adfreeze bond strengths are associated with mobilization of mechanical
interaction forces between the frozen sand and bar surface asperities. Lug
spacing for maximum adfreeze bond was shown to be a function of pressure
bulb overlap between lugs and formation of a void space behind lugs. Ex—
perimental relations presented showed the dependence of creep displacement
rates and/or adfreeze bond on lug size and spacing, sample dimensions,
water impurities, bar surface roughness, temperature, sand volume fraction,
and permitted development of empirical relationships for prediction of ad-

freeze bond strength.

 

TO MY FATHER WHO GAVE LOVE AND FATHERHOOD
THEIR IDEAL MEANING

T—f

ACKNOWLEDGEMENT

The author wishes to express a special appreciation to his major
advisor, Dr. 0.B. Andersland, professor of Civil Engineering, for his
encouragement, initiative, and aid throughout the author's doctoral
studies and for his guidance during the preparation of this thesis.

Thanks are also due the other members of the author's doctoral committee:
Dr. R.K. Wen, Professor of Civil Engineering; Dr. G.L. Cloud, Professor

of Metallurgy, Mechanics, and Materials Science; and Dr. C.O. Horgan,
Professor of Metallurgy, Mechanics, and Materials Science. Sincere
appreciation is also given to the author's father for his love and support
throughout the author's studies.

Thanks are also due the National Science Foundation, Division of
Engineering Research, and the Department of Civil and Sanitary Engineering
for their financial assistance which made this study possible at Michigan

State University. The author also wishes to extend thanks to the Univer-

sity of Technology in Baghdad for financial support.
Special appreciation is also given to Margaret L. Cridge for her love,

patience, and encouragement. Appreciation is also extended to the author's

close friend, Dr. Saad M. Alhir and his family, for their moral support.

 

 

 

LIST OF
LIST OF
LIST OF

CHAPTER
I.

III.

TABLE OF CONTENTS

TABLES ...................... . . .

FIGURES ................... . . . .

SYMBOLS ........................

INTRODUCTION .................. . . . . .

1.1 Ground freezing ............... . .
1.2 Ground reinforcement ..... . ..........
1.3 Scope of investigation ...............

REVIEW OF LITERATURE ...................

Ice adhesion ............ . . . . . . . .

Mechanical properties of frozen soil . . ......

Pile tests in frozen soils .............

2 1

2 2

2.3 Pile tests in ice ...............
2 4

2 5

Circular footing in frozen soil ...........

2.6 Frozen soil beam tests ..... , .........
MATERIAL PROPERTIES AND SAMPLE PREPARATION ...... .. .

3.1 Steel rods ................. . . . .
3.2 Ice samples ........ . ...... . . . . . .
3.3 Sand-ice samples ..................

EQUIPMENT AND TEST PROCEDURES . . . . . . . . . . . .

4.1 Equipment . . . . . . . . . . . . . . . . . . .
4.2 Test procedure . . . . . . . . . . . . . . . .

EXPERIMENTAL RESULTS . . . . . . . . . . . . . . . .
5.1 Constant displacement rate tests . . . . . . .
.1. Displacement (and loading) rate effects.
3%:
.1

1
2 Temperature effect ..............
3 Bar size and sample height effects ......
4 Lug size and shape effects ..........

78
8O

 

 

 

5.1.5 Bar surface roughness ...... . . . .
5.1.6 Water impurities and sand concentration . .
5.2 Constant load (creep) tests ...........
5.2.1 Load effect ........ . ..... . .
5.2.2 Temperature effect . . ..........
5.2.3 Lug size and position ...........
5.2.4 Sample diameter and sand concentration
VI. ANALYSIS AND DISCUSSION ..... . ..........
6.1 Load transfer mechanisms .............
6.1.1 Ice adhesion and sand friction ......
6.1.2 Lug bearing ......... . ......
6 1.2.1 Lug behavior in frozen sands . .
6.1.2.2 Lugs versus standard deformed bars
6.1.2.3 Long- -term lug bearing capacity.
6 1. 2. 4 Sample size and sand concentration
6 1. 2. 5 Lug bearing in ice ........
6.2 Creep versus constant displacement rate tests
6.3 Correlation of experimental results .......
6.3.1 Plain rods ................
6 3 2 Deformed rods ...............
6.4 Theoretical predictions .............
6.4.1 Bond for plain steel rods .........
6.4.2 Load prediction based on cavity expansion
theory ..................

6.4.3 Roughness criterion for steel rods

6.5 Applications ...................

6.5.1 Frozen soil reinforcement . . . . . . . . .

6.5.2 Pile or anchor capacity . . . . . . . . . .
VII. SUMMARY AND CONCLUSION . . . . . . . . . . . . . . . .
7.1 Test procedures . . . . . ........ . . .
7. 2 Load transfer mechanisms ......... . .
7. 3 Applications--reinforcement, piles, anchors
7.4 Recommendations for future work ....... . .

BIBLIOGRAPHY .................... . . .

Page
82
84

176

187
191
192
194
197
198
199

200
202

206
207

209
212

214

215
218

264
264
265
267
269

271

vi

Page
APPENDIX—-Data .................. . . . . . 281
A. Constant displacement rate tests .......... 282
B. Constant stress (creep) tests ........... 317

 

 

Table

5.1

5.2

5.3

5.4

5.5

6.1

6.2

6.3

6.4

6.5

A—l
A-2
A-3

LIST OF TABLES

 

Page
Title No.
Physical properties of ice, frozen sand specimens,
and steel rods for constant displacement rate
test ....................... 97
Summary of test results for constant displacement
rate tasts .................... 103
Temperature effect on the parameters used in
Equations 5.1 to 5.4 ............... 108
Physical properties of ice, frozen sand specimens,
and steel rods used in constant load (creep
tests ...................... 109
Summary of test results for constant load (creep)
tests ...................... 111
Comparison of frozen sand cohesion with the bond and
lug bearing capacity ............... L
Effect of lug size on creep rate and lug bearing in
frozen sand at -10°C ............... 223
Values of "pseudo" instantaneous displacement a. and
time to failure t at different loads for a 178 in. r
lug in frozen sand ................ 29

Comparison of the n parameter from unconfined com—
pression tests with that from bond tests . . . . . 226

Variation of ultimate load with rod surface roughness, 227

including lugs . . . . . . . . . . . ..... . .
Data for ice samples I—l to 1-21 (Ice 1) ...... 282
Data for ice samples I—22 to 1-27 (Ice 2) ..... 287

Data for ice samples I-28 to I—32 (Ice 3) .....

vii

 

Table
No.

A—4
B-l
B-2

viii

Title
Data for frozen sand samples ...........
Data for ice samples CI-l and CI—2 (Ice 3) .

Data for frozen sand samples ...........

Page

289
317
318

 

LIST OF FIGURES

 

Figure Page
No. Title No.
1.1 Frozen sand backfill with steel bars simulating

a reinforced simple beam for partial support of
a pipeline in areas of differential front heave

action .................. . . . 5
1.2 Ground reinforcement ............... 6
2.1 Adhesive strength of ice to different materials as

a function of temperature ....... . . . . 39
2.2 Effect of test type on the adhesive strength of

ice to polystyrene surface (after Jellinek

1957 b) .................... 40
2.3 Constant-stress creep test: (a) basic creep curve;

(b) true strain vs. time ............ 41
2.4 Creep strength vs. time to failure for frozen Ottawa

sand at -3.850C (after Sayles, 1973) ...... 42
2.5 Volume concentration of Ottawa sand and peak strength

(after Goughnour and Andersland, 1968) ..... 43
2.6 Results of strength testing on frozen soils at low

confining stresses ............. . . 44
2.7 Effect of confining stress on the strength of ice

and frozen soils . . . . . . . . . . . . . . . . 45
2.8 Effect of ice thickness and pile diameter on the

ice adhesive strength (after Frederking, 1979) . 46
2.9 Adfreeze strength for piles in ice as a function

of pile displacement rate (after Parameswaran,

1981) ..... ..... 47
2.10 Adfreeze strength, of silt—water slurry to an

8-in. diameter steel pipe pile, versus tempe— 48

rature (after Crory, 1963) ...........

 

 

 

 

Figure
No.

2.11

2.12

2.13

2.18

3.1

3.2

4.1

4.2

Title

Definition of terms used in the bond analysis of a
plain rod (pile) in frozen soil (after Johnston
and Ladanyi, 1972) . . . ..... . ......

Variation of adfreeze strength of piles in frozen
sand (solid lines) and contribution due to ice
adhesion present in sand (dashed lines) with
nominal pile displacement rate (after Parames—
waran, 1981) ..................

(a) Soil response to penetration by a flat cir-
cular punch, (b) and (c) Notations for
transformation of cavity expansion theory to
a deep footing problem (after Ladanyi and
Johnston, 1974) ................

Effect of footing depth on settlement of circular
test plates in ice at -2.3°C (after Vyalov et
al., 1973) ...................

Results of a loading creep test on a circular
plate (1.4 in. diameter, 18 in. deep) in frozen
sand (after Ladanyi and Paquin, 1978) .....

Creep deflection at the center of frozen sand beams
(after Klein and Jessberger, 1978) .......

Frozen Soil Beam; (a) Diagram of a Simply Supported
Beam; (b) Stress Distribution (after Klein and
Jessberger, 1978) ...............

Deflection creep curves for simply supported beams
of frozen sand at -10°C (after $00, 1983) . . .

Equipment for pull«out tests including a frozen
sand sample, steel rod with lug, and loading
frame immersed in the circulating coolant . . .

Bar configuration. (a) Plain rod. (b) Plain rod
with 90 degree lug. (c) Lug with 45 degree face
angle ................ . . . . .

Diagram of test system for constant displacement
rate tests ' 0 0 0 Q O O I O O i I V O U ' 1 C 0

Diagram of test system for constant load (creep)
tests . . . . . . . . . . ...........

Page
No.

49

50

51

52

53

54

55

62

63

68

69

 

 

 

Figure _ _ Page
No. . Title No.

5.1 Typical strip chart load and displacement records
for plain steel rods in ice and sand—ice speci-
mens ........... . .......... 118

5.2 Typical load-displacement curves at different
loading rates for the bond of frozen sand
(v5 = 64%) to a plain steel rod ........ 119

5.3 Effect of loading rate on the bond strength of
frozen sand (v = 64%) to a 5/8 in. diameter
plain steel rod at different temperatures . . . 120

5.4 Effect of nominal displacement rate on the bond
strength of frozen sand (vS =64%) to a 5/8 in.
diameter plain rod ............... 121

5.5 Typical load-displacement curves at different
displacement rates for a plain rod with a single
lug in frozen sand (v5 = 64%) ......... 122

5.6 Effect of displacement and loading rates on the
ultimate lug capacity in frozen sand (vS — 64%)
at —100c .................... 123
5.7 Typical load-displacement curves at different
temperatures for a plain rod in frozen sand r
= 64%) 124

ooooooooooooooooooo

 

5.8 Temperature effect on the bond strength of frozen
sand (vS =64%) to a plain steel rod, at several
loadingS rates ................. 125

5.9 Relationship between temperature and bond strength
of frozen sand (v 64%) to a 5/8 in. plain rod,
at several loading rates ............ 126

 

5.10 Comparison of strain hardening parameters, n based
' on constant loading rate and n' based on constant
displacement rate, for 5/8 in. diameter plain rod

in frozen sand (v5 = 64%) . . . . ........ 127
5.11 Temperature effect on the n parameter as defined by

Equations 5.1 and 5. 3 . ............ 128
5.12 Relationship between temperature and the n parameter

for a plain rod in frozen sand (vS =64%) . . . 129
5.13 Temperature effect on ice adhesion to a 5/8 in.

diameter plain rod ........... . . . . . 130

 

 

 

 

Figure . Page
No. T1tle No.

 

5.14 Typical load-displacement curves at different
temperatures for a plain rod with lug in
frozen sand (vS = 64%) .............. 131

5.15 Temperature effect on lug bearing, ultimate and
"initial yield" loads, in frozen sand (vS = 64%) . 132

5.16 Effect of rod size (diameter) on the bond strength
of frozen sand (v5 = 64%) to plain steel rods . . 133

 

5.17 Comparison of bond strength for two plain rod

diameters in frozen sand (vS =64%) at several

temperatures ................... 134
5.18 Relationship between the d/H ratio and bond strength

of frozen sand (vS = 64%) to plain rods ..... 135
5.19 Effect of sample height on bond strength for plain

steel rods in frozen sand (v5 = 64%) ....... 136
5.20 Load-displacement curves for a 3/8 in. steel rod with

four lug sizes in frozen sand (vS = 64%) at —10 C 137
5.21 Load- -displacement curves for a standard deformed

No.3 bar in frozen sand (vs — 64%) at —100C . . . 138
5.22 Effect of lug size (height) on the ultimate lug

capacity in frozen sand (vS = 64%) ........ 139
5.23 Typical load- displacement curves for plain rods

with two lug types, 450- -lug and 90°— lug, in frozen

sand (vS 64%) ................. 140
5.24 Relationship between temperature and lug bearing

capacity in frozen sand (v5 = 64% ........ 141
5.25 Roughness criterion (After Wright, 1955) ...... 142

5.26 Typical outputs for three types of steel surfaces as
recorded on the strip chart, with a list of their

properties . . ....... . ...... . . . . 143
5.27 Load- displacement curves for three types of steel
finish, in frozen sand (vS =64%) at —10° C. . . . 144

5.28 Effect of surface roughness on the bond strength of
frozen sand (vS = 64%) to plain steel rods . . . . 145

5.29 Bond strength dependence on ice sample preparation
method ...................... 145

 

 

 

X111

 

Figure
No_ T1tle
5.30 Bond strength dependence on water impurities for a
3/8 in. diameter plain rod in frozen sand (vS =
64%) at several temperatures ..........
5.31 Effect of sand fraction on bond strength for plain
rods, peak and residual at —10°C ........
5.32 Creep curves based on step loading a plain rod in
frozen sand (vS = 64%) at -10°C .........
5.33 Loading effect on the creep rate of a plain steel
rod in frozen sand (vS = 64%) at -10C .....
5.34 Creep curves for step loading of a plain 3/8 in.
rod with a 1/8 in. lug height in frozen sand
(vS = 64%) at —10°C ..............
5.35 Creep curves for step loading followed by unloading
a plain rod with one lug in frozen sand
VS=60 ........... . .......
5.36 Loading effect on creep displacement rate of a
plain rod with one lug in frozen sand (v5 =
64%) ......................
5.37 Creep curves for a plain rod in frozen sand (v5 =
64%) Step loading procedure ..........
5.38 Temperature effect on creep rate of a plain rod
in frozen sand (vS = 64%) ...........
5.39 Creep curves for step loading a plain rod with one
900 lug in frozen sand (v v5 = 64% ........
5.40 Creep displacement rates for a plain steel 3/8 in.
rod with a 1/8 in. lug in frozen sand vS = 64%)
at four temperatures . . . . ..........

5.41 Effect of lug height on the creepO behavior of a
plain steel 3/8 in. rod with 900 lugs in frozen
sand (vS =64%) at -100 C ......... . . .

5.42 Creep curves for step loading a standard No.3 bar
in frozen sand (vS —64%) at -10° C . . . . .

 

5.43 Load-displacement rate curve for a plain 3/8 in.
steel rod with a 1/8 in. lug at different
positions, in frozen sand (vS = 64%) at -10°C. .

 

 

 

 

Figure .
No. Title
5.44 Effect of sample diameter on creep rate of a plain
3/8 in. steel rod with a 1/8 in. lug in frozen
sand (v5 = 64%) at —10°C ............
5.45 Pull-out load dependence on soil cover for a plain
steel 3/8 in. rod with a 1/8 in. lug in frozen
sand (vS = 64%) .................
5.46 Creep curves for step loading plain bars with a 1/8
lug in frozen sand at —10° C .........
5.47 Creep curve for a plain 3/8 in. steel rod with a
1/8 in. lug in snow-ice (vS = 0) at -10°C . . .
5.48 Creep displacement rates for plain 3/8 in. steel
rods with a 1/8 in. lug in frozen sand at -10°C
and different sand volume concentrations . . . .
5.49 Load capacity dependence on sand concentration
during creep for plain steel 3/8 in. rods with
a 1/8 in. lug in frozen sands at —10°C .....
5.50 Comparison of lug loads at different sand concen-
tration for frozen sand at -10°C ........
6.1 Temperature effect on ice adhesion to different
materials ....................
6.2 Typical bond stress versus displacement curves for
plain steel piles (rods) in frozen sands . . .
6.3 Variation of bond strength with the ratio d/H for
ice and frozen sand (to piles of different
materials) ...................
6.4 Comparison of bond strength for different pile
types in frozen sandy soils ...........
6.5 Comparison of load-displacement curves for a plain
steel rod in ice and frozen sand . . ......
6.6 Comparison of ice adhesion and sand friction for a
3/8 in. plain steel rod embedded in ice and
frozen sand ...................
6.7 Comparison of lug bearing, sand friction, and ice

adhesion effects on the load—displacement curves
for steel rods .................

Page

No.

167

168

169

172

173

174

175

228

229

230

231

232

233

234

 

Figure

No.

6.8

6.9

6.18

6.19

 

6.20

ri—

XV

Title

Lug contribution for the initial bond rupture con—
tribution at different temperatures ......

Comparison of ice adhesion, sand friction, and lug
contribution for steel rods at different
temperatures and ultimate conditions .....

Lug contribution to the ultimate pull—out capacity
of a plain rod, with a single lug with different
lug heights, in frozen sand ...........

Comparison of frozen sand cohesion (at different
temperatures) with the bond and lug bearing
capacity ....................

Relationship between lug bearing capacity and
uniaxial compressive strength for frozen Wedron
sand ......................

Comparison of lug behavior in frozen Wedron sand
with that of a 1.4 in. diameter penetrometer in
frozen quartz sand at —60C ...........

Creep rate dependence on size (area) for lugs and
punches ....................

Pressure bulb overlap for consecutive lugs on a
3/8 in. diameter deformed bar .........

Time dependence of load relative to specified
creep displacements, 6f, at -10°C for a 3/8
in.)rod with a 1/8 in. lug in frozen sand (v5 =
64% ......................

Sand volume fraction effect on the unconfined
compressive strength of frozen Ottawa sand, and
the lug bearing capacity in frozen Wedron sand.

Bond strength for a plain rod in frozen sand, as
a function of displacement rates, for creep
tests and constant displacement rate tests. .

Lug capacity comparisons, creep tests versus
constant displacement rate tests, for different
test conditions ................

Modified relationship between bond strength and
the ratio d/H for plain steel rods in frozen
sand (vS = 64%) ................

Page
No.

235

236

237

238

242

243

244

245

246

247

 

 

Figure

No.

6.21

6.22

6.23

6.24
6.25

6.26

6.27

6.28

6.29

6.30

6.32

6.33

6.34

xvi

Title
_____.____.__._____________________.______.._____
Bond strength of plain steel rods in frozen sand

at —10°C as a function of sand fraction .

Comparison of Equation 6.2 with experimental data
for bond of plain rods in frozen sand (v5 = 64%)

Lug pull-out capacity as a function of temperature
for a displacement rate of 5 X 10'“ in./min.

Effect of lug size (area) on bearing pressure .

Lug capacity as a function of sand volume fraction
for a creep rate of 5 X 10’“ in./min ......

Load versus instantaneous displacement for a 3/8
in. diameter rod with a 1/8 in. lug in frozen
sand (v3 = 64%) at -1o°c ............

Time to 1 in. total displacement versus applied
load, for a 3/8 in. rod with a 1/8 in. lug in
frozen sand at —lO°C ..............

Comparison of experimental bond strength with the
theoretical predictions for a plain 5/8 in.
diameter rod in frozen sand (v5 = 64%) at
-10 C .....................

Comparison of n-values for bond tests with com—
pression tests on frozen Wedron sand ......

Comparison of Equation 2.16 with experimental
bond strength for a plain 5/8 in. diameter rod
in frozen sand (vS = 64%) at -10°C .......

Load predictions based on cavity expansion theory
for a single lug in frozen sand (v5 = 64%). . .

Zone of crushed sand observed around the lug, at
the end of creep tests .............

Correlation between profilometer and strip chart
readings of aspirity heights for different
steel surfaces .................

Roughness criterion (Figure 5.25) for a cold-.
rolled steel bar extended to include 1 1/8 in.
lug ......................

Page

248

249

250

251
252

253

254

259

260

261

 

 

xvii

 

 

Figure Page
No. Title No.
6.35 Ultimate pull-out capacity of steel rods in frozen

sand (v = 64%) as a function of rod surface
roughne s .................... 262
6.36 Conditions of equillibrium for a frozen sand beam

reinforced with a plain steel rod ........ 263

 

xviii

LIST OF SYMBOLS

Rod, pile, or footing radius, (d or B)/2,
Constant (Equation 5.5),

Constant (Equation 2.16), or beam width (Equation 2.20),
Constant (Equation 5.5),

Constant (Equation 2.16), or soil cohesion,

Long-term soil cohesion,

Rod (or pile) diameter,

Reinforcement depth in a frozen soil beam,

Flow value, (1 + sin¢)/(1 — sin¢),

Lug height, or frozen soil beam (cross-section) height,
Constant (Equation 2.3),

Constant (Equation 2.6),

Exponent (Equation 5.1),

Exponent (Equation 5.3),

Strain-hardening parameter, l/m,

1/m',

Confining stress, (01 + 03)/22

Cavity expansion pressure,

Overburden (confining) pressure,

Maximum shear stress, (01 - 03)/23 0* lug (or f°°ting) pressure,

"Proof” lug Pressure at e = 0°C and éc = 0.0005 in./min. (Equation

’

 

 

xix

"Proof“ lug pressure at temperature a and 3C = 0.0005 in./min.
(Equation 6.4),

Lug pressure, P1/A1,
Pressure under the first lug in a deformed bar,

Ultimate lug pressure, Pu/A],

Radial dimension in a cylindrical or spherical coordinate system,
Radius of a cylindrical plastic "cavity”

"Cavity” radius at the elastic—plastic boundary,

Radius of a spherical ”cavity”

Footing settlement,

Creep settlement rate,

Time,

Time to failure,

Time parameter (Equation 2.1),

Soil dependent parameter (Equation 2.1), exP(°o/Oi)
Vertical soil displacement at distance r,

Soil displacement at r=a,

"Cavity" expansion displacement,

Creep deflection at the center of a frozen soil beam,
Deflection rate at the beam center,

Sand volume fraction,

Horizontal coordinate,

Vertical coordinate, or lug position relative to the reaction plate,
Neutral axis position in a frozen soil beam,

Constant (Equation 2.7),

Lug area,

Rod area ,

 

P
co

C0
PC

XX

Footing Width, or equivalent footing width (= d + 2h),
Sample diameter,

Footing embeddment depth,

Soil modulus of elasticity,

Bond resisting force in tension zone,

Soil resistance in compression zone,

Resultant force in tension zone, (Fb + Ts),

Sample height,

Bond development length (embedded portion of rod in frozen sample
or beam),

Constant (Equation 2.9), U/R', OK,
Length of a standard chord,

Length of ”roughness" line,

Total length of standard chords,
Pull-out load,

Loading rate,

”Proof” load for any lug height, at 3C = 0.0005 in./min. (Figure 5.4)

Arbitrary loading rate, 1 lb./min. (Equation 5.1),

“Proof” load at e = 0°C and 8C = 0.0005 in./min. (Figure 6.23),
“Proof” load at temperature a and 8C = 0.0005 in./min. (Figure 6.
Load from constant load (creep) tests,

”Initial yield" load for deformed rods,

Lug load,

Long-term lug load,

Residual load for plain rods,

Ultimate load,

Load from constant 3,

 

Radius of beam curvature,

Universal gas constant,
Maximum and minimum aspirity heights observed on the profilometer,

Maximum and minimum aspirity heights as observed on the recorder
strip chart,

Temperature (0C or 0K),

Soil resistance in tension zone

Apparent activation energy,

Displaced volume of soil under the footing,
Displaced volume of an expanding ”cavity"
Wedging angle under the footing,

Angle of internal friction,

Long—term angle of internal friction,
Strain,

Creep strain,

Compression strain at the beam outer fiber,
Failure strain,

"Psuedo" instantaneous strain,
Instantaneous plastic strain,

Strain at the reinforcement level,
Instantaneous strain,

Strain rate,

Creep strain rate,

Arbitrary strain rate

Major principal strain rate,

 

—<o-<

—<o

xxii

Constant (Equation 2.3), or ratio (D-d)/2h,
Ratio 10 s/B (Equation 2.18), or ratio A1/Ar,
Constant (Equation 2.17),

Constant (Equation 2.4),

Constant (Equation 2.2),

Roughness factor,

Soil shear strain, du/dr,

Shear strain rate,

Arbitrary shear strain rate,

Absolute value of temperature below freezing,
Tangential coordinate,

Reference temperature, 1°C,

Stress,

Soil compressive strength,

Uniaxial compressive strength of soil at 0°C, and arbitrary
'strain rate EC,

Uniaxial compressive strength at temperature 6, and arbitrary strain
rate cc,

Long-term, or failure, soil strength,
Compression stress corresponding to ek.,
Instantaneous strength,

Soil dependent parameter (Equation 2.1),
Soil tensile strength,

Major and minor principle stresses,

Bond (or shear) stress,

Soil shear strength corresponding to ye,

Ice adhesive strength,

 

 

xxiii

Long-term bond strength,

Residual bond strength,

Frictional component of residual bond,
Ultimate bond strength,

Frictional component of ultimate bond,
Reciprocal of viscosity (Equation 2.2),
Rod (or pile) displacement relative to the frozen sample,
Creep displacement,

"Psuedo“ instantaneous rod displacement,
Rod displacement at rupture (Figure 5.1),
Rod displacement rate,

Creep displacement rate,

Arbitrary displacement rate,

Nominal (machine) displacement rate,

 

1.1 (in
The
eration
agent, 1‘
their c1
work car
light it
from 9
$0113 di
much m0,
general]
temperai
Projects
Usu
Competjt
bearing
ahigh w.
UneXPECtI
Construc:
and the 1

inside a

 

 

CHAPTER I
INTRODUCTION

1.1 Ground Freezing:

The principle behind controlled ground freezing is the use of refrig—
eration to convert in-situ pore water to ice. The ice becomes a bonding
agent, fusing together adjacent particles of soil or rock to increase
their combined strength and make them impervious. Excavation and other
work can then proceed safely inside of, or next to, a strong water-

tight frozen earth barrier. This frozen soil is analogous to perennially

 

frozen ground in permafrost areas. Artifically and naturally frozen

soils differ from unfrozen soils in that their mechanical behavior is

much more time and temperature dependent. The strength of frozen soils
generally decrease with time (due to creep) and increase with colder
temperatures. Thus, temporary ground freezing can be used on construction
projects.

Usually, one of the following conditions makes ground freezing more
competitive than other methods (Sanger and Sayles, 1978): a water-
bearing stratum is so deep that grouting is impractical; the lowering of
a high water table by pumping may damage adjacent structures; a sudden
unexpected flow of water into a tunnel or shaft has occurred during
construction and a watertight barrier is needed; permafrost is present
and the ground must remain frozen; a clear space (no bracing) is required

inside a cofferdam; or the site is for a below-ground cold storage tank

where the
for low—t6

Gener
provide is
shafts (Je
in this sc
competiti v
of creati 11
clear area

System (B

1.2 Groun

Previ.
that the t.
C0mllressiv.
”02911 501“
because of
“dated to
regions is
the pipem
501'] adjace

In 30”
3°”. the f
beam Under
“etessary 1

Steel rElnf

frozen Cant

 

 

 

where the frozen soil would serve as a watertight retaining structure
for low—temperature liquids.

Generally, controlled ground freezing has been profitably used to
provide temporary support for large open excavations, tunnels, and mine-
shafts (Jessberger, 1980; Klein and Jessberger, 1978). Recent improvements
in this science and its refrigeration techniques have made it a
competitive construction alternative. Ground freezing has the advantage
of creating no smoke, vibration, shock or undue noise. It also provides
clear areas within the excavation free of bracing and earth support

systems (Braun, et al., 1979).

1.2 Ground Reinforcement:

Previous work (Bragg and Andersland, 1980; Eckardt, 1982) has shown
that the tensile strength of sand-ice material is about one—fifth of its
compressive strength, particularly at high strain rates. Therefore, a
frozen soil mass (beam) subjected to an external bending moment may fail
because of a weakness in tensile strength. One special application,
related to natural gas pipeline construction in discontinuous permafrost
regions is shown in Figure 1.1 (a and 6). Operating temperatures for
the pipeline would be in the range of -800 to -12°C, thus granular bedding
soil adjacent to the pipeline would remain continuously frozen.

In some areas where ice lenses may form in the underlying unfrozen
soil, the frozen sand backfill behavior can be simulated by a simple
beam under a uniform load (Figure 1.1 c). In this case, it would be
necessary to use steel reinforcement in zones of tensile stresses.
Steel reinforcement would also be needed in other cases such as the

frozen cantilevered retaining wall shown in Figure 1.2 (a). Without

 

reinforce
compressi
Vertical i
and they 1
section.
reinforcec
frost (F11
be transit
Use c
Pipe piles
Successful
Luscher, 1
embedded 1'

Dull-out c

1-3 Scope
Time
the Struct
1°19” fol
and Strain:
temperatun
a"111313 a1
and “31151
Re matErie
We“ soil
studies are

With iOadt

 

 

reinforcement the frozen soil structural system would be limited to
compressive stresses associated with a curved or circular arch walT.
Vertical H—beams, could be easily installed as shown in Figure 1.2 (a),
and they would permit a frozen soil wall to be designed as a cantilevered
section. The space between H-piles could be designed as short non-
reinforced frozen soil beams. Steel piles are often installed in perma-
frost (Figure 1.2 b) so that large (compressive or pull-out) loads can

be transferred to the soil at a proper depth.

Use of corrugations can significantly improve the load capacity of
pipe piles in frozen granular soils (Figure 1.2 c). This technique was
successfully used on the trans-Alaska pipeline project (Thomas and
Luscher, 1980). In other cases, power—installed screw anchors deeply
embedded in warm permafrost (Figure 1.2 d), were used to provide a large

pull-out capacity (Johnston and Ladanyi, 1974) for power line anchors.

1.3 Scope of the Investigation:

Time and temperature dependence of frozen soil behavior means that
the structural analysis for a frozen earth ground suppOrt system would no
longer follow conventional analytic procedures. The geometry, stresses,
and strains at any point within the structure all vary with time and
temperature. In general, a non-linear viscoelastic—plastic finite element
analysis appears to be the future analytic method most effective for design
and analyses. These analytic methods require information on the sand—
ice material properties, bond adfreeze strength and creep behavior at the
frozen soil/reinforcement bar interface. Beam tests and analytical
studies are part of another project. This thesis is concerned primarily

with load transfer at the interface between frozen sand and the model

structura
Bond
force bet
includes
action be
bars depei
prinari 1y
This
friction,
roughness
tests incl
size, lug
diameter) ,
lni ti
Placement
Pull-out 1
01' from
”Mar DUI
creased in
C011111080115

b33811 0n CE

 

 

structural member.

Bond at the interface can be described as the shearing stress or
force between the reinforcement member and surrounding soil. Bond
includes three components: ice adhesion; friction; and mechanical inter-
action between frozen soil and the bar surface roughness. Bond of plain
bars depends primarily on the first two components. Deformed bars depend
primarily on mechanical interlocking for bond properties.

This experimental study involved measurement of ice adhesion, sand
friction, and mechanical interaction between frozen sand and surface
roughness of steel bars. Test variables affecting results of pull-out
tests included: temperature, bar displacement rate, loading rate, bar
size, lug size and shape, lug position, sample dimensions (height and
diameter), water impurities, steel surface type, and sand concentration.

Initially, adhesion of ice to the steel bar prevents slip. As dis-
placement increased friction and mechanical interaction contributed to the
pull—out load. When a single lug was added to the bar, bearing action
of frozen sand on the lug restrained bar movement giving rise to much
larger pull—out loads. Colder temperatures and high loading rates in-
creased the pull—out load. Considering lug bearing on frozen sand,
comparisons were made between observed and predicted pull-out loads

based on cavity expansion theory.

1m QCTFQCwn— wCO

. 5' ~|-.:.i...|...rlu...|ui Inn...|...||..l...||.| .|.|n |.'...l..l nI..| ..I..| .l. I .k

Fin.

. \4 @UMkiLIW UCJOLC

-

we ... Cu»

 

 

.H.H mczmwd

cowuom m>mma amocm mepcmcmmmwm mo mmmcm cw mam—mama m mo ucoaazm
Ewm cmucomcmmc m mcwpm_:Ewm mama mepm apwz Ppwwxmma acmm :mNocm

meucma Low Emma mfia
flea

cma pmmpm.i/K
LAT .m>mma pmogm mo mpcmoq
pm ampcoaasm Emma mnemormc<

 

 

"Ill'l'rl|'l"'llil

Emma mowimcmm

:a:an:aaa:a

mmoa mcwmmama

 

 

flay
Amv

m>mmmam mmcm
a pmoad m>mma pmocd
mmmcmF
mum compumm

m
\\\\ mo
0 “NU
[\lrklli \\ \\ \K\ \\\\ \ x s \K
. ... u .. . . .. .. . .. .... nrlnicma pmmpm
. . a .. /—U:mm CQNOLL

o b
O
o v
.

    

 

 

 

mW. mCPPmawa mmw

.. ...: ...... ....M ”Junk... “33.4.13“. flu.fl.|..u.....u.. n... 13.114 .l........ n\\\
\ \\ \\\ l\¢ \\\\V\Dmomt:muc:ogw+
m

 

\ \

 

Ch

 

 

freeze pipe‘ 'GSL
ll ///77—
cu
Frozen earth wall-— / . #5—- Steel H—beam
/ I
4'!

L x” you! __
Large open ; g. 2&3
excavation / "

/ I
m /'l
///’1// fill
'I
/
/'|
ll!
a II
II

 

 

 

 

 

(a)

Cantilevered frozen earth

 

Smooth pipe pile

Figure 1.2:

  
 

Corrugated pipe pile

wall with H-beam reinforcement

Permafrost

t
a? :2,

(d)

Screw anchor

 

Ground Reinforcement

conpl
the l
point
Postu'
consic
gators
Hosler
uniroz
I

to at ‘
termini
(Jenn
1960 a;
EXPEFin
Steel ,
muthe
Strengt
This in

“99] ll

 

 

CHAPTER II
REVIEW OF LITERATURE

2.1 Ice Adhesion:

Studies on the adhesive properties of ice normally deal with
complex physico-chemical phenomenon. There are numerous papers in
the literature which report on ice studies based on a fundamental view—
point. The existence of a liquid-like layer on the ice surface was
postulated by Faraday (1859; 1860). This concept has since undergone
considerable attention. Experiments carried out by several investi—
gators (Nakaya and Matsumoto, 1953; Hori, 1956; Jensen, 1956; and
Hosler et al., 1957) offer strong evidence for the existence of an
unfrozen transition liquid film on the ice surface.

It is currently agreed that liquid-like water exists on ice down
to at least -250C. However, it has been difficult to precisely de-
termine the nature, and properties of this unfrozen water
(Jellinek, 1967; Barnes et al., 1971). Jellinek (1957 a; 1957 b;

1960 a; 1960 b; 1960 c; 1970) carried out tensile and shear adhesion
experiments on ice frozen to various surfaces including stainless
steel, polystyrene, lucite, and fused quartz of varying surface
roughness. These results, summarized in Figure 2.1, show the adhesive
strength of ice to Various matenials as a function of temperature.
This figure indicates that the adhesive strength of ice to stainless

steel increases linearly with decreasing temperature until it becomes

 

larg

only
i nte
expe
occu
by H.
The

iinis
rougl

poly:

1957

obtai

conti

 

 

larger than the cohesive strength of ice.

Cohesive type breaks (i.e. within the ice crystals) were observed
only in the tensile experiments. Adhesive type breaks (i.e. at the
interface between the ice and solid) were observed in the shear
experiments. The transition from adhesion to cohesion appears to
occur at about —14°C. These results are very similar to those reported
by Hunsaker,et al. (1940) on the adhesion of ice to brass (Figure 2.1).
The transition temperature in the latter case was close to —12°C.

The adhesive shear strength was also strongly dependent on surface
finish, with strength decreasing significantly with decreasing
roughness. This is also shown in Figure 2.1, where the surface of the
polystyrene plate was much "smoother" than the stainless steel surface.
The adhesive strength obtained from tensile experiments (Jellinek,

1957 b) was observed to be about fifteen times greater than that
obtained from shear experiments (Figure 2.2).

To explain these results, Jellinek (1957 b) proposed that a

continuous liquid-like transition layer exists between the ice and the

material. For shear experiments, the adhesive strength was a function

 

of the interfacial strength, and consisted of a viscous contribution from

this thin film, as well as contribution due to surface roughness.
For the tensile adhesion experiments, the measured strength was con-
trolled by the ice cohesion as shown by the nature of the observed
breaks (mostly cohesive). Jellinek proposed that the ice/solid
interfacial strength was increased in the tensile mode due to the

reinforcing effect of surface "tension" at the perimeter of the

inten

Ii
unfroz
proper
induce
to dis
analys
been c
(1971)
sisten
Unfort
nechan
eXperi
unifon
0f the
caused
(1957 r

consisn

2-2 Me

Mec
Seviet
are a f
bOundar
0f Unfr

Ariders]

 

 

 

interfacial transitional water.

If the two-dimensional liquid model for absorbed water in
unfrozen soil were to be adopted (Martin, 1960), then these adhesive
properties may be readily explained: while it was relatively easy to
induce motions laterally parallel to the surface, it was very difficult
to displace the absorbed water normal to the surface. Other tests and
analyses related to the nature of the unfrozen water film in ice have
been carried out. These are summarized by Jellinek (1967) and Barnes, et al.
(1971). Generally, the results from these investigations are con-
sistent with the concept of a liquid-like transitional film in ice.
Unfortunately there are very few available data on this basic
nechanical property of ice and there is a lot of scatter in the
experimental results because of difficulty with control of the
unifonnity of the ice structure. Further, the extreme sensitivity
of the interfaces to any small impurities, such as solutes or gases,
caused poor reproducibility of the experimental results. Jellinek
(1957 b) duplicated each data point 12 times to obtain the reasonably

consistent variation shown in Figure 2.1.

2.2 Mechanical Properties of Frozen Soil:

Mechanical properties of frozen soils have been studied in the
Soviet Union as early as 1920 (Tsytovich, 1975). Most frozen soils
are a four—phase material; containing solid particles, ice, gas, and
boundary water in equillibrium with ice (Vyalov, 1965). The amount
of unfrozen water in frozen soil depends on the soil type (Dillon and

Andersland, 1966), temperature (Nerseova and Tsytovich, 1963),

 

grai
pres
uate

llers

facb
rate
501 )1
corn
coupe

quart

deper
rUptu
Press
C0ntr
in pa
for c
(Vial
With

8 up
(Obta'
Primal
and ii

Anders

 

IrIIIIIIIIII____________________________________——_

10

grain-to—grain contact pressure (Tsytovich, 1975), and hydrostatic
pressure (Chamberlain,et al., 1972). In frozen sands nearly all
water is frozen at 0°C under atmospheric pressure, as reported by
Nerseova and Tsytovich (1963) and Dillon and Andersland (1966).

The mechanical properties of frozen soils are dependent on nany
factors including tine, temperature, soil type, frozen and unfrozen
water contents, ice content, etc. Owing to the varying nature of
soils and experimental conditions, it is difficult to compare or
correlate the data obtained by different authors. However, one can
compare the results obtained for materials of the same type (such as
quartz sands) tested under similar conditions.

The strength of frozen soils may take on different meanings
depending on the engineering problem. It includes the concept of
rupture and that of excessive deformation (Andersland,et al., 1978).
Pressure melting of ice at points of contact with soil particles
contributes to plastic defbrmation of the pore ice and readjustment
in particle arrangement. .This mechanism appears to be responsible
for creep of frozen soil under constant stress and temperature
(Vyalov, 1965; Andersland,et al., 1978). The increase in deformation
with time, under constant stress and temperature, is represented by
a typical creep curve in Figure 2.3 (a). After an initial strain eo
(obtained immediately upon loading), the creep rate decreases over the
primary stage, remains approximately constant during the secondary stage,
and increases over the tertiary stage until terminated by rupture.

Andersland (1963) and Vyalov (1963) attribute this creep behavior mainly

 

 

partic
unent
(lysto
1965),
streng
theory
an ang
depend
that f
tenn s
Salles
Vyalov

rePtesc

Where 1
tr is 1
(01.031

t* equa

 

to the change in soil structure, otherwise, the creep rate would have
renained constant. Refreezing of liquid water is also considered to be
a healing mechanism which contributes to the time-hardening process
(i.e. decreasing creep rate) over the primary stage, Figure 2.3 (b).

The strength of frozen soil develops from interparticle friction,
particle interlocking, and cohesion. The latter is related to ice
cementing of soil particles and from the shear strength of ice
(Tystovich, 1958; Vyalov, 1965; Sayles, 1968). Tests by Vyalov (1959,
1965), Tsytovich (1966), and Sayles (1973) have shown that the shear
strength of frozen soil can be described, using the Mohr—Coulomb failure
theory, by a normal stress on the failure surface and two parameters;
an angle of internal friction and cohesion. Time and temperature
dependence is governed primarily by the cohesion. It is recognized
that frozen soil has a high instantaneous strength and a smaller long—
term strength. Test results (Figure 2.4) on Ottawa sand at -3.85°C by
Sayles (1973) show a reasonable agreement with Vyalov's long-term strength.
Vyalov (1959, 1963) suggested that the variation of strength can be

represented by:
0' = 00 - 2 00
f lnE(tfl‘t*)/td] ln(tf/to) (2'1)

 

where the parameters 00 and t0 depend on soil type and temperature,
tf is the time to failure, of is the long-term strength and equals
(01-03) as modified by Sayles (1973) to allow for triaxial conditions,

t* equals exp (00/01) and °i is the instantaneous strength.

Where (
Mdifi.

frozen

 

 

12

For long-term strengths Vyalov (1963) stated that the quantity t*
may be neglected. This appears to be applicable to frozen soils
where either primary or secondary creep dominates the deformation
behavior. Equation 2.1 with tf equal to infinity results in a long-term
strength equal to zero, which is not consistent with the idea of con—
tinuous strength at some finite time. Vyalov (1963) stated that the
idea is purely conventional, but in practice, after some long period of
time, the additional strength reduction is so insignificant and so slow
that the reduction can be neglected in engineering calculations. Long-
term strengths reported by Vyalov (1965) ranged from 18 to 37 percent
of the instantaneous cohesion for a frozen sandy silt. Vyalov (1963)
pointed out that with a stress level below the long-term strength the
strain rate would not change to the stationary (or secondary) stage, but
would remain in the primary stage where it continues to decrease until
it becomes zero. If a = of, then the rate goes to zero, according to the

following equation:

Q.
r4.»

= E = 5 (0' 0r)“ (2.2)
where u > 1 and g is the reciprocal of viscosity. Equation 2.2 is a
modified version of the primary creep model Vyalov (1965) has proposed for
frozen soils and which has been successfuly applied to various frozen soils
by Sayles (1968, 1974) and others. In this model, the strain is expressed as:
atA W (2.3)

)k

6:8(t)=(w(e+e
0

where 6 equals -T, 90 is a reference temperature, usually 1°C, and

in:

Note
thus
hard

iron

When
Proo

59) er

 

 

 

 

13

m, k, 1, m are soil parameters.

By differentiating Equation 2.3 with respect to time, the strain rate
é may be obtained:

8 = _%.( o k )l/m t(A—m)/m (2.4)

m( 6 + 00)

Note that the strain rate é decreases continuously with time, for A < m,
thus confirming that Equation 2.3 represents the primary (strain and time
hardening) creep model. The parameters in this equation may be obtained
from a series of log-log curve fits based on experimental data.

Secondary creep models for frozen soils have been proposed by Ladanyi

 

(1972) based on work in metals by Hult (1966), Odqvist (1966) and others.

One model is based on a power law which may be written as:

$F-”=° (/ )"
dt E 8c U Ocue (2.5)
where “cue and n are creep parameters, both dependent on temperature. The

proof stress 0 (Hult, 1966) is the uniaxial stress for an arbitrarily

cue
selected creep rate EC. Then, for a given material at a constant temperature,

the total strain e(t) is:
an=9+ 8 mm M'+té<mofl (2m
E k' k' C C '

where the first two terms represent the pseudo-instantaneous elastic and
plastic strains respectively (60 in Figure 2-3 a) 33 proposed by Ladanyi
(1972). Vyalov (1959) provided experimental data showing that for time
intervals greater than 24 hours these two terms together become less than
10 percent of the total creep strain. For engineering applications where

time intervals are greater than one day, use of the third term only in

Equation 2.6 gives good approximations of the total strains (Andersland,

et al.,

<
reactior
the temp
llkili, l
1970; Ho
Eyring (
by Mitch
narized

l
(oi role
strained
riers se
flow uni
0i suifi.
energy")
Derature

becomes (

E

Umg EqL

been done

where U 1'

Stan, an

and U/Rn

 

 

 

 

14

et al., 1978).

Several authors have attempted to apply the theory of absolute
reaction rates, commonly known as the rate process theory, to explain
the temperature—dependence of strength for frozen soils (Andersland and
Akili, 1967; Goughnour and Andersland, 1968; Andersland and AlNouri,
1970; Hooke, et al., 1972). The theory was originally formulated by
Eyring (1936) and later by Glasstone,et al. (1941). It was also applied
by Mitchell,et al. (1968) and Mitchell (1976) to unfrozen soil, and sum-
marized by Andersland and Douglas (1970).

The rate process theory is based on the assumption that flow units
(of molecular size) participating in a deformation process are con-
strained from movement relative to each other by virtue of energy bar-
riers separating adjacent equilibrium positions. The displacement of
flow units to new positions requires their activation through acquisition
of sufficient energy (thermal or mechanical called "the free activation
energy") to overcome an energy barrier. In analytical terms, the tem-
perature dependence of the creep rate é, when other factors are constant,

becomes (Ladanyi, 1972)!
e = A exp(-U/R' T) (2.7)

Using Equation 2.7 together with the power law (Equation 2.5), as has

been done for ice (Glen, 1955; Gold, 1970), one can write:

8 = & (o/oc)n = A exp(-U/R' T) (2.8)

where U is the apparent activation energy, R' is the universal gas con-
stant, and the other parameters are as defined before. The constants A

and U/R' can be found for different temperatures by plotting the

 

natur
of th

nanip

Oc6

where

streu

l
applie
to cre
to Hoe
the ac
corres
the th
but in
frozen
iron 0
be dis

DOUg)a

A
for a,
In this
Strain

absOrba

 

 

 

 

15

natural logarithm of the observed creep rates against the reciprocal
of the absolute temperature. Finally, with some mathematical

manipulations Ladanyi (1972) derived the following relationship:
oce= oCO exp(L'o / 273n(273—0))= ace exp(L'o / (273)2n) (2.9)

where L'= U/R'with units of temperature, a: (273—T) 0k = proof

, 6C0
stress at -273°k.

While the rate process theory has fundamental significance when
applied to reaction rates in pure chemical systems, its application
to creep of frozen soils is primarily empirical in nature. According
to Hoekstra (1969), Andersland and Douglas (1970), and Ladanyi (1972),
the activation energy calculated from Equation 2.7 does not necessarily
correspond to its physical definition, since in frozen soil not only
the thermal energy of the moving molecules changes with temperature
but in addition a gradual phase change also occurs. Moreover, in
frozen soil a mechanically activated process (moving a soil particle
from one equillibrium position to the next) is involved, which should
be distinguished from a thermally activated process (Andersland and

Douglas, 1970).

A more sophisticated ”tertiary creep” model which does account
for accelerating creep was proposed by Goughnour and Andersland (1968).
In this model, the plastic strain rate is composed of the sum of a
strain hardening term and a softening term which is dependent on the

absorbed strain energy. Since integration of their (strain-based)

enpi'
enpii
node'
(1961
strer
strer
rates
sligl
sand,
obser
(due
llilat
betue

titer

 

 

empirical model was quite difficult, the two authors proposed a separate
empirical relation in terms of time. Excellent correlations between their
model and data on Ottawa sand was reported. Goughnour and Andersland
(1968) studied the effect of sand concentration on the frozen soil
strength. They obtained a bilinear relationship between the peak
strength and sand concentration at different temperatures and strain
rates (Figure 2.5). For low volume fractions of sand there was a

slight linear increase in strength. When a critical volume fraction of
sand, about 42 percent, was reached a rapid increase in strength was
observed. At this point friction between sand particles and dilatancy
(due to particle interlocking) began to contribute to the shear strength.
Dilatancy must act against cohesion of the ice matrix and the adhesion
between sand grains and ice, thus creating an effect analogous to higher
effective stresses (Goughnour and Andersland, 1968).

The effect of strain rate and temperature on strength is also shown
in Figure 2.5. These two factors, strain rate and temperature, have
been studied by several investigators (Vyalov, 1959; Sayles, 1966,

1968; Sayles and Epanchin, 1966,; Goughnour and Andersland, 1968;

Alkire and Andersland, 1973; Ladanyi and Arteau, 1978; Eckardt, 1979;
Bragg and Andersland, 1980; Parameswaran, 1980; Parameswaran and Jones,
1980). Their experimental results showed that there was a general trend
of increasing strength upon increasing strain rate and decreasing
temperature below 0°C.

Other factors such as confining pressure (Andersland and Al—

Nouri, 1970; Chamberlain,et al., 1972; Alkire and Andersland, 1973)

 

St)
am
on

Fig

 

and sample size (Bragg and Andersland, 1980) on the strength of
frozen soils have also been investigated. A summary of the results
for strength testing by various investigators at low confining
stresses is shown in Figure 2.6. Strength increases almost linearly
with confining pressure. Note that the apparent friction angle
for frozen sands is similar to that for unfrozen sand; however, a
sigificant apparent cohesion exists in the frozen sand.

Chamberlain,et al. (1972) reported that at high confining
stresses (greater than 10 kips/sq.in.) the shear strength begins to
decrease (region II in Figure 2.7). The authors attributed the
decrease in strength to the suppression of dilatancy due to a
reduction in particle size caused by crushing under high pressures.
Pressure melting of ice in frozen soil is also expected when
hydrostatic or deviatoric stresses are applied. It develops from
stress concentrations on the ice component between soil particles
and from hyrostatic pressure on ice. Since global pressure melting
occurs at a pressure of 15 kips/sq.in. and more (region III in
Figure 2.7), the soil is analytically similar to unfrozen soil where
friction becomes the major component. Since the strength measured
in region III was relatively low, the observed quantity must be an
undrained shear strength at effective stresses much less than the
total stresses.

Strength tests on frozen soils are generally of two types: the
constant strain rateand constant stress (creep) tests. Odqvist

(1966) presents a theoretical discussion of the relationship between

 

 

18

these two test types. Since engineering strain (in the compression test)
is nearly equal to the true strain for small strains, and if primary
creep may be neglected, the peak stress (at failure) determined from
constant strain rate tests may be related to the creep parameters
determined from creep tests as shown below:

E =Ec(of/oc)n (2.10)
where of is the failure stress, 0c and n are the creep parameters
defined in Equation 2.6, and g is the applied strain rate. This
implies that the two parameters Cc and n should be equivalent for
constant strain rate and constant stress creep tests. Some authors
(Goughnour, 1967; Bragg, 1980) experimentally found that the constant
strain rate strengths of frozen sands were considerably higher than
the corresponding strengths in creep tests. Goughnour (1967)
attributes this disagreement to the possibility that higher plastic
strain energies would be consumed in strain—controlled (constant strain
rate) tests than would be consumed in stress-controlled (creep) tests,
considering the stress-strain curve. Other experimental data on
POlycrystalline ice at —7°C (Hawkes and Mellor, 1972) indicated that
no significant difference existed between the results of constant

stress creep tests and constant strain rate tests at low strain rates.

2.3 Pile Tests in Ice:
As a result of ice adhesion a floating ice cover can develop
substantial vertical loads on a structure to which it is frozen, due

to changes in water level. In the analysis of vertical ice loads on

 

 

piles, Lofquist (1951) assumed that ice has a viscoelastic behavior
with deformation proportional to the cube root of time. A means for
predicting vertical load was given, but no direct field or laboratory
measurements were available to confirm the prediction. An almost
similar analysis was made by Sanger (1969), giving average values for
the ultimate adfreeze bond strength for saturated fine sands, saturated
silts with ice lenses, and ice. Sanger cautioned that the strength
values should be checked by field tests for final design.

Stehle (1970) has reported that short-term tangential adfreeze
bond strengths of about 290 psi were observed for wooden piles
(2.75-in. diameter) in ice. The bond stength varied linearly with
pile settlement rate on a log-log scale SUQQEStlng a power 16W
similar to that in Equation 2.5. Measurement of vertical ice loads
developed on instrumented piles have also been reported by Doud (1978).
A maximum load of 18 kips on a 14 in. diameter steel pile was developed
by a 13.7-in. thick ice cover.

Frederking (1979) carried out laboratory tests to study the effect
of ice thickness, pile diameter and displacement rate on downdrag
loads developed on a vertical wood pile. These model piles, with
diameters ranging from 2-in. to 6-in., were used with ice thicknesses
from 1 to 8 inches. The pile load was observed to increase with
increasing displacement rate. Unit pile load also increased with
increasing ice thickness and decreased with increasing pile diameter
as shown in Figure 2.8. Observations of the failure process for that
test series showed that a combination of flexural and shear behavior

was present in all cases. The relative contribution varied with test

and
shear
Freda
ice 2

rate,

 

 

20

conditions, flexure predominating for the thinner ice sheets and
shear for the thicker ice. Combining the effects of all test variables,
Frederking (1979) formulated a composite empirical equation for the
ice adhesive strength to wooden model piles as a function of displacement
rate, ice thickness, and pile diameter.

Parameswaran (1981) conducted constant displacement rate tests
to evaluate the adfreeze strength of wood, concrete, (both 3—in. in
diameter) and steel H—section (4-in. wide flange) piles in ice at -600.
Displacement rates ranged between 4 X 10'6 and 4 X 10'3 in./min.
(10"t to 10'1 mm/min.). The test results, summarized in Figure 2.9,
include the three pile types. These curves show an almost linear variation
of adfreeze strength with displacement rate leading to the conclusion
that a power law, similar to Equation 2.5, can be derived relating
the adfreeze strength to the displacement rate. Note that the
adfreeze strength of smooth concrete piles in ice was independent of
the displacement rate.

To check this behavior Parameswaran (1981) carried out a few
tests using a smooth painted steel cylindrical pile in ice.
Preliminary results from these tests showed a rate independent of
adfreeze strength. For these piles (concrete and painted steel with
smooth surfaces) Parameswaran (1981) suggested that the adfreeze strength
probably is decided by the first major crack developed at the pile—
ice interface. Parameswaran (1981) used his test results to estimate

the contribution of ice adhesion to the total adfreeze strength of

 

var

and
Unit

Doss

 

21

various piles in frozen sand, as described in the following section.

2.4 Pile Tests in Frozen Soils:

The shearing resistance developed at the pile—soil interface has
two components: ice adhesion to the pile; and soil grain friction
at the pile-soil interface. End bearing is normally excluded from
design considerations because of the greater movement required to
mobilize end bearing resistance in comparison with much smaller
movements for adfreeze bond failure (Crory, 1963; Nixon and McRoberts,
1976; Nixon, 1978; Phukan and Andersland, 1978; Morgenstern,et al.,
1980; Weaver and Morgenstern, 1981 b). This point will be discussed
in more detail later.

Although measurements of the adfreeze strength between piles
and frozen ground have been taken since 1930, mainly in the Soviet
Union, sufficient information is not yet available to make it
possible to design for different foundation materials in various soils.
The results of early work in the Soviet Union from short—term tests
have been given by Tsytovich and Sumgin (1959), Tsytovich (1975), and
Vyalov (1965). Crory (1963), Crory and Reed (1965), Sanger (1969),
and Nixon and McRoberts (1976) have reported the results of pile load
tests in the permafrost regions of North America. Most of the available
adfreeze strength for piles were obtained from field tests of piles
in ice—rich soils (i.e. soils which possess a continuous network of
segregated ice, and usually with bulk densities less than 106 lbs./
cu. ft., as categorized by Weaver and Morgenstern, 1981 a).

Long-term pile creep tests were conducted in ice-rich frozen soils

by se
data

facto
and i
appli
loadi
facto
steanr
gives

of the

1
taking
0f ins
holes
dianet
lncrea

Or
of ful‘
Pipe 1
a diame
Pile an
moist”r
Flgure

that c)

 

22

by several investigators. An independent comparison of the published
data is difficult because the pile behavior depends on a number of
factors, the more important including pile type, type of frozen soil
and its temperature, method of installation, and the rate of load
application. Vyalov (1959) performed a large number of full-scale
loading tests in which he studied the effects of most of the above
factors. For timber piles (6-in. diameter) hand—driven into pre-
steamed holes, in both silty sandy loam and agrillaceous loam, Vyalov
gives the following empirical formula for the temperature dependence

of the long-term adfreeze strength, r1t:

a” (kg/cm2)= [176551155 - 0.3 (2.11)

Vyalov (1959) also proposed a series of correction factors,
taking into account the effect of different soil types and methods
of installation. For example, for piles driven into dry-augered
holes having a diameter 10 to 40 percent smaller than the pile
diameter, he recommended that T1t values from Equation 2.11 be
increased by 30 to 50 percent.

0n the other hand, Crory (1963) reports on the results of a series
of full scale tests carried out in Alaska on 8-in. (20 cm) diameter
pipe piles installed in frozen silty sand in dry-augered holes having
a diameter about 2-in. greater than the pile. The gap between the
pile and the soil was filled with a silt-water slurry (40 to 80 percent
moisture content). The adfreeze strength of this slurry is shown in
Figure 2.10 as a function of temperature. Crory (1963) also reported

that clay—water slurries have adfreeze strengths which approach that

 

of i
silt
In c
that

stre

by Jr
at ti
site

about
clay

incre
year,
secon
Pile

diffe
all t.
by an
deiorr
Plasti

in Fig

Shear
homoSte

1epath

 

23

of ice, and have an ultimate strength of approximately half that of
silt. The behavior of a frozen dry sand backfill was like thawed sand.
In contrast, Crory reports that low-temperature pile tests disclose
that a well-graded sand slurry, vibrated in place, has an adfreeze
strength 50 percent higher than its silt counterpart at the same temperature.

Long-term creep tests were also carried out on grouted rod anchors
by Johnston and Ladanyi (1972). A power auger (6-in. diameter) was used
at the Thompson site, whereas a hammer drill was used at the Gillam
site to make holes for the anchors. The grouted lengths ranged from
about 8.5 ft. to 10.5 ft. In both sites, the soil was ice-rich varved
clay and silt at temperatures from ~0.11°C to —.28°C. Step-wise
increasing loads, applied for time periods from 1 hour to more than a
year, permitted observation of all three stages of creep (i.e. primary,
secondary, and tertiary). The piles failed by slip along the soil-
pile interface. The slip failure occurred at different times under
different loads, but the total displacement was of the same order for
'all tests. A theoretical analysis of the test results was also made
by Johnston and Ladanyi (1972). In their analysis, the frozen ground
deformation around a pile shaft was idealized as shearing of a visco-
plastic half space medium around an elastic cylindrical rod as shown
in Figure 2.11, thus following the work of Nadai (1963).

Also, the analysis involved several assumptions including uniform
shear stress distribution over the rod (pile) length, weightless and
homogeneous soil, and a constant temperature over the embedded rod

length. Using a cylindrical coordinate system (r, e, and z), and

assurn
strai a
due tr
under
of a p
define
h as l

(1972)

.1,
in whi

creep

 

24

assuming a vertically loaded pile as shown in Figure 2.11, the

strains around the pile in the tangential e-direction will be zero

due to symmetry. Consequently each element of frozen soil deforms
under plain strain by simple shear. If u denotes vertical displacement
of a particle at radius r, the corresponding shear distortion y was
defined (Nadai, 1963) as (-du/dr), and similarly the shear strain rate
; as (—du/dr). In order to solve the problem, Johnston and Ladanyi
(1972) introduced the following stress—strain relationship:

i= term/TC)" (2.12)
in which Tc and n are creep parameters determined in a simple shear
creep test, and Tc is an arbitrary shear strain rate. From Figure
2.11, substituting (—du/dr) for )and ta(a/r) for Tin Equation 2.12,
then integrating, and noting that at r=a, o=ua, and at r=w, u=o,
gives:

da= i, many/(m1) (2.13)

Note that if there is no slip between the rod and the soil, i.e.
before failure, then s=oa, where Sis the rod displacement rate. Equation
2.13 shows how the time—dependent rod displacement is related to average
creep properties of frozen soil obtained from a simple shear test.

For a general state of stress and a plastic (incompressible) von Mises
material subjected to simple shear, under plane strain conditions,
Johnston and Ladanyi (1972) showed that the generalized creep equation
reduces to:

i: 3(n+1)/2 gc (T/Ocue)n (2.14)

where
tests
one f

1
This 1
estima
for t1
with e
same a
creep
(in uh
creep
the po

N
the th
in rea
Use of
to‘gra'
cOrlsen
the twc
tests,

Mr
W sugg
SGconda

theSe 1

 

 

25

where the proof stress “cue is obtained from uniaxial compression creep
tests at a constant temperature 6. Comparing Equations 2.12 and 2.14,
one finds that if Ocue is chosen equal to Tc’ then:
,C = 3(n+1l/2 éc (2.15)
This analysis assumes a constant n-value for a particular soil. The
estimated long-term adfreeze strength, based on this analysis, computed
for the rod anchors at the Thompson and Gillam sites compared well
with experimental data (Johnston and Ladanyi, 1972). Following the
same analysis, Nixon and McRoberts (1976) proposed a two-term power
creep law for adfreeze strength of piles in ice and ice—rich soils
(in which secondary creep dominates). The authors used existing
creep tests from many different sources to obtain the constants in
the power law.

Nixon and McRoberts (1976) and later Nixon (1978) showed that
the theoretical predictions provided by the two-term power law were
in reasonable agreement with test results from other published data.
Use of the same law for dense ice-poor soils (soils for which grain-
to-grain contact may influence soil behavior) would be overly
conservative, since in such soils primary creep dominates whereas
the two-term creep law was based on steady-state (secondary) creep
tests, with two deformation mechanisms.

Morgenstern et al. (1980) and Weaver and Morgenstern (1981a and 1981
b) suggest that in ice-poor soils displacement rates would not reach
secondary creep under loads normally encountered in the field. Since

these loads are probably smaller than the long-term strength of dense

(El

bor
day
frc
soi
tio
uhi

DUr

deVl

 

26

frozen soils, soil creep would stay within the primary stage,
region I in Figure 2.3. Therefore, Weaver and Morgenstern (1981 b)
used the primary creep law (Equation 2.3) and adopted the same
approach as that suggested in Figure 2.12, in order to recommend the
following relationship:

a/ atb = 3(“1”2 (r/ m(e+1)k

)C/ (c-l) (2.16)
where b=A/m, c=1/m, k, m, A, and m are the soil parameters defined
in Equation 2.3. In their analysis, Weaver and Morgenstern (1981 b)
assumed constant values for b, c, k, and m of 0.263, 1.32, 1.0, and
21.0, respectively, based on experimental data available in the
literature (Sayles, 1968)for frozen Ottawa sand. Using these values
in Equation 2.16, the two authors have suggested a design chart for
friction piles in frozen Ottawa sand.

Direct simple shear tests were also conducted on a variety of
reconstituted frozen soils and ice (Weaver and Morgenstern, 1981 a)
to investigate the load transfer process associated with adfreeze
bond to piles. Tests were performed at —1°C for durations up to 45
days to explore ultimate creep rates. These experimental data on
frozen soils show how creep rate tends to decrease with increasing
soil density except for dirty ice (ice with very low solid concentra~
tions and typical bulk densities between 59.0 and 62.4 lbs./cu. ft.)
which displays creep rates slightly higher than those observed for
Pure ice (Weaver and Morgenstern, 1981 a).

Weaver and Morgenstern (1981 b) later extended procedures

developed previously for predicting settlement and adfreeze strength

of pi
that
long-
rough
both

expre

where
a roug
assume
normal
reduce
tlpica
1.2 b)

A
betwee
the in
Concre‘
anlnnr
(1.2 my
"frost
increag
lncreas
(1978 a

Piles o

 

 

 

 

 

27

of piles in ice-rich soils to include ice-poor soils. They proposed
that the adfreeze strength of a frozen soil, T, is related to the
long-term shear strength, Tlt’ by T=KT1t, where K is primarily a

roughness parameter. The long-term shear strength I consists of

lt
both frictional, ¢lt’ and cohesive, Clt’ components and can be

expressed by the Mohr-Coulomb relationship, so that r becomes:
T : K(C]t + 0 tan ¢]t) (2.17)

where o is the normal stress on the shear plane and K'is primarily
a roughness-dependent parameter. Weaver and Morgenstern (1981 b)
assumed that the frictional omnponent may be neglected, since the
normal stress is typically less than 14.5 psi. Hence, Equation 2.17
reduces to e=Kc]t. Using published data the authors suggest

typical values for K ranging from 0.6 for plain steel piles (Figure
1.2 b) to 1.0 for corrugated pipe piles (Figure 1.2 c).

A systematic laboratory measurement of the adfreeze strength
between piles and frozen soils is not available. Laba (1974) measured
the instantanuous adfreeze force (frost grip) between frozen sand and
concrete under laboratory conditions. Frozen sand was pushed from
an inner cylindrical cavity in a concrete block at a rate of 0.0472 in.
(1.2 mm) per minute. These results indicate, in general, that
"frost grip" decreased rapidly with increasing porosity, that it
increased with increasing ice content in the sand and that "frost grip”
increased with decreasing temperatUre. Recently, Parameswaran
(1978 a, 1978 b, 1979) conducted laboratory experiments on cylindrical

piles of wood, concrete, and steel(cylindrical and H-section)in frozen

 

i Ct
sar

ice

was
1101)

Strr

that

but

 

 

 

 

 

Ottawa sand (14 percent moisture content). Each cylindrical pile
was 3¢in. in diameter. The H-section pile was a wide flange steel
H-beam with a‘44in. depth of section and 12-in. in length. Test
results indicate that the load dependence for steady-state creep
rate of piles in frozen sand agrees with the displacement rate de-
pendence of the adfreeze strengths determined by extrapolation of the
results of constant rate tests. In general, maximum adfreeze
strength developed with natural B.C. fir, and the minimum with
creosoted B.C. fir. Based on a linear variation of adfreeze strength
with either the creep or displacement rate on a log-log scale,
Parameswaran (1979) deduced a simple power law similar to that suggested
by Johnston and Ladanyi (1972).

More recently, Parameswaran (1981) estimated the contribution of
ice adhesion to the total adfreeze strength of similar piles in frozen
sand. Returning to Figure 2.9, Parameswaran reduced the values for
ice adfreeze strength by the ratio of the volume of water present
in sand, packed round the pile to a 7.5—in. hETth, to the
volume of water filled to the same height. This reduction factor
(0.2377) was used because the ice in a later study (Parameswaran, 1981)
was of a different structure from that of the polycrystalline ice
normally present in soil pores. Figure 2.12 shows the adfreeze
strength of piles in frozen sand (solid lines) and the contribution
to this strength by ice present in the sand (dashed lines). Note
that the dashed lines are the same as the solid lines in Figure 2.9;

but reduced by a factor of 0.2377. If the lines were not reduced

 

Phi
sul
pn
Ni}

Cl‘E

Dari

aCCl

 

 

29

the adfreeze strength for the concrete piles in ice would have been
much higher than that in frozen sand at displacement rates below

4.0X10'5 in./minute.

2.5 Circular Footing in Frozen Soils:

In permafrost regions deep circular footings and piles are the
most frequently used foundation types (Ladanyi and Paquin, 1978).
Loads applied to deep foundations or piles are partly carried by
lateral shearing forces and partly by end bearing. The ratio of
the two supporting forces depends on a number of factors including
soil type, footing shape, method of installation, embeddment ratio,
and the footing settlement. The complexity of the observed
phenomena sometimes obscures the contribution from the two separate
supporting forces and makes any clear comparison with theoretical
predictions difficult. Some authors (Nixon and McRoberts, 1976;
Nixon, 1978) show theoretically that for ice—rich soils with the
creep parameter 1.5<n<3, the ratio of end bearing to pile shaft load
is approximately in the ratio of pile diameter to its length.
Therefore, for a 20-ft. long and l-ft. diameter pile, only 5 percent
(1/20) of the total load would be due to end bearing.

A similar theoretical analysis made by Weaver and Morgenstern
(1981 b) showed that even for ice-poor Ottawa sand and an 82rft. long
by 8-in. diameter pile the fraction of load supported in end bearing
is only 1.1 percent. Note that these analyses are based on creep
parameter data from published sources. In the literature several

accounts have been made regarding field and laboratory pile load

 

test
labo
frOZl
hen

the I

Gibsoi
was f-
with 5
for Va
1975;

Pledk

 

 

 

30

tests. However, very few data are available concerning field or
laboratory tests on deep footings (including end bearing piles) in
frozen soil. Only limited comparisons have been made between exper-
imental data and theoretical predictions due to the complexity of
the problem.

To help fill this gap, Johnston and Ladanyi (1974) conducted
field tests on deep circular anchor plates in permafrost. Ten single
helix screw anchors, Figure 1.2 (d), were power-installed in layers
of frozen varved clay and silt with temperatures of —0.1°C to -O.33°C.
The anchors ranged in diameter from 8.15 in., and were installed at
depths ranging from 8 to 14 ft. Although the observed anchor pres-
sure varied non—linearly with displacement rate on a log-log scale,
the data compared reasonably well with the theoretical values pre-
dicted by Ladanyi and Johnston (1974) using cavity expansion theory
(Hill, 1950). The two authors used creep parameters determined from
pressuremeter tests on frozen soil at the site (Ladanyi and Johnston,
1973).

Since the early work on bearing capacity of foundations by
Gibson (1950) and Meyerhof (1951), where the cavity expansion model
was first mentioned in the literature, the theory has been applied
with success to the problem of bearing capacity of deep foundations
for various soil types by several authors (Ladanyi, 1963, 1967, 1973,
1975; 1973). In frozen soils, the theory can be used for

predicting creep settlements and time—dependent bearing capacity of

res

hal

ass
301'

the

 

 

 

 

 

31

deep circular footings (Ladanyi and Johnston, 1974), strip foot—
ings (Ladanyi, 1975) and laterally loaded piles (Rowely et al.,
1973, 1975). In the latter case, the solution was obtained by
analogy with the strip footing case.

According to observations made by Vyalov (1959) on circular flat
ended punches in undisturbed frozen silty clay and by Johnston and
Ladanyi (1974) on the screw anchors, it appears that in frozen soils
the Prandtl-Terzaghi bearing capacity model loses its experimetal
justification since the formation of Prandtl type slip surfaces was

not observed in those tests. There was, instead, a deformed zone

 

or ”cavity” clearly visible above the screw anchor's in the varved
soil when the anchors were excavated. Note that the term "cavity“
refers to an expanding spherical region which is formed as the
circular punch (footing) advances continuously into the soil. As
shown in Figure 2.13 (a) the cavity is always preceded and surrounded
by a plastic front which may have a spherical shape(under a circular
footing) or a cylindrical shape (under a long rectangular footing).
Also note that the penetration of a circular punch deep into the
soil produces a state of stress and strain very similar to that
resulting from the expansion of a spherical cavity in an elastic
half space medium (Hill, 1950). For transformation of a cavity
expansion to a deep footing problem, Ladanyi and Johnston (1974)
assumed that during penetration of a footing into the soil, a rigid
soil cone (circular footing) or wedge (strip footing) is formed at

the footing base (Figure 2.13 b) with the angle a close to 45°. Both,

 

for it
the pr
cohesi
indica
total
the au
settle
VS=Vu
soil ll
pressfl
(for a:
criter'
(1974)
settlen
§<
where r
Pressm
(for ¢>
c=
Where 8
f1's th
Us
Perletra

Strengt‘

 

 

32

for the cone and the wedge, statical considerations at failure give
the pressure q as a function of the cavity expansion pressure Pi’
cohesion c, and the friction angle a. Ladanyi and Johnston (1974)
indicated that in frozen soils the ultimate load is reached at a
total settlement close to 10 percent of the punch diameter B. Also,
the authors transformed cavity—expansion displacements ”i into footing
settlements s, by equalizing the displaced volumes in both cases, i.e.
VS=Vu in Figure 2.13 (c). This assumption implies that the frozen
soil inside the "cavity” or the cone is perfectly plastic (incom-
pressible). Using the generalized stress vs. strain rate relationship
(for an incompressible von Mises material), the Mohr-Coulomb failure
criterion, boundary and equillibrium conditions, Johnston and Ladanyi
(1974) deduced the following relationship between the footing creep
settlement rate s0 and the applied pressure q:

éc = (éC/zogue)(3/2n)”<q—po-nc)n (2.18)

where n=10s/B if s<0.1OB, and n=1 if s>o.IOB, po is the confining
pressure, and c is the triaxial test cohesion and is given by
(for ¢>0):

c=ocue(ef/éctf)/2f% (2.19)
where ef is the strain at failure, tf is the time to failure, and
f is the flow value and equals (1+sin¢)/(1—sin¢).

Using the same theory, Ladanyi (1976) has shown that the static
penetration test can also be used in frozen soils to determine their

strength parameters. A series of quasi-static and static (or

incre
site
study
with
tip c
resis
leadi
of pe
with
(1978
soil
(Hoff
that
“P to
numer
aDDea

compel

effeci
soils.
avails
3990.
settiE
from C

that w

 

 

 

incremental loading) penetration tests were conducted at.a permafrost
site near Thompson, Monitoba (T=~0.1OOC to -O.33°C). The field

study included stress and penetration rate-controlled tests, performed
with an electric penetrometer (1.4—in. dia.). The original conical
tip of the penetrometer was replaced by a flat disc. The point
resistance was plotted versus penetration rate on a log—log scale
leading to a general power law relationship for the rate dependence
of penetration resistance. The experimental data were in good agreement
with the solution obtained from the cavity expansion theory. Nixon
(1978) attempted two solutions; the Boussinesq solution for linear
soil behavior, and a numerical solution based on Hoff's analogy

(Hoff, 1954) to consider the soil non—linearity. Nixon (1978) showed

 

that cavity expansion theory tends to over—predict settlement rates
up to a factor of 2 for value of 1<n<3.5, when compared with the
numerical method. For values of n>3.5 the linear Boussinesq solution
appears to over—predict settlement rates up to a factor of 3, in
comparison to the two other methods.

The depth of embeddment of the footing appears to have a similar
effect on the settlement and bearing capacity as that in unfrozen
soils. Although no theoretical solution of this problem is yet
available (Ladanyi, 1981b), experimental data (Vyalov et al. 1973;
Sego, 1980) show that the depth of burial reduces the total creep
settlement to about one half when the relative depth D'/B increases
from 0 to 1.25 (Figure 2.14). On the other hand, Sego (1980) observed

that when a punch settles more than about 5 percent of its diameter,

the

dept
buri
1980
1978

sett
succ

load

test
Lada
Plan
a sai
rate
(Figi
tests
tests
Settl
long
mobil
attai
loadj

then

 

 

34

the soil enters into a viscoplastic region and the effect of burial
depth becomes much smaller. Under such conditions, no effect of
burial depth was observed between Dl/B = 0.5 and 3.3 in ice (Sego,
1980) and between D'/B = 5 and 15 in frozen sand (Ladanyi and Paquin,
1978). ‘

Little information is available in the literature on the effect
of loading history. Vyalov et al. (1973) presents information on the
settlement of 6.3—in. (16 cm) diameter punches, loaded in five
successive stages, on blocks of ice in Figure 2.14. In each stage the
load was increased and kept constant for 10 days. Data in Figure
2.14 show that a steady—state settlement rate was attained in these
tests after about 1 to 2 days. A similar conclusion was reached by
Ladanyi and Paquin (1978) when analyzing the results of their deep
plate (1.4'in. diameter circular punches) loading tests embedded in
a saturated frozen sand. They observed that a constant settlement
rate usually developed after 1 to 5 days under constant load .
(Figure 2.15). The instantaneous settlement was negligible in such
tests. It was also negligible in Ladanyi‘s (1976) static penetration
tests on frozen silt. Ladanyi and Paquin (1978) found that the
settlement rate remains a function of the loading history only as
long as the penetration resistance of the soil is not completely
mobilized, which happens after the total accumulated settlement
attains about 1/3 the footing diameter. Beyond that settlement, the
loading history effect appears to be erased and the settlement rate

then depends only on the applied load, as long as the soil remains

unc

reu
cons
in i
give
visc
1966
stud
abse
work
Jess

unde

Flgm
1976:
CvaE
on tt
behav
was t
Prand
Mende
compu

Both

 

 

 

 

35

unchanged.

2.6 Frozen Soil Beam Tests:

To investigate the multiaxial stress state and to verify the
results of analytical or numerical analyses based on properties or
constitutive laws, it is convenient to simulate frozen soil structures
in the laboratory. Examples of model piles and footings have been
given in previous sections. Although analytical solutions for
viscoelastic plain beams have been attempted (Hult, 1966; Odqvist,
1966) and also numerical solutions (Klein and Jessberger, 1978), model
studies on frozen soil beams (plain or reinforced) are generally
absent in the literature. An exception to this is the experimental
work published by Meissner and Eckardt (1976), as quoted by Klein and
Jessberger (1978), and more recently an experimental—analytical study
undertaken by $00 (1983).

An experimental creep curve for Emscher-Marl at —6°C is shown in
Figure 2.16 for a frozen simply supported beam (Meissner and Eckhardt,
1976). The dashed lines represent the theoretical creep deflection
curves, for Emscher-Marl at —10°C and Karlsrdher sand at —33°C, based
on the finite element method of analysis. Identical stress-strain
behavior was assumed in tension and compression. The uniaxial case
was transformed to the multiaxial state of stress based on the
Prandtl—Reuss equation and the von Mises flow rule (Hill, 1950;
Mendelson, 1968). The finite element program was also used to
compute the stress distribution across the beam depth, Figure 2.17.

Both the deflection rates and the stress distribution were in close

agr

Odo

Odc

eni
iai
Few
rai

cre

 

 

36

‘agreement with the analytical solution proposed by Hult (1966) and
Odqvist (1966). The stress distribution can be evaluated analytically
by satisfying the moment equillibrium equation given below (Hult, 1966;
Odqvist, 1966):

h‘
f o z(b dz) = O (2.20)
o
where ht b, and z are defined in Figure 2.17, and for identical soil
behavior in tension and compression the neutral axis is located at the
cross—section centroid, i.e. at 21 = h/2. Applying the elastic analog,
in which the strain rate e is made equal to the elastic strain e, then

0 can be expressed as:

o= ocue(e/éc)1/“ (2.21)

The assumption of identical soil behavior in tension and com-
pression does not appear to be always valid. Short-term creep tests
on frozen silt, at -10°C, showed that the ratio of its compressive
to tensile strength, oC/ot, was about 4 (Vyalov, 1965) at one hours time
to failure. The ratio was deduced from parabolic Mbhr-Coulomb failure
envelopes at different times to failure. For a 12-hour time to
failure envelope, the ratio was reduced to about 3.5, thereafter
remained almost constant. Hawkes and Mellor (1972) observed that the
ratio oC/ot increased from unity at strain rates below 10’6 sec.'1, in
creep tests, to about 8 at higher rates, in constant strain rate tests
on polycrystalline ice at -7°C. This is because for higher rates, ten-
sile strength ot changed very slightly as the strain rate é increased,

while the compressive strength oc continued to increase (Ladanyi, 1981).

 

 

 

stn
axis
at l
ShOl
Oven

Brag

whei

Subs

lnte

 

 

 

 

37

Haynes et al. (1975) reported that the ratio remained equal to
one, for a frozen silt at -9.4°C, until the strain rates exceeded
10'3 sec.'1. After that, the ratio started to increase UP to about
3.5, with increasing brittleness in tension. Haynes and Karalius
(1977) observed that the gap between the compressive and tensile
strengths of frozen silt did not only increase with the increase in
strain rate, but also with the decrease in temperatures. More
recently, Bragg (1980) reported that the ratio oC/ot is about 5,
obtained by extrapolation for frozen Nedron sand.

Based on the foregoing observations, a frozen soil beam subjected
to external bending moment may fail because of a weakness in tensile
strength. Therefore, it can no longer be assumed that the neutral
axis is located at the centroid of the beam cross-section, especially
at high strain rates. In this case, the position of the neutral axis
should be determined by satisfying the force equillibrium condition,

over the beam cross-section, given below (Hult, 1966; Odqvist, 1966;
Bragg, 1980):

Fc = Ft (2.22)
where FC is the resultant force in the compression zone, 30d 15 given AV:
z .
Fc=0f 1 ocue(e/eC)1/"(b d2) (2.23)

Substituting Z/R for e, where R is the radius of beam curvature, and
integrating gives:

FC = (n b ocue/ 1+n)(1/ R ec)l/“ (21)1+1/" (Z 24)
A similar expression may be derived, for the resultant fOrce Ft in the

tension zone, with 21 replaced by (hLzl) in Equation 2.24. The creep

 

 

 

par
pre
Equ
bot

and

Sdt

 

38

parameters n, Ocue’ and go may not be the same in tension and com-

pression. Substituting for FC and Ft by their corresponding expressions

Equation 2.22 can be solved for 21. Note that R would cancel from

both sides of Equation 2.22.

300 (1983) has included the effect of steel reinforcement (plain

and deformed bars) in his experimental study. Frozen beams of

saturated Wedron silica Sand (64 percent sand by volume) were tested

at -10°C under step loadings maintained for 7 to 20 hours. The two

types of steel reinforcement used included plain rods and rods with

a single lug at each end. Preliminary data from his tests are

summarized in Figure 2.18. As expected, the plain beam showed the

largest deflection when compared to the reinforced beams under almost

the same load. Although the ultimate load carrying capacity of

the beam with the deformed rod was larger than that of the beam with

the plain rod, the latter showed a smaller deflection and deflection

rate than the beam with a deformed bar. A discussion of these findings

relative to the results from the present study are included in Chapter
VI.

 

 

N:_0\Uv_ ..rz. wfiumcmium 0>wm®£u< ®U~

 

Ice Adhesive Strength, T1, kg/cm2

 

 

 

 

 

39
22 L’ I
' I
I
I
l a
18 ._ [Ice/brass adhesion
'(Hunsaker et al., 1940)
I
.— II
at.
14, _ ' Ice/stainless steel
' l(Jellinek, 1957b )
__ I
I
,l
10- I
I
l
t’ I
l
w I
6 I
I
F I
l
2 _ l
V l
lfit- l
l ‘ Ice/polystyrene
l (Jellinek, 1957b )
0.5 L ’ +
o -... 41..l..fl....1.1#141..l
0 -5 -10 -15 -20 —25 —30

Temperature, °C

Figure 2.1: Adhesive strength of ice to different materials
as a function of temperature.

 

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7.0 [:
6.0 _
5.0 '
N
5 _.
\
C7)
‘5‘40—
:5
:3 + From tensile experiments. Data
E; showed strength dependence on
5 3 0 sample volume, cross-sectional
i; area, and loading rate.
(1)
>
'S 2.0
G)
I
'0
< _
Cl)
.2
1.0“
‘7
0 5t— From shear experiments
I
0025—
0 1I‘lllllllllllljl'lllllllllllll
0 -5 -10 ~15 -20 —25 —30

Temperature, °C

Figure 2.2: Effect of test type on the adhesive strength of
ice to polystyrene surface (after Jellinek,

1957 b

 

41

 

 

 

Rate e

Strain

(b) Time

Figure 2.3: Constant—stress creep test:
(a) basic creep curve; (b) true
strain rate vs. time.

 

r

.+\ +\C._.\ i...” a tliwxx

Ammrokw>npv @950 u MMF

Rh

 

42

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omfi ONH ow ow o
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mE:_o> xa Ucmm “smegma

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1

 

 

3O "

25 -

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44

Leena

>0 .963qu

80—120 silica sand —24Oc c3=o.17MPas'1 Smith and

' . o ,= -l Cheatham
80 120 Slllca sand 16 C O 0.17MPas (1975)

Unfrozen OWS 20—30 (grained) _3 _1 Alkire and
Frozen OWS 20—30 —12 C 4.4x10 8 S=55% Andersland

Frozen ows 20-30 —12°c 4.4x10‘3s‘1 S=97z (1973)

Penn Sand 80—200 —8°c lO-Ss—l Perkins and Ruedrich
(1973)

ows 100—200 -1o°c 10‘3s'l Chamberlain et al. (1972)
ows 2-30 -3.9°c 3xlO—3s—l Sayles (1974)
OWS: Ottawa sand

Penn: Pennsylvania sand

S: Degree of saturation

 

unfrozen sand I

l l I 1 l n n l n n 1 n I 1 1 1 l

10 15 20 25

p: (01+o3 )/2, MPA

Figure 2.6: Results of strength testing on frozen soils at

low confining stresses

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47

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fij.~... . _u..-q. q a —.....~ - q HO
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a £HmC®LHW UNUOLuvU.‘

Adfreeze Strength, psi

10

20

3O

4O

'50

 

 

\\
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\ \.
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\Q—Ultimate strength
_, \’ (10 Kips/day)
\
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\
h; 1 l l l _l \ A
32 30 28 26

Ground Temperature, °F

Adfreeze strength, of silt-water slurry to
an 8—in. diamter steel pipe pile, versus
temperature (after Crory, 1963).

Figure 2.10:

49

O \

/l ‘ <1" I“ ,-/’"
/l du la ' ”
I

 

/ v=du/dr

/‘\\%assumed
I
l
l
l
l
l

 

/

load distribution

 

 

I
I
I
I
I
I
I

Z: P I

 

 

 

t
Loaded End .___J.

 

dr
Y‘
plan en. a @de
\

Definition of terms used in the bond analysis of a
plain rod (pile) in frozen soil (after Johnston and

Ladanyi, 1972).

Figure 2.11:

 

.Afimmfi .cwgmzmmEmLma

empemv mums “cmeuwFQmwv wpwa FMCWEo: 5pm: Amwce_ umcmmuv

scam cw “commas :oemwcum woe Op mzv coeuznenpcou new Amm:__
ueFOmv teem :mwoce ce mm_ea mo cpmcwapm wwaLevm we cowpmwcm> ”NH.N enamel

.CeE\EE .ouom pcmewanQmeo mFTQ

 

50
I
0
r4

   

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muwaucoou \m.m l
aeez- .<.< -
Use- as a: ..... i 9N
goo- we teem :mwocn Trill:

 

pdw ‘uibuaais eZBBJJPV

Flgu

 

 

 

a a Elastic-plastic
rc boundary; cylindrical
—e4q ”cavity”

 

 

I

  
  

  

Advancing
plastic zone

Elastic—plastic

 

 

 

 

 

boundary;
spherical
”cavity”
(a)
B
n S
t ‘1 V 7 W .-
IL I; 51.14% @9114 _
-Pntand> "\I ‘3'?“
/ ‘ 1v t2
/;>\ ’/<:\ Vu l‘ a 1”
K 43/
. . ~~ "J,
P_i Semlfsp'berlcal Jig/i7,_7§§7
cav1ty (Not Tu.
to a scale) 1
(b)

(C)

Figure 2.13: (a) Soil response to penetration by a flat circular
punch, (b) and (c) notations for transformation of
cavity expansion theory to a deep footing problem
(after Ladanyi and Johnston, 1974).

on: m.OwU rlIIIIIIIJHHMW
0.0

on: N.Dwo

 

52

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mzwn .mEeH

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an: N.oua

 

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mu ‘5 ‘1uawaiaiag

Six 0 +£0..c0...++QV FELFC.

Total Settlement, s, cm

     
 
  
  

 

 

T= —6°C

+q= 18.07 MPa
= 21 83 hours

s: 9.81x10'2 cm/min.

4.06 MPA

98.05 hours

1.83x10'4 cm/min.
I

 

   

 

 

 

 

a
7.57 MPa
48.05 hours
3 69x10'3 cm/min.
—————— ‘*‘~13 80 MPa
24.52 ho rs
,r 20.92 MPa 3.09x10‘ cm/min.
3.75 hours
0.155 cm/min
I l l l l l l
0 20 40 6O 80

Tine, hours

Results of a loading creep test on a
chtMarMam(14in.Mmmwm Min.
deep) in frozen sand (after Ladanyi and

Paquin, 1978).

Figure 2.15:

 

54

com

ae\ee m-onem.m (io\.\i

..E\EE muofixwamna o..l|.l.l. IIIII I.l..|ol..||.|.|.

 

.Awmmfi .mecmnmmww use
cew_¥ qummv mamas scam :mNoLe eo cwpcmu age we cove

meso; .maeh

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QIIU'IIIII AIIO‘

 

we .pcwswawaxm

 

“cmEmFm evened -1-- ;

0H

mm ‘A ‘uoiioaliag

EU 'mwX< FQLU.DOZ EOLns OUEMHWwQ

 

 

Distance from Neutral Axis, cm

’N

 

55

P= 4kN(900 lbs) b: o 1

 

 

 

 

0.15 m

 

 

 

 

 

 

 
  

 

 

  

 

 

 

——-———————a> x AA] —— -
1‘ 1g ///
I .Om J /
(a)
6
.e_.._____‘g___Finite element
3 _. I
0
-3 —
-6—
l J J I 1
-400 -300 ~200 -100 O 100 200 300 400
(b) Normal Stress at x=0.475m, ox, psi

Figure 2.17:

Frozen Soil Beam; (a) Diagram of a Simply Supported
Beam; (b) Stress Distribution (after Klein and
Jessberger, 1978).

a:

r

NIOHX

A)

I C0?U.UU~M¥.WQ

.: GUM.

 

Beam Deflection, v, x10“2 in.

 

 

   
 

 

56
I P P
10 .. Ix’T— “\\
21/ 3%” / A :1] V A \5\
— I<—_>‘ H II 10" loll u
2.75" “We—+434 l—
I<+——- 2L= 403——.———+>4
8 _ Rod (3/16"¢)
Lug (1/16” ht.)
I I I
6 .. II I
I .
I I Unreinforced Beam
P= 205 lbs.
- I (7: 0.002 in/hr.
plain rod and lug
4 I3: 200 lbs. 8
v: 0.002 in/hr. ,_
E] E] El El Cl
_ E]
I [3 E] E! El E
E! El '3 a E e o o
2 _ El '3 o 0 O O O

 

E= 226 lbs.
v= 0.001 in/hr.

0 l l l l l l l l I l l
O 2 4 6 8 10
Time, t, hours

Figure 2.18: Deflection creep curves for simply supported beams of
frozen sand at -10°C (after $00, 1983).

IOC
fro

USE

 

 

CHAPTER III

MATERIAL PROPERTIES AND SAMPLE PREPARATION

3.1 Steel Rods:

For most ice samples, a plain cold—rolled steel rod, 5/8 in.
diameter, was used. A few samples were tested with a 3/8 in. diameter
rod for comparison. For sand—ice samples, several diameters ranging
from 3/32, 6/32, 12/32, and 20/32 in. plain cold—rolled steel rods were

used. All of these rods included a threaded portion at both ends which

 

permitted easy attachment to an eye connector and hook at the bottom of
the coolant bath and mounting of a displacement transducer at the upper l
rod end (Figures 3.1 and 3.2 a). Lugs welded to the 3/8 in. diameter
rod, shown in Figure 3.2 b, included heights, h, of 1/16, 2/16, and 3/16
in. and a length of 1/2 inch. Lug heights were limited to 3/16 in. so
as to avoid a tensile failure in the 3/8 in. steel rod. Plain rods with
a single lug are referred to as deformed bars.
Lug face angles of 45 and 90 degrees, Figure 3.2 (c), were selected
to provide preliminary information on wedging between the frozen sand
and rod during pull-out. Rod surface roughness was considered by using
three steel surface types: ground-finish steel with a roughness factor
0 of 90 percent, cold-rolled with p = 625%, and shot—blasted with p =
1489%. The latter type surface was prepared from the ground—finish steel
by shot-blasting with commercially available, uniformly graded glass beads

with all beads passing the No. 60 U.S. standard sieve and retained on

57

 

IN

tes

alh

beti

at1
ail
Illa

IIEQ

 

 

58

the No. 100 sieve (size range of 0.0058 in. to 0.0097 in.). Shot—blasting
generally gave the steel surface a finely-textured grey color, whereas

the ground-finish steel displayed an almost white, shiney and highly
polished surface with minimum imperfections. The cold-rolled steel rod
surface was dark grey with very fine spots due to some imperfections which
originated during its manufacture. A standard deformed 3/8 in. diameter
rod with lugs spaced at 0.3125 in. on centers was also used in several

tests. The lug height for this rod was about 1/64 in. (0.015 in.).

3.2 Ice Samples:

Three ice types used in this study are referred to as Ice(1),
Ice(2), and Ice(3). For the first type, ice samples were prepared by
pouring distilled water directly into a cylindrical split steel mold with
a 6 in. inside diameter and 7 in. height. The water was poured to a
height of 6 in. i 1/8 in., in four layers. Before that, the mold was
first mounted on a 6 in. diameter by 1/2 in. high steel reaction plate.
This plate was attached, by screws, to a rectangular aluminum reaction
beam, 1 1/2 in. X 5 in. X 12 in., with a circular hole in the center to
allow for the steel rod embedded in the pull-out specimens. Clearance
between the rod and the hole for each rod size was maintained at 0.063
in. by use of a removable washer in the reaction plate.

To prevent leakage a commercial putty was used to seal the mold
at the edges and to minimize ice adhesion to the mold it was lined with
a thin sheet of plastic saran wrap. All samples were frozen and stored
in a freezer at close to -20°C for at least 12 hours. Directional

freezing took place from the base upward and from the sides radially to

 

the
in:
on
san
huh

WEI

ire

bini

SllOI
snow
and
nomi

IOe

Ion“
sawp
diff
ice

Cool;
samp‘
Perm!

Place

 

 

 

59

the center. The sample top was covered with a one inch layer of styrofoam
insulation to minimize sample damage due to delayed volume expansion
caused by change of water to ice in the central core of the sample. Ice
samples prepared in this way were clear with few or no entrapped air
bubbles. After removal from the mold, rubber membranes (double thickness)
were placed, as shown in Figure 3.1, to protect the ice sample during
immersion in the circulating anti-freeze/water coolant mixture.

For Ice(2), samples were prepared by first cooling the mold in the
freezer, placement of crushed ice in the mold, and pouring precooled
(T=0°C) tap water over the crushed ice. Very little ice melted when com-
bined with the precooled water. The crushed ice was also formed from
ordinary tap water, with a hardness close to 450 parts per million as
calcium carbonates. Ice(3) was prepared by placing dry natural fresh
snow in the precooled mold and pouring precooled distilled water over the
snow. The fresh natural snow was collected and stored in the freezer
and later, before use, was sieved on a No. 4 U.S. standard sieve, with
nominal grain diameter of 0.187 inch. Ice(2) and Ice(3) were cloudy due
to entrapped air bubbles.

These two ice types are both polycrystalline, similar to that which
forms in the pores of the frozen sand. The average density of all ice
samples was close to 56.2 lbs/cu. ft. (0.90 gms./cu. cm) i 1.2 percent
difference depending on the temperature (Pounder, 1967). The prepared
ice samples were transferred to a bath in which an anti-freeze/water
coolant mixture at the test temperature circulates around the protected
sample. All samples were cooled at least 10 hours before testing to
permit temperatures to equalize throughout the sample. Constant dis-

placement rate tests were carried out on ice samples labeled as

 

th

in)

exc
vol

tie

bath

90811

 

 

6O

I-l through I-32 in Appendix A. Constant load (creep) tests were

conducted on ice samples labeled as CI in Appendix B.

3.3 Sand-Ice Samples:

A commercially available sand produced by the Wedron Division of
the Pebble Beach Corporation of Wedron, Illinois, was used in this
investigation. The Wedron sand consisted of sub-angular quartz particles
with a specific gravity of 2.65. The sand gradation was uniform with
all material passing the No. 30 U.S. standard sieve and retained on the
No. 40 sieve (size range of 0.0232 to 0.0165 in.). The coefficient of
uniformity was about 1.50. The pull-out samples were, 6 in. in diameter
and 6 in. high, prepared in the same split mold used for ice samples.

A sand volume fraction of 64 percent was selected for samples,
except those prepared for the study of sand concentration effect. This
volume fraction is comparable to values normally encountered in the
field and insured the development of dilatancy and interparticle friction
in front of the lugs. The value of 64 percent sand by volume corresponds
to a dry density of 105.8 lbs./cu. ft., void ratio of 0.5625, and ice
saturation close to 97 percent. To insure this high degree of saturation,
the mold was partially filled with a pre—determined amount of distilled
water and sand was slowly poured into the water permitting air bubbles
to escape to the surface. The sample was tamped with a steel rod (0.25 in.
in diameter) to give the desired sand packing. The methods of sealing
and lining the mold, freezing the sample and transferring it to the coolant
bath for testing were similar to those described for the ice samples.

Preparation of frozen samples with sand particles in dispersed

positions, i.e. with sand concentrations lower than 64 percent, involved

 

WEI

les

wer

Dif

dia

Ire

 

— . ~~-‘Lx‘m.’ . ~.—. .. ‘-' .- . 7-. ..— —.—.. .....- v

 

61

use of sand chilled to below freezing and then carefully mixed with dry
natural snow (sieved through a No. 4 sieve). This sand—snow mixture was
then placed into the cold mold and distilled water, precooled to 0°C,
was poured over the mixture. Very little snow melted as a result of
pouring ice water over the snow and sand mixture. The resulting sample
contained uniformly dispersed sand particles. For each sample, a specified
amount of air-dry sand was pre—determined to give the required sand
concentration. After sample preparation, the excess sand was air-dried,
weighed, and deducted from the pre-determined amount to give the weight
of sand in the sample. Several sand concentration values, vs, were
considered; 64.0 percent, 52.8 percent, 48.6 percent, 44.5 percent,

32.3 percent, 17.5 percent, and 0 percent for the snow-ice sample.

Tests were also conducted on frozen sand samples of different
heights (1 in., 2 in., 3 in., 4 in., 5 in., and 6 in.). These samples
were prepared in the same manner, except that for samples with heights
less than 6 in., styrofoam plates 6 in. in diameter and about 1 in. thick
were added to the top of each sample to make a total height of 6 inches.
Different diameter samples were prepared using available molds with
diameters of 1.9375 in., 2.75 in., and 4 inches. These samples were

prepared in the same manner as described for the 6 in. sample diameter.

 

F"9Llr

 

 

Lead wire to force
transducer

      

Mounting bracket for dis-

placement transducer .
Thermister lead wire

[L L
I

w

r
77:-_--
T—L

 

 

 

——
_
-

 

l

circulating

coolant ‘ —-Membrane

 

l_--_

1r:-rfé--»

Rod

tug—~84

Frozen'

.‘ WVVWW A.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

  
  
 

I I
. I I
i ; Soil
- . I ISpecimen «e—Load frame
‘I ' ‘T—f‘s I. 9 T1
I i. .
I I I I \\\\ I , Reaction
1 l I l : I Beam
LaiJ L1 :lIr‘ \ L95]—
Membrane—ee—I Reaction
7 4' Plate

 

 

  
 
 

Removable

Eye connector for
Nasher

attachment to hook at

bottom of coolant tank .
Mounting bracket for

displacement trans—
ducer core

Figure 3.1: Equipment for pull-out tests including a frozen sand sample,
steel rod with lug, and loading frame immersed in the

circulating coolant.

n. 12 in. 1.5 in.
L4
"1 F , ‘ri P-

Q.
Rod diameter, d

 

 

 

[lllllllll : I JDIHII]
(b) Lug -i I;TI§

length,
3g in.

 

 

 

Lug height, h

, Lug face angle, 450

59(‘\
1/8 in. ‘ ____

 

(C)

Figure 3.2: Bar configuration. (a) Plain rod. (b) Plain rod with 90
degree lug. (c) Lug with 45 degree face angle.

 

CHAPTER IV
EQUIPMENT AND TEST PROCEDURE

4.1 Equipment:

Constant displacement rate and constant load (creep) pull-out tests
were conducted using the loading frame shown in Figure 3.1. The frame
consists of a 6 in. diameter by 1/2 in. thick sample reaction plate; upper
and lower 1.5 in. by 5 in. by 12 in. aluminum beams; two 5/8 in. diameter
16 in. long steel rods, for load transfer from the upper to the lower beam;
and eye connectors at the top and bottom. The frame stiffness was assumed
to be high compared to that of the reinforcement rods. Attached to the
upper beam was a 10,000 pound capacity load cell (Strainsert Model
FL10 —25PKT) for monitoring the applied load. Displacement of the steel
rod relative to the base of the frozen sample was monitored using a San-
born Linearsyn differential transformer (Model 585 DT—lOOO). This transducer
was supported by a bracket mounted at the top of the steel rod (Figure 3.1).
The LVDT core was supported by a small rod attached to the lower reaction
beam.

An insulated steel rectangular coolant bath, 14 in. by 14 in. by 16
in., helped confine the coolant liquid surrounding the frozen sample.

The anti-freeze/water coolant mixture circulates from a refrigeration unit
to the bath as shown in Figure 4.1. Equal parts of water and ethylene
glycol was close to the optimum combination for the coolant mixture. A

one inch external covering of styrofoam insulation helped maintain the

64

ten
par

cir

Sens

 

65

sample and bath test temperatures. A Graham electric motor, with variable
speed gear box, mounted on a 10,000 pound capacity Soiltest loading frame
(Figure 4.1) provided the constant displacement rate pull-out loads.
Displacement rates available from the variable speed gear box ranged from
about 10"5 to 10‘1 in./min. For constant load (creep) tests, a steel lever
system (Figure 4.2) with arm ratio of 8:1 was attached to the Soiltest load
frame. Dead weights, up to 550 lbs., were applied to give a maximum pull-
out load of 4,400 pounds.

A Sanborn two-channel recorder (Model 77028) with preamplifier (Model
8805A) was used to monitor loads and displacements for all tests. At
maximum sensitivity, one millimeter stylus deflection on the displacement
strip chart represented 0.0005 in. displacement for the transducer. This
sensitivity limited the range of creep data obtainable for small dis-
placements observed for ice adhesion to plain rods. For the load chart,
at maximum sensitivity one millimeter of stylus deflection corresponded
to 14 lbs. applied to the load cell. A Hewelett-Packard digital Logging
Multimeter (Model 3467A) and thermistor was used for monitoring sample
temperatures. The thermistor was embedded in the sample during pre-
paration with the lead wire entering through an opening in the upper
circular plate (Figure 3.1).

Measurement of the rod surface roughness was made using a Physicists
Research motor (Mototrace Type V, No. 680) which moves a tracer (Model
LA4-10) on the surface in question at constant speed of 0.221 in/sec.;

a profilometer and DC input (Type 0, Model 8, No. 538) with a dial gage
for reading the maximum and minimum surface aspirity heights. At maximum

sensitivity, a one micro-inch height of aspirity could be read on the gage.

 

SUl
rep
in

dat

 

66

was achieved or when excessive displacements were reached. For constant
load (creep) tests, incremental step loads were added until either failure
occurred, the ultimate load capacity of frame was reached (about 4,500
lbs.), or the displacement limits on the load frame was reached. Recorder
strips were labeled and the data for each test was transcribed to test
summary sheets.

For measurement of rod surface roughness, a duplicate surface for
each steel type was prepared on a flat plate, for easy tracer movement on
the surface. Measurements were made by connecting the tracer to the
profilometer, and the latter to the DC recorder. After warm-up, the
recorder stylus was adjusted to the chart center. Paper movement was
selected at 10 cm/min. and the attenuation knob was adjusted to keep the
trace on the recording paper for the highest asperity on all surfaces.

As the tracer moved on each surface, the maximum and minimum asperity
heights (in micro-inches) were read on the profilometer gage, while a
continuous record of the surface irregularities was being made on the
recorder strip chart. The corresponding aspirity heights (including the

maximum and minimum) were given in millimeters on the chart. For each

 

surface, these measurements were repeated several times for a better
representation of the surface roughness. A roughness factor, defined by
Wright (1955), was used as a criterion for the rod surface roughness. The
data and method of calculating the roughness factor are presented in

Chapter V.

 

 

th.
te:
lUs
Cor
was

red

67

A Sanborn one-channel chart recorder with DC input (Model 322) was

used to make a continuous chart record of the surface irregularities.

4.2 Test Procedure:

After sample preparation, placement in the test loading frame, and
cooling to the desired test temperature, several additional steps were
required for completing the test. The transducers were connected to
the recorder which was allowed to warm up for approximately 30 minutes
prior to testing. After the warm—up period, the stylus needles were
adjusted to a zero reading for load and displacement. For the constant
displacement rate tests the loading ram of the Soiltest load frame was
adjusted and attached to the upper eye connector in Figure 3.1, but with
no applied load. The sample temperature was observed and recorded. During
a constant load (creep) test, which normally took several days, the
coolant refrigeration bath maintained test temperatures to within j
0.10°C. Large room temperature variations did influence the coolant ’
bath temperature. For test temperatures warmer than -10°C, the variation
appeared to be within i 0.050C. In all cases, these temperatures
variations did not appear to have affected the test results.

The gear box controls were adjusted to give the desired displacement
rate, and the ram was engaged with upper eye connector. The strip
chart paper movement was selected according to the loading rate. As the
test progressed, the stylus needle deflections on the recorder was ad-
Justed, as needed, to keep the trace on the recording paper. All
constant displacement rate tests were continued until complete failure
was observed by sudden rupture in the case of plain rods. When deformed

rods (with lugs) were used, the test was stopped after an ultimate load

 

 

Ch
01‘)

Electric motor and gear
box for mechanical drive

   

Loading ram
Coolant Out ’

 
 
  
  
  

‘—‘
Insulation

    

~‘-‘-.

Cold bath

Pull-out specimen

  
  

""'l

l
Coolant .
Reservoir l

l
L._._.- .....J

    

l

 
    

*-7-
Coolant In

   

l'
I
l
l
l

    
 
  

Soiltest load Recorder

Refrigeration frame
Unit

Figure 4.1: Diagram of test system for constant displacement rate tests.

 

 

 

I

69

Load transfer
plate

     
 

Lever Arm (8:1 ratio

Transducer lead
res

   

Coolant
out

*-

    
  
 
 

Pull-out sample

 
 
  

 
 

Load frame
(Figure 3.1)

   
 
 

Coolant
In

  
 

Dead weights

Supporting plate

 
 

Hydraulic jack

 

 

Figure 4.2: Diagram of test system for constant load (creep) tests.

 

 

 

 

 

 

CHAPTER V

EXPERIMENTAL RESULTS

 

A summary of experimental data along with a brief discussion of
their implications are given in this chapter. The material is pre-

sented in two parts: constant displacement rate tests and constant

 

load (creep) tests.

5.1 Constant Displacement Rate Tests:

In this section, results are presented for a series of constant
displacement (or loading) rate pull-out tests conducted to determine
the influence of displacement rate, temperature, bar size, sample
height, water impurities, sand concentration, lug size and shape, and
bar surface roughness on the adfreeze bond strength characteristics of

frozen pull-out specimens. Physical properties of ice and sand—ice
specimens, including the steel rods, are listed in Table 5.1. Test

results for these samples are summarized in Table 5.2.

5.1.1 Displacement (and Loading) Rate Effects:

Pull-out tests were conducted on frozen sand samples with a sand
volume fraction close to 64 percent (dry unit weight of 105.6 lbs./
cu. ft.). The bar displacement rates ranged from about 0.0001 in./min.
to 0.1 in./min. The corresponding loading rates ranged from about
2 lbs./min. to 20,000 lbs./min. The pull-out specimens were tested at

temperatures ranging from -2°C to —26°C using two bar configurations; a

70

Na

in“

tin
ste

the

IGCO
perm

(Fig

the
the

lncn

OCCUI
and '
ObSel
n d(
Fate

rods.

 

 

71

plain cold—rolled 5/8 in. diameter steel rod, and a 3/8 in. diamter rod
with a single 1/8 in. high lug as shown in Figure 3.2 (b).

For the plain rods, typical load and displacement variations with
time are shown on the recorder strip chart in Figure 5.1. As the test
starts, the load increased with time at an almost constant rate while
the displacement showed little or no increase until an initial slip

occurred. At rupture a large reduction in the load was observed and in
some cases zero values were noted. The displacement before rupture, the
elastic portion, was in most cases too small (about 0.0005 in.) to be
clearly shown on the recorder. Due to equipment limitations, the
"pre-failure" displacements were estimated on the basis of very small
recorded movement. During testing the bond resistance or stiffness
permitted development of a nominal (machine) displacement rate An
(Figure 5.1) only after rupture had occurred.

The observed behavior may be attributed partially to stiffness of
the test system. In turn, the test system stiffness was dependent on
the rigidity of all system components and connections. As the load
increased, elastic strain energy was stored in the test system
(i.e. the test frame, loading ram, loads cells, etc.). Once rupture had
occurred and the loading rate decreased, the strain energy was released
and the corresponding slip and increase in displacement rate were
observed, as shown in Figure 5.1. Since the nominal displacement rate
An does not correspond to the ”pre-failure” loading stage, the loading
rate P, as defined in Figure 5.1, was used in place of 6“ for the plain
rods.

The time to failure for all plain rods ranged from a few seconds to

 

 

abm
tehi
wit)
depi
rods
str'
vs.-
sho»
ice

adhe

 

 

 

72

about two hours depending on the loading rate, and to some extent on
temperature. At colder temperatures the time was longer. For plain rods
with lugs, the time to failure ranged from one hour to about 12 hours
depending mainly on the lug size. It took longer to reach failure for
rods with larger lug sizes. Data for each test, as given on the recorder

strip (Figure 5.1), were transcribed to a data sheet from which a load-

 

vs.-displacement curve was plotted. Typical load—displacement curves are
shown for plain rods at different loading rates in Figure 5.2. Initially,
ice adhesion combined with sand friction prevents slip. After the

adhesion was broken and partial slip had occurred, a small residual value

for adhesion and friction controls the pullout load. At the slow rate

 

(35 lbs./min. in Figure 5.2) the "healing" effect due to ice recrystalli-
zation, was more pronounced than at the faster rate (1978.6 lbs./min.)
although an increase in load was also observed after failure at the latter
rate. The "healing” effect at slow rates may be attributed to the fact
that more time permitted more ice refreezing as compared to the fast rates.
The change in loading rates illustrate two features; the faster
rates produced a higher peak load and a larger initial slope within the
elastic range of the curves in Figure 5.2. The right hand scale of this
figure shows the bond stress, 1, calculated by dividing the load by the
rod surface area. This calculation assumes a uniform stress distribution
along the rod. The bond strength (ultimate value at failure, Tu) is
plotted versus the loading rate P, at different temperatures, on a log-
log plot in Figure 5.3. A consistent linear relationship between Tu and

P, on a logarithmic scale, appears to exist at all temperatures. Using

the least-squares fitting method, the relationship suggests a power law

 

 

oft

0i“

wher
para
chos
of n

temp
con
beth
Fl gu
anal

0i“

wher

lhes

diSp

0the

prob

Ol‘

 

 

73
of the form:
. . m
Tu : TC (P/PC) (5.1)
or
P = PC(tU/Tc)n (5.2)

where n = 1/m is the strain—hardening parameter in compression tests. The

parameter T corresponds to the "proof” bond stress at an arbitrarily

C

chosen loading rate, P taken equal to 1 lb./min. in this study. Values

c9

of m, n, and T listed in Table 5.3, indicate that these parameters are

c’
temperature dependent.

Despite the fact that the nominal displacement rate In does not
correspond to the pre~failure stage, a general linear trend was observed
between TU and an on a log-log scale at the same temperatures as shown in

Figure 5.4. Thus, a power law may be established using the least-squares

analysis, of the form:

Tu = Té (én/éc) (5-3)

07‘

an = ac (TU/TC)“ (5 4)
where m', n', 6,, and ré are equivalent to m, n, PC, and TC, respectively.
These parameters also appear to be temperature dependent.

The large data scatter made it difficult to analyze the effect of
displacement (and loading) rate on the adhesive strength of ice specimens.
The scatter may be attributed to the large variability in the ice material.
Other researchers (Jellinek, 1957 b; Gold, 1978) have noted the same
problem.

For the 3/8 in. diameter rod with a single lug, the displacement

(or loading) rate appears to have an effect on lug bearing capacity similar

 

to
CU
C6
Fi
mo
W6:
on
C0

fu'

the

hit
tha
and

POW

to

dis
but
the

is}

 

 

74

to that on adfreeze bond for plain rods in frozen sand. The load-displacement
curves for two displacement rates, Figure 5.5, both indicate a signifi-

cantly different failure mechanism from that for plain rods shown in

Figure 5.2. Initially, adhesion combined with friction and a slight
mobilization of bearing action on the lug prevents slip. After adhesion

was broken and partial slip had occurred, bearing action of frozen sand

on the lug restrained movement of the reinforcement bar and prevented

complete rupture. However, some reduction in the load may occur before

full mobilization of bearing action on the lug. This "initial yield”

was more noticeable at the lower rates (Figure 5.5). As the displacement

 

increased, the load was gradually transferred to the lug until full
mobilization of the ultimate lug bearing, Pu, was reached at about 5 mm
displacement. This observed lug performance shows a close similarity to
the bearing capacity behavior of frozen soils.

The displacement and loading rates are both plotted versus the
ultimate load Pu on a logarithmic scale in Figure 5.6. The data indicate
that an approximate linear relationship may exist between Pu and I, or PU
and P on a log-log scale. Using the least-squares fitting method, a
power law of the form given in Figure 5.6 may be suggested for either re-
lationship.

The two regression lines in Figure 5.6 appear to be almost parallel
to each other, implying that the same failure meChanism governs both the
displacement rate and the loading rate. This fact may be the result of
both lines representing data from tests on the same samples. Note that
the displacement rate an.at which the specimens were loaded in Figure 5.6,

is believed to be close to the true rate. The In value represents

 

 

the

OCI

5.]

Ice
to

rod

rat
tur
sho
cl S

sma

p10
Fig.
We]
SEtI
of 3
rate
may

atier

 

 

75

the slope of the straight line portion after the "initial yield" has

occurred, while the load was still increasing on the strip chart.

5.1.2 Temperature Effect:

The temperature effect on ice adhesion, sand friction, and mechanical
interaction was determined for a sand volume fraction of 64 percent.
Ice and sand-ice samples were tested at temperatures ranging from -2°C
to -26°C, using plain and deformed rods (with a single lug). For plain
rods typical load-displacement curves plotted at different temperatures
in Figure 5.7 indicate a behavior similar to those at different loading
rates (Figure 5.2). The bond strength Tu increased with colder tempera-
tures as it did with faster loading rates. The displacement curves also
show the same initial behavior characteristics of smooth bars followed by
a sudden drop in the load. The ultimate (peak) load developed at very
small displacements, about 0.0005 to 0.001 inches. Also note in Figure
5.7 that the ”healing" effect was more pronounced at temperatures colder
than -150C. At these temperatures several cycles of failure (represented
by immediate rupture) and "healing“ were observed.

Using the data in Figure 5.3, the ultimate bond strength Tu can be
plotted versus temperature T for different loading rates, as shown in
Figure 5.8. The loading rate indicated on each curve in Figure 5.8 is the
average of all loading rates for samples tested at the same machine
setting with a nominal displacement rate. For example, the average value
of 157.4 lbs./min. corresponds to samples tested at a machine displacement
rate of 0.0029 in./min. The data scatter for some points in Figure 5.8
may be attributed to small difference in loading rate, with a common

average value, pavg , indicated for each curve.

 

 

de

wh

pl

n I

Re

0f

Sh:

file!

At

the

fa;

Par

 

76

To reduce data scatter, values of Tu interpolated (Figure 5.3) to
the given average loading rate shown in Figure 5.8 were plotted versus
temperature in Figure 5.9. The data in this figure appear to be more
consistent and the curves are improved. The bond strength for the plain
bars, Figure 5.9, show a linear relationship with temperature at the
slow (5.6 lbs./min.) loading rate. At higher rates strength increased
more rapidly for the warmer temperatures, down to -50C. At colder
temperatures the increase in strength was similar to that for the slower
loading rates. For the temperature range between —50C and —ZOOC one may
deduce, using the least-squares fitting, the following linear relationship:

ru = a'+ b'o (5.5)
where e = —T, a'and b'are constants which depend on the loading (and dis-
placement) rates as indicated in Figure 5.9.

The temperature effect can be extended to include parameters n and
n' (defined by Equations 5.2 and 5.4) for plain rods in frozen sand.
Recall that the slopes of lines in Figure 5.3 were almost equal to those
in Figure 5.4 at the corresponding temperatures. Therefore, if the values
of n (Table 5.3) are plotted against the corresponding values of n' as
shown in Figure 5.10, one may conclude that n is the same whether displace—
ment rate or loading rate is used. For this reason values of n (or n') are
plotted against temperature in Figure 5.11 giving a nonlinear relationship.
At temperatures warmer than -ISOC the relationship appears to be approx—
imated by a straight line.

In general, values of n increased as temperatures decrease. The
rapid increase in n as temperature goes below —150C occurs because the n
parameter is cotangent of the angle at which the regression lines in

Fl'QUY‘es 5.3 and 5.4 are inclined from the horizontal axis. A

 

th

no

Se

all

no
st
tel

all:

 

 

77

semi-logarithmic plot of the n parameter versus temperature, shown in
Figure 5.12, gives approximately a linear relationship. The least-
squares fitting method gives

n = exp (1.224 + 0.0886 9) (5.6)

(for -26°C 5 T g -2°C)
where o = -T. The change in n value with temperature indicates a change
in the mechanism which controls failure of plain rods in frozen sand at
different temperatures.

Temperature effect on ice adhesive behavior was generally similar to
that observed for sand-ice material in Figure 5.7. More data scatter was
noted, as compared to the frozen sand, when the ultimate ice adhesive
strength was plotted against temperature as shown in Figure 5.13.
Sensitivity of ice adhesion to the same minor factors, mentioned earlier,
appears to be the main cause of scatter (Jellinek, 1967; Barns et al.,
1971; Gold, 1978). The data trend shown in Figure 5.13, despite the
noticeable scatter, indicates an almost linear increase of adhesive
strength with colder temperatures. This statement may be limited to
temperatures in the range of -SOC to -15°C. There was no data above ~50C
and below -150C the scatter was too large to give a reliable conclusion.
The line segment shown in Figure 5.13 with dashed lines indicates
limitations of the results. The linear relationship (solid line) given

in the figure was obtained using the least—squares fitting, where o = -T.

 

The temperature influence on lug behavior appears to be similar to
that for the displacement rate effect. Pullout sand-ice samples were
tested at different temperatures, using a plain rod with a 1/8 in. height

lug. Typical load-displacement curves are plotted in Figure 5.14. At low

 

wi

WEI

P0

to

re

90

SEI

in

Dle

C(lll

bar

 

 

 

temperatures, below —15°C, the load reduction due to "initial yield" was
less noticeable than at warmer temperatures. Only a reduced rate of load
increase was observed below ~150C. The curves generally show the same
initial behavior characteristics as for plain bars, followed by a drop or
reduced rate of load increase. Additional lug displacement continues to
mobilize frictional and dilatancy strength components of the frozen sand

in front of the lug leading to increased pullout loads. The ultimate

 

load Pu, plotted against temperature in Figure 5.15, showed a linear increase
with decrease in temperature at the given displacement rate. No adjustment
was required to Pu values, since displacement rates for individual data
points was close to the average displacement rate, éavg’ for all points.

The "initial yield" loads, designated Pjy in Figure 5.14, also appear

 

to vary with temperature. The data in Figure 5.15 show a nonlinear
relationship between Piy and T. The temperature effect on other lug shapes
(Figures 3.2 b and c) was also considered for comparison with the present
9OQ-lug shape. Further discussion on this aspect will be made in a later

section.

5.1.3 Bar Size and Sample Height Effects:

Experimental data presented in previous sections were based on
frozen specimens 6 in. in diameter by 6 in. high with a 5/8 in. diameter
plain rod. The diameter of all bars with lugs was 3/8 in. Starting with
the effect of rod size on bond strength of plain bars, four bar diameters
were selected for pull-out tests; 3/32 in., 3/16 in., 3/8 in., and 5/8 in.
The tests were carried out at two temperatures; —10°C and -20°C, for
comparative purposes. Note that all tests on plain rods, regardless of

bar size, sample size or temperature, showed the same load-displacement

ch
di

ra
to
al

al

re
ot
th
th
di

th

Cd

DO

at
Do
on
th
da

at

 

79

characteristics as summarized in Figures 5.2 and 5.7. Data from tests on
different bar sizes, summarized in Figure 5.16, show a nonlinear increase
in bond strength with increase in rod diameter, d. The dimensionless
ration d/H, where H is the sample height, provided a convenient number
for expressing the bar size effect on the bond strength. This ratio may
also be related to the bar stiffness relative to the total load applied
along a certain height H.

Data points with question marks in Figure 5.16 are those doubtful
results caused by possible leakage of coolant fluid into the sample, or
other equipment problems. For rod sizes larger than 3/8 in. (Figure 5.16)
the rate of strength increase has decreased. There was some increase in
the 5/8 in. rod strength (d/H = 10.4 percent) over that of the 3/8 in.
diameter rod (d/H = 6.25 percent). When tested at different temperatures,
these two rod diameters showed similar strengths (Figure 5.17). The
difference in strengths was consistently small at all temperatures and
can be neglected in comparison to the experimental scatter for some data
points.

Plotting the data of Figure 5.16 on a semi-logarithmic scale
(Figure 5.18) suggests a linear relationship between the ratio d/H and Tu
at both temperatures. Note that the colder temperature, -20°C, corres-
ponds to a higher rate of strength increase with increase in d/H. Based
on linear regression analysis two empirical expressions were deduced for
the two temperatures as given in Figure 5.18. These equations represent
data within the range of 2 percent < d/H < 10 percent.

The effect of varying sample height on bond strength of frozen sand,

at -10°C, has been considered for two bar diameters; 3/8 in. and 5/8 in.

 

 

ti

51'
Th
3/
th

th

lh

st

to
We
cu
be
Ii.
bu

on

 

 

 

80

Test results are summarized in Figure 5.19 for both bar sizes. The data
show that sample height appears to have a small effect on bond strength.

The strength increased slightly as height was increased from 1 to 6 inches.

5.1.4 Lug Size and Shape Effects:

A 90 -lug with different heights was welded to each of three iden-
tical plain, 3/8 in. diameter, cold-rolled steel rods. A fourth rod,
with no lugs (h = 0), permitted measurements to be made on four lug
sizes. Tests were conducted at -10°C and —20°C for comparative purposes.
The load-displacement curves for lug heights of 0, 1/16 in., 2/16 in., and
3/16 in., plotted in Figure 5.20, all show a typical behavior except for
the plain rod (with no lugs). The latter showed a significant drop in
the load after bond due to adhesion and friction was broken. A standard
3/8 in. diameter deformed bar was also tested in frozen sand at ~100C.

The hot-rolled steel rod (usually more brittle than the cold-rolled
steel) has lugs spaced at 0.3125 in. with equivalent lug height of 1/64
inch.

It was anticipated that the load would exceed the equipment capacity
for a 6 in. sample height with standard deformed bar. Therefore, tests
were carried out on sample heights of 2 and 3 inches. The load-displacement
curves for these two samples, presented in Figure 5.21, show a different
behavior from those in Figure 5.20. Generally, there was an "initial
yield “ as for other lugs, caused by rupture of adhesive and frictional
bond, followed by an increase in the load as bearing forces were mobilized
on the lugs. After a peak load was achieved, the load decreased mono-

tonically with additional displacement as shown in Figure 5.21. This

behavior appears to be more dependent on the lug height and spacing than

 

 

 

 

81

on the number of lugs in each sample. For comparison, the displacement
curve for a plain rod is also shown in Figure 5.21.

The ultimate load Pu is plotted versus the ratio of lug area A] to
the rod cross—sectional area Ar in Figure 5.22 for all lug sizes. This
ratio appears to give a more consistent variation of load than does the

lug height. Although the relationship shown in Figure 5.22, for both

 

temperatures, may not be linear, they can be reasonably approximated by

a straight line. The small data scatter with large loads suggests good
agreement between tests. It was convenient to represent two least-squares
regression lines in Figure 5.22 by the linear relationships given in the
figure for each temperature. The data points indicated by question marks

in Figure 5.22 were not included in the regression analysis. These data

 

are doubtful due to possible leakage of anti-freeze coolant into the
sample. The two points indicated for a standard.deformed bar were cal-
. culated as if they were a single lug, using the data from Figure 5.21.
The method of calculation will be explained in Chapter VI. Since these
two points are nearly identical with each other, and consistent with the
general trend of the other data, they were included in the regression
analysis.

Changing the lug shape from a 90 degree angle, Figure 3.2 b, to a
45 degree angle, Figure 3.2 c, did not appear to have a significant effect
on the experimental results. Typical load-displacement curves for both
types (Figure 5.23) show an almost similar behavior, although some
relatively minor differences may be noted. These differences were not
repeated when the two types are compared at several temperatures in

Figure 5.24. For practical purposes, the ultimate loads, Pu, for both lug

 

 

82

shapes are close enough to be considered equal.

The same conclusion may be drawn for the "initial yield" loads,
Piy shown in Figure 5.24. Generally, an almost linear variation of
ultimate loads, P“, with temperature may be noted. The linear relationship
shown in Figure 5.24 can be considered applicable to both lug shapes for
temperatures ranging from —2°C to -20°C. The "initial yield" loads Piy
appear to change rapidly in a non-linear fashion at temperatures above
-2°C. Below -2°C the loads appear to increase almost linearly with
colder temperatures. An empirical least-squares equation of the form
given in Figure 5.24 appears to be applicable to both lug shapes for

temperatures ranging from -2°C to —20°C.

5.1.5 Bar Surface Roughness:

 

The terms smooth and rough are often used to describe a certain surface
texture. Surface texture is considered as those aspects of fine surface
topography, or small-scale irregularities, that might be neglected in
comparison with any measurable lug height. Smoothness and roughness are
relative terms that have not yet been well classified or precisely de—
fined. Besides, there are no sub-divisions (or terms) to identify the
intermediate conditions between the two limits of rough and smooth
surfaces. These two characteristics depend to a large extent on the
degree of magnification to which the surface is exposed. At a higher
magnification there are greater irregularities, no matter how smooth
the surface appears to the naked eye.

Several roughness criteria have been suggested in the literature by
different authors. Each criterion appears to have its merits, and

requires a certain measurement technique. The one favored in this

st

V6

6C

de

Ch
gr
ch
ea

ch

do
su
st
di

f6:

 

 

 

83

study is the ”roughness factor” Q, defined by Wright (1955) as follows:

Lr ‘ Lt
p = -—-———— X 100

Lt (5.7)
where Lr = length of the "roughness line“ in Figure 5.25, and Lt =
total length of the standard chords, LC, forming the "unevenness line.”

From Figure 5.25, note that as the chord length LC decreases the
value of Lt increases when the "unevenness line" comes closer to the
actual surface. However, Wright (1955) tried different chord lengths
(8 cm, 4 cm, 1 cm, 1/2 cm, 1/4 and 1/8 cm) and compared the value of Lt
for each one with that of a series of four 8 cm chords, i.e. Lt = 32 cm.
He recommended a standard chord 8 cm in length. Four standard chords are
desired, whenever possible, to give Lt = 32 cm as shown in Figure 5.25.

In this study the profilometer and the DC recorder described in
Chapter IV were used for measuring roughness of three steel surfaces;
ground-finished, cold-rolled, and shot-blasted surfaces. The strip
chart recorder output for each surface type is shown in Figure 5.26. For
each surface represented in this figure, a single 8 cm standard chord was
chosen for practical reasons, hence a value of 8 cm for Lt in Equation
5.7. Note that the chord length of 8 cm, measured on the strip chart,
does not represent the actual tracer distance travelled on the steel
surface. The chord length, and the "roughness line," both depend on the
strip chart movement per unit of time, i.e. chart speed which was
different from that of the tracer on the actual surface. The roughness
factor Q is a relative number which may be used only for discriminating
the three surfaces. Tests were also carried out on three bars, each

with a 45°-lug but having different surface roughnesses; ground—finished

 

 

84

cold-rolled, and shot-blasted as for bars with no lugs.

The load-displacement curves for the rods with no lugs (Figure 5.27)
clearly indicate that the shot-blasted surface with the highest roughness
factor) showed the highest ultimate load. The residual load Pr for this
surface was also the highest compared to the other two surfaces. The
ultimate loads are plotted versus roughness factor on a logarithmic
scale in Figure 5.28. Since the data in this figure are almost linear,
they were approximated by a straight line to give the power law given in
Figure 5.28. Consideration of lug height as part of the roughness

definition will be discussed in Chapter VI.

5.1.6 Water Impurities and Sand Concentration:

The effect of water impurities on the adfreeze bOnd strength of
ice and frozen sand materials was included to provide a basis for
comparison. The initial objective was to evaluate the contribution of
ice adhesion to the total bond strength of frozen sand to plain steel
rods. Samples of crushed ice saturated with distilled water, prepared
and tested at several temperatures, provided a polycrystalline structure
similar to the ice which forms between sand particles. The crushed ice
prepared from ordinary tap water, contained about 450 parts per million
(ppm) in hardness as calcium carbonates.

The experimental results for snow ice (distilled water), crushed ice
(tap water), and ice layers (distilled water) are summarized in Figure
5.29. Data scatter makes comparisons difficult. Bands indicated by the
lines show that water frozen in layers has a significantly lower strength
as compared to crushed ice. The effect of bar size on bond strength

has been neglected based on data shown in Figure 5.17.

 

 

 

 

Ice structure and water impurities represent the only difference
between the three ice types shown in Figure 5.29. To separate the effect
of ice structure from that of water impurities, several samples of fresh
snow, saturated with distilled water, were frozen and tested. Test
results for Ice (3) shown in Figure 5.29 appear to give the highest
bond strengths. Comparing Ice (1) with Ice (3) only on the basis of
structure and ignoring minor difference in water impurities, one may
conclude that the polycrystalline ice possesses a higher adfreeze bond
strength than monocrystalline ice. A comparison of Ice (2) with Ice (3),
neglecting differences due to ice grain size, shows that water impurities
reduce the ice adfreeze bond strength. To verify this conclusion,
several frozen sand samples were fully saturated with tap water during
preparation and were tested. Test results are compared with those from
tests on samples saturated with distilled water in Figure 5.30. The
data clearly indicate the effect of water impurities on bond strength of
frozen sand to plain rods.

Constant displacement rate bond tests were also carried out on
samples with different sand concentrations. Preparation of samples with
a sand fraction less than 59% (by volume) was described in Chapter 111.
Test results, plotted in Figure 5.31, showed that the range in which
strength varied significantly is limited to a sand concentration of 45
percent to 64 percent sand by volume. This conclusion has important
implications relative to design and construction procedures for friction
piles in permafrost. It is believed that the sand concentration in most
naturally and artifically frozen soils would not fall below about 40
percent soil by volume. In addition, the scatter in test results for

samples with vS < 45 percent was large enough so that the bond strength

 

ad

ci

Sp:

 

86

was not clearly defined.

5.2 Constant Stress (Creep) Tests:

In this part of the study, pull-out specimens were tested by
applying a constant load while continuously monitoring the displacement
until either failure occurred or the equipment displacement limit was
exceeded. Most samples were tested by a step loading procedure. Each
load increment was maintained for a time long enough to achieve a steady—
state displacement rate (secondary creep). The effects of load, temperature,
lug size and position, sample diameter, and sand concentration on the
pull-out loads are considered in this section. Physical properties of all
ice and frozen sand samples tested under constant load are listed in

Table 5.4. Test results on these samples are summarized in Table 5.5.

5.2.1 Load Effect:

The data from Section 5.1 permitted a preliminary prediction of
the initial load increment to be applied to each sample. Tests on plain
rods did not show, in general, a clear indication of the actual creep
behavior for the small displacements due to equipment limitations. The
first four samples failed before any precise measurement of the creep
displacement could be taken under each step loading. Transducer limita-
tions t 0.0005 in. prevented more.accurate measurements for the small
displacements associated with creep for the ice adhesion component of
adfreeze bond. The failure displacement was often too small to be pre-
cisely measured with the available recorder at its maximum sensitivity.

However, some data was acquired by running additional tests on

specimens with plain rods. These data were, in most cases, obtained on

 

 

 

87

the basis of visual judgement using a magnifying lense to enlarge the
stylus movement on the strip chart. Typical creep curves for plain rods,
in sand-ice specimens tested at —10°C, are shown in Figure 5.32 (a and b).
For sample number CS-29, Figure 5.32 (a), an initial load increment of
250 lbs. did not appear to achieve the secondary creep stage. A second
load increment, which brought the total load to 530 lbs., was then
applied to the specimen. This load appeared to cause tertiary creep
failure almost immediately in the sample. The duplicate sample, number
08—30 in Figure 5.32 (b), was tested under loads as indicated on the
creep curves.

The displacements plotted versus time in both figures, were close
to the actual values measured during the load application. The creep
displacement rate ac was defined as the slope of the straight line portion
of each curve in Figure 5.32 ( a and b). Data points with the large
scatter, in this figure on plain rods and the subsequent tests, were
approximated by straight lines. The total loads are plotted versus their
corresponding displacement rates on a log-log scale in Figure 5.33. The
four data points appear to be consistent with each other and give a
straight line relationship. Based on data summarized in Figure 5.33 it
appears that the creep displacement rate of plain rods in frozen sand
increases with increase in applied load according to a power law similar
to that given by Equation 5.4.

For rods with lugs, the loading effect was more clearly shown in
contrast to that described for plain rods. The scatter in data points
was far less. A creep curve for step loading, on a 3/8 in. diameter

plain rod with a 900-lug, is shown in Figure 5.34 (a). Several points

 

 

 

 

88

may be featured in this figure. The creep displacements at all loading
stages were large enough to be measured on the recorder with reasonable
accuracy. These displacements were the result of bearing action of frozen
soil on the lug. The stress history appears to have some effect in re—
ducing (or almost eliminating) the primary creep, for loading stages
following the first load increment. 0n the other hand, the displacement
rate does not appear to be much influenced by the stress history, for

the continuous loading condition. To explain this, refer to Figures 5.34
(b) and (c), both of which show creep curves for two more duplicate
samples step-loaded at the same temperature (-10°C). Comparing the
displacement rate of 2.53 X 10'“ in/min. for the third load increment

(P = 2780 lbs.) in Figure 5.34 (a) with that of 2.5 x 10'” in./min. for
the first load increment (same total load, P = 2780 lbs.) in Figure 5.34 (b),
the small difference may be due to interpretations from the plot. Also,
the same comparison would hold for the displacement rates corresponding
to the last load (P = 3560 lbs.) in Figure 5.34 (b) and to the first load
(P = 3560 lbs.) in Figure 5.34 (c), respectively. The difference of

1.7 X 10’” in./min. (about 17 percent) may, again, be considered within

a possible experimental scatter.

Tests on other duplicate samples, Table 5.5, confirmed the same
behavior as shown for the previous three samples. However, when the
sample was unloaded, the creep rate appeared to be far more influenced
by the stress history than in the loading case. Sample 05-18 was step-
loaded from 1220 lbs. to 2380 lbs. as shown in Figure 5.35. After that,
and before total failure, the sample was unloaded back to 2100 lbs.,

1540 lbs., and to the initial load increment of 1220 lbs. As shown in

 

 

 

89

Figure 5.35, only the 2100 lbs. unloading stage showed some creep at a
constant rate of 4.25 X 10’5in./min., which was reasonably close to the
5.09 X 10'5 in./min. value for the loading stage.

When the sample was unloaded from 2100 lbs. to 1220 lbs. in two
steps, the final displacement remained almost constant with very negligible

decrease after 22 hours. This behavior may be attributed to densification

 

in front of the lug during the loading stages. It was also possible that
the 22 hrs., of unloading from 2100 lbs. to 1220 lbs., may not have been
long enough for the sample to reach complete recovery, and then to start
creeping again. Had the recovery occurred, the creep rates then would
have, probably, been slightly less than the corresponding rates of the

loading stages. Creep curves for other duplicate samples were also

 

plotted, with the creep rates summarized in Table 5.5. Creep data for
these samples are given in Appendix B.

Using the data in Table 5.5, the total load P applied at each stage
can be plotted against its corresponding creep rate SC on a log-log scale,
as shown in Figure 5.36. A linear trend on this plot describes the creep
behavior, at least within the range of experimental data. A straight line
was fitted to the data points, using the least-square analysis, and the
empirical power law indicated in Figure 5.36 can be written. Note that
the regression line in Figure 5.36 is ended with dashed lines as shown,

due to the limited data available at both ends.

5.2.2 Temperature Effects:
The temperature effect has been considered on a 5/8 in. diameter
plain rod, and on a 3/8 in. diameter rod with a single lug, both in

frozen sand. A clear conclusion, again, can not be drawn for the plain

 

 

9O

rod samples, because of difficulty with precise measurement of very small
displacements. Three different temperatures; -60C, -10°C, and -15°C, were
selected for comparison. At each temperature two duplicate samples with
the same plain rod (5/8 in. in diameter) were step-loaded until failure.
As observed for the two samples tested at -10°C (Figures 5.32 a-and
b) creep curves for step loading on samples at other temperatures showed
the same behavior. Figures 5.37 (a) and (b) show a similar behavior at
both temperatures with relatively small displacements (less than 0.002 in.)
for all loading stages which was followed by a sudden rupture. The last
segments of both curves appear to represent tertiary creep. Before
failure, the colder temperature, -15°C, reduces creep displacements at all

loading stages. It also reduces the total displacement at failure; for

 

example, compare the value of 0.004 in. for P = 1040 lbs. at -15°C
(Figure 5.37 b) with 0.012 in. for P = 640 lbs. at -69C (Figure 5.37 a).
The total.load P, applied at each loading stage was plotted against
the corresponding displacement rate SC on a log-log scale at ~6°C, -10°C,
and -150C, in Figure 5.38. The data points at all temperatures show
some scatter, especially at —6°C. Only two data points are plotted for
-60C, to remove any possible confusion in showing the temperature effect.
For the other temperatures, —lO°C and -15°C, the data are approximated
by straight lines only for comparison. The general trend of these lines
indicates, qualitatively, that a power law may govern the relationship
between applied loads and the corresponding displacement rates. Note
that other data points for -6°C may all be considered questionable, hence
they were not shown in Figure 5.38. Further tests on duplicate samples

did not appear worth while, because of equipment limitations for small

 

 

displacements.

The temperature effect on lug behavior remains to be covered. Iden-
tical pull-out frozen sand specimens were tested at -2°C, -6°C, and -150C.
Creep curves for step loading on these samples, Figure 5.39 (a to c),
showed a behavior similar to those at -10°C (Figures 5.34 and 5.35). The
data for tests at -2°C, -6OC, and -15°C are compared to those plotted for
-1OQC, in Figure 5.40. The creep displacement rates decreased at colder
temperatures for a given applied load. Under small loads, the displacement
rates appear to be overestimated at all temperatures. However, it is
believed that when the applied load is smaller than the long-term strength,
the steady-state (secondary) creep rate was not achieved within the few
days available.

The data points in Figure 5.40 are approximated by a straight line
for each temperature. The four lines are almost parallal to each other,
implying that the failure mechanism, of bearing action on lugs, is the
same at all temperatures. Assuming a linear relationship on a log-log
scale also implies that a power law of the form defined in Figure 5.40,
would govern the P versus SC relationship. Since the lines were assumed
parallel to each other, the parameter m in Figure 5.40 should be inde-
pendent of temperature. Only the "proof" load Pce is a function of
temperature. The value of PCB corresponds to an arbitrarily chosen

displacement rate 8C of.5 X 10.9 in,/min.

5.2.3 Lug Size and Position:
Creep pull-out tests were conducted on frozen sand specimens with
3/8 in. diameter rods having lugs heights of 1/16 in., 2/16 in., and

3/16 in. The creep behavior for different lug heights was similar to

 

 

 

 

that observed for the 2/16 in. lug height. For comparison the exper-

imental data are summarized in Table 5.5 and are also plotted in Figure
5.41. Seven tests on duplicate samples with a 2/16 in. lug height and
the corresponding data points gave a consistent linear trend as shown in
Figure 5.41. Three samples with a 1/16 in. lug height and two samples

No additional

with a 3/16 in. lug height were tested using step loading.

tests were conducted with the 3/16 in. lug height, since its maximum

 

bearing load would exceed the equipment loading capacity.

All data plotted as straight lines in Figure 5.41. The dashed line
in this figure was estimated from three data points for a creep test on
a standard 3/8 in. diameter deformed bar, with lug height of 1/64 inch.

The load P in Figure 5-42 was divided by 10 for transformation to a single

 

lug load in Figure 5.41. This procedure assumes that the applied load

was equally carried by all 10 lugs for the standard deformed rod in a

3 in. sample height. Although the regression lines in Figure 5.41 are not

exactly parallel to each other, it appeared that the difference in slopes

may have resulted from limited experimental data, especially for the 1/64

inch and 3/16 inch lug heights. The other two lines, for 1/16 in. and ,

2/16 in. lug heights, appear to closely parallel each other, implying that
the bearing action mechanism of frozen soil on any lug size is the same.

A power law, relating the load and creep rate, was deduced for each
lug size and it has the form given in Figure 5.41, where m was made the
same as that for the 2/16 in. lug height, assuming that the lines are

parallel. The "proof“ load PC is dependent on lug size, and its values

are given in Figure 5.41.

Up to this point all lugs tested were positioned at the mid-sample

93

height, i.e. half-way between the loaded end and the free end of the
sample such that z = H/2. The distance 2 equals 0 at the base plate
(loaded end). The effect of possible lug interaction with the base plate
on creep behavior was considered by placement of a single 2/16 in. lug
at z = H/4 and z = 3H/4. The alternate lug locations were obtained by
reversing the rod direction before sample preparation. Test results for
these two lug positions, summarized in Table 5.5, did not show any sig-
nificant difference in behavior when compared with that of the H/2 lug
position shown in Figure 5.43. All data points in this figure appear to
be close enough to neglect any minor differences.

The above results show that lug position does not have an important
influence, if the 2 distance is maintained beyond a height H/4 from the

base place. Since the data are consistent with each other, for all lug

 

 

positions, they were included in one regression analysis in order to
deduce a single empirical equation. This equation, shown in Figure5.43,

is a power law that is valid for the three lug Positions.

5.2.4 Sample Diameter and Sand Concentration:

In concrete structures, a minimum cover requirement is considered
necessary for development of the normal bearing action against lugs for
deformed bars (Mains, 1951; Ferguson, 1966; Perry and Thompson, 1966).

In pull-out tests used to evaluate the load transfer behavior a minimum
over requirement for the steel bars would seem necessary. In this
'nvestigation, the cover required to develop the bearing action on the
ugs is represented by a minimum sample diameter relative to the rod
ize and lug height.

Three additional sample diameters; 1.938irn, 2.781 in., and 4 in. be-

ide the 6 in. diameter were selected for this purpose. The choice of

 

 

94

diameters was based on availability of steel molds in the laboratory.
Sample preparation depended to some extent on sample diameter but was
similar to the standard procedure described in Chapter III. Creep
behavior for the step loading technique on the different pull-out specimens
did not differ from those mentioned earlier.

The pull-out load P is plotted against its corresponding creep rate
ac on a log-log scale in Figure 5.44, for tests on the different sample
diameters. The influence of sample diameter appears to be more significant
for diameters less than 4 inches. The loads, interpolated to creep rates

of 10‘” in./min. and 10-5 in./min., are plotted in Figure 5.45 against a

ratio A defined as:

x = (D-d)/2h (5.8)
where D is sample diameter, d is rod diameter, and h is lug height. This
ratio is believed to be some measurement of the frozen sand cover. The
load appears to increase nonlinearly as the ratio A increases (Figure 5.45)
until it reaches a value of about 13 where the load was no longer affected
by sample diameter. The ”critical“ sample diameter corresponding to
A = 13 is about 4 in. This conclusion indicates that test results,
obtained in this investigation on 6 in. diameter samples were not being
affected by sample diameter.

Another factor, briefly described in Section 5.1.6 on plain rods,
was sand concentration which is now considered in terms of lug behavior.
All samples tested at sand fractions less than 64 percent by volume were
Prepared by mixing a pre-weighed quantity of dry precooled sand with dry
fresh snow to give the desired sand concentration. Two samples of snow-

ice (saturated with distilled water) included the zero sand fraction for

 

 

 

 

 

 

95

creep tests. Samples with sand fraction below 50 percent showed a
different behavior from that observed for samples with 64 percent sand
concentration.

Creep curves for step loading on these samples are presented in
Figures 5.46 and 5.47, including the snow-ice samples. For the initial
load increment of 520 lbs. note that the displacement for all of these
specimens remained unchanged (almost zero reading on the recorder) for
a period of time. This time period appears to be a function of the sand
fraction; longer times for the lower sand concentrations. Thereafter,
the displacement increased with time in the normal manner. Further
discussion on this behavior is made in Chapter VI. The effect of sand
fraction on lug bearing action was similar to that observed for adhesion
and friction to plain rods (Figure 5.31) with constant displacement rate
tests. Loads interpolated to two different creep rates; 10’” in./min.
and 10‘5 in./min., using the load versus creep rate curves in Figure 5.48,
showed that they vary significantly in the range of about 43 percent to
64 percent sand (by volume) as shown in Figure 5.49. In this range,
frictional and dilatational strength components of frozen sand appear to
be more effective. Below a sand fraction of 43 percent the change in
strength was not significant.

To evaluate the amount of lug load reduction as sand fraction vs was
educed from 64 percent to 51 percent (by volume), consider loads at
elected creep rates. Loads for v5 equal to 64 percent (P64 percent) are
lotted against loads for Vs equal to 51 percent (P51 percent) in Figure

.50. An almost linear relationship between P64 percent and P51 percent

is noted for the range of creep rates selected. This result implies

 

 

 

 

 

96

that the amount of load reduction due to decrease in sand fraction was
almost constant and independent of the creep rate ac, at least within a
range of creep rates. As shown in Figure 5.50, values of P64 percent
equal 1.75 times P51 percent, i.e. about 43 percent reduction in the loads

occurred on reducing vs from 64 percent to 51 percent.

 

 

 

 

 

 

 

97

le 5.1: Physical properties of ice, frozen sand specimens, and steel
rods, for constant displacement rate tests.

 

 

 

 

 

 

 

ple Mater- Steel VS H D d h
o. ial surface (%) (10.) (10») (l”°) (l”°)
Ice(1)* CR“ 0 5.866 5.677 5/8 0i
-- .. 0 5.750 5.875 5/8 0
.. .. 0 5.688 5.875 5/8 0
.. u 0 5.438 5.938 5/8 0
.. .. 0 5.938 5.875 5/8 0
.. .. 0 5.938 5.969 5/8 0
.. .. o 5.938 5.875 5/8 0
.. .. 0 6.125 5.938 5/8 0
.. .. 0 5.875 5.938 5/8 0
0 .. .. 0 5.969 5.938 5/8 0
1 .. .. 0 5.938 5.938 5/8 0
2 .. .. 0 6.000 6.000 5/8 0
3 .. .. 0 6.000 6.000 5/8 0
4 .. u 0 5.906 6.000 5/8 0
5 .. .. 0 5.938 6.000 5/8 0
6 .. .. 0 5.938 6.000 5/8 0
.. .. 0 5.344 6.000 5/8 0
.. .. 0 5.906 6.000 5/8 0
9 .. .. 0 5.938 6.000 5/8 0
3 .. .. 0 5.000 6.000 5/8 0
.. .. 0 6.000 6.000 5/8 0
Rem... .. 0 6.000 6.000 3/8 0
.. .. 0 6.500 6.000 3/8 0
.. u 0 6.375 6.000 3/8 0
.. .. 0 6.125 6.000 3/8 0
.. .. 0 6.000 6.000 3/8 0
.. .. o 6.000 6.000 3/8 0
Ice(3)* " 0 6.197 6.000 5/8 0
.. .. 0 6.000 6.000 5/8 0
.. .. 0 6.375 6.000 5/8 0 (Continued)
.. .. 0 6.375 6.000 5/8 0
l

 

98

1le 5.1: (Cont'd)

 

 

 

1ple Mater- Steel Vs H D d h
lo. ial surface (%) (in.) (in.) (in.) (in.)
:2 Ice(3) CR 0 6.500 ' 6.00 5/8 0
. FS(DW)* CR 64.0 6.000 6.00 5/8 0
2 " " 64.0 6.000 6.00 5/8 0 l
1 " " 64.0 6 000 6.00 5/8 0 ’
1 " " 64.0 6.000 6.00 5/8 0
s " " 66.3 6.871 6.00 5/8 0
s " " 64.0 6.000 6.00 5/8 0
1 " " 64.0 6.000 6.00 5/8 0
3 " " 64.0 6.000 6.00 5/8 0
a " " 64.6 5.938 6.00 5/8 0
10 " " 64.0 6.000 6.00 5/8 0 1
11 " " 64.0 6.000 6.00 5/8 0 3
12 " " 64.0 6.000 6.00 5/8 0
13 " " 64.0 6.000 6.00 5/8 0
14 H " 64.0 6.000 6.00 578 0
.5 H " 64.5 6.000 6.00 5/8 0
.6 " " 64.0 6.000 6.00 5/8 0
7 " " 64.0 6.000 6.00 5/8 0
8 " " 64.0 6.000 6.00 5/8 0
9 H " 64.0 6.000 6.00 5/8 0
0 " " 64.0 6.000 6.00 5/8 0
1 " " 64.0 6.000 6.00 5/8 0
2 H " 64.0 6.000 6.00 5/8 0
3 H " 64.0 6.000 6.00 5/8 0
I " " 64.0 6.000 6.00 5/8 0
3 n " 64.0 6 000 6.00 5/8 0
s H " 64.0 6.000 6.00 5/8 0
' " " 64.0 6 000 6.00 5/8 0
n H 64.0 6.000 6.00 5/8 0
n " 64.0 6.000 6.00 5/8 0
1' " 64.0 6.000 6.00 3/8 0 + (Continued)
1' " 64.0 6.000 6.00 3/8 1/16

 

99

able 5.1: (Cont'd)

 

 

 

 

ample Mater- Steel Vs H D d h
.No. ial surface (%) (in.) (in ) (in.) (in.)
-32 FS(DW) CR 64.0 6.00 6.00 3/8 2/16
-33 " H 64.0 6.00 6.00 3/8 3/16 ,
-34 " H 64.0 6.00 6.00 3/8 3/16 1
-35 H H 64.0 6.00 6.00 3/8 0 I
-36 H H 64.0 6.00 6.00 3/8 1/16
-37 " H 64.0 6.00 6.00 3/8 2/16
-38 " H 64.0 6.00 6.00 3/8 3/16
-39 H H 64.0 6.00 6.00 3/16 0
-40 H H 64.0 6.00 6.00 3/32 0
-41 H H 64.0 6.00 6.00 3/16 0
-42 H H 64.0 6.00 6.00 3/32 0
-43 H H 64.0 6.00 6.00 3/8 2/16
-44 H H 64.0 6.00 6.00 3/8 2/16
-45 H H 64.0 6.00 6.00 3/8 2/16
-46 H H 64.0 6.00 6.00 3/8 2/16
-47 H H 64.0 6.00 6.00 3/8 2/16
-48 H H 64.0 6.00 6.00 3/8 2/16

-49 H H 64.0 6.00 6.00 3/8 2/16

-50 H H 64.0 5.00 6.00 3/8 0

-51 H H 64.0 4.31 6.00 3/8 0

-52 H H 64.0 6.00 6.00 3/8 0

-53 H H 64.0 3.00 6.00 3/8 0

.54 H H 64.0 2.00 6.00 3/8 0

55 H H 64.0 1.00 6.00 3/8 0

56 H H 64.0 1.00 6.00 3/8 0

57 H H 64.0 6.00 6.00 3/8 2/16

58 H H 64.0 1.00 6.00 3/8 0

59 H H 64.0 1.00 6.00 3/8 0 .

60 H H 64.0 6.00 6.00 3/8 0

51 H H 64.0 6.00 6.00 3/8 0

32 H H 64.0 6.00 6.00 3/8 2/16 (continued)
53 H H 64.0 6.00 6.00 3/8 0

 

 

100 1

)le 5.1: (Cont'd)

 

 

 

 

 

hple Mater- Steel Vs H D d h
19. ial surface (%) (in.) (in.) (in.) (in.)
54 FS(DW) CR 64.0 6.0 6.0 3/8 0
55 " H 64.0 6.0 6.0 3/8 0
56 H H 64.0 6.0 6.0 3/8 0 1
57 " H 64.0 6.0 6.0 3/16 0 i
58 " H 64.0 6.0 6.0 3/8 2/16
59 H H 64.0 6.0 6.0 3/8 0
70 " H 64.0 6.0 6.0 3/16 0
71 " H 64.0 4.0 6.0 3/8 0
72 " H 64.0 6.0 6.0 3/8 0
73 " H 64.0 6.0 6.0 3/8 0
74 H H 64.0 6.0 6.0 5/8 0 .
75 H .. 64.0 6.0 6.0 3/8 0 I
76 H H 64.0 2.0 6.0 5/8 0 V
77 " H 64.0 5.0 6.0 3/8 0
78 H .. 64.0 5.0 6.0 5/8 0
79 " H 64.0 3.0 6.0 3/8 0
zo H " 64.0 4.0 6.0 5/8 0
u '7 H 64.0 2.0 6.0 3/8 0
2 H H 64.0 3.0 6.0 5/8 0
3 H H 64.0 2.0 6.0 5/8 0
4 H '7 64.0 2.0 6.0 5/8 0
5 H GF** 64.0 6.0 6.0 3/8 0
6 " H 64.0 6.0 6.0 3/8 0
7 H H 64.0 6.0 6.0 3/8 2/16 (45°)+
3 H 38** 64.0 6.0 6.0 3/8 0
l H CR 64.0 6.0 6.0 3/8 2/16 (45°)
1 H SB 64.0 6.0 6.0 3/8 2/16 (45°)
H CR 64.0 6.0 6.0 3/8 2/16 (45°)
H H 64.0 6.0 6.0 3/8 0
H H 64.0 6.0 6.0 3/8 2/16 (450)
u n 54,0 6.0 6.0 3/8 0 (Continued)
5 n 54.0 6.0 6.0 3/8 0

101

ale 5.1: (Cont'd)

 

 

 

 

 

nple Mater- Steel Vs H D d h
V0. ial surface (%) (in.) (in.) (in.) (in.)
96 FS(DW) CR 64.0 6.00 6.0 3/8 1/8 (45°)
97 " H 64.0 6.00 6.0 3/8 1/8 (45°)
98 " H 64.0 +2.00 6.0 5/8 0
99 " H 64.0 4.00 6.0 5/8 0
100 H H 64.0 3.00 6.0 5/8 0
101 H H 64.0 5.00 6.0 5/8 0
102 H H 64.0 6.00 6.0 5/8 0
103 H H 64.0 2.00 6.0 5/8 0
104 " H 64.0 0.50 6.0 5/8 0
105 H H 64.0 5.00 6.0 5/8 0
106 H H 64.0 2.00 6.0 5/8 0
107 H H 64.0 5.00 6.0 5/8 0 .
108 H H 64.0 3.00 6.0 5/8 0 1
109 H H 64.0 6.00 6.0 5/8 0
110 H H 64.0 4.00 6.0 5/8 0
111 H H 64.0 5.50 6.0 3/8 0
112 " H 64.0 5.62 6.0 3/8 0
113 H H 64.0 6.00 6.0 3/8 0
.14 " H 64.0 6.00 6.0 3/8 0
‘15 FS(TW)* H 64.0 6.00 6.0 3/8 0
16 H H 64.0 6.00 6.0 3/8 0
17 H H 64.0 6.00 6.0 3/8 0
18 '1 H 64.0 6.00 6.0 3/8 0
19 FS(DW) H 64.0 6.00 6.0 5/8 0
20 H H 64.0 2.00 6.0 3/8 s08”r
21 H H 64.0 3.00 6.0 3/8 SDB
22 H H 64.0 6.00 6.0 3/8 1/16
23 F(SID)* H 42.0 6.00 6.0 3/8 0
:4 H H 29.6 6.00 6.0 3/8 0
' H H 53.8 6.12 6.0 3/8 0 (Continued)
6 H H 52.7 6.00 6.0 3/8 0

102

>le 5.1: (Cont'd)

 

 

 

 

 

iple Mater- Steel Vs H D d h
[9. ial surface (%) (in.) (in.) (in.) (in.)
.27 F(SID) CR 30.3 6.0 6.0 3/8

.28 " " 44.2 6.0 6.0 3/8

Ice - Ice layers (distilled water)

(1)
(2)- Crushed ice (tap water) saturated with distilled water
Ice(3)- Snow ice saturated with distilled water

DW)- Frozen sand (with distilled water)
FS(TW)- Frozen sand (with tap water)
F(SID)- Frozen sand (snow ice with distilled water)

CR- Cold-rolled steel surface
GF- Ground-finish steel surface
SB- Shot-blasted steel surface

Plain rods for h = 0

All 909-lugs, unless noted

(450)- 45 degree lug

SDB- Standard deformed bar, with h = 1/64 in.

 

 

103

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

19_§;§i Summary of test results for constant displacement rate

tests.

ple T 6n P “P Puff 5r

(or P)

p. ( °C) (in./min.) (lbs/min) ((15s“ ) (lbs ) (x10-3in.)
-10.1 4 24x10'3 140.0 930 190 0.90
-15.1 4.50x10'3 136.8 620 105 0.76
-15.1 4 40x10'3 130.0 607 200 1.50
-15.1 4.90x10'3 105.4 455 96 1.10
- 6.1 * * * * *

- 6.1 3.68x10’3 89.6 112 56 0.40
-20.0 3.90x10'3 136.4 500 175 0.001
-20.0 3.60x10‘3 115.7 216 130 0.28
-10 0 3.55x10'3 111.8 326 82 0.40

0 -10.1 3.32x10'3 91.6 336 94 0.20

1 -10.1 3.43x10'3 94.0 470 78 0.20 i

2 -26.1 9.00x10‘4 126.0 840 120 0.20 1

3 -26.7 4.00x10'3 143.2 1480 180 0.24

4 -26.4 3.50x10'4 4.9 440 ‘320 2.00

5 -26.6 8.00x10'4 22.5 990 140 0.80

6 -20.4 1.05x10'3 27.0 1040 90 1.40

7 -15.5 3.52x10'4 23.6 2830 200 5.70

3 -10.1 2.30x10'4 28.5 890 80 0.75

9 - 6.6 not recoded 22.1 1940 300 4.40

1 -10.1 2.60x10-4 26.4 1780 200 2.80

1 -20.1 6.15x10'4 23.0 1740 210 0.80

2 -10.1 5.70x10'4 21.6 540 30 0.20

1 -22.0 5.645(10’4 20.1 624 50 0.40
-20.0 5.105(10’4 17.2 560 32 0.40

~ -15 0 5.40x10’4 20.0 670 40 0.10

-6.0 5.655(10'4 16.1 210 16 0.10
- 2.0 5.80x10’4 19.1 105 10 0.10
-10.0 4.50x10'3 183.5 1560 70 0.75
-14.8 1.92x10'3 233.3 1435 30 0.50
-16.0 4.50x10”3 280.0 880 50 0.25

(Continued)

 

)le 5.2: (Cont'd)

104

 

nple T én g
16. ( °C) (in./min.) (lbs/min)
31 - 6.0 Fast rate 1290.0
32 -10.0 4.61x10’3 258.0
1 -10.0 2.40x10‘3 159.5
2 -10.2 5.001110"4 32.5
3 - 6.3 6.97x10'4 29.7
1 -14.6 3.67x10'4 30.7
5 -20.0 5.00x10'4 32.7
5 -26.7 6.38x10'4 37.5
7 -26.8 6.15x10'4 39.7
3 -27.1 1.32x10'3 208.7
9 -20.0 2.53x10"3 210.3
10 - 6.3 1.12x10'3 186.7
11 -15.1 3.86x10'3 204.6
12 -10.0 3.73x10'3 222.7
13 - 6.0 1.96x10'2 2133.3
14 -15.2 3.00x10‘2 2033.3
15 -10.0 2.13x10'2 1978.8
16 -25.6 4.00x10’2 2210.0
17 -20.0 1.32x10’2 2240.0
.8 -25.0 1.10x10'4 5.1
.9 —1o.0 1.66x10'4 4.5
:0 - 6.0 1.00x10"4 4.4
:1 —15.0 1.86x10'4 6.5
2 -20.0 1.33x10‘4 4.4
3 -10.0 3.83x10'3 268.1
4 - .0 4.00x10’3 100.0
5 - .0 3.71x10'2 1714.3
6 - .0 8.60x10’5 2.1
7 - .0 6.00x10"4 26.3
3 -20.0 3.90x10“2 2478.3

 

 

 

 

 

. P
(0:1PU) (oruPr) 6:3
(lbs.) (lbs.) (x10 in.)
430 Not recored
1030 70 0.001
1075 135 0.16
1300 280 2.20
860 400 1.60
2090 520 2.50
2810 575 1.80
2740 600 1.90
2800 450 0.80
3800 920 1.00
3260 580 0.80
1400 200 1.80
2660 430 1.20
2450 970 2.60
1920 330 1.40
3660 530 0.60
2770 440 0.40
4420 900 0.70
3360 750 0.20
3440 1080 2.90
1630 1200 1.80
720 380 0.50
2080 580 1.50
2770 1000 1.40
2480 540 1.50
450 240 4.10
720 320 6.10
210 190 3.60
460 350 2.40
3420 770 0.01

(Continued)

 

 

e 5.2 (Cont'd)
1le T é é Piy Pu 5r
0 . n. (or Pu) (or P ) -
L;_ ( ) (1n./m1n.) (lbs/min) (lbs.) (lbs?) (x10 3in.)
1 -15.0 LVDT OFF 220.9 2540 660 LVDT OFF
1 -10.0 8.00x10'4 29.3 880 140 0.80
. -10.0 5.20x10'4 29.7 1560 1040 1.20
E -10.0 5.33x10'4 33.4 2040 1520 2.00
1 -10.0 5.825(10'4 32.0 2190 1830 2.00
-10.0 4.21x10’4 31.3 2320 4800 2.80
5 -20.0 6.94x10“4 138.7 1040 300 1.00
5 -20.0 5.10x10'4 32.8 1820 2660 0.01
' -20.0 3.90x10'4 37.5 2360 3580 1.60
-20.0 4.10x10"4 35.8 3400 7020 0.60
1 -10.0 5.00x10'4 27.1 130 40 0.40
1 -10.0 4.32x10'4 5.3 80 20 1.20
1 -20.0 5.43x10‘4 15.8 530 50 0.40
a -20.0 4.50x10“4 7.4 152 80 0.20
1 -10.0 2.10x10'2 2220.0 2220 .4320 3.00
* -10.o 3.36x10'3 198.1 2080 4540 1.60
i -10.0 7.00x10'5 4.8 1120 1000 2.00
-15.0 2.80x10'4 30.6 2080 3800 1.80
-26.0 5.02x10’4 33.0 2540 6080 2.00
- 6.0 5.70x10'4 32.4 1200 2080 2.60
- 2.0 3.285710'4 17.0 440 1390 1.50
-10 0 5.63x10'4 28.4 610 120 0.40
-10 0 1.14x10'4 5.2 730 120 0.01
-1o.0 1.40x10'4 6.3 940 140 0.01
-10.0 5.63x10'4 24.3 54 130 1.00
-10.0 6.50x10'4 20.4 224 64 0.10
-10 0 1.44x10'4 9.6 48 28 2.00
-10.0 3.00x10’4 14.4 144 64 0.16
-10.0 4.435(10'4 34.7 1560 2920 0.20
-10.0 6.00x10'4 10.4 188 40 0-40
(Continued)

105

 

 

 

 

le 5.2: (Cont'd)

 

106

 

 

 

 

 

le T ' ' P' P
p _0 . 6n. P . (or1Pu) (oruPr) 6:3
0. ( ) (1n./m1n.) (lbs/min) (lbs.) (lbs.) (x10 in.)
9 -10.0 6.80x10"4 13.7 96 40 0.16
0 -15.0 1.255(10’4 33.8 760 200 0.10
1 -26.0 6.31x10'4 24.1 880 400 1.00
2 -20 0 4.65x10'4 29.0 1740 4580 0.10
3 - 2.0 2.25x10‘4 17.3 260 50 0.60
4 - 6.0 9.50x10'5 18.2 520 100 0.30
5 -15.0 6.31x10'4 22.6 880 220 0.10
6 -20 0 6.47x10'4 23.9 800 320 0.40
7 -20.0 LVDT OFF 16.3 480 210 LVDT OFF
8 -20.0 2.30x10‘4 17.7 1220 4580 0.10
9 -25.4 6.50x10‘4 21.8 1060 420 0.05
0 -10.0 5.87x10'4 12.3 344 120 0.40
1 -10.0 6.10x10‘4 19.5 495 95 6.00
2 -10 0 5.50x10'4 12.0 520 160 0.80
3 -15 0 4.675110'4 25.5 1170 470 0.40
4 -20.0 * * * * *
5 - 6.0 4.50x10‘4 11.6 625 175 0.10
6 * * ‘k ‘k * 'k
7 -10 0 4.10x10‘4 11.7 780 190 2.00
3 -10.0 * * * * *
9 -10.0 4.75x10'4 13.7 540 140 1.40
1 -10 0 * * * * *
1 -10.0 4.50x10'4 9.0 270 60 0.80
2 -10.0 * * * * *
1 -10.0 3101104 6.8 290 80 0.60
: -10.0 * * * * *
1 -10 0 3.24x10'4 20.0 530 180 0-10
1 -10.0 2.80x10"4 24.4 670 140 0.10
-10.0 5.93x10'4 23.2 2090 2730 0.40
-10.0 7.50x10'4 18.5 1740 730 3.20

(Continued)

 

 

 

e 5.2 (Cont'd)
1le T ' ' P' P
o . 6n. P . (orTPu) (orupr) 6E3
_;_ ( ) (1n./m1n.) (lbs/min) (lbs.) (lbs.) (x10 in.)
-10 0 5.501110“4 25.0 1560 3380 0.10
1 -10 0 5.39x10'4 30.0 2160 3400 1.00
-15.0 5.00x10‘4 23.0 2160 4360 1.50
-15 0 6.25x10'4 20.5 1250 280 0.80
-20.0 5.20x10'4 34.9 2720 4620 0.10
-20.0 5.85x10'4 18.0 1700 440 0.40
1 -26.0 6.671(10'4 22.2 1840 640 0.10
1 - 6.0 5.601110"4 27.6 1380 2040 3.00
- 2.0 5.67x10‘4 19.3 580 1520 4.50
1 -10.0 5.73x10’4 25.0 576 145 0.40
1 -10 0 6.67x10’4 36.9 1920 480 1.80
10 -10.0 5.60x10'4 31.2 1280 290 0.80
11 -10.0 6031104 44.2 3120 1220 2.00
12 -10.0 5.71x10’4 41.7 3380 1180 1.50
13 -10.0 5.70x10’4 28.1 980 . 260 2.40
1.4 -10.0 * ‘k 'k 'k *
5 -10.0 * * * * *
6 -10.0 5.86x10'4 25.0 576 120 1.60
7 -10.0 * * * * *
3 -10 0 5.80x10'4 26.5 530 130 0.01
9 -10 0 5.001110"4 35.0 1960 480 2.00
1 -10 0 5.561(10‘4 28.2 500 120 0.80
1 -10.0 6.671(10'4 20.0 560 40 0.50
2 -10 0 6.251110"4 21.2 640 80 0.10
1 -15.0 4.751110'4 25.2 1060 280 0.70
-15.0 4,00x10'4 25.8 1250 240 0.40
1 -10 0 4.201110‘4 17.3 580 30 1.00
-10 0 4.841110’4 20.3 460 40 0.40
-20.0 2.101(10‘4 24.6 1330 180 0.60
-20 0 2.801710‘4 22.0 810 170 0.10

107

 

 

 

 

 

(Continued)

 

108

able 5.2: (Cont'd)

 

. . p. P
amp16 6 . 6“. P _ (or1Pu) (oruPr) 6f3
JE1;_ _(__Q) (1n./m1n.) (lbs/min) (lbs.) (lbs.) (x10 in.)
-119 -10 0 1.17x10'4 33.4 1370 240 0.90
-120 -10.0 4.601(10'3 170.0 1420 2300 7.50
-121 -10.0 4.50x10'3 235.0 2100 4025 4.40
-122 -10.0 6.70x10'4 40.0 1600 1750 1.80
-123 -10.0 8.20x10'4 22.0 375 120 0.50
-124 -10 0 6.80x10'4 31.8 525 60 0.20
.125 -10.0 7.001x10'4 20.8 225 75 0.10
126 -10 0 8.401(10‘4 44.0 710 145 0.40
127 - 9.9 1.10x10‘3 31.6 525 60 0.80
128 -10.0 8.10x104 30 0 320 40 0.50

_________________________________________________~_____________
Piy for lugs, Pu for plain rods (both correspond to the small
displacement 6r)

Pu for lugs, Pr for plain rods (both at large displacements)
Data void (leakage of fluid or equipment defect)

1le 5.3: Temperature effect on the parameters used in Equations

 

 

 

5.1 to 5.4.
Tc* m n Td** m' n'
C) 12§;1 12511 ______

16 0.1945 5.40 17 0.2220 4.50
48 0.1585 6.31 62 0.1852 5.40
113 0.1087 10.40 137 0.1404 9.35
145 0.1000 11.40 167 0.1016 10.60
213 0.0400 23.00 240 0.0436 22.90
270 0.0314 31.00 308 0.0314 31.00

 

‘nterpolated from Figure 5.3 for PC — l lb./min.
‘4 in./min.

nterpolated from Figure 5.4 for °c — 10

 

109

e 5.4: Physical properties of ice, frozen sand specimens, and
steel rods used in constant load (creep) tests. Sample
height H = 6 in. and cold—rolled steel rods were used
for all specimens.

 

 

 

 

le Mater- VS 0 d h z
___ 1§fl____ .L;2_1 .Lifl;1 (in.) (in.) (in )
Ice(1)* 0 6 00 3/8 1/8** 3 0+
" 0 6 00 3/8 1/8
FS(DW)* 64 00 5/8 0** NA
H 64. 00 5/8 0 H
H 64. .00 5/8 0 H
'1 64. 00 5/8 0 H
H 64. 00 3/8 2/16
H 64. 00 3/8 2/16
H 64. 00 3/8 2/16
H 64. 00 3/8 2/16
H 64. 00 3/8 1/16
H 64. 00 3/8 2/16
H 64. 00 3/8 3/16
H 64. 00 3/8 2/16
" 64. 00 3/8 1/16

64.
” 64.
” 64.
" 64.
" 64.
” 64.
" 64.
" 64.

-vwWNO\U‘I-¥>Wl\>1—IO

.938 3/8 2/16
.00 3/8 2/16
.938 3/8 2/16
.00 3/8 2/16
.938 3/8 2/16
.00 3/8 2/16
.781 3/8 2/16

" 64.
" 64.
" 64.
” 64.

OOOOOOOOOOOOOOOOOOOOOOOOOOO
NOSI—IOSI—HOSHmmmmmmmmmmmmmmmmmmmon
O
O
(A)

\

(I)
N
\
...—J
OW

.CDzJ'IOOOU'IOU'IOOOOOOOOOOOOOOO

Continued)

Table 5.4: (Cont'd)

 

Sample
No.

68-28
CS-29
CS—30
CS—31
CS-32
CS-33
03-34
38-35
33—36
IS—37
IS-38
SS-39
6—40
S-41
S-42
S—43
S-44
S-45

 

Mater-

ial

FS(DW)

Vs

( 7

 

64

64.
64.
64.

0
0
0
0
0
.0
0
0
6
3
6

6
8
5
.0
0
0
0

6.
6.

6
6
6
6.
6
6

6
6
6
6.
l
4
6
6

4.
6

00

See Table 5.1 for abbreviations.
* All 90 degree lugs, h = 0 for plain rods.

Lug position from reaction base plate
l—Not applicable

lote: For sample cs-44, H = 3 in.)

508*
1/8

wwwmwwww
00000000

00 Z
. > .
O

 

111

Table 5.5: Summary of test results for constant load (creep)

 

 

 

tests.
tc

Sample 0T P U
__111;_ I__£;JL (lbs.) (in./min.)
CI-l -10.0 520 2.881110"3
CI-2 -10.0 200 2.701110“7
280 62.5x10'7
360 6.25x10'7
440 20811106
520 4.34x10‘6

Cs-1 -10 0 1420 TC

05-2 -10.0 1040 TC
CS-3 -10.0 750 4.41x10‘°
910 2.00x10‘5

cs-4 -10.0 520 TC
CS-5 -10.0 2300 1.00x10'4
2500 1.68x10'4
2780 25811104
3020 3.75x10'4
3180 4.46x10'4
:s-6 -10.0 2800 2.50x10‘4
3080 4.361110'4
3320 6.961110'4
3560 8.30x10‘4
:s-7 -10.0 3560 1.00x10'3
3780 1.66x10'3
4000 2.00x10'3
S-8 -10.0 4000 2.861110'3
5-9 -10.0 1400 4 67x10'4
1560 7.00x10'4
1800 1 0411103

2080 1.66x10-3 (Continued)

 

112

Table 5.5: (cont'd)

 

 

 

Sample T P 39
No. ( °C ) (lbs.) (in./min.)
cs-10 -10.0 1580 8.33x10'5
1860 9.38x10'5
2140 13311104
2300 1.31x10'4
03-11 -10.0 2000 1.61x10“5
2480 4.05x10’5
3040 6.71x1o'5
3600 9.07x10'5
4160 2.04x10'4
cs-12 -10.0 4000 LVDT OFF
cs-13 -10.0 850 8.00x10‘5
1200 2.12x10“4
1560 5.89x10'4
2600 7.181110'3
05-14 -10.0 4500 9.75x10‘3
cs-15 -10 0 3000 2.50x10'2
Cs-16 -15 0 1550 1 881410"5
2000 2 75x10“5
2500 5.391110"5
3000 1 06x10"4
3500 2 45x10‘4
4000 7 00x10"4
-5
:5-17 —10.0 2500 8 00x10 4
3500 1.00x10‘
2500* u 1 41x10‘5
-5
s-18 -10.1 1220 1.82x10 5
1540 1.77x10'
1820 3.64x10“5

(Continued)

 

113

Table 5.5: (cont'd)

Sample T P éc
No. ( OC 1 (lbs.) (in./min.)
CS-18 -10.1 2100 5.09x10'5
2380 6.94x10'5

2100 u 4.25x10'5
1540 U Negative value

1200 u 0
cs-19 -10.0 1500 9.271110‘4
cs-20 -10 0 1500 4.58110"5
2000 1.11x10'4
2500 3.85x10'4
3000 9.30x10'4
CS—21 -10.0 1000 3.20x10‘5
1740 2.92x10‘4
CS-22 -10.0 1500 4.13x10'5
2000 6.25x10‘5
2500 zoom"4
3000 5.221110“4
3500 1.28x10'3
:s-23 -10.0 1500 2.83x10'4
s-24 — 6.0 1020 1.97x10‘5
1500 4.57x10'5
1980 1.21x10’4
2500 5.40x10'4
3000 2.88x10'3
5—25 -10 0 1500 1.69x10‘4
1980 1.00x10'3
-6

1-26 -10.0 1020 3.89x10
1500 3.75x10‘5

(Continued)

 

_-__— ___‘_—‘ l ‘

114

Table 5.5: (Cont'd)

 

 

 

 

 

Sample T i P 6C
No. , ( °C ) (155.) (in./min.)
CS-26 -10.0 1940 3.751110"5
2500 1.94x10'4
05-27 -10.0 1020 4.551110"6
1500 1.13x10'5
2020 1.20x10‘4
2500 4.641110”4
3000 TC
CS-28 -10.0 1020 7.57x10“7
1500 2.961110"6
1980 64311106
2500 2.30x10‘5
3000 2.02x10"4
cs-29 -10.0 250 0
530 4.50x10’6
Cs-30 -10.0 250 57111107
410 2.051110“6
570 4.93x10”6
730 TC
:s-31 -10.0 160 =0
240 =0
360 =0
480 =0
600 =0
720 20
840 =0
960 =0
1080 TC
(Continued)

115

Table 5.5: (Cont'd)

 

 

 

 

 

Sample T P 6C
No. ( OC ) (lbs.) (in./min.)
05-32 -15.1 240 PC
400 =0
600 =0
760 =0
900 =0
1080 =0
1280 TC
cs-33 -15 1 400 1.101110"7
560 24417107
720 23251107
880 8.33x10'7
1040 4.96x10'°
CS-34 - 6.1 160 4.36x10—7
280 4.101(10‘7
400 1.251110”6
520 50051107
640 TC
cs-35 - 6.0 280 7.581110“7
400 5.80x10’7
520 1.00x10'6
640 8.30x10”7
720 1.00x10'6
800 9.60x10'7
920 2.20x10"6
1040 TC
CS-36 -10.0 520 1.66x10’5
1080 2.26x10T4
1560 1.22x10'3

(Continued)

 

116

able 5.5: (Cont'd)

 

 

 

 

ample T P éc
N0. ( °C ) (lbs.) (in./min.)
s-37 -10.1 520 7.05x10"5
1080 1.68x10‘3
S-38 -10.0 520 8 5011106
1080 2.441110”5
1520 2 68x10"4

1980 TC
s-39 -10 0 520 9.521110"6
1080 2.121110"4
1560 1.25x10"3
s-40 -10.0 520 5 38x10"6
720 7 00x10‘6
1080 7.37x10'5
1240 7 30x10"5
1560 2 8511104
5-41 -10.0 520 21.0110"4
600 5.00x10'4
680 12911103
5-42 -10.0 520 2.70x10'6
720 4.98110”6
1080 28011105
1240 3.57x10"5
1560 8.29x10'5
1840 2.50x10'4
2120 1.281110”3
-43 -10.0 1080 1.27x10’5
1440 2.831110'5
1840 5.00x10’5
2200 1.051110“4

(Continued)

117

able 5.5: (Cont'd)

 

 

 

ample T P 5c
No. ( °C ) (lbs.) (in./min.)
5-43 -10 0 2560 2.35x10'4
2840 5.62x10'4
5-44 -10.0 1080 5.60x10'6
1560 9.00x10'6
1980 3601110"5
2560 TC
5-45 - 2.0 520 1.10x10‘5
800 2.70x10‘5
1160 1.28110"4
1560 1.04x10'3

 

Primary creep
Tertiary creep
Unloading

 

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for the bond of frozen sand (vS = 64%

D 250
. A. -
_Pr/P =1978.6lbs./min. _
‘ (S-15)
_ E“ _200
- -‘-—_222.7 118.7an
\. (s- 12)
- (D
"\-
1 — I \ (Linear scale between _150 '5
, 35 0 lbs /min numbered points) °-
/1 (5-109) ‘5
1-EI ‘ - 3:
. 1 F
l S
_ 1 Plain rod, 3 = 5/8 in. 1100 E
- Frozen sand (v5 = 64%) 13
*l D=6in.,H=6in. ‘
) -' T = —10‘t
I \
l P \ \ A
1 o \
I 1 ES!“ \ A - 50
\ _\ \. __\
1 \ I O-O E! O A—A
' ‘d
I
I I I I a I I ' ‘ A ‘ TL 0
0 Si 1 2 3 4 5 6
Displacement, 6, X 10'2 in.
5.2: Typical load-displacement curves at different loading rates

) to a plain steel rod.

 

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121

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122

 

 

 

 

F (5-44)
ELIJ-—————--———-C1
/ 1% = 3.36 x 10'3
/El 1'n./m1’n.
E]
- E]
Pu (S-57)
<——— _____ —O--—-—O
/,O 5.33 x 10'”
in./min.
/ OI
_ /
I
I/
_ Eli] }3
, T = -100c
65/ h = 1/8 in.
C} / d — 3/8 in.
1 x (9’ v = 64%
1 s
7 q\ 1”
1 ‘®~®@’
1
l
,.
l
1
/ (Linear sca1e between numbered points)
I I I l E 1 I 1 I 4: I
O E' 0.01 0.02 0.10 0.20 0.30
C)

Disp1acement, 5 , in.

ure 5.5: Typica1 1oad-disp1acement curves at different disp1acement
rates for a p1a1n rod with a sing1e 1ug.in frozen sand
(v5 = 64%).

 

 

"- 1h,

 

123

 

 

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124

(Linear sca1e between numbered points)
I l I l l l l l

 

     
  
   
  
   

-1 210
P1ain rod, d = 5/8 in. ._
Pavg = 24.4 1bs./min.
:5an = 4.5 x 10‘L+ int/min - 170
Frozen sand (vS = 64%)

.. 140

I
I3
01
Bond Stress; T, psi

 

 

" 7O

.1

- 35
0

 

Disp1acement, 6, X 10'2 in.

' o n . t
1 5.7: T ica] 1oad-disp1acement curyes at differen 0
gure tggperatures for a p1ain rod in frozen sand (vS = 64%),

 

0
[ P1ain Rod
_ d = 5/8 in. [,0
v5 = 64% I I

P
avg.
(1bs./min)

 

 

Temperature, T, 0C

e 5.8: Temperature effect on the bond strength of frozen sand
(vS = 64%) to a p1ain stee1 rod, at severa] 1oading rates.

126

;. (VaTues of Tu adjusted to P avg' using data from
‘ Figure 5. 3)u

Best Fit Lines

- = +
Tu a be

     
   
  

avg a b
(Tbs./min.) (psi) (psi/0C)

_ / /1:1

- A 5«6 0 12.0
—- /// // .1 .1 28.0 0 12.6
- o //A E] 157.4 55 11.0

/ I/_ 0 2136.0 116 9,75

 

-/////.. Jun-I’ll! IIIIJIII

Temperature, T} 0C

5.9: ReTationship between temperature and bond strength of frozen
sand (v = 64%) to a 5/8 in. pTain rod, at severaT Toading

rates.

4O

32

24

a 5.10:

 

 

127

 

/
/
P1ain Rod ’/
d = 5/8 in. //
_ [[0 /’. —26 -
vS — 6.A /, éf
,/ . .
GII/ #1
WG)
/ -20 — E
/ E
‘(D
:S 45 E-
0)
/ I—
.o/ -15 _
-10 .—
1’
{’1’ —6 _.
‘D -2 -
/
2’
I J I I I 1 I l I
8 16 24 3? 40
n'

Comparison of strain hardening parameters, n based on
constant Toading rate and n' based on constant dis-
pTacement rate, for 5/8 in. diameter pTain rod in
frozen sand (vS = 64%).

 

 

128

PTain rod, d = 5/8 in. / 0
Frozen sand, vs = 64% /
/
O n /
1. n' /
a /
/
/
/
/
/

 

_ /
/
o /
. ,xa
/
/
- /
O //
/
_g,/0
”. ‘

1 1I1|11l1111I1111I11111
0 -5 -10 ~15 -20 -26
Temperature, T, 0C
2 5.11: Temperature effect on the n parameter as defined by

Equations 5.1 and 5.3.

 

 

129

[ PTain rod, d- 5/8 in.
Frozen sand (vS = 64%) ’9

Best Fit Line 0 ’
n = exp(1.224 + 0.0886 e)\:’//

 

o = -T /
/
./
/
/
I
/
/
/
_ / ii
_ o ,’
./
_ ,f/
O // . (Data from Figure 5.11)
- /
O [6
" /
o/
’l 14' I I I I I I I I I I 1 L; l I 441 I I I j
-5 -10 -15 —20 —25

Temperature, T, 0C

5.12: ReTationship between temperature and the n parameter for
a p1ain rod in frozen sand (v5 = 64%).

 

130

Pavg = 24.5 1bs./min.
6avg = 3.65 X 10-3
- Ice (Frozen in Tayers)
_ /
/
/
_ Best Fit Line /’

’q = 4.67 6 ~18 /

_ (9 BT/

 

 

_ O O?
(3
‘ O?
,4/8
—" ’- I I I I I I I I I I I I I I I 4
0 -5 -10 -15 -20

Temperature, T, 0C

re 5.13: Temperature effect on ice adhesion to a 5/8 in,
diameter pTain rod.

 

 

131

   
 

vS = 64%

d = 3/8 in, -26°C
h = 1/8 in.

6 = 4.5 x 10‘”

avg in./min.

  
 
  

-150C
O~-—-——---——O-- —
(5-46?

 

A.
_ ,’ -1n°c
' E]-""""B_-—B'—
:_ ,/ (5-57)
I]
' -606
ii [AT/E1 —20<:
I ' 0-0---_-----.0---_ _
y’fl/x’ <S-.9>
i- 37 ,0’
I
. _.0’
II Ov'o"O
.I I (Linear scaTe between numbered points)
’I I 'I I I I J I I I I I I J l J I I I I I
I—1 r—I N
0 E; C2 C? 0.10 0.20 0.30
DispTacement, 6, in.
e 5.14: TypicaT Toad-dispTacement curves at different temperatures

for a pTain rod with Tug in frozen sand (vS

 

132

 

 

[ El
/
_ /
_ aavg = 4.5 x 10‘“ in./min. /
d = 3/8 in. //
— h = 11/8 in. /
_ v5 = 64%
V)
—D.
E;
_ 0:3
2 E]
- m P vs T
'>5 u E]
_ Q'—
m“
'U
_ (U
3 E1
Piy Vi;,I:3L’C3
//0
E3 E1 I/IC)
/
" //’
/
_ /o
E] /
/ /
/ o
/’
_ /
/
— /
G
' // Temperature, T, 0C
L4 1 J I I I I I I 1 I I I I 4 I I I I I I I I I J
-5 -10 -15 —2O -25

yier” Toads,_in frozen sand (vs — 64% .

‘ure 5.15: Temperature effect on Tug bearing, uTtimate and ”initiaT

 

133

Rod Diameter, d, in.

3/32 6/32

12/32
I I |

20/32
|

 

 

 

 

 

_' PTain rods in frozen sand D = 6"
Favg = 24.2 Tbs./min. + d
6avg = 5.3 X 10‘L+ in./min. H — H
_I P
— T = 200C”
13 /’
_. ’1’
[I
1’
1... /
1”
" -100C
_. // ”' ”',,, ——’ C)
‘3/ E //
[3 o I—’
z’ o
- // ” (3 9
/’ (D 1”
_ E] z’
/ ,z'
/ /
-/ O o?
// I J I J J I I I I J
O 2 4 6 8 10
Ratio, d/H, %
re 5 16' Effect of rod size (diameter) on the bond strenath of frozen

sand (v = 64%) to piain stee] rods.
5

134

Pavg = 25.0 Tbs./min,
— éavg = 5.0 X 10“+ in./min.
- G d = 5/8 in.
0 d = 3/8 in. O
//
. o/ 0
1|
_ O
O
o I

 

[IIIIIIIIJIIIJIJIIIIIIIII

O -5 -10 ~15 -2O -25

Temperature,0C

3 5.17: Comparison of bond strength for two pTain rod diameters in
frozen sand (vS = 64%) at severaT temperatures,

 

135

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_ . _ . _ . _ _ _ _ _ o
o \11
o X
o1 \
\
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m.
m.
9
I .1?
a
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S
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oov

 

136

 

 

 

F T= 40°C
- Favg = 24.5 1bs./min.
... _L', - .
éavg 5 X 10 in./min.
(3 d 3/8 in.
- CD d 5/8 in.
Least-Squares Line
(3
' O 8 o o
C) c) i;
o . ‘
j I I J I I l I I l I I
0 1 2 3 4 5 6

e 5.19:

SampTe Height, H, in.

Effect of sampTe height on bond strenoth for pTain steeT
rods in frozen sand (vS

= 64%).

 

137

 

 

 

1.6 [
8 — A A
‘ z T = -100(:
O ._ ///// fiavg = 25 Tbs./min.
éavg = 5.0 X 10'“ in./min.
_ ZK//// H = 6 in.
2 _
.—-——'CP——‘-
‘ A /EJ"’ (3—57)
/ EIZ h = 2/16 in.
1 _ ZS—ik A//
‘V‘-—-—“’-— - - --‘V-—---
; '75 (5-122)
1 A; {v / h = 1116 in.
. ‘V E]
1 W/ -
I (Linear scaTe between numbered points)
_. I‘
I I \
I \
1. ~ .
’ \ (S—JO)
f EEID-C) {)_ PTain bar h = 0
I I l I I I I I l I I I I I I I I I
o_ . ..10 0.20 0.0

Disp1acement, 6, in,

9 5.20: Load-dispTacement curves for a 3/8 in. steeT rod with
four Tug sizes in frozen sand (v5 = 64%) at —100C.

 

 

 

138

  
   

 

 

0 E pm
- j \ T = ”100C
/ a 1?an = 2201bs./m1‘n.
5 /El \ (3an = 4.5 x10'31'n./m1'n.
. d = 3/8 in.
E] \ h - / '
/ E] — 1 64 1n.
9 ' \
. / \
D E’\
5 ' I]
\x
. 000‘s a\_
G I \. E] -191
) _ O Q \ ( )
I \ El H= 31h , ten 1ugs
/O \ \m
. , o\ \B
/O O\\ \E]
) I— 0/ @\\ \E
/ \
./ Q ‘0\
Elcf) (5-170) \\
) ’ H= 2 in. C2\‘\
/ Seven 1ugs \\
v \
; O . -
(S-30) (L1near scale between numbered p01nts)
H = 6 in.
No Iugs
’ J l L I I I I I I I I I I I I I I l I I I I I I I
H v—I N
0 8 O. O. 0.10 0.20 0.30 0.40
- C:- O

Disp1acement, 5, in.

re 5.21: Load—disp1acement curves for a standard deformed No. 3

bar in frozen sand (vS = 64%) at -100C.

 

8.4

7.2

4.8

rure 5.22:

 

139

= 25 1bs./min.
= 5.0 X 10‘“ in./min,
= 3/8 in.

Residuai ioads for
piain rods

Caicuiated for a standard
deformed No. 3 bar

40% /

 

T :
Pu = 0.65 (1 + 3,3'j;// ,’
I}?
Least-Squares /}3 ,’
Line [I
‘z
/'
E]? I’
/
/ /
/<— T = -10 F
P

 

C) I
1/54 9’
I / l l
,’ h=1/16 in. 2/16 3/16 m.
I
I I l I l J l J I
1 2 3

Ratio n = A1/Ar

Effect of 1ug size (height) on the uitimate 1ug

capacity in frozen sand (vS = 64%)

 

 

140

 

 

 

 

 

 

 

 

6 __
_ 900—1ug
450-1ug_
5 .—
T = —]00(‘.
. $an = 4.5 x1o'~ in./min.
= 64%
4 _
/A—A_
‘ /
3 _ f
/ ,
- / //
C)
A //
2 _ / o
l
/ I
/
B A/ ,0
' \ I,
l ‘0“ ,Q/ o 900-1.ch
1 :- 0-00’ A 450-Lug
I (Linear scaie between numbered points)
0 A . . . I A I I . I A I . . . . I
ES ES 23 »
0 C3 - - 0.10 0.20 0,30
Dispiacement, 6, in.
re 5.23: Typical 1oad——disp1acement0curves for piain rods with two

iug types; 450 -1ug and 90 —1ug, in frozen sand (vs= 64%),

U)
o

 

' 141

= 25 ibs./min. /
= 5.0 x10” in./min, /

6
vS = 64% /
- « h - 1/8 in. /

_' d = 3/8 in.

C a I 450—1ug I

O, [a goo—iug

(o = -T)
- ' +Pu= 1.05 (1+o_1ge)
for 200C< T< ~20C
El

8

- B a
— E/ /= o. 65 (1 + o 0966)
_ for 263 C< T< -2°C
)3 (Least— Squares Line)
‘ G.)

 

 

/
—/
[O
7 Temperature, T, 0C
I...I....I.J#I..JII.1.II.
-25

—5 -10 -15 -20

gure 5.24: Reiationship between temperature and 1ug bearing capacity

in frozen sand (vS — 64%).

 

 

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144

3.0 [—

_ T = -100C
: / '
2 5 Pavg 24 lbs. min
. — . = “L; . .
6avg 5.0 X 10 1n./min.
_ d = 3/8 in.
v5 = 64A
2.0 __ C) Ground-Finish steel, p = 90%
m C Cold-Rolled Steel, 0 = 625%
D.
E A: Shot-Blasted, p = 1A8 %
i
5
'U
D
_l

 

(Linear scale between numbered points)

 

 

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l \
l/ (\A pr
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ego-GIG —————— e—-——o——l ————
0 . l J I J I l l l J
o .1 1 2 3 4. 5 6 7 8 9

Displacement, a, X 10’2 in.

Ire 5.27: Load-displacement curves for three types of steel finish,
in frozen sand (vS = 64%) at ~100C.

 

 

145

  
    

 

 

4i
_ Plain rods, d = 3/8 in.
T = -100C
2 - I5: 24 lbs./min.
C)
Best-Fit Line
PU = 0.019 00.606
: C)
. <3
- ' ' l
- .0 '8
' 'E z 2
L? 0.: 0.0
'o E +—>
L
(D
l I I} I l I l l l l I l I J J
50 80 102 200 400 600 800 103 2000

Roughness Factor, p, %

ure 5.28: Effect of Surface roughness on the bond strength of frozen
sand (vS = 64%) to plain steel rods.

 

l0

146

/
100 lbs./min. // /‘\\\\\\\\

fl
'0.
m
<
to
II

    

\

(1) Ice layers (distilled water)
(2 Crushed ice (tap water)
(
l

)
3) Snow-ice (distilled water)

1 I I I I I I I j 1 I I

 

O -5 -10 ~15 -20 —25

Temperature, T, 0C

ire 5.29: Bond strength dependence on ice sample preparation method.

 

 

 

 

147

 

 

280 '_ _
_ Pavg = 25 lbs./min. C)
6 = 4.43 x 10‘“ in /min
av O O
240 — 9 O / /
200 _ ///
_ / 0
l/
160 — G //
/’
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40 .. O 0 Tap water
0 J I I I I I 4 I I I I I I J I I I I I I I l I I
O —5 -15 -20 -25 —30
Temperature, T, 0C
gure 5.30: Bond strength dependence on water impurities for a 3/9 in.

diameter plain rod in frozen sand (vS = 64%) at several

temperatures.

 

140

120

100

 

 

148

. . /
Plain Rod, d = 3/8 in. 9)
T = -100c lo
13 — ' /
avg - 24.5 lbs./m1n. /
. . . /
Gavg - 5.45 X 10 u in./min. 6?
Peak Strength, TU //t)
___ ————— — — — — — i O /
09

Range for ice

 

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0?
X/
Residual Strength, Tr¢ §,/’/

__________ X_———>(————x —— x
I l I I I l I I I I J I I I I I I I I I I I I I I
10 20 30 _40 50 60 70

Percent Sand by Volume

Tigure 5.31: Effect of sand fraction on bond strength for plain rods,

peak and residual at -1G°C.

 

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so _ S /
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3n _
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10 . J L . I I I j I J I I I I I
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freep Displacement Pate. 5C. in,/min,

Loading effect on the creep rate of a plain steel rod in
frozen sand (vS = 647) at —100C,

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. d = 3/8 in,

_ P = 4000 lbs.
_ h = 1/8 in, SC: 2_0 X 10-3
T = -100F in/min.
vS = 64% 3780 lbs,

1.F6 Y 1P'3 in.’min.

   

 

“ 3560 lbs.
1.0 x 10‘3 in./min.

[—I
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33

Time, t, X 102 min.

Figure 5.34 (Continued): (Cl Sam!)16 “0‘ “3’7.

 

 

 

 

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157

”.1

 

CS—Rl'.
Plain Rod
10 h d = 5/8 in.
,5 _ T = -60C '6
.'_ 6 X 10'7
? P(lbs.) (in,’minél
Cl
x 1 160 4.35
90 2 280 4.10
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0 2 4 6 8
Time, t, X 102 min,

re 5,37: Creep curves for a plain rod in frozen sand (v = 65%).
Step loading procedure, (a) Sample No, FS-34, 3t -60F,
(Continued)‘

LA)
0

gure

  

 

 

158

 

r‘q_3q 1
Plain 90d
d = 5’8 in,
'T=—1WC
V = 64%
s
“F'P = 1040 lbs.
ac = 4.95 x 10'6
. / .
in. min. 880 lbs.
8.33 X 10'7 in,/min.
l
4— — _
_ _ .(_ _. —

 

   

i 720 lbs.
2.8? X 10‘7 in,’min_

560 lbs.
2.44 X 10‘7 in/min.

400 lbs.
1.1 X 10‘7 in,’min.

I I I I I I I I I J I I J
7 fl 6 8 10 1?

Time. t. x 10? min,

5.87 (Continued): (b) Sample ”0. CS—33. —IFOC,

 

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168

 

 

 

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35 3i S 8
(- T N <" m
D
l I I I
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Cl
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1 ‘ /
/ o
/ /
/ -/
// D = Sample Diameter
‘ / // Plain rod with a lug
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- /
/ h = 1/8 in.
= _ o
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Pull-out load dependence on soil cover for a plain steel

iqure 5 45: _ .
3’8 in, rod with a 1/8 in, luo In frozen sand (VS _

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5C. x 10‘2 in,

Creep Diso1acement.

Figure 5,47:

48

J>
D

32

24

 

172

CI-]

        
 

Snow-ice. T = —]00F
P1ain rod with one 1ug
(d = R/P in, h = 1’8 in.)

P = 520 lbs.
5C: 2.88X]0_3‘

in./m1n.

40 80 170 160 200 240

Time, t. min,

Creep curve for a p1a1n 3’9 in stee] rod with a 1/8 in,
1ug in snow-ice (vS = 0) at -10°C.

 

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to .,I.Il . . l#CH

ocp um 92% :wNOLu
“gcwm; 03_ .ew m\H e epvz U01 .CI m\« megm :Im_a

sdm ‘d ‘pP’O‘l PSHddV

 

174

 

 

 

41
Plain stee1 3’8 in, rod
with a 1/8 in. 1uq
— T=~1nc
.33.—
‘i _ (Data from Figure 5.48)
0.
_U“ .—
I'D
.3 . 1|
8'
w 2 __
(5 'c _ -4 . .
6 — 10 in./min.
“U _ .
.9
Ti __
D.
<
l _.
__ Snow—ice
0 l l l l J I l I l 1 I I l l I
O 8 16 24 32 40 48 56 64

Percent Sand by Vo1ume. vS

Figure 5.49: Load capacity dependence on sand concentration durina
creep for piain stee1 3/8 in, rods with a 1/8 in. 1u0
in frozen sands at -10°C

 

 

 

 

175
- (D
/
I/
' Piain 3/8 in. rod with a /,’/
1/8 in. “.19. /
3 F Frozen sand (3
T = 100C ////
_ (D
8. ///
s: i 0
Q“ 2 _ P64% ... 1'75 P519:
'55 Vaiid for ac _<_ 4.0 X10-”in./min.
u L o
>U')
g: F- //// (Data from Figure 5,48)
1% _ o
3 /
/
1 t /
F /
/
~ /
I/
i
0 L l l 4 I l L l J_ n l l I
0 1 2 3

Load, P(vS = 51%), Kips

Figure 5.50: Comparison of iug ioads at different sand concentration
for frozen sand at ~100C.

 

 

 

CHAPTER VI
ANALYSIS AND DISCUSSION

1 Load Transfer Mechanisms:

Shearing resistance at the rod—frozen sand interface develops as a
sult of ice adhesion (or cohesion) and soil grain friction. When a
igle lug is added mechanical interaction caused by lug bearing against
a frozen soil adds to load transfer between the rod and frozen soil.

ice ice adhesion and particle friction appear to have a similar load

 

ansfer behavior, it was convenient to consider them in the same section.

3 bearing will be discussed in a later section.

[.1 Ice Adhesion and Sand Friction:

These two adfreeze bond components provide shear resistance to
l displacements relative to the frozen soil. Each components contribu—
In to the total resistance has been considered separately throughout
5 study. Ice pullout specimens (prepared in different ways) were
ted, using a 5/8 in. diameter plain cold-rolled steel rod. It was

erved that ice adhesion was sensitive to many variables including

 

perature, loading rate, ice structure, water impurities, stress and
erature history, and other factors. Control of all these variables
difficult due to the complexity of the load transfer phenomenon.
JItS of these tests, summarized in Figure 5.29, show the upper and
er limits of ice adhesion for various conditions. Only a brief

arpretation of these results will be given here, since adhesive

176

 

 

 

 

 

 

177

properties of ice deal with complex physico-chemical phenomena which are
outside the scope of this investigation.

Starting with the temperature effect, the results show a general
linear variation of ice adhesion with temperature (Figure 5.13) despite
the large scatter. This is in agreement with work reported by other
authors, as shown in Figures 2.1 and 2.2. Values of ice adhesion to
plain wooden piles obtained by Tsytovich and Sumgin (1959), appear to
vary almost linearly with temperature in the range of —7°C to -20°C
(Figure 6.1). Although no displacement or loading rate was mentioned,
these values are compared with Ice (1) in Figure 5.29. Ice (1) appears
to be most similar to that reported by Tsytovich and Sumgin. Further

comparisons with previous studies, unfortunately, are not as promising

 

due to the diversity in test conditions, ice types, and contact surfaces.
The increase in ice adhesion at colder temperatures has been attributed
to a reduction in thickness of the liquid-like layerat the ice inter-
face (Jellinek, 1957 b; Barnes et al., 1971).

Returning to Figure 5.29, the effect of water impurities on adhesive
strength may be noted by comparing Ice (2) and Ice (3). Tap water often
ontains 1 p.p.m. or more flourine, in addition to about 450 p.p.m. in
ardness as calcium carbonates. Jones and Glen (1969) observed a reduction
f about 35% in ice shear strength at -70°C due to increase in flourine
ontent from 0 to 0.05 p.p.m. and the reduction was about 60% when
lourine content increased to 0.5 p.p.m. The two authors attributed
he results to different types of physico-chemical defects and dislo-
ations in ice crystals caused by the hydrogen fluoride (HF) molecule

hich enters the ice lattice.

 

 

 

 

 

 

 

 

178

The effect of water salinity on strength of ice and frozen soils
have been investigated by many authors. Velli and Karpunina (1973) rec—
commend that the adfreeze strength of frozen saline ground to concrete
piles be taken with a factor of 0.75 of that of fresh water, based on
several in-situ tests. In fact the reduction in adfreeze strength due
to using tap water in Figure 5.30 was about 45 percent, which is far
more than 25 percent. More recently, OgataIet al.(1982) observed that
a 3 percent increase in water salinity of alluvial sand over distilled
water caused about 70 percent reduction in its unconfined compression
strength at ~200C. They also observed that the rate of strength de-

crease with increase in salinity was largest for a low salt concentra-

 

tion,in the range of 1 percent. This was attributed to the increase

in unfrozen water content due to the salt concentration. Creep defor-
mation in the sand was also observed to be largely influenced by the
increase in salt concentration. Similarly, Sego et al. (1982) showed
that an increase in salinity of frozen sand caused a non-linear decrease
in its proof strength, the exponent n from a power law, and a similar
decrease in Young's Modulus, at -7OC.

Temperature history is believed to have some effect on ice
adhesion. Its effect, when combined with the other factors, appears to
be far less significant on the sand-ice behavior. More consistent
data were obtained from pullout tests on frozen sand than those on ice.
All specimens were frozen at about -24°C prior to testing. Ponomarjov
(1982) experimentally showed that soil samples at first expand on
freezing (due to phase change of water). The expansion continued to

a maximum value at about ~80C (if the samples are fully saturated).

 

 

 

__— ' ""‘" "““‘°"""’ '-* ’~—' —‘

179

With a decrease of temperature below —8°C, thermal contraction reduces
the sample back to its original size at about -55°C. Ponomarjov
attributed this reverse deformation to the presence of gases and the
possibility of deformation of ice crystals when the gases change in
volume due to changes in temperature. Theoretical predictions made by
Ponomarjov (1982) compared well with his experimental data.

Considering these possible changes in the pullout specimens, a
5/8 in. diameter circular hole (in the ice at -24°C) would expand* by
6.5 X 10-5 in. if the temperature were increased to -2°C. A steel rod
of the same diameter, frozen in the ice, would contract by 1.5 X 10"5 in.

The net effect would be a small gap of 8 X 10"5 between the ice and rod.

 

If these assumptions are correct and the gap does exist due to tempera-
ture history, then the adfreeze bond would be weakened and the reinforcing
effect due to surface ”tension“ would be less. This weakening may also
+occur during the test due to differences in Poisson's ratio of steel
and ice. As the rod is pulled a slight reduction in its diameter will
differ from that in the ice. The effect of temperature history and that
due to differences in Poisson's ratio may not be as severe as they
would first appear. The net result may be a cancellation of effects.
The load—displacement curves for plain rods in ice and frozen
sand (Figure 5.2) are very similar. This behavior indicates a very
small displacement leading to rupture at ultimate load followed by a
small slip (Figure 5.1). Experimental work reported by Parameswaran

(1978 b) on model piles in frozen Ottawa sand showed a similar

 

*See Michel (1978) for the coefficient of thermal expansion
of ice.

 

 

 

180

load-displacement behavior. The Wedron sand appears to have properties
similar to the Ottawa sand except for possibly grain shape. As shown in
Figure 6.2, the ultimate adfreeze bond strength occurs at small dis-
placements, followed by immediate rupture. The displacements in

Figure 6.2 are normalized with respect to the rod (pile) diameter.

The 25 percent difference in strength was probably due to a difference

 

in soil types, rod surface roughness, and test conditions.

The creep behavior observed for plain rods in Figure 5.32 (a) and
(b) was similarly reported in the literature. Vyalov,et al. (1973)
stated that one of the model piles in ice (at -O.4°C) remained under a
pulling stress of 4.26 psi for 1200 hours without any noticeable

displacement, but then it suddenly pulled out.

 

The effect of rod diameter on frozen sand bond strength, Figure
5.18, does not agree with the findings reported by Frederking (1979)
; for wooden piles in ice (Figure 2.9). On the other hand, the results
showing effect of ice thickness on bond strength in Figure 2.9, agree
with a similar trend observed on frozen sand in Figure 5.19. The
theoretical work of Mohaghegh and Coon (1973) also supports an increase
in adhesive strength with increasing ice thickness for the case of
thick circular plates. ‘Although the effect of rod diameter does not
agree with Frederking's findings, it would agree reasonably well with
the general trend if one considers the combined effect of d and H in
terms of the ratio d/H.

Combining Figures 5.18 and 5.19 on one plot, for T = —10°C, gives
the relationship between Tu and d/H shown in Figure 6.3. A ratio of

d/H in the range of 6 percent to 10 percent appears to be an optimum

 

 

 

 

181

one in yielding the maximum bond strength. The dashed line represents
both lines in Figure 2.9 plotted versus log (d/H) in Figure 6.3.
Regardless of the numerical values, the general trend of decreasing
bond strength with increasing d/H, in the range shown, is the same

in both the present study and Frederking's study.

The temperature effect on bond strength of frozen sand appears to

 

agree reasonably well with field tests conducted by Crory (1963) on

8 in. diameter steel pipe piles in frozen sand slurries (Figure

6.4). Data from Figure 5.3, interpolated to a loading rate of 7 lbs./
min., is equivalent to the 10 kips/day indicated on the ultimate
strength curve of Figure 2.11. In fact, Crory's dashed line in

Figure 6.4 gives 50 percent higher values than those of the ultimate

 

strength curve of Figure 2.11. This is explained by Crory's (1963)
statement that a well-graded sand slurry, vibrated in place, has an
I adfreeze strength about 50 percent higher than its silt counterpart
at the same temperature. Differences between Crory's values and
those given in this study may be due to scale effect, test conditions,
and dry densities (sand concentrations).

Small differences in dry densities makes a significant difference
in the adfreeze strength, as shown in Figure 5.31. For sand concen-
trations greater than 50 percent (82.7 lbs. per cu. ft. dry density)
the frictional component of adfreeze strength becomes more significant.
The long-term adfreeze strength Tlt proposed by Vyalov (1959) in
Equation 2.11 for 6 in. diameter timber piles in frozen silty sandy
loam is also plotted in Figure 6.4 for comparison. At warm temperatures
(above -3OC) the agreement between the three studies appears to be at

its best.

 

 

 

 

 

 

 

.——.___’ . L...- .9" “‘2 _.' ' ‘ *')- ‘- 7 -— 7 7? Aq- 1. ,I_,

182

At this point, it is convenient to distinguish between the two
components, ice adhesion and sand friction, by plotting a typical
load-displacement curve for a test on ice and one for a frozen sand.
Figure 6.5 clearly shows the difference in ultimate and residual loads.
Crushed ice was chosen to simulate the polycrystalline ice believed to
exist in the frozen soil. Although the initial (at rupture) displace-
ment, or, for frozen sand in Figure 6.5 is slightly larger than that
for ice, the difference is only part of the experimental scatter.

Both materials, ice and frozen sand, showed essentially the same average
initial displacement, about 0.002 in., considering the difficulty with
precise displacementmeasurements on the recorder. Also, the initial
displacement for all plain rods in iceIand frozen sand specimens ap-
peared to be independent of temperature and loading rate. A slight
decrease in the displacement was observed at loading rates higher than
200 lbs./min., especially in ice specimens.

The initial displacement includes elastic bond displacement,
rod extension, deflection of the reaction beam shown in Figure 3.1,
and deformation of the frozen specimen. The latter was estimated to be
about 8.0 X 10‘6 in. for a maximum load of 3000 lbs. applied to plain
rods in frozen Wedron sand. Data from previous work (Bragg, 1980) was
used in the calculations. The corresponding rod extension was esti-
mated to be 3.26 X 10'“ in. using Hook's law and a steel elastic modulus
of 30 X 106 psi. The reaction beam deflection (as a simply supported
beam) was estimated to be 3.0 X 10"5 inch. The total deflection due
to the system stiffness (about 3.5 X 10'” in.) appears to be negligible

in comparison to the average 0.002 in. initial displacement. This

 

 

 

 

.. _.___-_,_. -4.

 

183

displacement increased with increasing rod surface roughness as shown
in Figure 5.27. The increase was probably due to interaction between
sand particles and steel surface roughness, and which requires larger
displacements for mobilization.

The greater adfreeze bond strength for frozen sand, compared to
ice, develops because the presence of rigid sand particles adjacent to
the steel requires that a large volume of the ice matrix be involved
in the failure process. Another comparison between ice adhesion and
sand friction can be made using data for frozen sand in Figure 5.17
and data for crushed Ice (2) in Figure 5.29. The best-fit lines for
data from both figures are reproduced in Figure 6.6 with residual
loads for ice and frozen sand (from Table 5.2) plotted against tempera-
ture. At ultimate loads the frictional component ruf contributes about
50 percent to 60 percent of the total adfreeze strength for the frozen
sand. The frictional component becomes more significant for residual
loads, Trs. The frictional contribution Trf to residual loads ranges
from 78 percent to 84 percent. Note that these percentages are some—
what high due to the presence of water impurities in the crushed
Ice (2), which was formed from ordinary tap water.

The choice of crushed ice, for comparison in Figure 6.6, was
made instead of the snow-ice because the latter had the highest data
scatter as shown in Figure 5.29. Also, the crushed ice strength
values in Figure 5.29 are intermediate between the two extreme types,
Ice (1) and Ice (3). The error, of under-estimating the adhesive
strength for polycrystalline ice (in sand pores), is believed to be

compensated by ignoring the contact area between sand grains and steel

 

 

 

 

184

rod. However, Lambe and Whitman (1969) reported that measurements of
the contact area between sand particles (about 0.06 mm in. diameter)
showed that typically it is approximately 0.03 of the total area. For
quartz, with a hardness of about 1100 kg/mmz, the stress on an asperity
(in contact with the steel rod) must exceed 1,500,000 psi to produce
plastic deformation. The frictional component is, therefore, only
slightly affected by neglecting the contact area of sand grains in the
calculation.

The adhesion and friction components also vary with the displacement
rate at a given temperature along with surface (or pile) type as shown
in Figure 2.13. In this figure, Parameswaran (1981) showed that the
frictional contribution varied from about 84 to 93 percent of the total
adfreeze strength of frozen sand to steel piles. This over-estimated
percentage (at ultimate loads) was made when Parameswaran under—estimated
the percentage of ice contribution. He simply multiplied the ultimate
values of ice adfreeze strength (solid lines in Figure 2.9) by the
porosity of frozen sand (0.2377) in order to obtain the dashed lines in
Figure 2.12. In doing so, he had attempted to simulate the polycrystalline
ice which exists in the sand pores. The ice represented in Figure 2.9
may not be polycrystalline ice since it was prepared by freezing the
distilled water which had been poured around the pile to the required

height,

 

.1.2 Lug Bearing:
The addition of a single lug, by welding, to a plain 3/8 in.
iameter rod permits bearing action of frozen sand on the lug to

ontribute to the total pull-out load. To illustrate the relative

 

 

 

185

contribution of each component, load-displacement curves for ice and
frozen sand, both with a plain rod, and a third curve for frozen sand
with a plain rod and single lug, are plotted in Figure 6.7. Two stages
considered for comparison include the "initial yield" (failure) and

the ultimate load (large displacement) conditions. The lug contribution
has now been added to the ice adhesion and sand friction discussed
earlier. The displacement curves for specimens with lugs show initial
behavior characteristics similar to those observed with plain rods.
After the initial rupture, mobilization of lug bearing prevents the

load from dropping to the residual value.

When the "initial yield" loads, P , for rods with lugs are compared

W
with the ultimate loads. Pu’ for plain rods at different temperatures
(Figure 6.8) the lug contribution ranged from about 25 percent at -26°C
to about 75 percent at -2°C. Considering the ultimate condition at
large displacements where the total load was transferred almost entirely
to the lug (Figure 6.7) the adhesion and friction become negligible.
Comparing the ultimate lug loads Pu with the residual loads Pr for frozen
sand from Figure 6.6, the lug contribution (P1 in Figure 6.9) ranges
from 90 percent at -26°C to 100 percent of the total load at 0°C. Similar
conclusions can be drawn on comparing lug contribution with the residual
load for different lug sizes (Figure 6.10). In Figure 6.10 lug contribu-
tion increased from an average of 85 percent of the total load for
h = 1/64 in. to 95 percent for h = 3/16 in. at temperatures of -10°C
and -20°c.

It is of interest to compare the pull-out loads for plain rods

and lugs with the uniaxial compressive strength “U for frozen sand. To

 

 

 

 

 

 

—7— -- -1 -.-“--- _ ... _- _.. _ .

186

do so, strength data from Bragg (1980) are compared with the ultimate

lug pull-out capacity Pu’ the "initial yield" loads Pi , and the ultimate

y
bond strength Tu in Table 6.1. Values for the uniaxial compressive

 

strength correspond to an average strain rate é of 3 X 10'6 per

avg.
sec. in Bragg‘s (1980) Figure 5.9. This strain rate corresponds to a
nominal machine displacement rate In of 5.0 X 10'” in./min., calculated

by multiplying é times £0 (= 2.82 in.), where to was the length of

avg.
uniaxial test samples in Bragg's (1980) study.

Note that comparisons based on equal displacement rates in both
types of test (bond and uniaxial comparison tests) do not mean equal

strain rates. Differences in sample size, test conditions such as

 

applied confining pressures, and other factors onld not yield equal
strains in both tests. Equal displacement rates were chosen only to
make the comparisons under similar conditions. There is some similarity
in both tests because frozen soil experiences compression under the lug

and also in the uniaxial compression test. Comparisons are made for

 

lug movement at the same rate as the loading ram in the compression test
(neglecting any differences due to machine stiffness). The data in
Table 6.1 are plotted in figure 6.11. Although the relationships are
not linear, they are approximated by straight lines which pass through
the origin, except for the ru-vs-c relationship. By analogy to Equation
2.17 the ratio of bond strength, including lug bearing, to the frozen
sand cohesion (C.= oU/Z) ranges from 0.190 for plain rods to

0.390 for lugs. Values of‘k suggested by Weaver and Morgenstern (1981 b)

 

in Equation 2.17 ranged from 0.6 for plain piles to 1.0 for corrugated
iles, with c-values evaluated on a long-term basis. W0 standard

efinition for "corrugated” piles was mentioned by Weaver and Morgenstern

   

 

 

 

 

 

 

 

187

(1981 b).

For comparison purposes the cohesion (c = ou/Z) was defined on the
basis of a horizontal Mohr-Coulomb envelope (a = 0). Should this
envelope be inclined to the horizontal at an angle ¢, values of c
would be slightly reduced, thus increasing the values of K for this
study. A comparison also can be made between the lug bearing capacity
q1 (lug contribution P], in Figures 6.9 and 6.10, divided by the lug
area A1) and the uniaxial compressive strength cu. Such a comparison,
shown in Figure 6.12, was based on equal displacement rates for both
tests. As shown in Figure 6.12 the relationship is not linear but can be
approximated by a straight line. This relatidnship indicates that the
ultimate lug bearing capacity is about seven times the uniaxial com-

pressive strength of frozen sand.

6.1.2.1 Lug Behavior in Frozen Sands:

Lug behavior in frozen sand may be considered analogous to a deep

 
  
   
  
   
  
  
  
  

plate (or footing) in frozen ground. Experimental work reported by
Ladanyi and Paquin (1978) using a standard penetrometer (1.4 in. dia-
meter) in a frozen quartz sand (e0 = 1.034, Vs: 49.15 percent, Cu = 2.27)
provides comparative data for use with the lugs. The penetrometer placed
at a depth of 18 in. can be classified as a deep footing. Data sum-
marized in Figure 2.15 have been transformed into punch stress versus
creep rate on a log-log scale in Figure 6.13 and compared with data from

Figure 5.40 for a 1/8 in. height lug (equivalent punch diameter equals

 

d + 2h = 0.625 in.). Both tests were conducted at -6°C. Both
relationships are almost linear, especially for the penetrometer. The

two lines are almost parallel indicating that a similar mechanism

 

 

 

 

=.-—-fi-' “4“?” ' _‘cc '- A

188

appears to control the creep rates in both cases.

‘The main reason for a difference in creep rates (or strength) is

 

believed due to the difference in void.rat105 (or sand fractions) 0f the
two sands. The importance of sand fraction on lug bearing loads has
been shown in Figure 5.49. A reduction in the sand fraction from
64 percent to 49 percent at -100C resulted in a load reduction close to
49 percent for creep rates of 10‘“ in./min. or 10"5 in./min. A larger
reduction would be anticipated at warmer temperatures such as ~60C.
Additional factors responsible for the difference in applied stress
shown in Figure 6.13 involve the punch diameter, 0.625 in. for the lug
versus 1.4 in. for the penetrometer. The significance of this
difference in diameters is shown in Figure 5.41 and from Vyalov et al.
(1973). When the P-values (in pounds) from Figure 5.41 were normalized
with respect to the lug area A] to give q1 = P/A1 ( as punch pressure
in psi) the solid-lines in Figure 6.14 were obtained to help explain
the difference between curves in Figure 6.13. These lines clearly
show that creep rate increases with an increase in lug size for the
same stress level. This result is also supported by the experiments
conducted by Vyalov,et al. (1973) on punches of three different sizes
loaded in ice at -2.3°C. For comparison the results of those experiments
are included in Figure 6.14. Their data indicates that as the punch
diameter was increased from 1.26 in. to 12.6 in. (10 times larger) the
punch pressure decreased about 50 percent to 70 percent.

For the present study, increase in lug height from 1/16 in. (0.50
'n. equivalent punch diameter) to 3/16 in. (0.75 in. punch diameter)

aused a reduction in pressure of about 17 percent to 20 percent,

 

 

 

 

189

depending on the stress level. The difference in punch sizes indicated
in Figure 6.13 may, therefore, have contributed about 25 percent to the
total reduction in strength shown. Note that the comparisons made in
Figure 6.13 are valid for the range of creep rates indicated, which do
not necessarily coincide with the same range of strain rates at which
the uniaxial compression tests on both sands were conducted.

Ladanyi and Paquin (1978) indicated that when trying to relate the
behavior of a deep footing (punch) with that of a representative soil
sample under triaxial test conditions, the problem most difficult to
solve is usually that of selecting the strain rate in the test that would
be representative of the average strain rate of the soil during pene-
tration. Since the strain rate around a footing decreases continuously
with distance, various interpretations are possible. The one considered
by Ladanyi and Paquin (1978), relates the penetration rate of a circular
footing with the time to failure of a soil eIement located below the
base in the line of penetration, by a semi—empirical expression, the
same as proposed by Ladanyi (1976). Based on that expression the two
authors estimated that a range of penetration rates, from about 2 X 10-6
to 1.1 X 10“3 in./min., for their punch size of 1.4 in. diameter
(Figure 6.13), would be equivalent to a strain rate of about 1.8 X 10"8
to 8.3 X 10’6 per second. In fact, both ranges are lower than the lowest
applicable creep or strain rates in their experiments.

The difference in depths of punch embeddment in Figure 6.13, do not
iPpear to be a significant factor. As noted in Section 2.5, when a punch
;ettles more than about 5 percent of its diameter, the soil enters into

, viscoplastic region and the effect of burial depth becomes negligible.

 

 

 

 

 

 

 

190

Under such conditions, no effect of burial depth was observed for depth

to diameter ratiolJVB between 3.3 and 5.0 in ice (Sego, 1980) and between

 

5 and 15 in frozen sand of Ladanyi and Paquin (1978). The decrease in
creep rate, noted in Figure 2.16, with the increase of burial depth is,
probably, due to the fact that the settlement did not exceed 1 percent of
the footing diameter.

Since the lug penetration, or displacement, 6, exceeded the equiv-
alent footing diameter (more than 200 percent) the burial depth did not
appear to have any effect on the lug creep rate. This may be observed

from Figure 5.43, where the lug position 2 was changed from 1.5 in.

 

(D'/B= 4.5/0.635 = 7.2) to 4.5 in. (D'/B= 2.4). The equivalent depth

of embeddment D' was measured from the sample top surface (at z = 6 in.).
It is believed that some reinforcing effect by the steel rod to which

the lug is welded, may be involved in eliminating the effect of burial
depth on the lug creep rate. The steel rod would prevent any side tilting
of the lug, which might occur at shallow depths. Lug position was also
studied relative to possible interaction between the lug and the base

plate. Data summarized in Figure 5.43 also imply that with the lug

position as close as 1.25 in. (1.50 less 0.25 in: of total creep displace—

 

ment) to the base plate, the creep rate was not affected. Due to
equipment limitations, the distance 2 was not made smaller than 1.25 in.
This minimum distance, which showed no interference between the lug and

reaction base plate, corresponds to aIJUHBratio CI 7.6 (= 4.75/0.525)

and to a z/B ratio of 2 (= 1.25/0.625).

 

 

 

 

191

6.1.2.2 Lugs versus Standard Deformed Bars:

The measured peak loads for a standard deformed bar (d = 3/8 in.,

h = 1/64 in., lug spacing = 0.3125 in.) shown in Figure 5.21 were trans-
formed into single lug loads as indicated in Figure 5.22. For example,
the peak load for the 3 in. sample height (about 4000 lbs. in Figure 5.21)
was divided by the number of lugs (10) after subtracting the residual

load of 60 lbs. (= 120 X 3 in./6 in.) for the corresponding plain rod,
i.e. P = (4000 - 60)/10 = 396 pounds. The latter was then added to the
120 lbs. residual load, giving a total of 516 lbs. load for a 6 in.
sample height with a single lug. Similar calculations were applied to
the 2 in. sample height in Figure 5.21 in order to obtain the two data
points indicated in Figure 5.22. Although these points agree well

with experimental data points for other lug heights, there is a dif-
ference in the load-displacement behavior for the standard deformed bar
(compare Figure 5.20 with 5.21). The decrease in load after about 0.04
in. displacement (Figure 5.21) appears to be due to : (1) the lug inter-
action effect, and (2) the small ratio of lug height (0.015 in.) to the
maximum particle size (0.023).

Little or no interaction between the reaction base plate and the
nearest lug was anticipated because of adequate clearance between the
deformed rod and removable washer in the support plate. Besides, if
the lug spacing to lug height ratio were to be considered it would be
large enough (0.3125/0.125 = 21) to prevent any interaction between
the lugs. The interaction should, therefore, be explained in terms of
the ratio of lug spacing to equivalent punch width, which is in this

ase less than one (0.3125/0.405). Such lug interference is illustrated

'n Figure 6.15. As shown in this figure, when the soil stress under the

 

 

‘vngr-.\- sitim' '4 I» .1 - .

192 ‘

top lug (No. 1) is go, the stress under the second lug would be 1.27qo

using Boussinesq's stress distribution. Soil stresses under successive

 

 

 

 

 

 

 

 

 

 

 

 

 

lugs can be evaluated in a similar manner as shown in Figure 6.15. It
has been estimated that a maximum stress of 1.37 qO occurrs under the
sixth lug, where its value remains constant for subsequent lugs.

If the above assumption is reasonable, failure would initiate in
the soil under the bottom lug and move upwards, due to lug interaction.
As the lug penetrates into the soil, it leaves a gap equal to the dis-
placement, thus weakening the support for the next lug. When the
ultimate load is reached, the gap would have reached its limiting size,

the soil support would decrease and hence the load for any further

 

displacement (beyond 0.04 in.) as shown in Figure 5.21. In addition to
lug interaction effect, when the lug height is smaller than the particle
size the frictional and dilatational components may not be fully mo-
bilized in front of the lug. A creep test on the 3/8 in. diameter

1 standard deformed rod (Figure 5.42) indicated a reasonable agreement

I with tests on other lug heights as shown in Figure 5.41. In general,

- standard deformed rods tolerate smaller displacements at failure than

i single lugs (compare 0.04 in. in Figure 5.21 with 0.30 in. in Figure

5.20).

6.1.2.3"L0ng+Term Lug Bearing Capacity:

The data shown in Figure 2.4 suggest that frozen soil strength
decreases exponentially with time to failure. The relationship described
by Equation 2.1 indicates that the long—term strength (when tf goes to
infinity) is zero. Vyalov (1963) pointed out that after some long

period of time the additional strength reduction is so insignificant

 

 

 

 

Jh—._" v _‘ A...

193

that it can be neglected in engineering calculations. The same con-

clusion applies Us the lug behavior in frozen sand. A plot of load versus
time to failure was prepared, for all creep tests on single lugs at

-10°C, and is shown in Figure 6.16. In this figure, the time to

failure tf was defined as the time required to achieve a certain displacement
since complete tertiary creep failure was observed only on the last

load increment.

Excessive displacements (such as 0.25 in. or 1.0 in. as shown in
Figure 6.16) may be considered as the basis for defining failure. Values
of tf corresponding to a certain load increment were obtained by extra-
polation. The displacements 0.25 in. (or 1.0 in.) were divided by the
creep rate corresponding to that load increment, after subtracting the
”pseudo" instantaneous elastic displacement 61 andvalues of t.1C are listed in
Table 6.2 for different load increments on a 3/8 in. rod with a 1/8 in.

lug height in frozen sand at -100C. The data summarized in Figure 6.16
indicate that the long-term load capacity was about 32 percent of the
instantaneous capacity for 0.25 in. allowable displacement, and about
38 percent for a 1.0 in. maximum allowable displacement. Thus, assuming
about 4500 lbs instantaneous strength for both cases, any load less than
about 1800 lbs. in Figure 5.40 would not reach a secondary creep rate
stage within the experimental times used.

Creep rates for loads below the long-term strength would continue
in the primary stage and attenuate until the rate approached zero at
infinite time. Creep rates for initial loads below 1800 lbs. in
Figure 5.40 would be over-estimated if they were considered as secondary
creep rates, even though they were measured at the end of that step

loading. Also, at the small initial loads the adhesion and friction

 

 

 

 

 

 

 

—>——_ - ‘3- 1"-“ 1‘. a x. .- T‘V' . _l ,- ‘ 1 A‘v :-

 

194

contribution along the rod was significant (Figure 6.8), thus giving
some errors in creep measurements, especially at colder temperatures
(below -IOOC). However, at large displacements the lug contribution
becomes far more significant (Figure 6.7) and neglecting the residual

loads would not cause significant errors.

6.1.2.4 'Sample Size and Sand Concentration:

The effect of sample diameter on the creep behavior of lugs in frozen
sand has been considered in Figures 5.44 and 5.45. Note that sample
diameters greater than about 4 in. would not affect the pull-out capacity
for the bar and lug sizes used in this study. The reduction in strength
at diameters smaller than 4 in. may be attributed to (1) the contrib’
ution of sample deformation to the total creep, and (2) the increased
effect of stress concentration under the lug area.

For sample diameters greater than 2 in., it was estimated that
sample deformations under a 4000 lbs. load would be negligible in
comparison to soil creep under the lug. Therefore, stress concentration
under the lug area is believed to be the main factor in reducing the load
for sample diameters less than 4 inches. For these smaller diameters,
visual inspections of samples after failure indicated that rupture
had occurred on planes inclined at about 600 from the horizontal, thus
suggesting that the soil samples may have failed due to stress concen-
tration. The latter also produces high tensile stresses in the
tangential e—direction with possible failure in tension.

The effect of sand concentration on lug bearing, summarized in
Figures 5.48 and 5.49, may be compared with similar effects on the

uniaxial compressive strength summarized in Figure 2.5 for Ottawa sand.

 

 

 

 

 

 

 

 

 

 

195

The average grain size of Ottawa sand used by Groughnour (1967) was
0.0306 in., slightly larger than that of the Wedron sand used here
(about 0.0138 in.). Besides, the Ottawa sand grains were well—rounded
compared to the sub-angular shaped Wedron sand grains. Other properties
are believed to be almost the same for both sands. To make possible
a closer comparison, one curve has been chosen from Figure 2.5; that
corresponding to -12.03°C and a strain rate of 1.33 X 10-” per min.
‘This rate appears to correspond to a nominal (machine) displacement
rate of about 3 X 10-“ in./min., based on a sample length of 2.26 in.
used by' Goughnour (1967). Load values were interpolated from Figure
5.48, for 6c of 3 X 10'” in./min. and different sand concentrations.
These loads were increased by about 13.5 percent to allow for the
decrease in temperature from -10°C to -12.03°C. Using Figure 6.9 the
ultimate loads in Figure 5.24 were reduced by 240 lbs. to account for
the residual shear loads along the rod. A residual load of 90 lbs. at
—12°C for the ice sample (vS = 0) was considered in the calculations.
The adjusted loads were divided by the lug area to obtain the lug
bearing pressure q]. Values of q] are plotted versus sand concentration
vS in Figure 6.17. The unconfined compressive strengths cu, show bi-
linear relationships with Vs' Although the lug bearing may not vary
linearly with sand concentration (Figure 5.49) it was approximated by
two straight lines whose intersection point gives a critical sand con-
centration close to 43 percent sand by volume. This point, which is quite
close to the 42 percent observed by' Goughnour (1967), appears to

indicate a change in the failure mechanism for frozen sands. This

sand concentration corresponds to a porosity of 0.580 which is close to

 

 

 

 

 

 

 

 

196

0.476 for the loosest possible packing of uniform spheres (in cubical
array) listed by Harr (1962). Therefore, at the critical volume
concentration, sand grains just begin to make contact as the soil
behavior becomes dependent on the closeness of adjacent sand particles.

Goughnour (1967) reported that samples of frozen Ottawa sand with
more than 42 percent sand by volume showed a more rapid initial volume
decrease (under compression tests) followed by a volume increase which
progressed at an increasing rate, and was similar to the volume
increase of dense unfrozen sand (Bishop and Henkel, 1962). Goughnour
suggests that three mechanisms may serve to strengthen samples made with
sand and ice versus those of pure ice. The first is associated with
sand volume concentration in the sample and is probably caused by
virtually all ice plastic deformation. This mechanism may explain the
behavior represented by the first line segments in Figure 6.17.

The second mechanism is associated with grain to grain contact
and is the result of friction at these contacts. The third mechanism
is associated with volume increase of dense samples and has a counter-
part in unfrozen sand. The sample volume increase is impeded by the
ice matrix. Thus the ice matrix effectively exerts a confining pressure
on the sand particles. The effect of this ice exerted pressure is to
increase the effective stress on the sand skeleton, thus increasing
its resistance to deformation similar to unfrozen soil. This "equivalent“
confining pressure continues to increase until dilatancy levels off at

a critical void ratio or the limiting ice cohesion is overcome.

 

 

 

 

 

 

197

6.1.2.5 Lug Bearing in Ice:

The lug behavior observed under the initial load increment in
Figures 5.46 and 5.47, for samples with a sand fraction below 50 percent,
appears to support Goughnour's (1967) first mechanism. He suggested
that virtually all of the plastic deformation is accomodated by the ice
matrix. This appears possible only when the sand friction and dilatancy
do not significantly contribute to the total resistance. As shown in
Figure 6.17 these components begin to develop at a sand fraction close
to 42 percent. There may be some reinforcing effect by the sand grains
which would impede the deformation mechanism of pure ice. Below the
critical sand concentration (about 42 percent), the lug behavior ap-
pears to be mainly influenced by the ice deformational characteristics.
In this case, seven deformation mechanisms have been identified by
Gold (1963) in observing surface features of ice during deformation.
These include slip bands, grain boundary migration, kink bends, dis-
tortion of grain boundaries, crack formation, cavities, and
recrystallization.

Since ice creeps under very small stresses (Vyalov, 1963) the
behavior observed in Figure 5.47 cannot be attributed to these seven
mechanisms. The initial delay in the creep of snow-ice under pressure
appears to be caused by bond (shear resistance) along the rod. This
bond was mobilized at much smaller displacements as compared to lug
bearing. .The delay time observed in Figures 5.46 and 5.47 for the
520 lbs. initial load increment can be considered as the time to
failure of the bond resistance under that load. Since the creep

displacement at bond failure would have been too small to measure

 

 

 

 

 

 

 

198

with the available equipment, no precise correlation can be made between
the sand fraction and time to failure (or the initial delay time).

From a careful consideration of the snow-ice behavior shown in
Figure 5.48 one may conclude that a residual bond resistance of about
200 lbs. along the bar may have caused higher loads in some of the snow-
ice samples as compared to the sand-ice samples. Otherwise, there is
no reason why the data points would not be consistent with the dashed
line extrapolated from the data point of test number CI-I (Figure 5.47).
This data point appears to be a coincidental point in which the residual
bond resistance may have been the least.

For snow-ice and sand—ice (sand fraction below 50 percent) samples
prepared by pre-Cooling the mold and the rod below 0°C a thin layer of
ice was observed around the rod soon after adding pre-cooled distilled
water. For those samples the bond resistance along the rod may be
due mainly to ice adhesion. That is probably why the initial delay
time was not much affected by the sand fraction. Since the snow-ice
was expected to have the lowest lug bearing compared to the sand-ice,
the bond contribution in the latter may also be less significant than
in the snow-ice. No correction for the residual bond (due to ice
adhesion) was made in Figure 5.48, since it should be negligible in
sand~ice and no accurate measurements for its value in snow-ice was

possible due to equipment limitations.

6.2 Creep versus Constant Displacement Rate Tests:
Comparison of strength data from constant strain rate tests and data
from creep tests was made by several authors (Goughnour, 1967; Hawkes

and Mellor, 1972; Bragg, 1980). In this study, the adfreeze bond

 

 

 

 

___.— — 9‘2-"&.g —n._r ~I.;-~~l~ "' .3 —

199

strength for plain rods in frozen sand from constant displacement rate
tests has been compared with the strength from creep tests as shown in
Figure 6.18. There appears to be a reasonable agreement between the
results from both tests. Assuming that the observed displacement rates
(6n and 6C) are representative, there appears to be a transition from
creep data (slow rates) to constant displacement rate (high rates).
The change in slopes of the lines in Figure 6.18, at about 10'“ in./min.,
indicates a change in the strength mechanism upon transition from slow
creep rates to the fast displacement rate for both temperatures.

For lugs, the displacement rates for both test types were approx-
imately in the same range (Figures 5.6 and 5.36). The lug capacity

from creep tests is plotted against lug capacity from constant

 

displacement rate tests, at different temperatures, lug sizes, and creep
rates, as shown in Figure 6.19. The data in Figure 6.19 indicate that
creep tests produce lug capacities which are slightly higher than

those obtained from constant displacement rate tests. It is believed
that this higher strength was a result of some densification and
stiffening of the frozen sand in front of the lug during creep. The
difference is, however, small and might be neglected for practical

purposes.

6.3 Correlation of Experimental Results:

The empirical equations deduced in Chapter V are combined in this
section in order to give a composite equation which relates the lug
capacity to the test variables. A similar equation has also been

derived for the adfreeze bond strength of plain rods.

 

 

 

 

 

200

6.3.1 Plain Rods:

The ultimate bond strength Tu of frozen sand (64 percent sand by
volume) with a plain steel rod (5/8 in. diameter) was observed to depend
on the nominal displacement rate én (machine speed) as shown in Figure
5.4 (and Equation 5.3). Bond dependence on the loading rate P was
illustrated in Figure 5.3 and expressed by Equation 5.1, where TC and
m (= l/n) are the temperature dependent parameters shown in Figures 5.11
and 5.12, respectively. Since measurements of loading rates were more
reliable than the displacement rates, it seemed reasonable to use
loading rates for further correlations. Substituting 10.5 a for Tc in
Equation 5.1 (Figure 5.9) and with PC = 1 lb./min. gives the following
relationship:

10.5 6 Im (6.1)

Tu (psi)
exp - (1.224 + 0.0886 6) as shown in Figure 5.12, and

where m = 1/n
e = -T. Since Equation 6.1 was derived for a 5/8 in. diameter rod (d),
6 in. sample height (H), rod roughness p of 625 percent, and sand fraction
vS of 64 percent, these parameters may also be included in this equation.
To do so, consider the combined effect of the ratio d/H using the
data in Figure 5.18. The two regression lines are combined by plotting
Tu/G versus d/H on a log-log scale. This eliminates the temperature
effect and gives the empirical relationship shown in Figure 6.20. This
procedure implies that at a given ratio d/H the bond strength Tu is a
linear function of temperature a. This assumption appears to be true
for all loading rates P below 6.0 lbs./min. as shown in Figure 5.9.

Assuming that the optimum ratio of d/H is 10 percent (Figure 6.3) and

neglecting the effect for d/H z 10 percent, the value of Tu in Equation

 

 

 

201

5.1 can be multiplied by (d/10H)S, where s = 0.692 if d/H g 10 percent
and s = 0.0 if d/H z 10 percent. The roughness factor Q can be
included by multiplying Equation 6.1 by (p/625)0'35“, using the
relationship in Figure 5.28.

The sand fraction vS was included in Equation 6.1 by replotting the
data in Figure 5.31 on a log-log scale. An experimental power law of
the form shown in Figure 6.21 can be deduced, applicable for v a 45

5
percent. Combining all test variables into one equation leads to

TI (psi) = 10.5(P°)"‘ (II/625W3‘5‘*<vs/64)2I246 (d/IO H)5 e (6.2)

where m = exp - (1.224 + 0.0886 6); e = -T, 0C; P = loading rate, lbs./
min.; p = roughness factor, percent; 5 = 0.692 for d/H s 10 percent
and 0.0 for d/H 2 10 percent; vS = sand fraction, percent.

Equation 6.2 would be limited to test conditions used in this
study, i.e.

2°C g_e s 26°C

1 lb./min. g e g 2000 lbs./min.

2 percent 3 d/H g 40 percent

45 percent 5 v s 64 percent

s
p s 1500 percent
Equation 6.2 appears to compare reasonably well with experimental
data as shown in Figure 6.22. A comparison of Equation 6.2, derived
from constant displacement rate test data, with creep displacements
of plain rods could not be made because of equipment limitations. More

experimental work is needed on creep of plain rods using equipment

capable of displacement measurements to at least t 10'8 inches.

 

 

 

 

 

202

6.3.2 Deformed Rods:

Test variables affecting lug behavior can also be combined into
one Equation. The relation between load and creep rate shown in Figure

5.36 gives the expression:
' = 'C ' 0.203
q (p51) qce (6 (5c) (6 3)
where qCe is the temperature dependent lug pressure, corresponding to an
arbitrarily chosen creep rate 6C of 0.0005 in./min., and can be expressed
as:

e 0.6
qce qc0 (1 + 6“90) (6.4)

where qCO = 3718 psi at 0°C, and 90 is a reference temperature equal to
1°C. Equation 6.4 was derived by plotting values of PCB from Figure
5.40 and values of Pu from Figure 5.24 against the temperature function
(1 + 6/00) on a log-log scale as shown in Figure 6.23. For small
temperature intervals, the exponent in Equation 6.4 can be made equal
to unity, and the power law reduces to the linear expression given in
Figure 5.24. The power law (Equation 6.4) will be used instead since
it is applicable over larger temperature ranges (Vyalov, 1963; Ladanyi,

1972). Substituting PCe from Equation 6.4 and 0.0005 in./min. for BC

into Equation 6.3 gives:

q (psi) = 3718 (1 + e)0-6 (SC/0.0005)°~203 (6.5)
Next consider the effect of lug size. Note that the load Pu (pounds)
in Figures 5.22 and 5.41 increases with an increase in the ratio Al/Ar
(or in the lug size). Also note that, when the same load was normalized
with respect to the lug area A1 for each lug size (lug pressure q1 in
PSI), the effect was reversed as shown in Figure 6.14. The applied

lug pressure in Figure 6.14 appears to increase with a reduction in lug

 

 

 

 

 

 

203

height. Since the rod area Ar is constant (d = 3/8 in. for all lugs)
the ratio Al/Ar (data from Figures 5.22 and 5.41) may be plotted versus
lug pressure q1 on a log-log scale, as shown in Figure 6.24.

Using the least-squares line on Figure 6.24 gives:

q (psi) = 15600 (A1/1.78 Ar)-°-191 (6.6)

Note that Equations 6.5 and 6.6 were derived for frozen sand with

64 percent sand by volume. Frozen sands with less than this percentage
of sand give lower lug pressures as shown in Figure 5.49. Equations

6.5 and 6.6 can be modified to include the sand fraction. Recall that
for a sand fraction below about 42 percent (Figure 6.17), the soil would
be in a very loose condition and would represent an ice—rich permafrost.
The emphasis here has been placed on sand fractions above 42 percent
where a small change in the sand content would result in a significant
change in the frozen strength.

It was shown in Figure 5.50 that the ratio of P51% to P64% is almost
constant, about 0.571. This has two implications: (1) the reduction
in lug bearing (about 43 percent) was almost independent of creep rate
and (2) the rate of change of lug bearing with change in sand concen-
tration, i.e. slope of the lines for vs>’42 percent in Figure 5.49, is
dependent on the creep rate. The latter condition complicates the
solution. It means that for every value of creep rate a certain
empirical equation must be used to account for the sand concentration
effect. Although several procedures can be made to avoid this compli-
cation, plotting the load P versus sand fraction vS on a log-log scale

appears to be the simplest procedure.

Choosing an arbitrary creep rate, say 0.0005 in./min., the load

 

 

 

204

P can be interpolated from Figure 5.48, and plotted against the cor-

responding vS values on a log-log scale as shown in Figure 6.25. The

linear relation suggests a power law of the form:

q (psi) = 15790 (vs/54)2.6su

(valid for vS a 45 percent sand by volume) (6.7)

The overall effect of different factors may be considered by combining
Equations 6.5, 6.6, and 6.7 to give the following composite empirical

equation:

q (PSI) = 0.2288 (1 + a)0-6 (éc)o.203 (A1/Ar)-0.191 V52.654

(6.8)
The range of experimental data limits Equation 6.8 to the following
conditions:

20C 3 e s 20°C

10-5 in./min. : éc g 10-2 in./min.
0 g Al/Ar g 3

v5345 percent sand by volume

The remaining question involves a relationship between the applied lug
bearing pressure q and time to failure tf. Most creep curves for sand-
ice specimens did not reach tertiary creep within a relatively short

time period, say less than 24 hours, except at large loads (more than

4000 lbs.). In constant displacement rate tests, the time to failure

varied from a few hours up to about 36 hours. Besides, the failure

was clearly defined by two peaks (see Figure 5.20). In creep tests,

it appears reasonable to use a certain limiting creep displacement 6f

as a basis for creep failure criterion in the frozen sand. This has

the advantage of limiting the total displacement to acceptable values

 

 

 

 

205

in the design.
From the displacement behavior for a constant stress creep test

(first load increment in Figure 5.34) it follows that:

5f = 5i + tf If (6-9)
where 61 is the pseudoeinstantaneous displacement. A small part of this
displacement is elastic, which may be neglected, and the major part is
plastic (irreversible). Equation 2.6 shows that the plastic strain
(second term) is related to the applied stress by a power law. By
analogy, a similar power law may be deduced by plotting the measured

displacements 61 versus the applied lug load P (Figure 6.26). Dividing

the empirical equation shown in this figure by the lug area gives:

8i = 8.3 X 10‘8 ql"+29 (6.10)

By rearrangement of variables in Equation 6.8 and some calculations,

the creep displacement rate 0C is given by:

'C = —2.95 0.941 —13.07 4.93
6 8980 (1 + e) (Al/Ar) vS q (6.11)

Substituting for a, and 6C from Equations 6.10 and 6.11, respectively,
into Equation 6.9, and rearrangement of terms gives the time to failure
tf:
5f - (8.3 X 10'8) ql'h‘29
tf =
8980 (1 + e)-2.95 (A1/Ar)0.941 VS-13-07 qho93 (6.12)

 

To compare Equation 6.12 with some of the experimental data, the
"failure" or limiting displacement 5f should be defined. Values of
0.25 in. and 1.0 in. were earlier selected as convenient limiting dis—

placements, a The time to failure tf was observed for these limiting

f0
displacements and the corresponding applied load was noted. For step

 

 

 

 

 

206

loading creep curves, an extrapolation procedure was used to estimate
some values of tf. The data are summarized in Table 6.3 and plotted in
Figure 6.16. This figure shows that lug bearing pressure decreased with
time until it reached a finite value q1t after a fairly long time.
Neglecting 6,, Equation 6.12 predicts that the lug bearing capacity

will go to zero as tf approaches infinity. Figure 6.16 suggests that
the long-term lug bearing capacity, q1t, does exist since infinity is
far beyond the usual service life. Note that when the applied lug
pressure is lower than qlt’ the secondary creep rate becomes practically
zero. For these reasons, it appears reasonable to replace q in Equation
6.12 by (q-q1t). In this case the lug bearing capacity q would tend to
a finite value q1t as tf tends to infinity.

The data in Figure 6.16 may now be compared with Equation 6.12.

Only data points corresponding to 1.0 in. limiting displacement are re-
plotted in Figure 6.27 on a log-log scale. Experimentally derived data
and failure times predicted by Equation 6.12 (dashed lines) are in

reasonable agreement.

6.4 Theoretical Predictions:

This experimental study has been primarily concerned with pull-out
loads for a steel bar embedded in a small frozen sand sample. A model
pile or anchor section is preferred when they save money and when
questionable assumptions are required for analytical methods. There
are some limitations imposed on straight-forward application of the
results from model tests to full-scale structures. Rocha (1957) suggests
that the following conditions must be fulfilled; the geometry, stresses,

strainssdisplacements, time required, Poisson's ratio, modulus of

 

 

 

 

 

 

207

elasticity and other properties of the model should be proportioned to
those of the field structure. The main purpose here is to adopt some
of the analytical methods, using the basic creep parameters of

frozen Wedron sand (Bragg, 1980) and the available theories, in order
to predict the pull-out capacity of plain and deformed rods. The
failure mechanisms of plain rods are different from those of lugs. The
two cases are, therefore, treated separately. A theoretical attempt
has also been made to relate both cases to the roughness criterion

defined in Equation 5.7.

6.4.1 'Bond for Plain Steel Rods:

The stress and displacement analysis for frozen soil surrounding
a plain rod (Figure 2.12) reviewed in Chapter II may now be adopted
for comparison with the present experimental data. Equation 2.13 suggests
that creep parameters from simple shear tests can be used to predict
the bond vs. displacement rate relationship for a plain rod in frozen
soil. Since some data (Bragg, 1980) from unconfined compression tests
on frozen Wedron sand are available, the shear strain rate Ic from
Equation 2.15 may be used, with the proof stress TC replaced by Ocue’

in Equation 2.13 to give:

399”2 ( / )“/(n-1) (6.13)

c Ta Ocue

For given values of a, n, e , and Ocue the relationship between

C

Ta and 0 can be established for a given temperature. Consider, for
example, the experimental lines in Figure 6.18 for the -10°C, and a 5/8
in. diameter plain rod (radius a = 5/16 in.). Using the available data

(Bragg, 1980) on frozen Wedron sand, with n = 11.765, éc = 6.0 X 10‘5

 

 

 

 

 

 

 

 

 

208

min.’1, and “cue = 1875 psi at —100C, Equation 6.13 was plotted on a log-
log scale in Figure 6.28. The theoretical solution provided by Equation
6.13 significantly over-predicts the bond strength in comparison with the
experimental lines in Figure 6.28. This large difference between the
predicted and experimental strength values may be partly attributed to
the difficulty with precise measurement of small rod displacements due
to equipment limitations. Also, the difference between the predicted
and experimental values appears to depend on the proper choice of

o and éc for use in Equation 6.13. For example, using 0 = 7067 psi

cue
and éc = 60 min.'1, as given by Bragg (1980), in Equation 6.13 would

CUB

reduce the predicted bond strength by about 13.5 percent.

Equation 6.13 was derived (Johnston and Ladanyi, 1972) for piles
in permafrost ground with boundary conditions slightly different from
those applicable to the pull-out specimen. For example, the effect of
soil stresses at the base plate on the strains around the rod were
ignored. Derivation of Equation 6.13 clearly implies that the n-value
for bond and compression tests is the same, thus assuming the same failure
mechanism for both tests. Experimental data (Table 6.4) showed that
the n-value from creep bond tests is about 31 percent of that from
compression tests, and its value from bond tests with constant displace-

ment rates is about 88 percent of that from compression tests, as shown

 

 

in Figure 6.29. Therefore, the failure mechanism on the rod-soil interface

is not the same as that on a shearing surface within the frozen soil.
Jellinek (1957 b) observed that at temperatures warmer than -14°C

(Figure 2.1) the adhesive strength of ice was smaller than its cohesive

strength. At temperatures colder than -14°C he observed that all the

breaks were cohesive, i.e. within the ice crystal, thus indicating

 

 

 

209

that the adhesive strength of ice to stainless steel at these temperatures
was higher than its cohesive strength. In sand-ice materials, sand
friction becomes a contributing factor to the total bond strength.

The introduction of sand friction would make it difficult to assume
whether or not the adhesive and cohesive mechanisms in sand-ice are the
same, without further experimental work.

It may be of interest to compare Weaver and Morgenstern's (1981 b)
Equation 2.16 with some of the creep data on plain rods shown in Figure
6.30. Information on the time t, elapsed after load application, and
the corresponding creep displacement 6C for use in Equation 2.16 were
obtained from the creep curves in Figure 5.32(a and b). Weaver and
Morgenstern (1981 b) assumed constant values for the creep parameters
b, c, k, and m on the basis of available data on frozen Ottawa sand.

As shown in Figure 6.30, the predicted (dashed) line considerably
under-estimatesthe strength for frozen Wedron sand. This might be
attributed to differences in particle surface features of Wedron and
Ottawa sands. Also, calculations show that a slight change in the
value of c, according to Equation 2.16, would significantly affect the
results. No primary creep data on frozen Wedron sand are available,

to show whether or not the c-value remains constant with temperature as

assumed by Weaver and Morgenstern (1981 b).

6.4.2 Load Prediction_Based on Cavity Expansion Theory:

Lug bearing forces can be predicted on the basis of cavity expansion
theory. The solution proposed by Ladanyi and Johnston (1974) for
frozen soils (Equation 2.18) was based on an analogy for an expanding

spherical cavity in a linear elastic-plastic infinite medium (Hill,1950).

 

 

210

The theory uses experimentally determined creep parameters for frozen
soils, thus modifying the model to be used for a non-linear viscoelastic-
plastic medium. With this procedure it appears that creep settlement

and bearing capacity of frozen soils under deep circular footings may be
predicted.

A comparison between the experimental data (Figure 5.40) for the
lug behavior and Equation 2.18 can now be made. The lug area may be
considered as the difference between two concentric circular footings.

The inner circle corresponds to the rod cross-sectional area. The

outer one corresponds to a footing with diameter d + 2h (or B), where

h is the lug height. For given values of et, Ocua’ n, p0, and n the
relationship between q and 6C can be established for a certain temperature.
To apply Equation 2.18 unconfined compression test data from previous
work (Bragg, 1980) on frozen Wedron sand was used. Bragg (1980) observed
that the failure strain 5f was almost constant throughout the tests.

At strain rates below 10'5 sec‘l, its value was about 0.045 independent
of temperature. At higher rates the value of cf at all temperatures
decreased slightly with strain rate indicating a more brittle behavior.
For comparison, a constant cf value equal to 0.045 was assumed for use

in Equation_2.19 at all temperatures. Values of Ocue’ c, n are temp-
erature-dependent according to Bragg (1980). Using the condition of

n = 1 if 6C a 0.1 B, the value of n was assumed to be 1.0, since the
lug settlement 6 in all creep tests was larger than 0.1 B. The value

of p0 was set equal to zero in all calculations, since no confining
pressure was applied to the pull-out specimens.

Results from these comparisons are summarized in Figure 6.31. The

 

 

 

211

predicted loads, P in this figure, were obtained by multiplying g,
calculated from Equation 2.18 for each case, by the lug area. Note
that the friction angle a was assumed to be constant, and equal to 300
for all temperatures and strain rates. This assumption seems reasonable
since Ladanyi (1981) reported that the friction angle for most frozen
soils is almost equal to that of the corresponding unfrozen soil. The
predicted loads (dashed lines) in Figure 6.31, generally, over-estimate
the loads at creep rates below about 10‘“ in./min. and appear to under-
estimate the loads at higher rates. The reason appears to be a difference
between n-values from compression tests and those from bond tests. If
values of n from bond tests, were used in Equation 2.19 the predicted
loads would be significantly improved.

Since the experimental lines in Figure 6.31 are parallel to each
other, they clearly indicate that the failure mechanism for a deep lug
(or footing) in frozen sand is the same for all temperatures from
—2°C to -15°C. A single value of n (4.93) can be assumed for these
lines. Other reasons for a difference between the predicted and experi-
mental lines in Figure 6.31 may include a difference between the actual
lug shape and the assumed circular footing in the theory, or the
presence of the rod which may prevent the formation of a complete semi-
spherical plastic "cavity" zone under the lug. Assuming a perfectly
plastic (incompressible) material within the cavity may not be entirely
compatible with the observed behavior. Uniaxial creep compression tests
on frozen sand reported by Bragg (1980) showed that the volumetric strain
increased with the creep strain. It was not clear whether reducing the
dilatancy by application of confining pressure would have any effect

on the volumetric strain. Visual inspections after thawing of all

 

 

pull-out specimens lead to the conclusion that continuous crushing of
sand particles had occurred as the lug continued to penetrate through
the soil.

Crushing of sand grains was clearly shown by the observation of a
thin layer of powdered quartz radially pushed into the frozen soil all
along the deformed zone behind the lug (Figure 6.32). It appeared that
prior to crushing of grains complete melting of frozen water must have
occurred under the high pressure (estimated at 25,000 psi for sample
08-14). The condition in front of the lug may correspond to stage II
in Figure 2.7. Pressure melting may have occurred locally around the lug
as the lug continued to penetrate into the soil. A sudden failure was
not characteristic of the sample behavior, hence global pressure
melting did not occur.

The cavity expansion theory assumes a maximum limit for the footing
displacement 6C of 0.1 to 0.2 B. The experimental lines were based
mostly on the step loading procedure. Under the initial load increments
creep rates often did not reach the steady-state, as shown earlier in
Figure 5.40, thus under-estimating the strength. Under the same loads
the total displacements did not exceed 0.2 8 (Figures 5.39 a to c). At
higher loads (over 2500 lbs.) the displacements far exceed 0.2 B
without reaching tertiary creep (or failure), thus violating the

assumption of a maximum 6 of 0.2 B in the theory.

6.4.3 "Roughness CriteriOn for Steel Rods:

The pull-out capacity of plain rods increased with increasing rod
surface roughness as shown in Figure 5.28. If the roughness factor
defined in Equation 5.7 is modified, lugs can be considered a part of

the rod surface roughness, as a large-scale asperity. In this case,

 

 

 

 

213

the question arises as to how much roughness a single lug will add to
the plain rod. To help explain the procedure used to define a single
lug "roughness" the actual heights of asperities SP (in micro-inches)
observed on the profilometer for different steel surfaces have been
listed in Figure 5.26. Also listed are the corresponding heights SR
as measured on the strip chart recorder (in millimeters). The latter
record was used as a basis for measuring the roughness factor of plain
rods.

To consider the lug as a large-scale asperity 0n the rod surface,
a correlation was made between values of SP, minimum and maximum in
Figure 5.26, with the corresponding values of SR as shown in Figure 6.33.
The relationship in this figure was linearly extrapolated to include
the actual lug height h, which is equivalent to SP in micro—inches on
the profilometer and to SR in millimeters on the recorder. For example,
if h = 1/8 in., the equivalent roughness SP = 125000 micro-inches on the
profilometer and SR = 52090 mm on the recorder based on the relationship
in Figure 6.33. This procedure is also illustrated in Figure 6.34. The
information in this figure was calculated for a cold-rolled steel rod
with a single lug (h = 1/8 in.) frozen into a 6 in. high sample.

Assuming that the tracer stylus could be moved 6 in. along the rod,
the equivalent length Lt in Figure 5.25 would be 2720 mm (Figure 6.34)
on the recorder strip chart. This length was calculated in proportion
to the standard chord Lc of 80 mm used earlier for plain rods. A chord
length of 80 mm on the strip chart was observed to correspond to 0.1767
in. on the actual surface. The observation was based on a tracer speed

of 0.221 in./sec. on the surface roughness, and a corresponding strip

 

 

 

214

chart recorder speed of 10 cm/second. The total "roughness line"
length Lr in Figures 5.25 and 5.26 equals (2 X 5209 + 272 X 58/8) or
12,390 cm for the cold-rolled steel rod in Figure 6.34.

The roughness factor 0, based on Equation 5.7, equals {(12390 — 272)
X TOO/272} or 4455 percent. In the same manner, the roughness factor
for rods with surface irregularities, including lugs, can be estimated.
The effect of lug height on the ultimate loads shown in Figure 5.22,
can be transformed in terms of roughness effect. Values of roughness
factor p which correspond to different lug heights are listed in Table
6.5, with their ultimate loads. The data, plotted on a log-log scale
in Figure 6.35, show a reasonable agreement with the data summarized
in Figure 5.28. At small values of p, below 100 percent, the ultimate
load appears to be less influenced by surface roughness. This behavior
was attributed to the fact that bond strength of frozen soil does not
totally depend on mechanical interaction between the soil grains and
rod surface roughness. Even on a perfectly smooth surface, 0 = 0 percent,
there will be some ice adhesion which may give a finite value for the
ultimate load. For this reason, the best fit line was extrapolated as

shown by the dashed curve in Figure 6.35.

6.5 'Applications:

Design of frozen soil structures, which involve reinforcement,
piles, or anchors, requires that suitable criteria be available for
selection of allowable adfreeze bond stresses and displacements. Bond
strength and creep rates are dependent on the frozen soil material
properties, temperature, and surface properties (type and roughness) of

the reinforcement member, pile, or anchor. To allow for complex geometry

 

 

 

 

 

215

within the structure and that stresses and strains may vary with time and
temperature, a non-linear viso-plastic finite element analysis appears to
be the most effective method.

In this section, a simplified analysis is explored for a reinforced
frozen soil beam, under pure bending moment. The analysis is based on
certain assumptions which require further verification by means of beam
tests. Also, a preliminary analysis considers some of the experimental

pull-out test results for application to piles and anchors in permafrost.

6.5.1 Frozen Soil Reinforcement:

There is general agreement that the behavior of frozen soil in tension
is quite different from that in compression at strain rates higher than -
10‘5 sec.’1 (Vyalov, 1962; Hawkes and Mellor, 1972; Haynes et al., 1975;
Haynes and Karalius, 1977; Bragg; 1980; Eckardt, 1981). For frozen Wedron
sand, it was observed that the ratio of compressive to tensile strength,
oc/ot, was about 5. Therefore, a frozen soil beam may fail because of
a weakness in tensile strength. Also, it was noted in Figure 2.18 that
beam deflections can be greatly reduced with reinforcement.

For a simple beam subjected to pipeline load as shown in Figure
1.1 (c), or the laboratory beam model in Figure 2.18, the analysis can
be simplified by neglecting tensile resistence of the frozen soil.
Consider, for example, the beam in Figure 2.18 which failed under a total
load increment P of 433.1 lbs., when reinforced with a 3/16 in. diameter
plain rod (800, 1983). The corresponding external bending moment was
calculated to be 5630 inch—pounds. It is of interest to compare this
bending moment with that estimated from an analytical procedure. A

cross-section for this beam is shown in Figure 6.36 (a). Using

 

 

 

 

 

216

Bernoulli's theorem (a plane section remains plane after bending) a
linear strain distribution was assumed throughout the beam depth as shown
in Figure 6.36 (b). Assuming the equillibrium condition shown in Figure
6.36 (c) and neglecting the soil resistance in tension, the position of
the neutral axis 21 can be determined using Equation 2.22.

Now calculate the resultant force FC in the compression zone using
Equation 2.24 and available data on frozen Wedron sand (Bragg, 1980).
For a temperature of -10°C these data include the proof stress Ocue =

1875 psi at éc = 6 X 10'5 per minute, and n = 11.765. The radius of

curvature R can be replaced by zl/ec, where e is the compression strain

c
at the outer fiber. Substituting these parameters into Equation 2.24, the
compressive force FC can be expressed in terms of 21 and ac ( with b =

2.75 in.) as follows:
FC = 10858 21 eCO-OBS (6.14)

The resisting force in the tension zone is controlled by the steel
tensile strength or by the bond resistance. In reinfcrced concrete, where
creep is almost negligible in comparison to frozen soil creep, it is nor-
mally designed such that bond slip does not occur and failure is governed
by the steel tensile strength (Hughes, 1976). In frozen soil, where there
is no generally accepted design philosophy, bond creep is significant and it
cannot be neglected. Interaction between the adfreeze bond and tension of
steel rods depends on different factors including time, temperature, and
rod surface roughness. While the bond resistance varies significantly
with these factors, the tensile resistance remains almost constant for a
given rod cross-section. Based on the available experimental data, the

following calculations show that adfreeze bond resistance for the 3/16

 

 

1 MW ~

 

 

 

 

217

in. diameter plain rod (Figure 6.36 a) controls and that it was small in

comparison to tensile resistance in the rod.

The bond resisting force Fb (Figure 6.36 c) is given by:

' m
Fb n L d TC (of/tf 6C) (6.15)

where L ‘ 20 in. is the bond development length (Figure 2.18), d

3/16 in.
is the rod diameter, of = 0.06 in. is the bond displacement of the rod at
failure (Figure 5.32 b), and tf = 5040 min. is the time to failure for the
beam test (Soc, 1983). The creep parameter Tc, 6C, and m, obtained from
Figure 5.33 include values of 26.5 psi, 10"6 per min., and 0.37, re-
spectively. Substituting these values into Equation 6.15, gives Fb = 333
pounds. Comparing this value with that due to the rod tensile resistance
of 1100 lbs. (based on 40000 psi yield stress) suggests that bond failure
would be more probable than tensile failure. Equillibrium requires that

Fb = Fc’ and hence:

10858 21 eC0°°85 = 333 (6.16)

According to Equation 6.16, the value of 21 = 0.082in. for ac = 10'5,
and 21 = 0.04 in.for cc = 0.045 (failure strain based on Bragg‘s (1980)
data). In both cases the value for 21 is small compared to the depth
d' of 2.5 in. in Figure 6.36 (a). Therefore, any value for ac in the
range of 10-5 to 0.045 would not result in any significant error. With
an average value of 0.06 in. for 21, and (d' - 21) = 2.44 in. (Figure 6.36
b), the internal resisting moment would be 812.5 in.-lbs (= 2.44 in. X
333 lbs.). It is clear that the adfreeze bond would be inadequate to
counter act the applied bending moment of 5630 inch-pounds. For this

particular beam, a significant soil contribution must exist in the tension

 

 

 

 

218

zone.
It was anticipated that introduction of the frozen soil tensile re-
sistance Ts would shift the neutral axis upward, as shown in Figure 6.36
(d), thus further complicating the analysis. For beams reinforced with
deformed rods, which appear to tolerate larger displacements than plain

rods (Figure 5.20), the strain 5 at the reinforcement level (Figure 6.36

r
b) would be large enough to cause failure of frozen soil in tension before
bond failure occurs. In this case, neglecting any soil tensile forces

may introduce less error than the case of using plain rods. More ex-
perimental data on the beam behavior is needed before proceeding with the

analysis.

6.5.2 Pile or Anchor Capacity:
Pile installation in permafrost often involves dry augered holes

with a mixture of soil and water used to fill the annulus around the

pile (Figures 1.2 b and c). Piles and anchors develop their load supporting
capacity when the soil-water slurry is solidly frozen in place ( Anders-
land et al., 1978). Pile capacity is primarily dependent on the long-term
adfreeze bond strength of the frozen slurry, the pile type and surface
roughness. There appears to be a general agreement that pile design in
ice-rich soils should be based on limiting creep deformations, and in
ice-poor soils design should be based on allowable adfreeze strengths
(Weaver and Morgenstern, 1981 b). End-bearing at the pile tip is nor-
mally neglected, particularly in ice-rich soils.

Steel rods embedded in frozen sand specimens can be viewed as a

section of a model pile. It appears reasonable to consider the experimental

data herein in terms of potential field behavior of piles in permafrost.

 

 

 

219

Consider the following example involving a steel pipe pile (1 ft. diameter
and 25 ft. length) to be used as a vertical support in a warm permafrost
with an average temperature of -2°C. A sand slurry backfill, carefully
controlled as to density, is placed around the pile as shown in Figure
1.2 (b). The permissible pile displacement is limited to 1 in., and the
maximum active layer thickness is predicted to be 5 ft. during a 20 year
service life for the structure. The allowable axial load is required for
the pile (cold-rolled steel surface). The pile capacity is to be based
on limiting creep deformation to 1 inch. For the 20 year service life
and pile deformation the creep rate will be 8.5 X 10"8 in./min. This
rate corresponds to the slow rate range in Figure 6.18. Since very
limited creep data are available in this range, it will be assumed that
the constant displacement rate power law (Equation 5.3) is applicable.
Using the data in Table 5.3, Equation 5.3 bcomes Tu = 12 a
(10000 an)°°222. Since no loading rate P was given, Equation 6.2 could

be written in terms of the displacement rate an as follows:

Iu (psi) = 12 (10000 an)0'222(p/625)0°325 (vs/64)2'2”6 (d/IOH)S 0
(6.17)
Substituting d/H = 5.0 percent (1/20), s = 0.692, vS = 60 percent,

0 = 625 percent (cold-rolled steel), in = 8.5 x 10'8 in./min., and e =

 

20C into Equation 6.17, the adfreeze strength Tu = 5.436 psi gives an
allowable axial pile load of 49183 pounds. Note that the effective em-
beddment depth H was taken as 20 ft by neglecting the top 5 ft (the
active layer). Also, any downdrag loads, were ignored in the example.
Corrugations (lugs) have been used on pipe piles (Figure 1.2 c) to

increase load capacity on the trans-Alaska gas pipeline project (Thomas

 

 

220

and Luscher, 1980). The lugs were spaced at 1 ft, with a lug height

h = 3/4 inch. If similar lugs are used on the 25 ft. piles, there would
be 21 lugs in the effective 20 ft. embeddment. Load-displacement curves
reported by Thomas and Luscher (1980) for corrugated piles showed a
behavior similar to that observed for the lugs (Figure 5.14) with no
"initial failure" and no drop in load at the ultimate condition (Figure 5.21).
This behavior appears to indicate no lug interaction, described in
Figure 6.15 for the standard deformed rods. Using Equation 6.8, with

IS = 60 percent, A1/Ar = 0.2656, 6° = 8.5 x 10"8 in./min., and a = 2°C,
gives q = 7105 psi for a single lug on a corrugated pile, and the pile
load becomes 213,445 lbs. per single lug. The total allowable axial load
for 21 lugs through the 20 ft. depth bearing layer would be 4,482,339
pounds. This load compared with the 49183 lbs. for plain piles is about
91 times higher. Such a load, however, would exceed the ultimate com-
pressive strength bbout 550000 lbs) for a pile cross-section with a

3/8 in. wall thickness. The lug height can be reduced and the spacing

can be increased and still obtain a load capacity equal to that of the
steel pipe. Note that the confining lateral earth pressure was neglected
in the analysis. Had the confining pressure been considered, higher loads
(for plain and corrugated piles) would be obtained.

Anchors are also used in permafrost for resisting uplift forces in

 

various types of structures. The adfreeze strength of grouted rod anchors,
or pile anchors, may be estimated from pull-out tests on plain rods in the
same manner described for the pipe piles, using Equation 6.2. Screw anchors
of the type shown in Figure 1.2 (d) can be simulated by the lug behavior
observed in this study. The bearing capacity of a single helix (screw)

anchor can be estimated using Equation 6.8.

   

 

221

The foregoing examples indicate that the scale effect (due to
data transformation from a small—scale laboratory model to a large-scale
field structure) can be eliminated by applying the dimensionless ratios

d/H and Al/Ar‘ However, when considering the load-displacement curves,

the pull-out load can be normalized with respect to the rod surface area,
to give a nominal bond stress, and the corresponding displacement can be
normalized with respect to the rod radius to give the average shear strain,
assuming a uniform stress and strain distribution along the rod length.
Further model studies and dimensional analysis are required in order to

make possible a multiple regression analysis for further applications.

 

 

 

 

 

222

Table 6.1: Comparison of frozen sand cohesion with the bond and
lug bearing capacity.

 

T 00 c=ou/2 q1 Pu/nH dt Pjy/fiHdT tuft
( °c ) (2531 (psi) .1921 (psi) (psi) (£11
- 2 1100 550 7200 200 113 27
- 6 1700 850 10692 297 174 78
-10 2100 1050 15012 417 230 127
-15 2500 1250 19872 552 297 191

 

* Data from Bragg (1980), values of Cu correspond to éavg =
3.0x1076 sec."1 (Bragg's Figure 5.9).

i d = 3/8 in., H = 6.0 in.
if For a 3/8 in. diameter plain rod.

(Other data from Table 5.2 for Savg = 5.0x10-4 in./min. and
Pavg = 25 lbs./min.)

 

 

223

frozen sand at ~100C.

Table 6.2:
Sample h
No. in.)
08-9 1/16
CS-13 1/16
CS-15 1/16
CS-5 2/16
CS-6 2/16
CS-7 2/16
CS—8 2/16
CS-lO 2/16

A
l
(sg.in.)
0.08590

0.08590

0.08590

0.19635

0.19635

0.19635

0.19635

0.19635

ql
129;).
16298
18161
20954
24214

9895
13970
18161
30268

34924

11714
12732
14158
15381
16196

14158
15686
16909
18131

18131
19251
20372

20372

8047
9473

6Cx10-

(in./min.)

5

 

46.
70.
104.
166.

21.
58.
718.

2500.

10.
16.
25.
37.
44.

25.
43.
69.
83.

OKONO

OOOV

mmoooooo

003030

100.0
166.
200.

286.

(Continued)

Effect of lug size on creep rate and lug bearing in

 

 

 

 

 

 

224

Table 6.2: (Cont'd)

 

S l - -
amp e .h A, q1 6Cx10 5

No. (in.) (sq.1n.) (psi) (in./min.)
CS-10 2/16 0.19635 10899 13.30
11714 13.10
CS—14 2/16 0.19635 22918 975.00
CS-18 2/16 0.19635 6213 1.82
7843 1.77
9269 3.64
10695 5.09
12121 6.94
CS-11 3/16 0.33134 6036 1.61
7485 4.05
9175 6.71
10865 9.07
12555 20.40
05-17 3/16 0.33134 7545 8.00

10563 10.00

 

225

Table 6.3: Values of ”psuedo” instantaneous displacement 61
and time to failure tf at different creep loads
for a 1/8 in. lug in frozen sand, (6f is the
assumed failure displacement).

 

 

Sample P 61 tf(min.) for:
No. ribs.) (m) 6f= 1.0 in. 6f= 0.25 in.
cs-5 2300 0.081 9188 1688
2500 0.088 5431 967
2780 0.097 3270 605
3020 0.100 2400 400
3180 0.106 2004 322
05-6 2780 0.080 3680 680
3080 0.084 2100 380
3320 0.088 1310 233
3560 0.092 1094 190
05-7 3560 0.084 916 166
3780 0.090 548 96
4000 0.093 454 77
08-8 4000 0.081 320 59
CS—lO 1580 0.028 11669 2665
1860 0.032 10314 2319
2140 0.034 7263 1624
2300 0.035 7366 1641
CS—14 4500 0.120 95 13
CS-18 1220 0.022 53736 12527
1540 0.029 54859 12486
1820 0.036 26483 5879
2100 0.038 18978 4243
2380 0.039 13876 3069

 

226

Table 6.4: Comparison of the n parameter from unconfined

compressirn tests with that from bond tests.

 

T
n n n' n”

18:; C b b

— 2 3.30 5.40 4.50 --

— 6 8.70 6.70 5.40 —-

-10 11.67 10.40 9.35 2.70

-15 12.19 11.40 10.60 4.80

 

“b
"b

”6

From compression tests (Bragg, 1980)
From bond tests, Figure 5.3
From bond tests, Figure 5.4

From creep bond tests, Figure 5.38

Table 6.5: Variation of ultimate load with rod surface roughness,
including lugs.

 

Type of Lug** h p Pu
steel shape .11fl;1 (_%_)_ (lbs.)
CR* 45° 1/64 1085+ 1050
CR 90° 1/16 2540 1740
CR 90° 2/16 4455 2920
CR 90° 3/16 6370 4800
CF 45° 2/16 3920 2730
CR 45° 2/16 4455 3380
SB 45° 2/16 5319 3400

 

* CR: Cold—rolled steel surface
GF: Ground-finish steel surface
SB: Shot—blasted steel surface

*

*

For description of lug shapes, see Figure 3.2

i Data for a 3/8 in. standard deformed bar reduced to a
single 1/64 in. lug. The 6 in. rod length was assumed
to have the same roughness as that of the cold-rolled

steel surface.

 

228

   
   
 

300
S C)
avg = 3.65 X 10‘3 in./min.
280— V Stehle (1970) /
(H-section, 3 in., 6 in.,
and 12 in.) :
240 _ Tsytovich and Sumgin
(1959) (Ice to plain
'§_ - ////// wooden piles, long-
; term basis)
'T 200 _
I;
4.)
CD
8
L
a 160 - /o
.“>’ . /
g /
g 120 — /
§ /
/
/ / ”
80 ‘0.“‘Frederking (1979) /’
/ (Ice to 3 in. dia. wooden
// piles)
40 V
(Ice(1), Figure 5.29, 5/8 in.
./ plain steel rod)
//
0 ~./. . l I I I I I a . I I I

 

 

igure 6.1:

Temperature, T,°C

Temperature effect on ice adhesion to different materials.

140

'5

D.

l:\

I?

U)

OJ

L

4.:

U) 60

'U

C

O

CO
40
20

7igure 6.2:

 

80 I!

 

S

 

229

. = 2.53 X 10'3 in./min.
= -6.2°C

ample (S-10)

(Plain steel rod,

5/8 in., H = 6 in.)
A
[I
/l
[I
/ I
I l
/I
’ I

(Painted steel pile, d = 3 in.,

 

l

l H = 7.5 in., Ottawa sand)

I \

I \

' \

I - o—o——o——-

llf\\

V \ ________________
I l J l l J l 1 I I
1 2 3 4 5 6

No

rmalized Displacement, 26/d, Percent

Typical bond stress versus displacement curves for plain
steel piles (rods) in frozen sands.

 

AmPMeprmE “cmcwwmwu mo
mmFTQ cpv ucwm :mNoem new wow to; :\e cape; 65p sews spacmepm vcog mo cowumwem> Hm.o wezmwd

eeeecma .I\e .oepam

 

     

moH m m a N NoH w o v m CH m w e N H
—-_u—-_ q —d-—_-__ - —..u~.- a - O
\.
\
c III/ \ I ow
.eee\.ee mIoH x ec.m I e .uoeII/./ . m.
up mowIzocm :e mo_ea cocoozv xx - ”W
o / 0 al..
mw Amnmfiv mewxsmwmgd IIIIIIY/ Amfi m mczmwd Eocm mpmov w
62 // I ow mw
UI
Amfi.m beamed n1
Eoem mumev . .
/ c .m
.eee\.ee .IoH x mH.m I e I . I ONH .
.ucmm :wNocm Cw mvoc _wwpm III.
oooHI I H
E 08H

 

Bond Strength, psi

gure 6.4:

. — graded sand slurry)

p_I
03
O

|_.I
N
O

 

to
L0
....

Pavg. = 7 lbs./min. C)
---— Crory (1963)
(8 in. steel piles in well-

(Data interpolated
from Figure 5.3,
d = 5/8 in.)

C)

” // /”

 

 

/ ,.a"’ Vyalov (1959)
_ / // (Equation 2.11, for
/ 6 in. timber piles
"I, 0/ in frozen silty sandy
[/1 . . I I I I I 1 . I . . I 10aIm)I I
—5 -10 -15 -20 —25

Temperature, 00

Comparison of bond strength for different pile types in
frozen sandy soils.

1200

1000

800

 

[ T = -10°C
Pavg. = 25.5 lbs./min.
- (Plain steel rod, d = 5/8 in.)

 
    
 
   

-+-Frozen sand (vS = 64 percent)

Crushed ice

 

 

 

600 T
400
(Linear scale between numbered points)
200
\ #1—21 _Residual loads, Pr
.L_.___>ae-.-e-—.—-.-’—.--.--.-.- .
0 0.01 .0.04 0.06 0.08 0.10
Displacement, 6, in.
a 6.5: Comparison of load-displacement curves for a plain steel

rod in ice and frozen sand.

UVIIU Jul cllytll, pbl

320 [

280"

240 —

160 —

120 .

80 —

 

 

I‘J
(A)
L.)

‘——---- Tu vs. T for frozen sand

T vs. T for Ice

 

/
/
.___ _ Tr vs T for frozen sand _ 1’
————— Tr vs. T for Ice 2"],

Pavg. = 24.6 lbs./min.

= 'LI' '
°avg. 5 X 10 in./min.

L
*—

Sand Friction. Tuf

Ice adhesion "’,.e
[I ”f/.' 4
’//,//’/// ’ ”,,.—" Sand Frictgon,
rf

, ’ .ec’T”’ Ice Adhesion~\\

/

_‘

 

' 1- r'.‘T j" 771': I_1' . . I . . 1

 

Temperature, T, 00

Comparison of ice adhesion and sand friction for a 3/8 in.
plain steel rod embedded in ice and frozen sand.

 

234

I Gavg. = 6.4 X 10.1+ in./min.

   
  
    

Pavg. = 28.5 lbs./min.
_ T = —10°c
d = 3/8 in.

(Linear scale between
numbered points)

 

Lug bearing

Plain rod with lug in frozen sand (S-57)
_ h = 1/8 in. I

i
i‘L—Plain rod in frozen sand (S—30), h = O A

“/ Plain rod in ice (I-22), h = 0

   
  

 

 

Sand Friction

\ x. Ice Adhesio A

\ ~o— —.— ——"——_—)(I_
H ‘ "TQI 'I'.:-'Q_ T'LL— —l-—I- +—."' T"?".‘ L.
l 8 ‘ 0.1 .02

' 0,10 0.20 0.30
Displacement, 6, in.

'e 6.7: Comparison of lug bearing, sand friction, and ice ad-

hesion effects on the load-displacement curves for
steel rods.

 

 

235

   
   
 
  
   
 

 

 

4.2!-
' : “14' '
°avg. 5.0 X 10 in./min.
3.6 __ Pavg. = 24.6 lbs./min.
3.0 —
I”
m h /
,2— Plain rod with lug
if“ d=3/8 in.,h=1/8 in.
g? (Figure 5 24, for Lug Contribution,
/
.3 Piy vs T) v’ if
m 1.8 - '
3
D_
. - Plain rod
U“)

E d=3/8 1.11., h=0
_J 1'2 (Figure 5.17, for
- ,r Pu vs. T)

./
- / Adhesion and
6 n
/ friction
OL(.1.1....II...I.JI.I.JI-II
0 -5 -10 -15 -20 -25
Temperature, T, 0C
gure 6.8: Lug contribution for the initial bond rupture condition at

different temperatures.

 

6.6 [
Pavg. = 24.6 lbs./min.

: -LI’ '
°avg. 5.0 X 10 in./min.

. -————-- Frozen sand with lugs

4 6 _ —-—- Frozen sand] Plain rod
' -——— Ice only

G.

= 3/8 in.
3'6 7 h = 1/8 in.

 

Ultimate and Residual Loads, Pu and Pr’ kips

 

    
  
 
  
 
 
   

  
 

Lug contribution,

 

 

2.6
. P1
- Ultimate loads, Pu
(Figure 5.24)
1.6 ‘
/
0.6)-
. u -—””’
./I/
Residual loads, Pr _,,””'
0.3— _/
. Sand Friction
-—""”’ Ice Adhesion __
’/ __ ____-———— ""'—'
0 —————-.--r‘—'."'."."'II...I....I.r...|
0 -5 -10 -15 —20 —25

Temperature, T, 00

igure 6.9:

ultimate conditions.

Comparison of ice adhesion, sand friction, and lug contri-
bution for steel rods at different temperatures and

Ultimate and Residual Loads, Pu and Pr’ lbs.

7200

6000

4800

3600

2400

1200

Lug Height, h, in.

1/64 1/16 2/16 3/16
I I I I
. = _u n a
I °avg. 5.0 X 10 in./min. 1F//'
Pavg = 24 6 lbs /min

= 3/8 in. /////

d
/ /

' Least- -Squares Lines //,

for Pu in Figure 5. 2

~200C/
// o
_ -100Cr/’ Lug Contribution
/’
/

 

 

 

 

 

_ ’/ Residual Loads, P
44 3:3;
I A II.

Al/Ar Ratio

6.10: Lu contribution to the ultimate pull—out capacity of a
plgin rod, with a single lug With different lug heights,

in frozen sand.

 

238

 
 
  
   

1.0 ”
[ d = 3/8 in., h = 3/8 in., H = 6 in.
3% _ savg. = 5.0 X 10'“ in./min.
g: ‘8 Pavg = 24.6 lbs./mi'n.
I
!=
\
3
D- _
U .6
C
f5
4 —El—— (Pu/find) = 0.390 c /
U
I .—
ii” .4 ___—()—-- (Piy/an) = 0.227 c
3; I Tu = 0.190 C/ /
:3

  

 

 

Cohesion, c, Ksi

re 6.11; Comparison of frozen sand cohesion (at different tempera—
tures) with the bond and lug bearing capacity.

 

    

[ Pavg. = 24.6 lbs./min.

éavg. = 5.0 X 10‘” in./min.

15‘ A
(D
_ C h = 3/16 in.

3"

= 1/16 in.

3'

= 2/16 in.

Lug Bearing Capacity, q], Ksi

  

 

 

10_ q] = 6.87 Cu ___—v
/
/
- /
C)
/
5 ’ l l I I
-2°c -60c -100c -150c
/
- /
47 (Data for h = 2/16 in. listed in Table 6.1)
7
O / . I . I l l 1 I I I J I
0 0.5 1.0 1.5 2.0 2.5 3.0

Uniaxial Compressive Strength, cu, Ksi

'igure 6.12: Relationship between lug bearing capacity and uniaxial
compressive strength for frozen Wedron sand.

 

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I
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' (d = 3/8 int;\‘*(

 

 

 

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gmk
Load (Not to a scale)

re 6.15: Pressure bulb overlap for consecutive lugs on a 3/8 in.
diameter deformed bar.

 

243

 

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15.2 .
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6.8 -

4.0 _

0.8

Compressive Strength Cu and Lug Bearing q], Ksi

C)
4:.

 

(Data from Figure 5.48 adjusted for temperature
difference and residual loads)

   
  
 

Goughnour (1967), cu vs. vS

(é = 1.33 X 10"+ min.'1, data from

Figure 2.5)

  

J n l

 

igure 6.17:

10 20 3O 40 50 60
Sand Volume Fraction, Vs’ percent

Sand Volume fraction effect on the unconfined compressive
strength of frozen Ottawa sand and the lug bearing capa-
city of frozen Wedron sand.

 

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:igure 6.19: Lug capacity comparisons, creep tests versus constant
displacement rate tests, for different test conditions.

247

ru/e, psi/0C

 

 

20 O
{ Gavg. = 5.0 X 10'1+ in./min. C)
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10 .. \
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- Figure 5.19)
8 - E] C)
E]
6 -
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3 _ Best-Fit Line
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(Data from Figure 5.18)
2 -
1 . l I l J 1 l n I l . 1 1
1 2 3 4 6 8 10 20 3o

Ratio d/H, percent

Figure 6.20: Modified relationship between bond strength and the
ratio d/H for plain steel rods in frozen sand,
(v5 = 64%).

 

 

 

 

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Figure

249

 

 

F Plain rods , h = O
L
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. Pavg. = 25 lbs./min.
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Temperature, T, 0C

6.22: Comparison of Equation 6.2 with experimental data,
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Frozen sand
Steel rod ———-—L—a» «\k‘
Initiai position

Void

Fina] position '9’ 2:57
A 1.. ..

. .1 ‘.

/ Zone of crushed sand

 

 

 

 

 

 

Figure 6.32: Zone of crushed sand observed around the lug, at the end
of creep tests.

 

 

 

 

 

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CHAPTER VII

SUMMARY AND CONCLUSIONS

The results and conclusions based on this investigation are sum-

marized in the following sections: test methods; load transfer mecha-

nisms; applications-- reinforcement, piles, anchors; and recommendations

 

for future work.

7.1 Test Methods:
Measurement of adfreeze bond between frozen sand and steel bars
has involved use of constant displacement rate and constant load (creep)

pull-out tests. Test temperatures included the range of -2 CC to -26 OC

 

and were closely controlled (:_0.05 0C) by immersion of protected samples
(membrane enclosed) in a circulating and refrigerated coolant liquid.

A specially constructed load frame supported the frozen sample while one
end of the steel bar was attached to an eye connector and hook at the
bottom of the coolant tank. An external loading system permitted pull-
out of the bar at a constant displacement rate or bar creep movement
relative to the frozen soil for a constant pull-out load. The following
conclusions refer to sample preparation methods and test procedures:

1. Small rod displacements prior to bond failure and slip req-
uires a high sensitivity for measurement equipment. Transducer sensiti-
vity of at least :10'5 in. is desirable to show elastic movement prior
to rupture in constant displacement rate tests and creep movement in

constant load tests.

264

 

 

 

 

 

 

265

2. Loads selected for creep pull—out tests are limited by the
high immediate bond failure stress and a lower long—term bond strength.
Smaller loads do not develop steady—state creep rates needed for predic-
tion of bar pull—out capacity. Vyalov's (1963) equation may be used for
prediction of the long-term bond strength corresponding to a specified
rod displacement.

3. The step loading procedure provided creep rate data for several
pull—out loads on each sample. The initial time interval and load must
satisfy long-term strength requirements to insure development of steady-
state creep.

4. The ultimate load capacity of plain rods was dependent on rod
and sample size as represented by the ratio of rod diameter to sample
height, d/H. An optimum d/H ratio, on the order of 6 to 10%, appeared to
give the largest adfreeze bond strengths for plain rods in frozen sand.

5. For 3/8—in. diameter rods with 1/8-in. high lugs, frozen soil
diameters of 4 in. or more were required to avoid any influence of sample
diameter on lug bearing capacity. Lug height h, rod diameter d, and
sample diameter D may be used to define a ratio A = (D-d)/2h for which
values should be at least 13 in order to avoid any influence of sample

diameter on lug bearing and pull—out loads.

7.2 Load Transfer Mechanisms:

Adfreeze bond between frozen sand and embedded structural members
(reinforcement, piles, or anchors) include: (1) ice adhesion, (2) fric—
tion with sand particles, and (3) mechanical interaction of frozen soil
with surface roughness of the structural member. These bond components

involve different load transfer mechanisms. Ice adhesion is the result

 

 

 

 

266

of increased attraction between water molecules and the surface of the
structural member at freezing temperatures. Surface type (steel, wood,
paint, etc.) and contaminants (salts, minerals, etc.) in the ice signi—
ficantly influence adhesion forces at the interface and cohesive forces
in the ice. Friction with soil particles is dependent on the friction
angle between the soil particles and structural member, and is propor-
tional to the normal stress that is pushing the particles against the
member. Surface roughness of the structural member introduces particle
dilatancy and reorientation effects adjacent to the surface. An increase
in surface roughness by the introduction of lugs on the surface permits
bearing forces to interact with frozen soil. The result was a significant
increase in load transfer as surface movement mobilized mechanical inter-
action forces. The following conclusions refer to load transfer mechanisms:

1. Ice adhesion was mobilized at very small displacements with
rupture typically at less than 0.002 inches. The presence of sand part—
icles increased the ice matrix volume involved in rupture and thereby
increased the adfreeze bond. An increase in surface roughness by use of
lugs added significantly to the adfreeze bond with no observed change in
displacements required for initial rupture.

2. Friction between sand particles and the structural member was
negligible as shown by the small residual loads. Unconfined frozen sand
samples contribute little normal stress on sand particles adjacent to the
structural member.

3. An increase in surface roughness, by the addition of a single lug,
gave load—displacement curves showing the initial rupture of ice adhesion

followed by mobilization of bearing forces on frozen soil. Residual ice

 

 

 

 

 

 

 

267

adhesion and friction for larger displacements were negligible in contrast
to the larger bearing forces.

4. Use of multiple lugs, a deformed steel bar as used for concrete
reinforcement, showed an increase in the ultimate pull-out load in a
direct ratio to the number of lugs after residual adhesion and friction
loads were subtracted. A modified load—displacement curve suggested a
possible relationship between ultimate load, lug size and spacing, and
maximum particle size.

5. Sand density (or sand volume fraction) significantly influenced
the adfreeze bond strength for a given creep rate, bar and lug size. For
ice—rich samples (sand volume fraction < 42 %) the pull—out load depends
primarily on plastic deformation of the ice matrix. For ice-poor samples
(sand volume fraction > 42 %) the adfreeze bond and pull—out load incr-
eased abruptly with mobilization of sand dilatancy and particle reorient-
ation effects along with soil bearing forces on the lugs.

6. Colder temperatures (range of —2 0C to —20 0C) and larger
displacement rates both increased the adfreeze bond components. Expe-
rimental relationships were developed which account for temperature, bar
displacement (or creep) rate, and other test variables.

7. A roughness factor defined by Wright (1955), provided excellent
correlation with increased bond or pull-out loads for plain rods. An
increase in surface roughness, the addition of lugs, extended this corr-

elation suggesting possible future design applications.

7.3 Applications-- Reinforcement, Piles, Anchors:
Structural members (reinforcement, piles, anchors) are embedded in

frozen soil to take tensile forces or to transfer and distribute compre-

 

 

 

 

 

 

268

ssive loads in such a way that stability of the structure is maintained.
In all cases load transfer between the frozen soil and the structural
member involves the adfreeze bond and possible slip at their common inter-
face. Installation in frozen ground commonly involves dry augered holes
with a mixture of sand and water used to fill the annulus around the pile
or anchor. Load supporting capacity develops when the sand—water slurry
is solidly frozen in place. Load capacity is dependent on the long-term
adfreeze bond strength of the frozen sand in the zone adjacent to the
structural member. Several conclusions can be made relative to design
considerations for placement of these structural members:

1. The dry density (or sand volume fraction) of the sand-water
slurry should be as high as possible to insure a high adfreeze bond str-
ength. Vibration of the slurry during placement to insure a relatively
high dry density appears to be a desirable requirement in construction
specifications.

2. The use of multiple lugs on the structural member will signi—
ficantly increase the long—term pull-out load. The lugs increase both
the initial ice adhesion and friction and are responsible for develop-
ment of bearing forces on the lugs at larger displacements

3. Lug spacing should be such that forces which develop in front
of the lugs do not overlap with pressure zones from adjacent lugs. A
preliminary analysis based on Boussinesq's equations suggests that a lug
Spacing should be at least twice the lug width (d+2h). The design should
also allow for formation of a void space behind the lug with length equal

to the allowable pile or anchor displacement.

4. Constant load (creep) pull-out tests using pile and soil materials

 

 

 

269

of the same type as proposed for a project will help an engineer to
account for the many variables which affect adfreeze bond strength.
Field tests are needed to correlate test results on model pile sections

with full size piles.

7.4 Recommendations for future work:

Additional research is needed on several aspects related to the

bond and slip of steel bars in frozen sand. Several specific problems

are outlined below:

1. The influence of confining pressure on bond components should
be investigated. The available predictions are mostly based on the ass-
umption of soil THCOWPPESSlblIlto’ or volume constancy. The effect of
confining pressure on the volumetric strain of frozen sand (which char-
acterizes the dilatational behavior of soil in front of the lug) does
not appear to have received the attention it warrants.

2. Use of standard deformed bars for reinforcement appears to
provide a significant capacity at fairly small displacements. Further
tests are needed on these bars in order to provide design information on
the most appropriate lug spacing and lug size.

3. In order to precisely evaluate the bond stress along steel
bars, it will be necessary to monitor the stress distribution along the
bars during pull-out tests. Redistribution of stresses with time, during
creep tests, is also anticipated. Measurement of stresses and strains
along the bar can be achieved by using a split rod which accommodates a

number of strain gages.

4. Laboratory investigation of the performance of reinforced beams

would provide valuable information on the time dependent interaction bet-

 

 

 

 

 

 

 

 

 

 

 

270

ween the reinforcement and the frozen sand beam. Properly instrumented
beams would help define the strain distribution in the beam. Such infor-
mation would be a significant contribution to developing design proced-
ures for field applications of reinforced frozen earth structural elements.
4. Further model studies and dimensional analysis are required in
order to make possible a multiple regression analysis for further appli-

cations.

 

 

 

 

 

 

BIBLIOGRAPHY

 

BIBLIOGRAPHY

Reference Code:

ASCE: American Society of Civil Engineers

CGJ: Canadian Geotechnical Journal

HRB: Highway Research Board

ICP: International Conference on Permafrost

ISGF: International Symposim on Ground Freezing
JACI: Journal of the American Concrete Institute
JSMFD: Journal of the Soil Mechanics and Foundations Division
NAS: National Academy of Sciences, Washington, D.C.
NRC: National Research Council

RR: Research Report

SR: Special Report
USA CRREL: U.S. Army Cold Regions Research and Engineering Laboratory

USA SIPRE: U.S.AArmy Snow, Ice and Permafrost Research Establishment

271

 

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Rowely, R.K.; Watson, G.H.; and Ladanyi, B. (1973) "Vertical and Lateral
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tion, Yakutsk, U.S.S.R., pp. 712-720.

Rowely, R.K.; Watson, G.H.; and Ladanyi, 8. (1975) "Prediction of Pile
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(1969) "Foundations of Structures in Cold Regions," Monograph

Sanger, F.J.
III-C4, USA CRREL, Hanover, N.H.

Sanger, F.J.; and Sayles, F.H. (1978) ”Thermal and Rheological Computations
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26, pp. 311-337, (Edited

opments in Geotechnical Engineering, Vol.
Co., New York.

by H.L. Jessberger) Elsevier Scientific Publ.

Sayles, F. H (1966) ”Low Temperature Soil Mechanics," Technical Note 143,
UA CRREL, Hanover, N. H.

Sayles, F.H. (1968) "Creep of Frozen Sand," Technical Report 190, USA
CRREL, Hanover, N.H.

Sayles, F.H. (1973) “Triaxial and Creep Tests on Frozen Ottawa Sand,"
Proc. 2nd ICP, North American Contribution, Yakutsk, U.S.S.R.,
pp. 384-391.

Sayles, F.H. (1974) "Triaxial Constant Strain Rate Tests and Triaxial
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Hanover, N.H.

Sayles, F.H.; and Epanchin, N.V. (1966) "Rate of Strain Compression Tests
on Frozen Ottawa Sand and Ice,” Technical Note 158, USA CRREL,
Hanover, N.H.

Sego, D. (1980) "Deformation of Ice Under Low Stresses,“ Ph.D. Thesis,
University of Alberta, Edmonton, Alberta.

and Banasch, R. (1982) "Strength and Deformation

Sego, D. C. Shultz, T;
3rd ISGF, Hanover, N.H.,

Behavior of Frozen Saline Sand," Proc.
pp. 11-17.

Smith, L.L.; and Cheatham, J.B. (1975) "Plasticity of Ice and Sand Ice
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Soc, 5. (1983) _
Structures," Ph.D. thesis (in preparation),

sity, East Lansing, Michigan.

”Experimental and Numerical Analysis of Frozen Sand
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"...—.- .fl-x: _- . ..r_

280

TSytovich, N.A. (1975) "Mechanics of Frozen Ground," McGraw-Hill Book
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APPENDICES

 

281

APPENDIX - Data:

A. Constant Displacement Rate Tests.
B. Constant Stress (Creep) Tests.

\

 

 

 

 

282

Table A—1: Ice Samples I—l to I—21 (Ice 1)

Ice (1) samples were formed from distilled water in four layers (6 in.
sample height). Average density equals 56.1 lb./cu.ft. A 5/8 in. diam-
eter plain steel rod was used with samples I—l to I—21.

 

 

 

Sample I—l: Sample I-3:
T = —1o.1°c T = —15.1°c
H = 5.866 1n H = 5.688 in.
D - 5 677 1n D = 5.875 in.
L = 5 685 In L = 5.719 in.
P = 140 lbs./min. 13 = 130.0 lbs./min.
5n= 4.24x10—3 in./min. 6n= 4.40x10_3 in./min.
t (sec.) P (lbs.) 5 (XIO—Bin.) t (sec.) P (lbs.) 6 (xlO_3in.)
225 430 0.24 60 52 0.04
295 630 0.25 120 187 0.07
340 760 0.50 180 330 0.11
360 820 0.60 255 557 0.77
400 930 0.90 280 607 1.50
400 360 10.00 286 370 5.91
406 310 11.00 292 310 7.75
412 290 12.00 300 270 8.85
418 270 12.75 315 250 10.16
436 250 14.00 375 230 14.75
460 \ 240 15.50 435 218 19.31
475 230 16.50 495 210 23.72
670 200 30.00 615 200 32.32
895 190 46.50
Sample I—2: Sample I—4:
T = -15.o°c T = —15.1°c
H = 5.750 in. H = 5.438 in.
D = 5.875 in. D — 5 938 in.
L = 5.625 in. L = 5.375 in.
1>= 136.4 1bS,/min. i>= 105.4 lbs./min.
. _ . -3 . .
6n: 4.50x10 3 in./min. 6n= 4.90x10 1n./m1n.
_ -3.
t (sec.) P (lbs.) 5 (x10 31...) t (sec.) P (lbs.) 5 (x10 1n.)
2 40 0.010 220 390 0.37
20 85 0.025 259 455 1.10
65 150 0.080 259 220 9.20
110 200 0.135 265 170 11.26
170 360 0.210 280 150 13.32
230 510 0.282 295 135 15.10
272 620 0.760 310 126 16.46
272 220 7.640 370 110 21.72
278 185 9.110 430 100 26.65
290 165 10.580 505 96 32-53
305 152 11'900 Note: At t=0, P=O and 6=O, unless
350 130 15.440 —— noted.
395 120 19.120
575 105 32.350

 

 

 

283

Table A—1: (Cont'd)

Sample I-S:

(Leakage, data discarded)

 

 

 

 

Sample I—6: Sample I—8:
T = —6.1°C 1 = -2o.o°c
H = 5.938 in. H = 6.125 in.
D = 5.969 in. D = 5.938 in.
L = 5.969 in. L = 6.125 in.
P = 89.6 lbs./min. P = 115.7 lbs./min.
6n= 3.68xlo"3 in./min. én= 3.60x10_3 in./min.
t(sec.) P(lbs.) a (x10'31n.) c (sec.) P (lbs.) 6 (x10'31n.)
6O 90 0.037 100 200 0.04
75 112 0.404 112 216 0.28
81 95 1.03 118 196 1.08
87 82 1.765 124 172 2.01
93 77 2.500 130 160 2.60
135 67 5.300 160 144 4.80
195 63 9.120 220 140 8.80
255 60 12.870 380 135 18.00
315 58 16.690
435 56 23.900 Sample I—9:
0
T = —10.0 c
§EEE£E_£:ZL H = 5.875 in.
T = -20.0°c D = 5.938 in.
H = 5.938 in. L = 5.834 in.
D = 5:900 in' P = 111.8 lbs./min.
L = 5.812 in. . _3
' , 5n= 3.55x10 in./min.
P = 136.4 lbs./m1n. 3
8n= 3.905.1073 in./min. t (sec.) P (lbs.) 6 (x10 in.)
—3. 75 112 0.01
t (sec.) P(lbs.) 6 (X10 in.) 120 216 0.20
220 500 0.001 175 326 0-40
220 278 5.000 175 100 6-00
235 253 6.786 240 90 10.60
250 240 8.214 330 84 16.80
280 222 11.785 450 82 23.00
325 207 15.000
385 195 19.285 Eeeele_1:192
590 175 32.142 T = _10.10C
H = 5.969 in.
D = 5.938 in.
L = 5.938 in.

(Continued)

 

 

 

284

Table A-1: (Cont'd)

 

 

 

 

 

 

Sample I-lO: (Cont'd) Sample 1-13:
P = 91.64 le./min. T = -26.7OC
6n= 3.32;;10'3 in./min. H = 6-0 in-
D = 6.0 in.
t (sec.) P (lbs.) 6 (X10_3;'Ln.) L = 6.0 in.
90 134 0.01 P = 143.2 lbs./min.
165 240 0.12 6n= 4.02.10'3 in./min.
220 336 0.20
220 120 5.60 t (sec.) P (lbs.) 8 (xlO-3in.)
250 112 8.80 ’
310 100 13.60 90 170 0-01
415 94 19 40 430 1040 0.08
' 620 1480 0.24
620 190 40.00
S l I—ll:
-5EEL51——7;—- 640 220 42.14
T = —10.1 c 645 220 42.54
H = 5.938 in. 730 195 43.10
D = 5.938 in. 840 200 45.00
L = 5.938 in. 940 180 65.00
p = 94.0 lbsé/min. Sample I-l4:
6n= 3.43x10 1n./m1n. T = —26.4OC
—3. H = 5.906 in.
£;£§EELL P (lbs.) 6 (X10 in.) D = 6.000 in.
250 ‘ 370 0.01 L = 5.910 in.
270 424 0.08 - _ .
300 470 0.20 P — 4.9 lbsg/min.
300 145 6.00 6n= 3.5x10 in./min.
320 105 10.00 _3
350 90 12.00 t (min.) P (lbs.) 5 (x10 in.)
380 85 14.88 0.0 0 0.0
455 78 18' 84.8 2140 3.6
84.8 320 38.6
Sample I—12: 90.0 440 42.0
T = —26.1°c 90.5 330 43.0
H = 6.0 in. 96.0 420 44.0
D = 6.0 in. 96.5 325 45.4
L = 6.0 in. 101.7 400 46.2
' . 102.2 320 47.4
P = 126 lbs./m1n. 106.5 390 48.2
6n= 0.9x10_3 in./min. 107.7 320 49.0
3 111.4 375 49.8
t (sec.) P (lbs.) 6 (x10 in.) 113.0 320 50.4
116.0 355 51.0
110 165 0-08 118.4 320 51.8
400 840 0.20
400 140 5.60

580 120 8.20

 

 

 

285

Table A—1: (Cont'd)

Sample I—15:

—26.6°c
5.938 in.
6.0 in.
5.938 in.

22.5 lbs./min.

T
H
D:
L
P

t (min.)

18
32
44
44
70

8.0x10'4

340
680
990
160
140

Sample I-16:

in./min.

P (lbs.) 8 (XIO—Bin.)

0.40
0.45
0.80
12.80
33.60

t (sec.) P (lbs.)

Sample I—l7: (Cont'd)

6 (x10_3in.)

 

 

1 = —20.4°c

H = 5.938 in.

D = 6.0 in.

L = 5 938 in.

P = 27 lbs./min.

6n= 1.053.10—3

t (min.) P (lbs.) 6 (x10‘31n.)
10.0 320 0.8
25.0 680 1.2
38.5 1040 1.4
38.5 100 20.0
48.0 90 30.0

Sample I—17:

 

T = -15.5°c
H = 6.344 in.
D = 6.0 in.
L = 6.344 in.
P = 23.60 lbs./min.
6n= 3.52:.10‘4 in./min.
—3.
t (min.) P (lbs.) 6 (X10 1n.)
10.0 180 1.40
17.5 320 1.60
31.5 665 1.62
44.7 1010 1.68
57.8 1350 1.90
71.0 1710 2.22
84.2 2060 2.62

(Continued)

97.7 2380 3.34
102.2 2480 3.48
119.8 1830 5.70
119.8 200 18.00
125.2 300 18.80
127.3 200 20.80
131.5 250 21.60
134.5 200 23.20
Sample I—18:

T = —10.1°c

H = 5.906 in.

D = 6.0 in.

L = 5.906 in.

P = 28.5 lbs./min.

6n: 2.3x10‘4

t (min.) P (lbs.) 6 (x10_3in.)

7.5 360 0.40
24.5 750 0.42
31.2 890 0.75
31.2 80 8.80
40.3 80 10.80
Sample I—19:

T = -6.6°c

H = 5.938 in.

D = 6.0 in.

L = 5.969 in.

P = 22.1 lbs./min.
3n: **

t (min.) P (lbs.) 6 (x10'3in.)

43.
62.
80.
82
84.
85.
85.

**

O
O
.0
O
5
5

0

1030
1470
1890
1920
1940
1890

300

OOJ-‘WNl—IOO
DDOOON-L‘N

p—a

Not recorded due to lack
of recording paper.

 

 

 

 

Table A—1: (Cont'd)

Sample I-ZO:

T = —10.1°c
H = 5.0 in.
D = 6.0 in.
L = 5.0 in.
P = 26.4 lbs./min.

0n= 2.6x10-4 in./min.

t (min.) P (lbs.) 6 (x10'3in.)

60.0 1300 0.4
76.5 1720 1.2
79.0 1760 1.6
81.0 1780 2.8
83.0 60 19.0
100.0 245 20.6
108.0 200 26.0

Sample I—21:

T = —20.1°c

H = 6.0 in.

D = 6.0 in.

L = 6.06 in.

P = 23.0\1bs./min.

Sn: 6.15x10'4 in./min.

t (min.) P (lbs.) 6 (x10'3in.)

43.0 810 0.4
75. 1740 0.8
75. 100 17.
84. 200 17.

5

5 8
0 8
95.8 450 18.6
96.0 220 22.4
100.0 260 23.2
104.0 220 25.0
112.0 210 28.6

 

 

 

287

Table A-2: Ice Samples I—22 to I—27 (Ice 2)

Ice (2) samples were formed from crushed ice (tap water) and saturated
with distilled water. Average density equals 56.1 lb./cu.ft. A 3/8 in.
diameter plain steel rod was used with samples I—22 to I-27.

Sample I—22:

 

 

 

Sample I-25:

 

 

 

T = -10.OOC T = _15'000
H = 6.0 in. H = 6.125 in.
D = 6.0 In. D = 6.00 in.
L = 6.0 in. L = 6.125 in.
P = 21.6 lbs./min. P = 20 lbs./min.
6n= 5.7x10—4 in./min. 6n: 5,4x10_4 in./min.
. -3. —
t (mln.) P (lbs.) 6 (x10 1n.) t (min.) P (lbs.) 6 (x10 3:6.)
25 540 0-2 33.5 670 0.1
25 0 11-6 33.5 64 14.0
29 40 12.4 50.0 48 23.0
34 30 15.2 70.0 40 34.0
60 30 30.0 104.0 40 52.0
183 30 99.0
Sample I—26:
Sample I—23: T = —6.OOC
T = -22.0°c H = 6.0 in.
H = 6.5 in. D = 6.0 in.
D = 6.0 in. L = 6.0 in.
L = 6'5 in- P = 16.1 lbs./min.
P = 20.1 lbszl/min. (.311: 3.65X10—4 in./min.
5n= 5.64x10 in./min. _3 t (min.) P (lbs.) 5 (x10—3in.)
t (min.) P (lbs.) 5 (x10 in.) 13 210 0_1
31 624 0.4 13 48 3.5
31 16 13.6 14 32 5.5
34 72 14.0 16 16 8.0
74 50 36.8
M
Sample I—24: T = —2.OOC
T = —20.OOC H = 6.0 in.
H = 6 375 in. D = 6-0 }n'
D = 6.00 in. L = 6'0 1n'
L = 6.375 in. f = 19.1 lbs./min.
P = 17.2 lbs./min. 0n= 5,8){10—4 in./min.
. _ , -3.
6n= 5.1x10 4 in./m1n. 3 t (min.) P (lbs.) 6 (x10 in.)
t (min.) P (lbs.) 6 (x10 in-) 5 5 ,105 0.1
32.5 560 0.4 5.5 20 4.0
7 0 10 6.0
32.5 16 12.4
30 0 10 19.3
35.0 50 13.0
44.0 32 18.4 Note: At t=0, P=O and 6:0, unless
76.0 32 34°6 _’_—— noted.

 

 

 

288

Table A—3: Ice Samples I—28 to I—32 (Ice 3)

Ice (3) samples were formed from snow—ice saturated with distilled
water. Average density equals 56.1 lb./cu.ft. A 5/8 in. diameter steel
rod was used with samples I-28 to I—32.

Sample I-28:

T = -10.0°c

H = 6.197 in.

D = 6.00 in.

L = 6.197 in.

P = 183 5 lbs./min.

8n= 4.5x10-3 in./min.
6 (min.) P (lbs.) 6 (x10—Sin.)

 

1.0 150 0.01
2.0 300 0.02
3.0 460 0.03
4.0 640 0.04
5.0 840 0.05
6.0 1060 0.06
7.0 1300 0.25
8.0 1500 0.40
8.5 1560 0.75
8.5 100 15.00
10.0 70 15.00
10.2 70 18.00
11.0 ‘ 70 21.00
13.0 70 28.75
15.0 70 36.25
17.0 70 45.60

Sample I—29:

 

T = -14.6°c
H = 6.0 in.
D = 6.0 in.
L = 6.0 in.
P = 233.3 lbs./min.
6n= 1.922(10_3 in./min.
1 (min.) P (lbs.) 6 (x10-3in.)
6.15 1435 0.5
6.15 0 20.0
7.25 140 20.5
7.50 105 24.0
8.25 100 27.0
9.00 100 31.0
12.00 70 37.0
15.00 40 43.0
18.00 30 48.8
Note: At t=0, P=0 and 6=O, unleSS
noted.

Sample I-30:

T = —16.0 C

H = 6 375 1n
D = 6 00 1n
L = 6 375 in

P = 280 lbs./min.
0n= 4.58x10—3 in./min.
t (min.) P (lbs.) 6 (x10-Bin.)

2.75 880 0.25
2.75 70 13.00
3.00 70 13.20
4.00 70 18.50
5.00 60 24.75
8.00 60 41.60
11.00 50 56.60
14.00 50 70.35

§§TPle_1:211

T = —6.0°c

H = 6.375 in.

D = 6.00 in.

L = 6.375 in.

P = 1290 lbs./min.

0n= **

t (sec.) P (lbs.) 6 (x10—Bin.)

20 430 ** Rupture at

fast rate

§29212_1:§§1

= —10.0°c
— 6.5 in.
= 6.0 in.
= 6.5 in.
P = 258 lbs./min.

6n: 4.615x10'3 in./min.
t (min.) P (lbs.) 6 (x10—316.)

HUSH—3
|

4.00 1030 1.0
4.00 0 24.0
4.75 105 25.0
5.50 70 29.0
7.00 70 37.2
11.00 70 60-0
24.00 70 120.0

 

 

 

289

Table A-4: Frozen sand samples.

Sample S—l: Sample S—3:

T = ~10.0°c T = -6.3°c

D = 6.0 in.* D = 6.0 in.

H = 6.0 in.* H = 6.0 in.

L = 6.0 in.* L = 6.0 in.

vs= 64.0 2 * vS= 64.0 x

d = 5/8 in.** d = 5/8 in.

h = 0 (plain rod) h = 0

P = 195.5 lbs./min. P = 29.7 lbs./min.

6n= 2.0x10'3 in./min. 6n= 6.967x10'4'in./min.

t (min.) P (lbs.) 6 (x10-3in.) t (min.) P (lbs.) 6 (x10—316.)

 

 

 

 

 

3.34*** 640 0.16 20.5 640 0.8
4.58 890 0.22 29.0 860 1.6
5.50 1075 0.27 31.0 830 3.0
5.50 236 8.27 34.0 700 6.2
5.67 190 9.10 39.0 570 10.6
5.84 175 9.50 39.0 500 14.8
6.34 160 10.70 65.0 400 26.0
7.34 145 13.10
8.34 135 15.50 Sample S-4:
0
* Average values, unless noted. ; = 610 in?
** Coidgrolled steel, unless H = 6.0 in.
no e . _ .
*** at t=0, P=O and 6=0, unless Vs: ZA?01;'
noted.
d 5/8 in
Sample S—2: h 0
T = —10.2°c P = 30.7 lbs./min.
D = 6.0 in. 6n: 3 67X10—4 in./min.
H = 6.0 in
L = 6-0 in- t (min.) P (lbs.) 6 (x10—Bin.)
vs: 64.0 2
57 1790 0.8
d = 5/8 in- 65 2060 1.5
h = 0 68 2090 2.5
' = - 69 200 50.6
P 32.5 lbsg/min. 78 560 51.0
6n= 5.0x10 in./min. 80 590 51.6
_3 84 550 54.4
t (min.) P (lbs.) 6 (X10 in.) 87 530 56.0
21 700 0.8 99 520 60.0
29 980 1.0
38 1240 1.4
40 1300 2.2
41 1280 3.0
41 240 18.2
46 280 19.6
56 280 24.6

 

 

 

290

Table A-4: (Cont'd)

 

Sample S—5: Sample S—6: (C0nt'd)
o
T = ‘20-1 C t (min.) P (lbs.) 6 (x10—3in.)
H = 5.871 in.
D = 6.00 in. 91.0 800 34.5
L = 5.781 in. 91.3 500 38.5
Vs: 66.3 Z 97.0 680 39.3
99.5 570 41.7
d = 5/8 in. 104.0 610 43.3
h = 0 107.0 600 44.7

P = 32.7 lb . ' .
3 /mln Sample S—7:

 

 

 

6n 5.0x10 1n./m1n. T = -26.8OC
. —3. H = 6.0 in
t (mln.) P (lbs.) 6 (x10 in.) D = 6.0 in

28.0 680 0.20 L = 6.0 in.

47.0 1360 0.30 vs= 64.0 2

65.0 2080 0.35

82.0 2730 0.55 g ; 3/8 in

85.0 2810 1.00 .

85.6 2810 1.80 P = 39.7 lbs.£min.

85.6 0 42.90 ' — . .
114.0 860 43.50 Gn— 6.154x10 1n./min.
ii; 5 2:8 23:33 t (min.) P (lbs.) 6 (x10 316.)
124.0 560 50.00 20.50 720 0.40
127.0 ~ 470 52.40 36.75 1400 0.70
132.0 500 53.40 53.00 2080 0.76
135.0 575 54.20 70.50 2800 0.80
140.0 575 56.6QI 70.5 90 31.50

90.50 810 31.70
Sample S—6: 91.00 440 36.00
0 95.50 560 37.20
T = '26'? C 96.50 430 39.20
H = 6.01m. 100.50 510 40.40
D = 6-0 }n- 102.50 440 42.00
L = 6.0 1n- 106.00 480 43.20
Vs= 64-0 2 109.00 450 44.80
d = 5/8 in. 112.50 470 46.20
h = 0 116.00 450 47.60
P = 37.5 lbsL/min. Sample S—8:
6n: 6.38x10‘ in./min. T = —27.1°c
-3. H = 6.0 in.
t (min.) P (lbs.) (S (X10 1n-) D = 6.0 in.

20.0 680 0.6 L = 6.0 in-

37.5 1360 0.8 Vs= 64-0 Z
55.0 2040 0.9 d = 5/8 in

73.0 2740 1.9 h = 0

73.0 110 33.7 . ,

(Continued) P = 20.8.2 lbs./m1n.

(Continued)

 

 

 

291
Table A—4: (Cont'd)

 

 

 

 

 

Sample S—8: (Cont'd) Sample S—lO: (COnt'd)
' -3 . . —
0n= 1.32x10 1n./m1n. t (min.) P (lbs.) (3 (x10 3in.)
. —3. 6.0 1180 0.2
t m1 . P lb . 1 .
4;) 48“) Leg; 7.5 1400 1.8
7.00 960 0.01 8.0 220 11.4
15.50 3150 0.40 10.5 200 14.1
18.25 3800 1.00 13.0 200 17.0
18.50 10 50.00
25.00 1240 50.00 Sample s-11;
25.50 1320 50.40 _ 0
25.75 920 54.00 g : g13.1 C
26.50 950 55.40 D ‘ 6'0 :H'
1. . ‘ . I1.
3 00 920 60 00 L = 6.0 in.
Sample S—9: VS: 64'0 Z
T = -20.00C d = 5/8 1n., h = 0
H = 5.938 in. P = 204 6 lbs./min.
E : 2:88 :2: 5n= 3.86x10_3 in./min.
Vs= 64-6 z t (min.) P (165.) 6 (X10—3in.)
d f 5/8 in° 7.50 1440 0.8
h — 0 13.00 2660 1.2
P = 210.3 lbs./min. 13 00 220 29-0
. . _3 . 14.75 610 29.8
(Sn: 2.53X10 in./m1n. 15.25 450 33.6
— 19.00 430 46.2
t (min.) P (lbs.) 6 (X10 3in.)
11.00 2160 0.8 §EBB£E_§:£ZL
15 50 3260 1.2 T = —10.0°c
15.50 240 35.6 H = 6_0 in.
18 50 960 36.4 D = 6.0 in_
18.75 680 40.4 L = 6.0 in_
28.00 580 67.2 vs= 64_0 %
Sample S—lO: d = 5/8 in.
--—-‘-7;-- h = 0
T = _603 C ‘ / .
H = 6.0 in. P = 222.7 lb:. min.
D = 6-0 in- 6n= 3.73x10_ in./min.
L = 6.0 in. _3
VS= 64.0 Z t (min.) P (lbs.) 6 (x10 in.)
d = 5/8 in. 6,5 1490 0.4
h = 0 11.0 2450 2.6
' = . 11.0 590 20.0
P 186.7 16:./mln. 11,5 480 21.6
6n= 1.12x10 in./min. 14,5 470 32.8

(Continued)

 

 

 

 

292

Table A—4: (Cont'd)

 

Sample S—13: Sample S-15: (Cont'd)

" ' O“ ' T*—_“_—‘_‘—_'

T = -6.0 C P = 1978.6 lbs./min.

H = 6.0 in. ' —

D = 6.0 in. 6n= 2.13x10 2 in./min.

L = 6.0 in. , _

VS: 64.0 2 t (616.) P (lbs.) 6 (x10 316.)
= . 1 4 2770 0.4

g = 3/8 1n' 1 4 430 30.7

, 1.5 780 31.6

P = 2133.3 lbs./min. 1 6 550 33.6

5n= 1.964x10_2 in./min. i g :28 23'?

t (min.) P (lbs.) 6 (x10'3in.)

 

Sample S—16:

 

 

0.90 1920 1.4 _ _ o
0.90 490 23.5 g = 623.:nC
1.00 440 24.5 D = 6.0 in.
1.05 360 30.0 L = 6.0 in.
2.40 330 52.0 V = 64 0 7'
S O 0
Sample S—14: d = 5/8 in
o h =
T = -15.2 C .
H = 5.969 in. P = 2210 lbs./min.
D = 6'00 in. 6n= 4.0x10_2 in./min.
L = 6.00. in.
vs= 64-2 Z t (min.) P (lbs.) 6 (x10—316.)
d = 5/8 in- 1.60 3680 0.4
h = 0 2.00 4420 0.7
P = 2033.3 lbs./min. 2-00 800 16-4
° _2 2.40 1390 16.4
5n= 3.0x10 in./min. 2.55 1630 17.8
_3_ 2.60 1280 23.8
t (min.) P (lbs.) (X10 ln-) 3.80 900 72.0
1.6 3280 0.1
1.8 3660 0.6 Sample S—17:
1.8 480 40.0 T = —20.0°c
2.0 850 40.8 H = 6.0 in.
2.0 600 45.0 D = 6.0 in.
3.0 530 75.0 L = 6.0 in.
Vs: 64.0 %
Sample S—15: .
--——'--7;- d = 5/8 in.
T = -10.0 C h = 0
H = 6.0 in. ° .
D = 6.0 in. F = 2240 lbfé/mln'
L = 6.0 in. 6n= 1.32x10 in./min.
VS: 64‘0 A (Continued)
d = 5/8 in.

h
(Continued)

 

 

 

Tabl

e A-4:

(Cont'd)

 

 

293

gample S—17: (Cont'd)

 

 

—_..-. -«._... .— r > ..—..

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Sample S-19: (Cont'd)

t (min.) P (lbs.) 0 (x10-316.)
360 1630 1.80
375 1580 3.60
440 1200 14.40
Sample S—20:

T = -6.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

VS= 64.0 %

d = 5/8 in.

h = 0

P = 4.4 le./min.

6n= 1.0x10-4 in./min.

t (min.) P (lbs.) 6 (x10-3in.)
145 640 0.1
165 720 0.5
210 640 4.8
280 500 13.6
375 380 23.2
Sample S-21:

T = -15.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

Vs: 64 .0 c:1)

d = 5/8 in

h = 0

P = 6.5 lbs./min.

6n= 1.86.110"4

t (min.) P (lbs.) 6 (x10'3in.)
165 1020 0.25
260 1860 0.50
320 2080 1.50
330 2000 5.50
332 720 15.50
335 540 21.80
370 620 23.80
470 580 31.50

t (min.) P (lbs.) (x10-316.)

1.50 3360 0.2

1.50 160 14.0

1.85 1020 14.8

1.90 860 17.5

3.00 750 30.0

Sample S-18:

T = -25.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

Vs: 64.0 Z

d = 5/8 in.

h = 0

P = 5.06 lbs./min.

6n= 1.1..10‘4 in./min.

t (min.) P (lbs.) 6 (x10-3in.)

340 1600 0.3

495 2400 0.9

600 \2980 0.9

680 3440 2.9

680 1080 43.0

838 2460 43.0

860 2520 45.0

865 1100 48.6

995 1080 60.0

Sample S-19:

T = —10.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

vs= 64.0 x

d = 5/8 in.

h = O

P = 4.53 lbs./min.

6n= 1.661410.4 in./min.
_3.

t (min.) P (lbs.) 6 (x10 in.)

172 830 0.20

285 1350 0.45

355 1630 1.30

(Continued)

 

 

 

Table A—4:

 

Sample S—2

—20.0O

{II
II II

(Co

2:
C

6.0 in.

D = 6.0 in.
L = 6.0 in.

64.0 Z

<.‘
U)
ll

’11- 3‘
II II II

66= 1.34x1

6 (min.) P (163.) 6 (x10—316.)

410.
630.
647
647
650.
710.

OOP—‘OOO

5/8 in.
0

0-4

2550
2770
2650
1200
1000
1000

3:

Sample S—2

T = —10.00
H:

C

6.0 in.

D = 6.0 in.
L = 6.0 in.

vS 64.0 2

d
h

5/8 in.
0

nt'd)

4.4 lbs./min.

in./min.

0.
1.

4.
21.

24.
29.

P = 268.1 lbs./min.

6n= 3.83x1

t (min.) P (165.) 6 (x10—316.)

0-3

2400
2480
240
660
660
560
540

4:

8.80

9.25

9,25

11.00
11.30

11.75

16.00
Sample S—2
T = -2.0°c
H = 6.0 in.
D = 6.0 in.

= 6.0 in.

L
vS= 64.0 Z
d

= 5/8 in.
(Continued)

in./min.

O.

1.
30.
31.
33.
35.
51.

4

4
2
0
6
4

1

5
O
O
O
0
0

Sample S—24: (

h
P

t (min.)

.50

00\IO\U1-L\-L\LON

25
00
50
25

.25

25
75

O

Cont'd)

100 lbs./min.
6n= 4.0x10'3 in./min.

 

P (lbs.)

300 0.1
390 0.8
450 2.4
450 4.1
410 7.5
330 12.3
280 16.7
240 22.7

Sample S—25:

6n

—2.0°c

6.0 in.
6.0 in.
6.0 in.

64.0 Z

5/8 in.

1714.3 lbs./min.

3.71x1

0-2

in./min.

6 (x10—3in.)

t (min.) P (lbs.) 6 (x10—316.)

O.
O.
0.
0.
0.
0.

25
35
42
45
55
90

520
640
720
640
370
320

Sample S—26:

T
H
D
L
Vs

P
6n

d
h =

ll 1! II

II

-2.0°c

6.0 in.
6.0 in.
6.0 in.

64.0 2

5/8 in.

0

0.5
2.9
6.1
10.3
15.0
28.0

2.1 lbs./min.

8.6x10

-5

(Continued)

in./min.

 

 

 

Table A-4: (Cont'd)

Sample S—26:

t (min.) P (lbs.)

35
65
100
150
250

150
190
210
190
190

Sample S—27:

T:

L'"
I

n:

-2.0°c
6.0 in.

= 6.0 in.
_ 6.0 in.
= 64.0 2

5/8 in.
0

26.3 lbs./min.

295

(Cont'd)

6 (x10-3in.)

....
oxoowv—‘o
O‘DO‘DN

6.0X10—4 in./min.

t (min.) P (lbs.) 8 (x10'316.)

O).
:1
H

Sample S—29:

T = -15.0°c

H =6.0 in.

D = 6.0 in.

L = 6.0 in.

V3= 64.0 %

d = 5/8 in.

h = 0

P = 220.9 lbs./min.

**

t (min.) P (lbs.) 6 (x10—3in.)

 

 

 

13.5
17.5
21.0
32.0

50.0

Samp

420
\ 460
448
380
350

le S—28:

T:

:1:
ll

-20.o°c
6.0 in.
6.0 in.
6.0 in.
64.0 Z

= 5/8 in.
0

2478.3 lbs./min.

3.9x10"2 in./min.

(min.) P (lbs.)

3420
320
1060
840
770

6 (x10-3in.)

0.01
39.00
41.00
46.00
74.50

11.5 2540 **
11.5 400

13.0 760

15.0 660

20.0 660

** Not recorded, channel was OFF.
Sample S—30:

T = -10.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

vs= 64.0 Z

d = 3/8 in.

h = 0

P = 29.3 1bs./min.

6n= 8.0x10_4 in./min.

t (min.) P (163.) 6 (x10—3in.)

30.
30.
31.
33.
39.
94.

T
H
D
L
Vs:

ll

OOOOU‘IO

880 0.8
240 7.4
200 8.8
160 11.2
140 16.0
140 60.0

§3921£_§:§11

—10.0°c
6.0 in.
6.0 in.
6.0 in.
64.0 %

(Continued)

 

 

 

 

 

 

Table A—4: (Cont'd)

 

 

 

 

 

 

 

Sample S—31: (Cont'd) Sample S—33: (Cont'd)
d = 3/8 in- t (min.) P (lbs.) 6 (x10-3in.)
h = 1/16 in.*
. 68.5 2190 2.0
P = 29.7 lbs./min. 72,0 1870 6.0
6n= 5.23310"4 in./min. 76'0 1840 9'2
87.0 1830 15.6
. -3
t m1 . P lb . ' .
( n ) ( S ) 6 (X10 1n ) Sample S-34:
52.5 1560 1.2 o
53.0 1320 5.2 T i E10.0 C
54.0 1040 10.4 g : 6'3 :2'
66. 1 . _ '
0 040 16 6 L = 6.0 in
* A 9OO—lug, unless noted. VS: 64'0 °
d = 3/8 in.
Sample S—32: h = 3/16 in.
T = —10.0°c P = 31.3 lbs./min.
H = 6.0 in. ' _ -4 . .
D = 6.0 in. 6n— 4.212x10 in./m1n.
:3: 24301;. t (min.) P (163.) 6 (x10—316.)
. 74 2320 2.8
g : :4: PD' 85 2330 8.6
. “ 1n' 127 2800 26.0
P = 33.4 lbs./min. 177 3480 45.0
' - 357 4800 132.0
6n= 5.34X10 4 in./min. 360 4440 136-0
-3. 450 4800 160.0
t ( min.) P (lbs.) 6 (x10 in.) 640 4800 240.0
61 2040 2.0
62 1800 5.6 M
64 1580 9.2 o
T = —20.0 C
72 1490 14.8 H = 6.0 in.
78 1520 18.0 D = 6.0 in.
L = 6.0 in.
§2E212_§:§§: vs: 64.0 8
O
T = -10 0 c d = 3/3 in,
H = 6.0 in. 1’] = 0
D 6.0 in. .
L = 6.0 in. P = 138.7 leA/mln.
Vs= 64-0 Z 6n= 6.944x10_ in./min.
d = 3/8 in. —3,
h = 3/16 in. t (min.) P (lbs.) 6 (x10 in.)
P = 32 16 . in. 7.5 1040 1.0
~ 3 :2 8.0 420 9.0
6n= 5.82x10 in./min. 26.0 300 21,5
(Continued)

 

 

 

Table A-4: (Cont'd)

 

Sample S—36:

T
H
D
L

Vs:

II II

{TD-1
I! ll

. PU.
ll

6n=

t (min.) P (lbs ) 6 (x10-316.)

—20.00

C

6.0 in.
6.0 in.
6.0 in.

64.0 2

3/8 in.

1/16 1

32.8 lbs./min.
5.1x10 4

n.

 

 

in./min.

T:
H:
D—
L

07‘ "U‘ D‘D—
" H II

n:

t (min.) P (lbs.) 6 (x10‘3in.)

—20.0O

Sample S-38:

C

6.0 in.
6.0 in.
6.0 in.

— 64.0 Z

3/8 in.
3.16 in.

35.8 lbs./min.

4.1x10

—4

in./min.

 

95
103
145
185
225
265
305
345
385
425
465
505
590

3400
3360
3900
4441
4921
5382
5722
6000
6280
6480
6620
6740
7020*

0.
6.
18.
36.

46
70
87
104

118.
135.
152.
167.
200.

OOO‘DUDOU‘IU‘INV-‘bba

 

 

55.5 1820 0.01
59.0 1650 4.60
65.0 1610 8.70
105.0 2000 26.00
145.0 3380 45.00
185.0 2440 65.00
205.0 2500 76.00
240.0 2600 94.40
275.0 2680 113.40
310.0 2660 130.40
Sample S-36:
T = -20.0°c
H = 6.0 in.
D = 6.0 in.
L = 6.0 in.
V3= 64.0 %
d = 3/8 in.
h = 1/8 in.
P = 37.5 lbs./min.
6n= 3.9x10'4 in./min.
_3-
t (min.) P (lbs.) 6 (x10 in.)
63 2360 1.6
69 2330 5.7
110 2840 22.4
150 3340 37.4
170 3580 45.0

* This load caused rod yielding

 

Sample S—39:

T = —10.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

vs= 64.0 2

d = 3/8 in.

h = 0 p

P = 27.1 lbs./min.

6n= 5.0x10—4 in./min.

t (min.) P (lbs.) 6 (x10—316.)
4.5 130 0.8
6.0 64 1.6
9.0 45 3.2

17.0 40 7.2

 

 

 

 

 

Table A—4: (Cont'd)

Sample S-40:

T = —10.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

Vs= 64.0 Z

d = 3/32 in.

h = 0

P = 5.3 lbs./min.

6n= 4.32x10"4 in./min

t (min.) P (163.) 6 (x10—316.)

'P:

Sample S-42:
7.4 lbs./min.

(Cont'd)

6n= 4.5}(10-4 in./min.

 

 

8 56
15 80
17 45
42 20 1

Sample S-4l:

—20.0°c
6.0 in.
= 6.0 in.

0')
ON
.5_\ .
O
N

II II

‘D‘Du <F‘UIIIH
0‘1
O
H
:5

15.8 lbs./min.
5.43x10'4 in./min

OVI'U
:5
II I

t (min.) P (163.) 6 (x10‘3in.)

33.5 530
33.5 90
34.0 70
37.0 65
45.0 50
48.0 50 1

Sample S—42:

T = —20.0°c
H = 6.0 in.
D = 6.0 in.
L = 6.0 in.
Vs: 64.0 Z
d = 3/32 in.
h = 0

(Continued)

0.
1.
1.
2.

0
2.
4.
5
8
1

O‘CDNOm-b

CON-b

t (min.) P (lbs.) 6 (x10—3in.)
20.5 152 0.2
22.0 100 2.6
29.0 80 6.0
33.0 80 7.8

Sample S-43:

 

T = -10.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

vs= 64.0 2

d = 3/8 in.

h = 1/8 in.

P = 2220 lbs./min.

5n= 2.1x10”2 in./min.

t (min.) P (163.) 6 (x10—316.)
1.0 2220 3.0
1.5 2760 13.0
2.0 3140 24.0
2.5 3540 34.0
4.0 4040 74.0
5.5 4320 97.0

(If IT'U {Iii-i
|| II II II II

On "U' 5‘0.-
II

-10.00

198.1 lbs./min.

Sample S—44:

C

3.36X10_3 in./min.

(Continued)

 

 

 

Tab

 

 

 

le A—4: (Cont'd)

Sample S—44: (Cont'd)

6 (min.) P (163 ) 5 (x10‘3in.)

1.
5.

10.
11
25.
36.
46.
62
67.
94.

50 2080
.75 2080
25 3220
50 3900
00 4220
.00 4540
00 4540
50 4540

Sample S—45:

0‘)

II II

“*U'D‘Qa <HUZEH

O?
:3
II

t (min.) P (163.) 6 (x10—316.)

235
280
300
340
390
470
520
550

-10.0°c
6.0 in.

= 6.0 in.
= 6.0 in.
= 64.0 Z

3/8 in.
1/8 in.

4.8 lbs./min.
7.0x10m5 in./min

1120
940
880
880
940
940

1000

1000

Sample S—46:

T =
H =
D =
L =
VS:
d =
h =

P:

66:

—15.0°c
6.0 in.
6.0 in.

30.6 lbs./min.
2.8x10’4

(Continued)

37
65

94.
150.
167.
260.

2

9.
11.
14.
18
24
27
28

in./min.

DOOOOOOO

OMOJ-‘OOGJO‘

t (min.)

68

72

83
123
163
203
225
330
380
760

P (lbs.)

2080
2000
2000
2470
2880

3200

3360

3720

3800

3800

Sample S—47:

Sample S-46: (Cont'd)

6 (x10' ' .

U'IH

11

N
00
0000000000)

44
62
73.
82.
94
200

 

T = —26.0°c

H = 6.0 in.

D - 6.0 in.

L = 6.0 in.

vs= 64.0 Z

d = 3/8 in.

h = 1/8 in.

P = 33 lbs./min.

6n= 5.02x10—4 in./min.

t (min.) P (lbs.) 6 (x10'3in.)
77 2540 2.0
100 2880 10.0
120 3200 16.0
140 3520 24.0
160 3760 32.0
180 4000 38,4

282 4980 82.0

382 5440 129.0

480 5600 185.0

630 5760 257.0

820 6080 352.0

T =
H =
D =
L =
VS=
d =
h =

P:

—6.0°c

§§TPls_§:£§1

6.0 in.
6.0 in.
6.0 in.

64.0 Z

3/8 in.
1/8 in.

32.4 lbs./min.

(Continued)

 

 

 

Table A—4:

 

 

(Cont'd)

 

 

(Continued)

Sample S—48: (Cont'd)

8n= 5.7x10-4 in./min.

t (min.) P (lbs.) 6 (X10—Bin.)
37 1200 2.6
65 1120 18.0

165 1680 68.0

270 1920 119.0

370 2080 176.0

382 2080 240 0
Sample S—49:

T = —2.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

Vs= 64.0 Z
d = 3/8 in.

h = 1/8 in.

P = 17 lbs./min.
6n = 3.283x10"4 in./min.
t (min.) P (lbs.) 6 (x10—3in.)
26 440 1.5
36 540 5.0
46 600 10.0
56 660 15.0
66 700 20.0
76 760 24.0
126 1020 47.0
240 1390 87.0
330 1390 114.0
592 1390 200.0
Sample S—SO;

T = -10.0°c

H = 5.0 in.

D = 6.0 in.

L = 5.0 in.

vs= 64.0 2

d = 3/8 in.

h = 0

P = 28.4 1bs./min.

6n= 5.63x10_4 in./min.

Sample S-SO: (Cont'd)

t (min.)

21.5
21.5
22.0
34.0
50.0

Sample S—51:

t (min.) P (lbs.) 6 (x10—3in.)

140
140
145
280
420

T
H
D =
L

{I'D-1

Ova V'Uq

n:

t (min.) P (lbs.) 6 (x10—Bin.)

148
148
205

-10.0°c
5.312 in.
6.00 in.

5.2 lbs./
1.14x10'4

730
220
160
120
120

Sample S-52:

—10.0°c
6.0 in.
6.0 in.
6.0 in.
64.0 X

= 3/8 in.
0

120
120

P (lbs.) 6 (x10—3in.)

610 0.
280 6.
180 7.
6.
5.

GOOD-b

No—n

min.

in./min.

0.01

7.00
10.00
26.0
42.0

6.3 lbs./min.

1.43.10“4

940
220
140

in./min.

0.01
12.00
20.00

 

 

 

 

 

 

 

 

 

(Continued

)

Table A—4: (Cont'd)
Sample S—53:

T = —10.0°c

H = 3.0 in.

D = 6.0 in.

L = 3.0 in.

vs= 64.0 %

d = 3/8 in.

h = 0

P = 24.5 lbs./min.

6n= 5.63X10—4 in./min.

t (min.) P (lbs.) 6 (x10-3in.)
22.0 540 1.0
22.5 510 2.0
22.5 170 6.0
23.0 140 8.0
39.0 130 17.0
Sample S—54:

T = — 0.000

H = 2.0 in.

D 6.0 in.

L = 2.0 in.

vs: 64.0 2

d = 3/8 in.

h = 0

P = 20.4 lbs./min.

6n= 6.5x10_4 in./min.

t (min.) P (lbs.) 6 (x10—316.)
11.0 224 0.1
11.5 144 1.8
14.0 96 4.5
30.0 64 14.5
40.0 64 21.0
Sample S-55:

T = —10.0°c

H = 1.0 in.

D = 6.0 in.

L = 1.0 in.

Vs: 64.0 Z

d = 3/8 in.

h = 0

P = 9.6 lbs./min.

 

Sample S—55: (Cont'd)

6n= 1.43x10'4 in./min.

t (min.) P (lbs.) 6 (x10—Bin.)
5 48 2.0
9 48 7.0

25 28 30.0

Sample S—56:

T = -10.0°c

H = 1.0 in.

D = 6.0 in.

L = 1.0 in.

Vs: 64.0 Z

d = 3/8 in.

h = O

P = 14.4 lbs./min.

dn= 3.0X10—4 in./min.

t (min.) P (lbs.) 6 (x10—Bin.)

10 144 0.16

12 96 0.80

18 94 2.60

20 64 3.2

Sample S—57:

T = —10.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in

vs= 64.0 2

d 3/8 1n

h = 1/8 1n

6 = 34.7 lbs./min.

6n= 5.433x10—4 in./min.

5.191231 P (lbs.) 6 (x10—3in.)
45 1560 0.2
48 1240 4.2
53 1060 10.0
58 1040 14.0
63 1040 16.5

105 1281 49.2

220 1640 89.6

300 2062 118.2

(Continued)

 

 

 

 

 

M (Cont'd)

 

 

 

 

Sample S—57: (Cont'd) Sample S—60:
_ _ o
t (min.) P (lbs.) 6 (x10 31...) T ‘ ‘15-? C
—"“‘—_‘_ ‘_“““—‘—‘_ H = 6.0 In.
350 2522 152.7 D = 6.0 in.
420 2800 191.6 L = 6.0 in.
565 2920 270.0 Vs: 64.0 Z
660 2920 322.0
d = 3/8 in.
Sample S—58: 5 = 0
T = “1.0-00C 1'3: 33.8 lbsll/min.
H = 1-0 in. 6n= 1.25X10_ in./min.
D = 6.0 in. 3
L = 1.0 in. t (min.) P (lbs.) 5 (x10 in.)
= 64. : ""“"‘ “"‘"“' "“““““‘
VS 0 / 22.5 760 0.1
d = 3/8 in. 23.0 280 6.0
h = 0 37.0 200 16.0
P = 10.2 lbs./min. 82'0 200 22'0
8n= 6.0x10—4 in./min. Sample S—61:
_ _ o
t (min.) P (lbs.) 6 (X10—3in.) : ;6 g6ig C
18.5 188 0.4 D — 6.0 in
22.0 160 1.2 L = 6.0 in
30.5 ~128 6.0 Vs: 64.0 Z
32.0 80 10.0 .
34.0 64 13.0 g = 3/8 1“
38.0 48 15.2 .
42.0 40 17.6 P = 24.1 le./min.
. -4
= . 16 10 ' . ' .
Sample S—59: 6n 6 3 x 1n /m1n
T = _lO.OOC t (min.) P (lbs.) 5 (x10-3in.)
H = 1.0 1n. 36.5 880 1.0
D = 6.0 ln. 37.0 560 5.0
L = 1.0 1?. 50.0 400 15.0
vs= 64.0 4 57.0 400 19.0
d = 3/8 in.
h = 0 Sample S-62:
. o
P = 13.7 lbs./min. T = -20.0 C
- _3 . H = 6.0 in.
6n= 6.8x10 in./m1n. D = 6.0 in.
—3, L = 6.0 in.
32.933211 P (1bS-) 5 (X10 ln') vs: 64.0 2
7 96 0.16 d = 3/8 in.
9 64 1.80 h = 1/8 in.
25 40 10.00 . .
30 40 14.00 P = 29.0 lbs./mln.

6n= 4.65x10‘4 in./min.
(Continued)

 

 

 

Table A-4: (Cont'd)

 

 

 

 

Sample S-62: (Cont'd)

t (min.) P (lbs.) 6 (x10—3in.)
56 1700 0.1
60 1740 2.0
65 1760 4.5

240 3840 73.0

290 4170 95.0

340 4300 118.8

410 4460 152.2

470 4580 178.0

603 4580 240.0

Sample S—63:

T = -2.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

VS= 64.0 Z

d = 3/8 in.

h = 0

P = 17.3 lbs./min.

6n= 2.25361074 in./min.

t (min.) P (lbs.) 6 (x10—316.)

15 260 0.6

16 260 1.0

17 160 1.8

18 100 2.8

20 80 3.6

34 50 7.2

Sample S-64:

T = —6.OOC

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

vs= 64.0 2

d = 3/8 in.

h = 0

P = 18.2 lbs./min.

6n= 9.5x10_5 in./min.

—3.

t (min.) P (lbs.) 6 (x10 1n.)

28.5 520 0.3

29.0 140 1.0

40.0 100 2.0

48.0 100 2.8

—15.0°

s= 64.0 2

5‘9.- <F'UCUH
I

Pd.
N

6.316x

O).
:3
II

t (min.)

39.
39
39
41.
47.
60.

OOOUVOO

T =-20.0°c

Sample S—65:

C

6.0 in.
6.0 in.
— 6.0 in.

3/8 in.
O

10‘

P (lbs.)

880
460
300
240
220
220

6:

Sample S—6

H = 6.0 in.
D = 6.0 in.
L = 6.0 in.

64.0 Z

<1
03
ll

5‘0.-
II II

PU.

6n= 6.47x1

t (min.)

33.5
34.0
35.0
45.0
52.0

3/8 in.
0

0—4

P (lbs.)

800
480
400
320
320

7

§23212_§:2_1

—20.00

m
H II

C

6.0 in.

D = 6.0 in.

1:“
II

vs: 64.0 x

6.0 in.

d = 3/16 in.
= 0

(Continued)

4

22.6 le./min.

in./min.

= 23.9 lbs./min.

5 (x10-3in.)
0.

1

6.6
10.6
12.2
15.8
24.0

in./min.

1-->--

k04>®®0
OUIOO-D

6 (x10'3in.)

 

 

 

 

 

 

Table A—4: (Cont'd)

 

 

 

 

Sample S—67: (Cont'd) Sample S—69: (Cont'd)
P = 16.3 lbs./min. —3
. t (min.) P (lbs.) 5 (x10 in.)
6n= ** ————--- --————- —————————————
50.0 510 9.2
. —3. 52.0 480 11 2
t (m1n.) P (lbs. 6 10 . '
———*————‘ ) -$§--3£L2' 55.0 450 13.2
29.5 480 *6 64.0 420 19.0
30.0 270
46.0 210 Sample S—70:
_________________________________ -—----7;-
** Not recorded T = “10-0 C
H = 6.0 in.
Sample S—68: D = 6.0 in.
T — —20.0°c L f 2&0019'
H = 6.0 in. Vs‘ ° °
D = 6.0 in. d = 3/8 in.
L = 6.0 in. h = 0
VS: 64'0 A P = 12.3 lbs./min.
d — 3/8 in ' _ -4 . .
h _ 1/8 in 5n— 5.867x10 1n./m1n.
P = 17.7 lbs./min. t (min.) P (lbs.) 6 (x10-3in.)
6n= 2.3x10‘4 in./min. 28 344 0.4
_3 29 120 6.4
t (min.) P (lbs.) 6 (x10 in.) 30 120 7.2
50 1220 0.1 45 120 16'0
150 2660 18.0 S
S- l:
250 3140 34.0 ~3§3l5——7§—-
400 3840 60.0 T = —10.0 c
600 4580 112.0 H = 4.0 in.
620 4580 120.0 D = 6.0 in.
L = 4.0 in.
Sample S—69: Vs: 64.0 X
T = —25.4°c d = 3/8 in.
H = 6.0 in. h = 0
D = 6'0 in- P = 19.5 lbs./min.
L = 6.0 in. . -4 ,
Vs: 64.0 % 6n= 6.1x10 in./m1n.
d = 3/8 in. t (min.) P (lbs.) 6 (x10—Bin.)
h = 0
. 25.5 495 6.0
P = 21.86 lb80/min. 26.0 145 7.6
9n= 6.5x10’4 in./min. 28.0 130 10-0
—3. 41.0 95 18.0
1'. (min.) P (lbs.) (X10 ln') 47.0 95 21.6
48.5 1060 0.05
49.0 640 6.60

(Continued)

Table A—4: (Cont'd)

Sample S—72:

T
H
D
L
Vs

d
h
P

66

t (min.) P (lbs.) 5 (x10—Bin.)

II

-10.00

C

6.0 in.
6.0 in.
6.0 in.

64.0 Z

3/8 in.
0

12 lbs./min.

5.5x10

—4

in./min.

 

 

 

35.5 520 0.8
35.5 120 17.6
40.0 160 19.6
60.0 160 30.6
Sample S—73:**

T = -15.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

VS= 64.0 Z

d = 3/8 in.

h = 0

P = 25.5 1bs./min.

5n= 4.67x10_4 in./min.

t (min.) P (lbs.) 6 (x10—3in.)
46 1170 0.4
52 1000 0.4
54 1000 0.6
70 1540 1.6
70 280 21.6
79 500 23.6
81 470 25.2
93 470 30.8

 

** Sample loaded, unloaded, and
reloaded before rupture.

Sample S—74:

T
H
D
L

—20.00

C

6.0 in.
6.0 in.

6.0 in

(Continued)

Sample S—74:
64.0 Z
5/8 in.
0

VS:

D‘D...

P
6

n:

t (min.) P (lbs.) 6 (x10—3in.)

25.8 le./min.
**

(Cont')

 

 

118
133
143
153
193
230
280
370

3040
3180
3240
3240
2880
2200
1640
1240

*‘k

 

** Not recorded, equipment failure

Sample S—75:

H
II II

<1
(0
II

D.-
II H

"U
ll

6n=

t (min.)

44
55
55
60
69

T:
H:
D:
L

V5

d
h

ll

—6.0°c
6.0 in.

= 6.0 in.
= 6.0 in.

64.0 X

3/8 in.
O

11.64 1bs./min.
4.5x10_4

P (lbs.)

510
625
160
175
175

§29212_§:Z§;

—10.0°c

(Continued)

in./min.

6 (x10- '

0.1
3.2
14.4
17.2
21.2

 

 

 

Table A-4: (cont'd)

 

Sample S—76: (Cont'd)
P = 14.7 lbs./min.
6n= **

t (min.) P (lbs.) 6 (x10-Bin.)
50 1736 **

76 1090

84 1090

124 900

..__...__________________.____~_
** not recorded, equipment failure
Sample S—77:

T = —10.0°c
H = 5.0 in.
D = 6.0 in.
L = 5.0 in.
64.0 2

= 3/8 in.
0

= 11.7 lbs./min.

4.1x10—4 in./min.

5 (x10—3in.)

2.0
17.2
19.0
28.0

t (min.) P (lbs.)

67 780
67 220
68 190
90 190

Sample S—78:

Leakage, Data void

Sample S—79:

T = —10.0°c
H = 3.0 in.
D = 6.0 in.
L 3.0 in.
vS 64.0 Z

d 3/8 in.
h — O

P = 13.7 1bs./min.

6n= 4.75x1o‘4
(Continued)

in./min.

Sample S-79:

t (min.) P (lbs.) 6 (x10—Bin.)

39.5 540 1.4
40.0 130 14.4
41.0 140 15.4
57.0 140 23.0

(Cont'd)

Sample S—80:

Leakage, data void.

Sample S—81:

T -10.0°c
H
D
L

II ll 11 II
ON
.0.

S

<
03
11

5‘0.-
" 1|
Ob)
\
00
H
:1

9.0 lbs./min.
4.53410“4

07! 17d.
ll

:1
ll

in./min.

t (min.) P (lbs.) 6 (x10—3in.)

30 270 0.8
30 90 8.8
32 60 11.4
44 60 16.8

§EE212_§:§ZL

—10.0°c
3.0 in.
6.0 in.
3.0 in.
64.0 Z

= 5/8 in.
- 0

H:
D:
L:

<
(D
II

= 8.2 lbs./min.
én= **
t (min.) P (lbs.) 6 (x10—316.)

53 430 **
80 430

 

** Not recorded, equipment failure.

 

 

 

 

 

 

Table A—4: (Cont'd)

 

 

 

 

Sample S—83: Sample S—85: (Cont'd)
T = —10.0°c -3
H = 2.0 in. 1: (min.) P (lbs.) 6 (X10 in.)
D — 6.0 in. 26.5 530 0.1
L = 2-0 1n- 26.5 220 4.8
VS= 64.0 A 30.0 180 7.6
h = 0
. Sample S-86:
P = 6.8 lbs./min. o
, _4 T = ‘10.0 C
5n= 3.1X10 in./min. H = 6.0 in.
_3 D = 6.0 in.
t (min.) P (lbs.) 6 (x10 in.) L = 6.0 in.
42.5 290 0.6 VS: 64'0 Z
43.0 80 9.6 d = 3/8 in.
80.0 80 21.0 h = 0 (Ground-Finish)
Sample S—84: P = 23.3 1b:./min.
T = —10.OOC 5n= 4.0X10 in./min.
H = . ' . -
D = §.8 :2. t (min.) P (lbs.) 6 (x10 3in.)
L = 2.0 in. 0.0 100 0.01
VS= 64.0 Z 24.5 670 0.1
_ . 24.5 170 6.8
g ; 8/8 1“' 30.0 140 12.0
. 35.0 140 14.0
P = 4.45 le./min. 43.0 140 17.0
én= ** 113.0 140 45.0
— S l S-87:
t (min.) P (lbs.) 6 (x10 3in.) -EEEL£L——7;‘"
T = -10.0 C
55 240 ** H = 6.0 in.
, . D = 6.0 in.
4* Not recorded, equ1pment failure L = 6.0 in.
= 64.0 Z
Sample S-85: vs
0 d = 3/8 in.
T = ‘10.? C h = 1/8 in.
H = 6.0 1n. (450—lug, Ground—Finish)
D = 6.0 in. .
L = 6.0 in. P = 23.1 lbs./min.
vs= 64.0 2 5n: 5,93x10‘4 in./min.
d = 3/8 in. t (min.) P (1bS.) 6 (x10—3in.)
h = 0 (Ground—Finish)
, 35 810 0.4
P = 20 lbs./min. 160 2090 30.2
6n= 3.24x10'4 in./min. 260 2530 78.6
360 2730 129.6

Co ti d
( n nue ) 405 2730 154.6

 

Table A-4: (Cont'd)

 

 

 

SamPle S—88: Sample S-90: (Cont'd)
o
T = —10.0 c vs= 64.0 2
H = 6.0 in. _ .
D = 6.0 in d : 3/8 Pn‘
L = 6.0 1n h — 1/8 in.
VS: 64.0 Z . (45°-lug, Shot‘BlaSted)
d = 3/8 in. ?= 30 lbs./211n.
h = 0 (Shot-Blasted) 6n= 5.39x10' in./min.
P = 18.5 1 . ' . —
, ES /mln t (min.) P (163.) 6 (x10 3in.)
6n= 7.5x10 4 in./min.
72 2160 1.0
. -3, 75 1960 6.0
t (min.) P (lbs.) 6 (x10 in.) 85 1960 11.2
94 1740 3.2 185 2920 52.8
94 1600 9.6 370 3400 146.0
95 1140 19.2 520 3400 225.0
97 1070 24.2
100 990 28.3 Sample S-91:
160 730 69.3 0
T = -15.0 C
175 730 80.8 H = 6.0 in.
D = 6.0 in.
Sample S—89. L = 6.0 in.
T = —10,o°c vs: 64.0 2
g = 2'8~:n‘ d = 3/8 in. O
= n. -
' = 1 . 45 -l
L = 6.0 in. F /8 1n ( ug)
Vs: 64.0 Z P = 23 1bs./min.
= 3/8 in. 6n= 5.0x10—4 in./min.

d
h = 1/8 in. (45°—1ug)

- t (min.) P (lbs.) 6 (X10—3in.)
P = 25 lbs./min.

 

' -4 . . 85 2160 1.5
6n = 5.5x10 1n./m1n. 95 2160 8.0
-3. 195 2920 36.0
t (min.) P (lbs.) 6 (x10 in.) 350 3840 84.0
15.6 1440 0.1 560 4360 182.0
16.0 1560 9.0 680 4360 242.0
66.0 2300 26.5
166.0 3160 70.0 Sample S—92:
271.0 3380 121.0 T = —15.OOC
350.0 3380 165.0 H = 6.0 in
523.0 3380 260.0 D 6.0 in
L - 6.0 in.
sample 3'90: VS: 6400 z
T = ‘10.OOC d = 3/8 in
H = 6.0 in. h = 0
D = 6.0 in. , .
L = 6.0 in. P = 20.5 13s./m1n.
(Continued) ( Contlnue )

 

 

 

Table A—4:

Sample S—92:

 

6n= 6.25x10

t (min.) P

(lbs.)

 

(Cont'd)

(Cont'd)
-4

 

 

in./min.

6 (x10-3in.)

 

61 1250 0.8
61 95 20.0
70 320 20.4
71 320 21.6
73 300 23.2
85 280 31.2

100 280 39.7

120 280 52.2

Sample S—93:

T = —20.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

VS= 64.0 %

d = 3/8 in. o

h = 1/8 in. (45 —lug)

i = 34.9 lbs./min.

6n= 5.2x10‘4 in./min.

t (min.) P (lbs.) 6 (x10—3in.)
75 2720 0.1
80 2600 5.0
90 2600 9.6

190 3800 42.0

340 4480 116.0

420 4600 160.0

450 4640 175.0

Sample S—94

T = —20.0°c
H = 6.0 in.
D = 6.0 in.
L = 6.0 in.
vs= 64.0 %
d = 3/8 in.
h = o

f = 18 lbs.
6n= 5.85x10

(Continued)

/min.

in./min.

Sample s—94:

80 1440 0.
95 1700 0.
95 300 25.
103 480 25.
104.5 520 26.
108.0 520 29
110.0 460 32
126.0 440 41
136.0 440 48.

Sample S—95:

(Cont'd)

T = -26.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

Vs: 64.0 2

d = 3/8 in.

h = 0

f = 22.2 lbs./min.

6n= 6.67x10'4 in./min.

t (min.) P (lbs.) 6 (x10—3in.)

OOWO‘OOOO‘bN

t (min.) P (lbs.) 6 (x10—3in.)

83 1840 0.1
83 240 28.0
103 800 29.0
103 680 32.0
105 600 34.0
115 640 39.0
145 640 59.0

Sample S—96:

T = -6.0°c
H = 6.0 in.
D = 6.0 in.
L = 6.0 in.
Vs: 64.0 Z
d = 3/8 in.
h = 1/8 in.
é = 27.6 lbs
6n= 5.6x10_4
(Continued)

(45°—lug)
./min.

in./min.

 

 

 

Table A-4: (Cont'd)

Sample S-96: (Cont'd)

 

t (min.) P (lbs.) 6 (X10—3in.)

50
65
75
175
275
375

iI'aI-i
II II

n-

Ov w- r a < H U

t (min.)

30
80
160
200
320

1380 .0
1280 0
1280 0
1760 4
1960 0
2040 0
Sample S—97:
-2.0°c
6.0 in.
= 6.0 in.
6.0 in.
64.0%
= 3/8 in. 0
1/8 in. (45 —lug)
19.34 lbs./min.
5.67x10_4
-3.
P (lbs.) 6 (x10 1n.)
‘ 580 4.5
1040 23.0
1440 61.5
1520 81.5
1520 149.5

 

 

Sample S—98:

—10.00c
2.0 in.
6.0 in.
2.0 in.

T:
H
D
L:
vs
d
h
P

6n=

II II

II II

t (min.) P (lbs.) 6 (x10—3in.)

23.5
23.5
25.0
40.0

64.0 Z

5/8 in.
O

25 lbs./min.

5.73x10—4

576

170
145
145

H
l

"U‘ 'J‘Q. <
[D
II H II

II

6n=

t (min.) P (lbs.) 6 (x10—Bin.)

52
52
58
62
80

T:
H:
D:
L:
vs=

d:

5/8 in.
0

36.9 le./min.

 

Sample S—99:

— 10.000
‘ 4.0 in.
= 6.0 in.
= 4.0 in.
64.0 X

6.67x10'4

1

h = 0

II

P

6n=

t (min.) P (lbs.) 6 (x10—3in.)

41
41
42
54
62

T =
H =
D =
L =
VS—
d =
h =

31.22 lbs./min.
5.6x10'4

920
320
500
480
480

Sample S—100:
—10.0°c

1280

5.0 in.
6.0 in.
5.0 in.
64.0 Z

5/8 in.
0

(Continued)

460
300
290
290

§39212_§:1911
—10.0°c

in./min.

Table A-4: (Cont'd)

Sample S—101: (Cont'd)

P = 41 lbs./min.

6n= 6.03x10'4 in./min.

 

Sample S—103:

 

(Cont'd)

6n= 5.7}(10—1+ in./min.

 

 

t (min.) P (lbs.) 6 (x10—3in.)
t (min.) P (lbs.) 6 (X10—3in.) 35.5 980 2.4
76.0 3120 2.0 35.5 440 8.8
36.0 290 10.8
77.0 2920 4.0
40.0 260 13.4
77.5 1980 15.0 60 0 260 24 8
78.0 1620 20.5 ' -
79.0 1340 24.0
82.0 1250 27.5 §EEElE_§:lQi;
116.0 1220 48.0 Leakage, data void
Sample S-102: Sample S—105:
T = —10.0°c T = -10.0°c
H = 6.0 in.- H = 5.0 in.
D = 6.0 in D = 6.0 in
L = 6.0 in L = 5.0 in
vS= 64.0 2 vS= 64.0 Z
d — 5/8 in d = 5/8 in
h h = 0
P = 41.73 lbs./min. P = 35.6 lbs./min.
6h: 7.17.;10'4 in./min. 6n= 6.25x10_4 in./min.
— -3.
t (min.) P (lbs.) 6 (x10 3in.) t (min.) P (lbs.) 6 (x10 1n.)
81.0 3380 1.5 81 2880 2
83.0 2980 5.0 82 2520 6
83.5 1700 23.5 83 1320 21
85.0 1300 29.5 88 1080 27
88.0 1180 32.5 96 1080 32
130.0 1180 56.5

Sample S-103:

T = C

H = 2.0 in.

D = 6.0 in.

L = 2.0 in.

Vs: 64.0 Z

d = 5/8 in.

h = 0

P = 28.1 lbs./min.
(Continued)

—1o.0O

EEaTfl£;§:lgé:

T
H

D _

L

v:

S

"U'D‘Da

6n=

(Contiued)

—10.00

C

2.0 in.
6.0 in.
2.0 in.

64.0 Z

5/8 in.

0

25 lbs./min.

5.86x1

0—4

in./min-

 

 

 

312

Table A-4: (Cont'd)

 

Sample S-106: (Cont'd)

 

 

 

 

 

 

 

 

t (min.) P (lbs.) 6 (x10-Bin.) Sample S-109: (Cont'd)
23 576 1.6 L = 6.0 in.
23 210 7.2 vs= 64.0 2
24 120 8.0
d = 5/8 in.
39 12 .
0 16 8 h = 0

Sample s-107: P = 35 lbs./min.
T = -10.0°c 6n= 5.0x10’4 in./min.
h = 5.0 in. _3
D = 6.0 in. t (min.) P (lbs.) 6 (x10 in.)
t ; 240012’ 56 1960 2.0

S ' ° 56 320 19.0
d = 5/8 in. 65 640 23.0
h = 0 67 480 26.0
P = 38.9 lbs./min. 1:; :28 §g°g

n= 6.2x10-4 in./min.

030

Sample S-llO:

 

 

 

 

 

 

. —3.
t (min.) P (lbs.) 6 (x10 in.) T = -1O.OOC
52.5 2040 3.0 H = 4.0 in.
52.5 460 20.0 D = 6.0 in.
57.0 . 500 23.0 L = 4.0 in.
74.0 500 33.5 vs= 64.0 %
d = 5/8 in.
Sample S-i08: h = 0
T = '10'0 C P = 28.2 lbs./min.
H = 3.0 in. , _4
D = 6.0 in. 6n= 5.56X10 in./min.
L = 3.0 in. -3
vs: 64.0 % t (min.) P (lbs.) 5 (x10 in.)
d = 5/8 in. 17 500 0.8
h = O 17 170 4.0
’ . 20 140 8.0
P = 26.5 le./m1n. 30 130 14.0
6n= 5.8x10—4 in./min. 60 120 31°C
66 120 34.0

t (min.) P (lbs.) 6 (x10-3in.)

Sample S-111:
20 530 0.01

 

o

20 160 5.60 T = -10.0 C
21 130 6.40 H = 5.5 in.
41 130 18.00 D = 6.0 in.

L = 5.5 in.
Sample S—109; Vs: 64.0 %
T = -10.0°c d = 3/8 in.
H = 6.0 in. h = 0
D = 6'0 in. . (Continued)
(Continued)

 

 

 

Table A—4: (Cont'd)

Sample S-111: (Vont'd)

P = 20 lbs./min.

6n= 6.67x10_4 in./min.

t (min.) P (lbs.)

 

28 560
28 320
40 40
52 40
Sample S-112:
T = —10.0°c

H = 5.625 in.

D = 6.000 in.
L = 5.626 in.
vs= 64.0 2

d = 3/8 in.
h = 0

P = 21.2 lbs./min.
= 6.25x10‘4 in./min.

6

D

t (min.) \P (lbs.)

30 640
30 240
31 100
42 80
47 80

Sample S—113:

T = -15.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

vs= 64.0 2

d = 3/8 in.

h = 0

P = 25.2 lbs./min.

6 (X10-3in.)

0.5
8.5
16.5
24.5

6 (x10—3in.)

0.1
9.5
12.0
19.0
22.0

6n= 4.75x10_4 in./min.

t (min.) P (le-)

42 1060
42 280
58 280

6 (x10—3in.)

0.7
14.0
21.6

313

Sample S—114:

L—‘UCCH
llll

<1
U}

l
0"
4.x.
0
N5

3/8 in.
— 0

5‘04

P = 25.8 lbs./min.

6n= 4.0x10-4 in./min.

t (min.) P (lbs.) 6 (x10—3in.)

27 690 0.40
35 430* 0.16
38 430 0.18
63 1250** 0.24
63 280 14.20
64 220 14.40
68 250 15.80
78 240 19.80
_.__.___.__._______.______._____

* Decreasedeue to unloading
** Reloading

Sample S—115:

 

T = ~10.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

vs= 64.0 2

d = 3/8 in.

h = 0

P = 17.3 lbs./min.

6n= 4.2X10_4 in./min.

t (min.) P (lbs.) 6 (x10—3in.)
33.5 580 1.0
33.5 160 7.4
34.0 80 9.8
74.0 30 26.6

108.0 30 33.8

 

 

 

 

 

 

 

 

 

314

Table A—4: (Cont'd)

 

 

 

 

 

 

Sample S-116: Sample S-ll8: (cont'd)
T = -10.0°c h = 0
H = 6.0 in. ° .
D = 6.0 in. P = 22 lbs.ém1n.
L = 6.0 in. 6n= 2.8x10 in./min.
Vs: 64.0 Z _3
d = 3/8 in. t (min.) P (lbs.) 6 (x10 in.)
h = 0 37 810 0.1
' _ - 37 330 4.4
P 20.3 lbfg/mln. 40 200 8.8
6n= 4.84x10 in./min. 52 170 14.0
75 170 20.4

t (min.) P (lbs.) 6 (x10-Bin.)

 

Sample S-ll9:

 

 

 

 

 

 

 

 

22.5 460 0.4 o
22.5 170 4.4 T = -10.0 C
23.0 100 7.0 H = 6.0 in.
33.0 40 16.6 D = 6.0 in.
61.0 40 25.4 L = 6.0 in.
4 Vs: 64.0 Z
S' 1 S-1 .
amp e 017, d = 5/8 in.
T = -20.0 C h = 0
H = 6'0 Pn' P = 33.4 le./min.
D = 6.0 in. . _4
L = 6.0 in. 6n= 1.17X10 in./min.
Vs= 64'0 Z t (min.) P (lbs.) 6 (x10—3in.)
d = 3/8 in. 41 1370 0.9
h = 0 41 50 7.2
P = 24.6 le./min. 50 280 7.5
° —4 , 56 240 8.2
6n= 2.1x10 in./m1n. 68 240 9.6
-3.
t (min.) P (lbs.) 6 (x10 1n.) Samlpe S-120:
54 1330 0.6 _ o
T - -10.0 C
54 280 12.8 H = 2.0 in.
59 250 16.0 D = 6.0 in.
87 210 23.0 L = 2.0 in.
113 180 27.2 VS= 64.0 Z
160 180 36.9
d = 3/8 in.**
Sample S-118: h = 0.015 1n.**
T = -20.0°c P = 170 lbsé/min.
H = 6'0 in. 6n= 4.6x10- in./min-
D = 6.0 in.
L = 6-0 in. ** Standard deformed bar
Vs: 64-0 Z (Continued)
d = 3/8 in.

(Continued)

Table A-4: (Cont'd)

Sample S—120:

 

(Cont'd)

 

t (min.) P (lbs.) 6 (x10—3in.)
2 560 0.9
4 1060 2.5
5 1260 4.0
6 1420 7.5
8 1630 13.0

10 1870 19.0

12 2030 25.0

14 2190 32.5

16 2270 40.0

17 2300 43.8

20 2300 56.5

24 2200 75.0

29 2080 100.0

53 1800 130.0

40 1610 155.6

45 1540 180.0

60 1310 250.0

81 1050 325.0

Sample S-121:

T = —10.0°c

H = 3.0 in.

D = 6.0 in.

L = 3.0 in.

Vs: 64.0 Z

d = 3/8 in.**

h = 0.015 in.**

P = 235 lbs./min.

6n= 4.5X10_ in./min.

t (min.) P (lbs.) 6 (x10—3in.)
4 1102 0.5
8 2100 4.4

10 2678 8.8

12 3220 13.5

14 3378 18.5

16 3762 25.0

18 3890 31.0

21 4025 46.3

24 3940 60.5

28 3590 89.4

32 3100 117.0

36 2720 143.5

40 2470 174.0

(Continued)

Sample S—121: (Cont'd)

t (min.) P (lbs.) 6 (x10—3in.)

46 2300 207.0
52 2120 242.0
60 1980 282.0
68 1840 327.0

 

 

 

 

 

 

** Standard deformed bar

Sample S—122:

T = —1o.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

vS= 64.0 2

d = 3/8 in.

h = 1/8 in.

P = 40 lbs./min.

5n= 6.7x10_4 in./min.

t (min.) P (lbs.) 6 (x10—3in.)

20.0 770 0.3
40.0 1600 1.8
41.0 1570 2.5
43.0 1300 7.5
46.0 1080 15.0
50.0 1120 17.5
55.0 1150 20.0
60.0 1208 23.0
70.0 1310 29.5
90.0 1450 41.2
120.0 1560 61.2
160.0 1698 88.0
170.0 1750 116.5
190.0 1750 166.5
285.0 1750 230.0

Sample S—123:

P = 22 lbs./min.
6n= 8.2x10'4

(Continued)

in./min.

Table A—4: (C

ont'd)

Sample S—123: (Cont'd)

t (min.) P (lbs.) 6 (X10—3in.)

15.0
15.0 113
17.5 105
21.0 96
5
5

330

29.
32.

Sample S-124:

T = —10.0°c
H=D=L=6
d = 3/8 in.
h = 0
v = 29.6 Z
P = 31.8 lbs.

6n= 6.8x10"4

S

t (min.) P (lbs.) 6 (x10—3in.)

16.5 525
16.5 60
20.0 ‘60
30.0 60
46.0 60

.0 in.

/min.

in./min.

Sample S-125:

T = -1o.o°c
H = L = 6.125
D = 6.0 in.
d = 3/8 in.
h = O
v = 53.8 Z
P = 20.8 lbs.

7.05.10“4

t (min.) P (lbs.) 6 (x10—3in.)

10.5 22
11.0 10
21.0 8
28.0 7
34.0 7

in.

/min.

in./min.

5
0
5
5
5

0.5
4.5
6.2
10.0
17.0
21.0

0.20
8.25
10.75
19.62
30.63

0.1
2.5
10.5
15.1
19.6

Samp
T
H
d
h
VS

6 =

II II II

II II

6n= 8.4x10'4 in./min.

16.2
16.2
18.0
24.0
29.0

Samp

EH

<1D‘Qa

S

lad!

1'1—

O)‘

t (min.) P (lbs.) 6 (x10—3in.)

16.6
17.0
18.0
21.0

Samp

le S-126:

—1o.0°c

D = L = 6.0 in.
3/8 in.

0

52.7 %
44 lbs./min.

710
112
145
145
245

le S—127:

—10.0°c

D = L = 6.0 in.
3/8 in.

0

= 30.3 Z

31.6 lbs./min.
1.1x10'3

525
135
60
60

~___3513t12§;

in./min.

0.4
9.2
10.0
15.5
19.2

0.8
7.5
8.6
11.9

t (min.) P (lbs.) 6 (x10—3in.)

 

Vs: 44 . 2 %

P = 30 lbs./min.

6n= 8.1}(10-4 in./min.

t (min.) P (lbs.) 6 (x10—3in.)
10.6 320

11.0 120

13.0 75

16.0 40

19.0 40

317

Table B—1: Ice samples CI—l and 01-2 (Ice 3)

 

Sample CI—l: Sample CI—2: (Cont'd)
T = —10.0°c -3
H = 5.87 in. ELilP§il .E_QEEEQ_ §E£§19~_12;2
D = 6.00 in. 280 550 1.25
L = 5.87 in. 600 1.00
d = 3/8 in.(Cold-rolled) 690 1.00
h = 1/8 in.(9OO-lug)
360 0 1.00
P (lbs.) t (min.) 60(x10'3in.) :88 3';§
520 20 0.20 300 1.00
40 0.75 500 1.00
60 1.75 710 1.00
138 ::88 440 0 1.00
120 18 50 300 1'00
140 83.50 400 1‘12
160 141.00 :88 1'33
180 213.50 750 2'00
200 286.00 1000 2‘50
220 361.00 1100 3'00
240 443.50 1200 3'00
1300 3.00
Sample CI-Z:
T ;_ 1P717517 1380 3.50
H i 60'... 520 0 3'50
D = 6.0 in. 100 4.00
L = 6 0 in 200 4.50
- . . 300 5.00
d = 3/8 in. 400 6 00
h = 1/8 in. 500 7:25
P (lbs.) t (min.) 6C<x10 3m.) 228 1188
100 0.01 720 24.00
200 0.10
300 0.50
400 0.50
500 0.10
600 0.10
700 0.25
1000 0-25
1100 0.50
1255 0.40
280** 0 0.40
100 0.75
200 1.00
400 1.12
(Continued)

** Step loading procedure

318

Table B—2: Frozen sand samples

Sample CS—1:**

—10.o°c-
6.0 in.
6.0 in.
6.0 in.
5/8 in.
0, VS= 64.0 2

I] II II

D‘Qur‘Uth-J

** Instant rupture at t=0.75 min.,

P=1420 lbs., and 6C=0.0002 in.

Sample CS—2:

= -10.0°c
= 6.0 in.
= 6.0 in.
6.0 in.
= 5/8 in.
= 0 , Vs= 64.0 Z

D‘D-F‘UCL‘H
II

P (lbs.) t (min.) 6C(x10'3in.)

1040 l 0.1
2 0.7 rupture

Sample CS—3:

 

T = —10.0°c

H = 6.0 in.

D = 6.0 in.

L = 6.0 in.

d = 5/8 in.

h = 0 , vs= 64.0%

P (lbs.) t (min.) 6c(x10 3in.)

750 0.75 0.10
34.00 0.15

910 34.00 0.20
48.00 0.21
66.00 0.84
70.00 1.40

(rupture)

Sample CS-4:**
T = —10.0°c
H = D = L = 6.0 in.
d = 5/8 in., h = O , VS= 64.0 A
** Instant rupture at
P = 800 lbs.

Sample CS—5:

T = —1o.0°c

H = D = L = 6.0 in.

d = 3/8 in.

h = 1/8 in. (900-lug)
Vs= 64.0 2

2300 1
2

3

4

12
21
40
6O
80
120
160
200
300
400
600
800
1100
1420

2500 1425
1520
1640
1820
2030

2030
2220
2420
2620
2820
2920

2920
2960
3020

3020
3120
3220

2780

3020

3180

9
13
15
17
27
33
43
50

P (lbs.) t (min.) 6C(x1o'3in.)

319

 

 

Table B-2: (Cont'd)
Sample CS—6: Sample CS-7:
T = -1o.0°c T = —1o.0°c
H = D = L = 6.0 in. H = D = L = 6.0 in.
d = 3/8 in. d = 3/8 in.
h = 1/8 in. h = 1/8 in.
vs= 64.0 % vs= 64.0 2
P (lbs.) t (min.) 60(x10‘3in.) P (lbs.) t (min.) 6C(x10'3in.)
2800 1 15 3560 5 61
2 20 10 81
3 24 25 121
4 26 45 156
5 38 65 185
6 30 100 229
7 31 150 286
10 33 200 349
20 45 250 411
30 53 275 444
40 61 300 476
60 71 405 625
80 81 3780 405 626
100 89
420 664
150 109 460 751
200 125 500 822
- 250 139 545 916
300 $2: 4000 580 1019
350 615 1106
400 177
500 203
CS—8:
585 225 §23219__7;__
T = -10.0 c
3080 585 229 H = D = L = 6.0 in.
700 277 h = 1/8 in.
800 321 .
= .0 A
856 351 V5 64
, —3.
3320 865 353 P (le-> t (m1H-) 5C<X10 1n')
900 383 4000 1.5 33
1000 451 3,0 45
1100 518 5.0 54
1200 589 8.0 64
3560 1200 591 12-0 g:
1250 641 15.0
1300 687 gg'g iéi
833 -
1480 45.0 204
60.0 252
80.0 309
100 0 369

(Continued)

 

Table B-2: (Cont'd)

Sample CS—8:

(Cont'd)

320

Sample CS—9: (Cont'd)

3.

 

 

(Continued)

. -3. —
P (lbs.) t (min.) 6C(x10 in.) P (lbs.) t (min.) 6C(x10 in.)
4000 150 514 1800 700 710
200 654 780 788
240 771
275 874 2080 780 790
790 815
800 840
Sample CS—9: 810 860
T = _10 00C 820 875
H = D = L = 6.0 in.
d = 3/8 in. Sample CS—lO:
h = 1/16 in. T _ 10 00C
VS: 64'04 H=D=L=6.0in.
. -3. d = 3/8 in.
P . . C .
(lbs ) t (min ) 6 (x10 1n ) h = 1/8 in.
1400 1 15 vs= 64.0 2
2 20 _3
5 32 P (lbs.) t (min.) 6C(x10 in.)
10 40 1580 5 9
20 54
10 13
40 74
20 23
70 95
- 6O 33
100 115
100 36
120 128
145 40
137 161
137 178 1860 195 46
170 198 230 52
200 210 280 56
245 228 350 63
1560 245 230 2140 350 65
270 252 400 72
300 275 450 78
342 300 505 86
1800 342 302 2300 505 88
345 308 550 93
350 320 600 100
360 338 700 113
370 350 800 126
400 398 950 146
450 450 1050 153
500 500 1280 159
550 563
600 612
650 670

 

 

321

Table B—2: (Cont'd)

 

 

 

 

Sample CS—ll: Sample CS—12:
T = “lo-000 T = —10.0°c
H = D = L = 6.0 in. H = D = L = 6.0 in.
d = 3/8 in. d = 3/8 in_
h = 3/16 in. h = 1/8 in.
VS: 64.0 7° VS: 64.0 7;
P (lbs.) t (min.) 6C(x10 3in.) P (lbs.) t (min.) 6C(x10_3in.)
2000 1 2.5 4000 5 96
2 4.0 10 107
5 4.5 15 163
10 6.0 20 188
20 7.0 25 212
40 8.1
70 10.1 Test stopped, LVDT OFF
100 11.1
160 12.5 Sample CS-13:
230 13.2 T = _10.Ooc
2480 230 13.3 H = D = L = 6.0 in.
250 16.8 d = 3/8 in.
290 21.7 h = 1/16 in.
350 26.7 vs= 64.0 2
400 30.8
\ 500 36.8 P (lbs.) t (min.) 6C(x1o'3in.)
600 42.0
700 46.4 850 g é'g
785 49.5 18 12_5
3040 785 50.2 38 19.6
800 53.0 80 28.8
850 58.6 120 33.5
900 63.3 170 37.7
970 68.5 230 44.0
1040 72.7 300 49.0
3600 1040 73.3 345 51.5
1070 . 76.1 1200 345 52.0
1120 84.9 370 64.0
1160 88.9 400 74.0
1200 92.3 450 86.0
1270 98.5 500 97.0
585 114.6
4160 1270 98.9
1300 108.4 1560 585 117.1
1350 120.4 600 133.6
1400 131.6 618 149.6
1500 151.6 1220 417.1
1600 172.4 1300 467.3
1455 542.1

(Continued)

322

Table B—2: (Cont'd)

 

 

 

 

Sample CS-13: (Cont'd) Sample CS—lS: (Cont'd)
P (lbs.) t (min.) 6C(x10-3in.) P (lbs.) t (min.) 6C(x10_3in.)
2600 1455 592.0 3000 5,2 504
1458 652.0 6.0 607
1465 727.0 7.0 645
1480 829.6 8.0 670
1500 989.6 9.0 708
1520 1117.0 10.0 745
1540 1214.6 10.0 rupture
Sample CS—14: Sample CS-16:
T = —10.0°c T = —15.0°c
H = D = L = 6.0 in. H = D = L = 6.0 in.
d = 3/8 in. d = 3/8 in.
h = 1/8 in. h = 1/8 in.
vs= 64.0 2 Vs= 64-0 Z
P (lbs.) t (min.) 6C(x10‘3in.) P (lbs.) t (min.) 6C(x1o'3in.)
4500 0 31 1500 5 1.75
2 75 10 3.12
5 112 20 5.38
10 156 50 10.00
\ 20 231 100 15.00
32 312 150 18.75
47 400 200 20.88
48 475 300 25.00
57 578 400 28.50
72 725 500 30.25
89 890 600 32.50
90 950 706 34.25
100 1080 2000 706 34.38
730 38.50
Sample CS—15: 770 41'25
—‘_—____—7;“— 800 43.12
T = -10.0 C 850 46.00
H = D = L = 6.0 in. 900 48.75
d = 3/8 in. 1000 53.12
h = 1/16 in. 1100 56.50
vS= 64.0 % 1200 60.00
_3_ 1300 62.50
P (lbs.) t (min.) 6C(x10 1n.) 1400 65.00
3000 0.5 140 1500 67.50
1.0 280 1554 68.38
1.5 375 2500 1554 68.60
3.0 420 1600 71.40
4. 470
(Continued)

(Continued)

 

 

 

 

 

 

323

Table B—2: (Cont'd)

 

 

 

 

 

Sample CS—16: (Cont'd) Sample CS—17: (Cont'd)
P (lbs.) t (min.) 5C(x10—3in.) P (lbs.) t (min.) 6C(x10 3in.)
2500 1650 76.40 3500 350 91.0
1700 79.40 400 101 0
1800 86.40 500 116-0
1950 95.65 650 136-0
2134 104.40 800 151-0
950 166.0
3000 2134 105.65 1100 181.0
2150 109.40
2200 119.40 2500 1100 181.0
2300 133.15 Unloading 1210 182.0
2450 151.15 1445 186-0
2600 167.65
2750 184'15 Sample CS—18:
2938 203.15 '——_“":;—“‘
T = 10.0 c
3500 2938 205.15 H = D = L = 6.0 in.
3000 225.15 d = 3/8 in,
3100 255.15 h = 1/8 in_
3250 295.15 Vs: 64.0 2
3400 325.10
3600 375.15 P (lbs.) t (min.) 5C(x10_3in.)
3730 415.10
. 1220 20 6.5
4000 3730 415.60 60 12.5
3800 473.65 100 16.7
3880 525.15 150 20.2
3935 615.15 200 22.5
300 26.5
§29219_9§:111 238 3322
T = —10.0°c 725 35.0
E : 3/5 5 _ 6‘0 1n' 1540 725 36.5
‘ 1“' 760 38.4
h = 3/16 in. 800 39.9
vs= 64'0 Z 900 43.2
P (163.) t (min.) 6C(x10 3in.) 1238 ggjg
2500 10 17.5 1350 52.8
20 25.0 1450 54.6
40 33.8 1820 1450 55.2
70 40.0 1495 58.2
100 47.0 1550 60.7
150 53.8 1600 62.0
200 58.8 1700 66.4
275 65'0 1850 74.0
3500 275 66.0 2000 80.0
300 79.5 2155 86-4

(Continued) (Continued)

Table B—2: (Cont'd)

Sample CS—18: (Cont'd)

324

P (lbs.) t (min.) 6C(X10-3in.)

 

2100 2155
2200
2300
2500
2700
2830

2380 2830
2900
3000
3150
3300
3560

2100 3560
Unloading 3650
3700
3800
3910

1540 3910
4100
4355

1220 1 4355
4500
5100

Sample CS—19:

T = —10.0°c

H = L = 5.954 in.
D = 2.781 in.

d = 3/8 in.

h = 1/16 in.

vs: 64.0 Z

P (lbs.) t (min

1500 20 70.
60 120.
100 157.
150 207.
200 251
300 345.
400 426.
550 563.
700 713.
(Continued)

88.
91.
98.
108.
117.
125.

127
134
140
152
162
180

180.
184.
186.
190.
191.

191.
191.
190

190
190
190

000 CNN NNNNO O\I\JNN\I NNNNNN

0
0
5
5
.2
0
2
8
8

.) 66(x10‘3in.)

 

Sample CS—l9: (Cont'd)

 

P (lbs.) t (min.) 6C(x10'31n.)

1500 800 807.5
900 1032.5
1000 1148.8
1140 1317.6

Sample CS—20:

T = —10.0°c

H = D = L = 6.0 in.

d = 3/8 in.

h = 1/8 in. (at z = 4.5 in.)**

vS= 64.0 2

** For all samples z = 3.0 in.,
unless noted.

P (lbs.) t (min.) 6C(x10—3in.)

1500 8 10.0
18 15.0
38 21.2
68 27.5
100 32.5
150 38.1
200 43.8
300 51.2 ‘
400 57.5 ‘
500 62.5
650 70.0
800 76.8
1000 85.0
1225 93.0
2000 1225 95.0
1260 103.0
1300 110 0
1400 122.5
1500 135.0
1650 150.5
1800 166.8
2015 187.5
2500 2015 190.0
2060 215.5
2150 252.5
2300 307 5
2500 382.5
2675 452.0
(Continued)

 

 

 

 

325

Table B—2: (Cont'd)

 

 

 

Sample CS—ZO: (Cont'd) Sample CS—22: (Cont'd)
- '3. d = 3/8 in.
P (le-) t (mm) 6‘3 x10 1 .
“"”‘—" ‘“'——-—’- 7-J;---511- h = 1/8 in. (at z = 4.5 in.)
3000 2675 452.5 vs: 64.0 Z
2735 530.0
2800 605.0 P (lbs.) t (min.) 6C(x1o'3in.)
2900 748.8 ““““ "““*——' ‘-———‘—*————-
2960 883.8 1500 10 11-2
2960 rupture 20 15-8
40 21.0
70 27.0
§§E21£L£Etiéli 100 31.5
0 150 37.2
T = -10.0 c
H = L = 6.0 in. 200 42.2
D = 1.938 in. 300 48-4
= . 550 60.3
d 3/8 1n.
_ . 700 66.0
h — 1/8 in.
V = 64.0 7 850 73.4
5 ° 1000 78.4
P (lbs.) t (min.) 6C(xlO-3in.) 1120 81.0
1000 10 12.5 2000 1120 83.4
20 17 5 1200 94.7
‘ 1300 104.4
40 27.0
1500 119.1
70 36.3
1700 131.6
100 42.5 1840 140.4
150 50.0
200 56.8 2500 1840 143.3
300 67.0 2000 176.6
450 77.0 2200 221.6
600 83.8 2400 259.1
800 92.5 2600 301.6
1000 99.2 2700 321.6
1190 105'0 3000 2700 322.0
1740 1190 106.0 2850 409.5
1220 125.0 3000 492.0
1260 141.2 3200 602.0
1320 161.2 3370 685.0
1400 183'7 3500 3370 687.0
1500 213.7 3500 874.5
1650 263'7 3700 1137.0
1800 313-7 3900 1387.0
2000 406-2 4060 1612.0
2200 543.7
2200 rupture
w
ffififiébiifiiiéli T = —10.0°c
T = —10.0°c H = L = 6.0 in.
H=L=D=6.0 in. D=10938 1n.
(Continued)

(Continued)

Table B—2: (Cont'd)

Sample CS—23: (Cont'd)

 

d = 3/8 in.
h = 1/8 in.
Vs= 64.0 Z

P (lbs.) t (min.) 6C(x10‘3in.)

1500 10
20
4O
70

100
200
300
450
600
800
1000
1100
1250
1250

Sample CS—24:

T = —6.0°c
H = L = D F 6.0 in.
d = 3/8 in.
h = 1/8 in.
vs= 64.0 Z

P (lbs.) t (min.) 60(x10'3in.)

1020 20
60
100
150
200
300
450
600
800
1000
1115

1500 11150
1160
1200
1300
1400
1550
1700

26

38.
55.
72.
85.

122
155

200.
242.
325.
437.

512
682

rupture

7
14
17
19

22.
30.
35.

39
43

47.
49.

51.
57.

61
67
73

81.
87.

(Continued)

U1U1U‘IOU‘IOOU‘IOUIONN

WNO\\0\I\OO WWCNOO-DDKOCOOOH

Sample CS-24: (Cont'd)

K

P (lbs.) t (min.

1500 1840

1980 1840
1900
2000
2150
2300
2500
2625

2500 2625
2660
2700
2800
2950
3100
3300
3400

3000 3400
3440
3500
3600

§225££Ll£ii§ia

) 6C(x10'3in.)
93.

96.
110
128
145.
163.
188.
203.

204
226
248
293.
383.
487.
640.
728

778
890
1102.
1390.

0000 00000000 OOOU’IU‘ILflN O

 

T = —10.o°c

H = L = 6.0 in.

D = 1.938 in.

d = 3/8 in.

h = 1/8 in.

vs= 64.0 Z

P (lbs.) t (min.) 6C(x10 3in.)

1500 20 30.0

60 47.5

100 60.0
150 73.8
200 85.0
300 103.8
400 120.0
550 145.0
700 170.0
900 205.0
1100 241.0
1200 258.6

1980 1200 259.0
(Continued)

327

Table B-2: (Cont'd)

 

 

 

 

Sample CS—25: (Cont'd) Sample CS—26: (Cont'd)
P (lbs.) t (min.) 6C(x10"31n.) P (lbs.) t (min.) 6C(x10‘3in.)
1980 1240 290.0 1500 400 31.5
1300 350.0 500 33.0
1350 400.0 600 35.5
1400 465.0 740 38.0
1490 655.0
1490 rupture 1940 53 :g'g
100 47.5
Sample CS-26: 200 55.0
_ _ o 400 62.5
g _ DIQ-E E 6 0 , 500 75.0
‘ ‘ " ° 1“' 600 80 0
h = 1/8 in. (at z = 1.5 in.) 800 88'8
Vs= 64-0 z 900 92.5
P (lbs.) t (min.) 6C(x10‘31n.) 2500 0 97.5
1 .
1020 0 0-5 163 128 g
:8 i“: 200 140.0
4'0 300 160.0
70 4-5 450 186.2
\ 100 - 600 215.0
150 6-0 700 235.0
200 6-5 820 260.0
250 7.5
300 8.0
400 9.0 Sample CS—27:
500 10'0 T = 10.000
650 11'2 H = 5.594 in.
800 12'5 D = 2.781 in.
1000 14°C L = 5.594
1250 15.5 d = 3/8 in.
1500 16.5 h = 1/8 in.
1800 17-8 v = 64.0 2
2100 18.5 S 3
2400 20.0 P (lbs ) t (min.) 6C(x10 in.)
2800 21.0
1020 0 0-8
1500** 0 22.0 20 4.5
50 24.0 40 6.5
100 25.0 70 8.2
200 28.0 100 9.2
300 30.0 150 10.0
** Step loading procedure 258 ii.g
(Continued) 300 12.2

(Continued)

328

Table B—2: (Cont'd)

 

 

Sample CS—273 (Cont'd) Sample CS—28: (Cont'd)
P (lbs.) t (min.) 6C(x10 316.) P (lbs.) t (min.) 6C(x1o‘31n.)
1020 400 13.0 1020 50 1.0
500 13.8 100 1.4
650 14.5 200 2.2
800 15.1 300 2.8
1000 15.8 500 3.0
1310 16.8 800 4.0
1500** 0 16.9 1200 4'2
30 18.2 1460 4.5
60 18.8 1500 0 4.5
90 19.0 50 5.0
140 19.8 100 5.2
190 20.8 200 6.0
290 21.1 350 6.5
340 22.2 600 7.5
520 23.7 890 8.1
2020 0 23.8 1980 0 8.1
20 28.8 50 8.6
70 38.8 100 9.5
120 47.5 200 9-8
170 53.8 350 11.0
‘ 270 67.6 480 11-6
370 80.0 2500 0 11.8
520 98.8 50 14.0
660 115.0 100 15.4
2500 0 115.2 200 18.2
50 143.8 300 21.0
110 172.5 450 24.2
260 238.8 600 27.7
410 308.8 750 31.0
560 381.3 960 36.0
775 502.5 3000 0 37.2
3000 0 505.5 50 47.2
15 550.0 100 56.0
15 rupture 200 73'4
300 94.2
400 118.4
Sample CS—28: 500 219.7
_ o 600 317.2
1T1= 1132.8 in. 700 422.2
D = 4.0 in. 780 557 2
d = 3/8 in.
h = 1/8 in.
vs= 64.0 2

(Continued)

 

Table B—2:

(Cont'd)

Sample 08—29:

H:
d:

:1"
ll

VS:

P (lbs.) t (min.) 6C(x10_3in.)

—10.0°c

D = L = 6.0 in.
5/8 in.

0 (plain rod)
64.0 Z

 

250

520

50
100
200
300
450
600

1770

0
50
100
200
250
300
350
400
‘ 420
440
440

Sample CS—30:

.500
.825
.075
.275
.000
.000
.000

.000
300
500
200
750
500
680
380
500
11.000
rupture

mumbwwmmw NNNr—nr—IOO

Sample CS—30: (Cont'd)

 

 

 

T = -10.0°c

H = L = D = 6.0 In

d — 5/8 in

h =

vS= 64.0 7

P (lbs.) t (min.) 6C(x10—3in.)

250 50 0.25
100 0.50
200 0.87
400 1.00
600 1.12
900 1.50
1200 1.75
1400 2.00
1620 2.00

410 0 2.00
100 2.20

(Continued)

P (lbs.) t (min.) 6C(x10—31n.)
410 200 3.00
500 3.50
730 3.50
570 0 3.50
100 3.50
300 3.75
500 4.50
720 5.82
730 0 5.82
50 6.57
100 7.57
140 9.07
160 10.32
180 13.32
180 rupture
Sample CS—31:
T = —10.0°c
H = D = L = 6.0 in.
d = 5/8 in.
h = 0
vs= 64.0 2
P (lbs.) t (min.) 6c(x10—31n.)
160 50 0.50
100 0.68
200 0.88
300 1.12
400 1.37
1480 1.50
240 0 1.50
100 1.50
200 1.50
1665 1.50
360 0 1.50
100 1.62
400 1.50
820 1.50
480 0 1.50
200 1.50
400 1.50
800 1.50
1000 3.12

(Continued)

330

Table B—2: (Cont'd)

 

Sample CS-31: (Cont'd) Sample CS-32: (Cont'd)
P (lbs.) t (min.) 6C(x10 3in.) P (lbs.) t (min.) 5C(xlO—3in.)
480 1000 1.50 240 400 1.75
1180 1.50 740 1.75
600 0 1.50 925 4'25
60 1.88 925 1.75
100 1‘50 1100 1.75
200 1.50 1220 2'50
785 1.50 1220 1.00
1400 1.00
720 0 1.50 1455 2.00
400 1.50 2850 2.00
:23 é'gg 400 0 2.00
620 3.25 100 2'00
200 2.00
840 0 3.25 450 2.00
:88 €83 600 O 2.00
' 100 2.00
425 3.00 200 2.12
425 1-62 1040 2.12
700 1.62
960 o 1-62 "’0 108 iii?
5 100 1-62 200 2.50
200 1-12 480 2.50
560 1.50
560 1.62 920 0 2.50
680 1.62 50 2.88
100 2.50
1080 0 1.62 400 2.50
40 1.62 580 2.50
60 3.25
80 6.25 1080 0 2-50
80 rupture 100 2'50
200 2.75
400 3.00
Sample 05-32: 1280 3.00
O
T = -15.0 c . 1280 0 3.00
H = D = L = 6.0 in. 100 3.50
d = 5/8 in. 200 3.80
h = 0 300 4.35
vs: 64.0 2 400 5.00
_3 500 6~25
P (lbs.) t (min.) §E_gs!1_39111 600 7.50
1.00 640 8'75
240 138 1_50 665 11.50
200 1.75 665 rupture

(Continued)

 

331

Table B—2: (cont'd)

 

Sample CS—33: Sample CS—34: (cont'd)
T = -15.1°c _ C _3
H = D = L = 6.0 in. P (le-) t (mln-) 6 (x10 in.)
d = 5/8 in. 160 50 0.38
h = 0 100 0.50
Vs= 64-0 Z 200 0.80
P (lb ) ( _ 6C _3. 760 1.00
s. t min.) (x10 in.) 280 0 1.00
240 0 0.01 100 1.00
100 0.01 710 1.25
:28 3:3; 400 0 1.25
200 1.50
400 0 0.01 775 1.50
:88 6'46 520 0 1.50
1310 0.50 200 1'55
600 1.60
560 0 0.50 700 1.65
:88 8'33 640 0 1.65
610 1.00 200 1'72
300 3.50
720 o 1.00 400 6.00
100 1.00 450 7.80
. 200 1.25 500 12.00
400 1.38 510 14.00
1270 1.50 520 17.75
525 22.50
880 108 1:55 525 rupture
400 1.80
760 1.81 Sample cs—35:
1040 0 1.81 T = -6.0°c
100 2.30 H = D = L = 6.0 in.
150 2.55 d = 5/8 in.
200 4.00 h = o
200 rupture vs= 64.0 Z
P (lbs.) t (min.) 6C(x10 3in.)
Sample CS—34:
o 280 50 0.50
T = -6.1 C 100 0.75
H = D = L = 6.0 in. 200 1.00
d = 5/8 in- 400 1.20
h = 0 530 1.25
vS 64.0 A 400 0 1.25
(Continued) 100 1.25
200 1.44

(Continued)

332

Table B—2: (Cont'd)

 

 

Sample CS—35: (Cont'd) Sample CS-36: (Cont'd)
P (165.) t (min.) 6°(x10‘3in.) P (lbs.) c (min.) 6C(x10‘3in.)
400 400 1.50 520 200 14.0
860 1.75 300 18.0
450 22.8
520 0 1.75
100 2.00 600 26'7
300 2.25 1388 32'?
50 2. °
0 35 1250 40.1
640 0 2.35 1500 44.5
100 2.60 1750 48.0
400 2.68 2010 52.0
660 .
2 70 1080 0 52.0
720 0 2.70 50 74.5
100 3.00 100 92.0
200 3.08 200 123.2
400 3.28 300 152.0
600 3.50 400 178.3
820 3.60 500 203.2
800 0 3.60 650 239.5
800 275.8
100 3.80
1000 321.1
200 4.10
730 4 30 1250 377.7
. ' 1400 411.7
100 4.50 1560 0 437.7
200 4-80 50 507.8
:00 g»;g 100 567.7
00 -20 150 627.5
500 5- 200 690.2
1040 0 5.20 250 760.2
40 5_70 280 810.3
60 6.58
70 10.80
S— :
70 rupture §29219_9_6;Z_
T = —10.1 c
H = L = 5.625 in.
Sample CS—36: D = 6.0 in.
T — -10.0°c d = 3/8 in'
H=D=L=6.01n 11:0
d = 3/8 in Vs= 32-3 2
h = 1/8 in c -3.
vs: 44.5 : £L112251. 2.193211. §_1§19__3911
_3 520 200 0-5
P (lbs.) t (min.) 6°(x10 i 375 1.0
520 50 5.0 415 5-0
100 8.5 480 18.1
(Continued)

(Continued)

333

Table B-2: (Cont'd)

 

 

Sample CS-37: (Cont'd) Sample CS—38: (Cont'd)
P (lbs.) t (min.) 6C(x10 3in.) P (lbs.) t (min.) 6C(x10'31n.)
520 530 26.5 1080 300 76.25
600 37.8 450 90.00
650 45.5 600 99.25
800 63.3 800 112.50
1000 87.1 1000 125.75
1200 110.9 1250 138.72
1400 134.7 1500 151.68
1700 158.4 1600 156.88
2000 181.2 1800 164.63
2200 195.2 2050 170.63
2370 207'3 1560 0 172.13
1080 0 213.2 50 188.12
50 300.0 100 201.38
100 383.2 200 228.13
150 485.2 300 251.40
200 535.7 450 290.60
250 613.2 600 330.60
300 703.2 800 385.62
1000 443.13
_m.____ 1. 1233 322:8
T = -10.0°c 1590 668.10
H = L = 3°875 ln‘ 1980 0 668.15
D = 6-0 %n' 50 730.62
d = 3/8 }n- 100 793.12
h = 1/8 1“- 150 868.12
VS: 51.6 Z
P (lbs.) t (min.) 6C(x10'31n.) 833219_g§:39;
520 100 0.01 T = —10.0°c
290 1.25 H = L 5.875 in.
350 3.50 D = 6.0 in.
400 6.25 d = 3/8 in.
500 10 00 h = 1/8 in.
650 14.38 vs= 48.6 Z
800 17.50 C -3.
1000 21.80 311922 _t_<_m_i.n_> W
1250 28.75 520 100 1.25
1080 0 33.12 300 5.00
50 47.00 400 7.60
100 55.00 600 10 80
200 66.88 800 15-1

(Continued) (Continued)

334

Table B—2: (Cont'd)

 

 

Sample CS-39: (Cont'd) Sample CS-40: (Cont'd)
P (lbs.) t (min.) 6C(x1o'3in.) P (lbs.) t (min.) 6C(X10_3in.)
520 1000 18.9 520 600 21.5
1250 23.2 800 24.0
1500 27.6 1000 27.1
1800 31.8 1250 29.0
2100 37.0 1500 31.5
2400 40.8 1800 32.8
2700 44.5 1965 34.0
3015 47.5 720 0 35‘2
1080 0 51.0 50 36.5
50 67,0 100 37.8
100 81.2 200 40.2
200 106.0 300 41.5
300 129.1 450 44.0
450 151.0 600 46-5
600 192.2 800 49.0
800 234.8 1000 51.5
1000 276.0 1200 54-0
1200 316.0 1425 57-7
1400 361.0 1080 O 57.8
1560 0 362.0 50 62-8
. 50 424.5 100 65.3
100 487.0 200 7l~5
150 537.0 300 76-6
200 612.0 450 82-8
280 712.0 600 89-1
260 837.0 800 96-5
1000 104.0
1200 111.5
Sample CS—40: 1410 119,0
0
T = -10-0 C 1240 0 120.3
H = L = 6.125 in. 50 124 0
D = 6.0 in. 100 127.8
d = 3/8 in. 200 134.8
h = 1/8 in. 300 142.8
vs= 52.8 2 450 152.8
' c _3 600 164.0
P (lbs.) t (min.) 6 (x10 in.) 745 175.3
520 50 6.0 1560 0 175.5
100 9.0 50 190.5
200 12.8 100 205.5
300 15.2 200 234.5
450 17.8 300 263.0
, 450 318.0
(Continued) 600 393,0
700 454.2

820 585.5

335

Table B—2: (Continued)

Sample CS—41: Sample CS—42:
T = -1o.0°c T = -10.0°c
H = D = L = 6.0 in. H = L = 6.0 in.
d = 3/8 in. D = 1.938 in.
h = 1/8 in. d = 3/8 in.
vs= 17.5 2 h = 1/8 in.

v = 64.0 T

P (lbs.) t (min.) 6C(x10"3in.) S

 

P (lbs.) t (min.) 6°(x10'3in.)

 

280 100 0.12
200 0.25 520 100 1.0

300 0.50 200 1.0

600 0.50 300 1.0

935 0.20 500 1.5

360 0 0.20 615 3.0
100 0.25 720 0 3.0

200 0.50 200 3.3

400 0.50 500 3.5

500 0.75 800 4.0

740 1.00 920 5.0

1222 6:33 1080 0 5.0

50 8.2

520 0 0.50 100 12.0
100 1.50 200 18.0

‘ 200 2.62 300 21.5

300 4.50 400 25.5

400 7.75 500 28.5

500 11.00 615 31,5

938 :3'32 1240 0 31.7

840 61:00 100 37-0

200 40.9

600 0 61.00 300 44.0
100 111.00 400 47.7

200 161.00 550 53.4

300 211.00 700 58.4

400 261.00 850 59.6

500 311.00 1035 65.9

600 367'00 1560 0 65.9

750 448.00 100 75.9

680 0 448.00 200 83.4
100 555.00 300 91.5

200 655.00 450 104.0

300 780.00 600 117 1

400 918.00 750 128.4

500 1038.00 940 144.6

1840 0 144 6

50 158.4

(Continued)

336

Table B—2: (Cont'd)

 

 

 

Sample CS-42: (Cont'd) Sample CS—43: (Cont'd)
P (lbs.) t (min.) 6C(x10—3in.) P (lbs.) t (min.) 6C(x10'3in.)
1840 100 170.9 1840 100 87.0
200 192.1 200 95.8
300 217.0 300 103.2
400 233.4 450 112.0
500 267.0 600 122.0
665 317.0 800 133.0
2.20 0 317.0 iggg 123-3
50 364.5 '
100 426.4 2200 0 155.0
150 492.0 50 160.0
210 604.5 100 168.8
210 rupture 200 175.0
300 187.5
M :38 3‘23?
T = -10.00C 780 237.5
H = = . ' .
D = 1: 0 :n0 in 2560 0 237.8
_ ' , ' 50 253.0
d " 3/8 ln. 1
h = 1/8 in 00 271.0
V = 64 0 7- 200 293.0
s - 3 300 317.0
. —3. 400 340.0
P (lbs.) t (min.) 6C(x10 in.) 500 363.0
1080 50 6.8 600 385.0
100 13.0 670 403.0
200 21.0 2840 0 405.0
300 27-0 50 433.0
450 31-8 100 485.0
600 35-8 200 509.0
800 40-8 300 563.0
1000 44-5 400 613.0
1150 46-4 500 663.0
50 52.0
100 54.5 _ .
300 63.2 T = 10.0 C
450 68.5 H = L = 3.0 in.
600 72.0 D = 6.0 in.
680 75.0 d = 3/8 in.**
h = 0.015 in.**
1840 0 76.0 V = 64 0 7
50 83.2 S ' ° (Continued)

(Continued) ** Standard deformed bar

337

Table B—2: (Cont'd)

 

 

Sample CS—44: (Cont'd) Sample CS-45: (Cont'd)
P (lbs.) t (min.) 5C(x10 3in.) P (lbs.) t (min.) 6C(x10‘3in.)
1080 20 2.5 520 40 9.3
40 5.0 100 16.0
80 7.8 200 21.9
130 9.8 300 26.5
200 10.8 450 31.5
260 13.0 600 35.0
350 14.9 800 40.0
450 16.8 1000 43.5
600 18.0 1205 45.8
:38 :3'2 800 0 46.5
1130 21.8 50 52.5
100 55.8
1560 0 21.8 200 61.2
40 24.2 300 66.2
100 26.1 450 72.7
150 28.0 800 88.3
250 30.0 1000 95.8
400 34.8 1130 99.3
600 35'5 1160 0 100.0
800 38.0
50 115.3
\ 1000 40.2 100 123 8
1200 42.0 2 '
1365 43.5 00 141°C
300 156.0
1980 0 43.5 450 174.3
100 48.7 600 192.0
200 53.0 800 216.0
350 58.8 1000 241.0
500 64.2 1200 266.0
700 74.2 1400 292.3
950 88.0 1535 313.0
1270 124.2 50 403.0
2560 0 124.2 100 465.0
5 144.0 200 572.0
5 rupture 300 671.0
360 743.0
400 778.0
Sample CS-45: 455 896.0
T = —2.0°c
H = D = L = 6.0 in.
d = 3/8 in.
h = 1/8 in.
vs= 64.0 2

(Continued)

 

*****

 

 

 

.4

STATE

 

1293

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1') 3 O 3 0 8 2 0 7 51