NAT'L INST. OF STAND & TECH Reference

A 11 101 7 7 7 7 ( 5

NBSIR 82-2484

internal Strain, Deformation, and

Failure of Large Scale Pullout

Tests in Concrete

NBS

Publications AlllDb 2 2 2 2 a b

U.S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Building Technology

Washington, DC 20234

March 1982

Issued May 1 982

1

‘ S. DEPARTMENT OF COMMERCE

IONAL BUREAU OF STANDARDS

100

,1156

32-2484

1932

NBSIR 82-2484

INTERNAL STRAIN, DEFORMATION, AND

FAILURE OF LARGE SCALE PULLOUT

TESTS IN CONCRETE

national bwmw

OF STANDARDS

LIBRARY JUL 7 1982

/DO

William C. Stone

U S. DEPARTMENT OF COMMERCE

National Bureau of Standards

National Engineering Laboratory

Center for Building Technology

Washington, DC 20234

March 1982

Issued May 1 982

U S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director
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ABSTRACT

A study was performed to obtain detailed experimental data on crack propagation

and internal strain distribution for the pullout test method for non-destructive

strength evaluation of concrete. A 12:1 scaled-up pullout test was designed,

using a commercial pullout insert for the prototype dimensions, and was instrumented with small waterproof embedment strain gages so as to obtain internal

strain profiles at critical locations. Two large scale specimens were tested

with apex angles falling at the upper and lower bounds currently recommended

in ASTM C-900. Two dimensional axisymmetric finite element analyses were performed for the two experimental specimens and the results were compared with

measured strains for load stages below the onset of internal cracking.

The results showed good correlation between the analytical and experimentally

observed strains. The experimental data indicate that internal cracking, and

the formation of the failure surface, are principally governed by the tensile

strength of the concrete. The failure surface appears to be formed at 65 percent of ultimate load. Beyond this point, it is likely that the entire load is

carried by the mechanism of aggregate interlock. Ultimate failure occurs when

all aggregates mechanically bridging the failure surface pullout from the

retaining cement paste. It is likely that the pullout test measures the shear

strength of the cement paste or mortar which binds the concrete together.

Keywords: Concrete; crack propagation; failure surface geometry; failure

theory; finite element method; internal strain; laboratory testing;

large scale models; mathematical mode; pullout test; stress contours.

iii
EXECUTIVE SUMMARY

A study was performed at the National Bureau of Standards to obtain detailed

experimental data on crack propagation and internal strain distribution for

method for non-destructive strength evaluation of concrete. A 12:1 scaled up

pullout test was designed, using a common commercial pullout insert for the

prototype dimensions, and was instrumented with small waterproof embedment

strain gages so as to obtain strain profiles at critical locations. Two large

scale concrete specimens were tested with apex angles falling at the upper and

lower bounds currently recommmended in ASTM C 900. Two dimensional axisyrametric

finite element analyses was performed for the two experimental specimens and

the results were compared for load ranges below the onset of internal cracking.

The results of these tests indicate that there are two principal internal crack

systems for the pullout test. Radial cracks form parallel to the vertical

planes extending radially from the disk stem (r-z plane). These cracks initiate randomly along the circumference of the pulling stem at the top concrete

surface and propagate away from the stem as a curved front towards the primary

failure surface. Circumferential cracks form the primary failure surface—the

outer surface of the pullout "cone”. These cracks begin at the disk edge and

propagate towards the reaction ring as the load is increased. Correlation

between experimental and analytical studies have indicated that the circumferential failure surface very nearly follows the principal compressive stress

trajectory from the disk edge to the reaction ring. This means that the principal tensile stresses act almost normal to the failure surface. Experimental

data show that the compressive strains along this path are insufficient to cause failure of concrete by crushing, and additionally, that the tensile

strains normal to the failure surface do substantially exceed the limiting tensile strain of concrete. Circumferential cracking is thus primarily governed

by the tensile strength of the concrete. This, however, is not the principal

strength property of concrete that is measured by the pullout test.

A study of the discontinuities (sharp changes in slope, or large strain excursions) in the load strain histories of the embedded strain gages revealed three

distinct phases in the failure sequence of a pullout test. These occur between

30 to 40 percent of ultimate, 60 to 70 percent of ultimate, and from 80 to 100

percent of ultimate load. The three phases have been hypothesized in this

report to constitute the initiation and propagation of circumferential cracking

along the failure surface, completion of circumferential cracking along the

failure surface, and a progressive ultimate failure by degradation of aggregate

interlock across the failure surface. The second of these—that of completion

of circumferential cracking—was experimentally shown to take place via propagation beginning at the disk edge and ending at the inside edge of the reaction

ring. Completion of circumferential cracking occurs at approximately 65 percent

of ultimate load, regardless of variation in apex angle. Beyond this load, the

upward force is resisted by embedded sized aggregate particles mechanically

bridging the circumferential failure surface. An idealized discrete failure

model developed in this study, based on the assumption that failure occurs when

all such aggregate particles shear through the retaining cement paste, gave

ultimate load estimates that were encouraging. It is likely that the pullout

test directly measures the shear strength of the cement paste which binds the

concrete together.

iv
ACKNOWLEDGMENTS

The author is greatly indebeted to Dr. Nicholas J. Carino for his persistent

help in the experimental investigation and for his advice and counsel both in

the planning of the project as well as the writing of this report.

Many thanks are extended to the technical staff of the laboratory, including

Mr. Frank Rankin, Mr. Tom Ruschell, and Mr. Randy Williams, without whom the

experimental investigation could not have been carried out.

Lastly, thanks are due to Dr. Kyle A. Woodward for making available the program

TEXD used for data acquisition and reduction during the laboratory tests.

v
TABLE OF CONTENTS

ABSTRACT iii

EXECUTIVE SUMMARY iv

ACKNOWLEDGMENTS v LIST OF TABLES viii

LIST OF FIGURES ix

1. Introduction . » 1

2. Procedure 9

2 . 1 General 9

2.2 Instrumentation 9

2.2.1 Micro Embedment Strain Gages 9

2.2.2 Slip Gages 18

2.3 Data Acquisition 25

2.4 Specimen Propagation 28

2.5 Loading System 37

2.6 In Place Strength of Concrete 41

2.7 Test Procedure 46

3. RESULTS 49

3.1 Load Deflection Curve 49

3.2 Shape of Failure Surface 53

3.3 Comparison of Experimental and Analytical Strain

Distributions 53

3.3.1 Axial Strain 62

3.3.2 Radial Strain 62

3.3.3 Circumferential Strain 62

3.3.4 Vertical Strain Along Disk Edge 62

3.3.5 Radial Strain on Vertical Line Above Disk Edge 70

3.3.6 Radial Strain Along Disk Face 70

3.4 Elastic Behavior 70

3.4.1 Stress Contours 70

3.4.2 Principal Stress Profiles Along Failure Surface 76

3.4.3 Principal Stress Trajectories 76

3.5 Crack Propagation Analysis 88

3.5.1 Post Cracking Behavior 88

3.5.2 Discontinuity Histograms 94

3.5.3 Crack Propagation Sequence 100

vi
TABLE OF CONTENTS (Continued)

Page

4. Discussion 107

4.1 Failure Theory 107

4.2 Aggregate Interlock Model 108

5. Summary and Conclusions 115

5.1 Failure Sequence 115

5.2 Internal Strains 115

5.3 Principal Stresses 117

5.4 Failure Surface Shape 117

5.5 Failure Mechanism 118

5.6 Recommendations for Further Research 118

REFERENCES 120

vii
LIST OF TABLES

Page

Chapter 2

2.1 Linear scale factors for direct modelling 10

2.2 Concrete mix specifications for as-placed material 37

2.3 Cylinder strength results 44

Chapter 3

3.1 Ranges in percent of ultimate load which mark pronounced changes ... 49

3.2 Ranges in percent of ultimate load of primary discontinuity zones observed by embedment gages 100

viii
LIST OF FIGURES

Introduction

1.1 Schematic representation of the Pull-out test

1.2 Ottosen's finite element cracking analysis of the Pull-out test

1.3 Radial and Circumferential cracks in the Pull-out test

Chapter 2

2.0 Scale comparison between prototype and macro model

2.1 Steps in the fabrication of micro embedment gages

2.2 General view of instrumented pull-out disk showing, top to bottom:

RV,RN (radial gages), TC (triaxial compression, radial) and CW

(circumferential gages)

2.3 Close up view of RV (right) and RN (left) gages prior to casting

2.4 ET (edge tension) gages measure vertical strain along side face

of disk

2.5 ANE (axial, north east quadrant) gages

2.6 Gages are wired to piano wire grid using flexible copper wire

2.7 Typical end detail for a circumferential gage string

2.8 Typical assembled gage strings ready for mounting on disk

2.9 Overall view of instrumented disk prior to casting

2.10 Slip wire details at disk connection

2.11 External hardware for slip gage

2.12 Alternate system for monitoring vertical displacement of disk

2.13 Pull-out insert specifications

2.14 Assembled loading tendon and pull-out insert (specifications)

2.15 Mating disk and loading head (with hydraulic jacks) prior to welding

2.16 Assembled loading tendon and pull-out insert (as built)

2.17 Supplemental reinforcement specifications

ix
LIST OF FIGURES (continued)

2.18 As built supplemental reinforcement

2.19 Perforated angle frame used to support gage strings. Cross channel

frame is used to stabilize and level disk prior to casting

2.20 Checking embedment gage resistance prior to casting

2.21 Jacking-Counter Pressure system (schematic)

2.22 Compression frame specifications

2.23 Jacking head and ram cluster specifications

2.24 Calibrating Jacking-Counter Pressure system

2.25 Seating Jacking-Counter Pressure system on hydrostone grout prior

to leveling

2.26 Pre-tensioning loading tendon strands

2.27 Placing C.I.P.P .O.C.s in freshly cast specimen

2.28 Slave cylinders in temperature controlled water bath

2.29 Temperature histories in Pull-out specimen # 1

2.30 Temperature histories in Pull-out specimen # 2

2.31 Graphics terminal used to monitor real time strain profiles

during testing of Specimens 1 and 2

2.32 Acoustic emmission device and amplifier used to provide advance

warning of impending failure

2.33 Overall view of lab test set-up

Chapter 3

3.1 Vertical disk displacement vrs load, Specimen 1, 2 a = 70

3.2 Vertical disk displacement vrs load, Specimen 2, 2 a = 54

3.3 Normalized load deflection plot

3.4 Experimental macro pull-out "cones"

x
LIST OF FIGURES (continued)

3.5a Measured failure surface for Specimen 1, 2 a = 70

3.5b Measured failure surface for Specimen 2, 2 a = 54

3.6 Typical embedded strain gage history

3.7 2D axisymmetric finite element mesh for Specimen 1

3.8 2D axisymmetric finite element mesh for Specimen 2

3.9 Axial strain along (idealized) failure surface : experimental

vrs FEM at P = 50 kips ; Specimen 1

3.10 Axial strain along (idealized) failure surface : experimental

vrs FEM at P = 50 kips; Specimen 2

3.11 Radial strain along (idealized) failure surface : experimental

vrs FEM at P = 50 kips; Specimen 1

3.12 Radial strain along (idealized) failure surface : experimental

vrs FEM at P = 50 kips; Specimen 2

3.13 Circumferential strain along (idealized) failure surface: experimental

vrs FEM at P = 50 kips; Specimen 1

3.14 Circumferential strain along (idealized) failure surface: experimental

vrs FEM at P = 50 kips; Specimen 2

3.15 Vertical strain along disk edge: experimental vrs FEM at P = 50 kips;

Specimen 1

3.16 Vertical strain along disk edge: experimental vrs FEM at P = 50 kips;

Specimen 2

3.17 Radial strain along disk face: experimental vrs FEM at P = 50 kips;

Specimen 1

3.18 Radial strain along disk face: experimental vrs FEM at P = 50 kips;

Specimen 2

3.19 Radial strain along vertical line above disk edge: experimental

vrs FEM at P = 50 kips; Specimen 1

3.20 Radial strain along vertical line above disk edge: experimental

vrs FEM at P = 50 kips; Specimen 2

3.21 Maximum principal stress contours for Specimen 1

xi
LIST OF FIGURES (continued)

3.22 Maximum principal stress contours for Specimen 2

3.23 Minimum principal stress contours for Specimen 1

3.24 Minimum principal stress contours for Specimen 2

3.25 Sigma Max (maximum principal stress) along failure surface

(from axisymmetric FEM analysis)

3.26 Sigma Min (minimum principal stress) along fasilure surface

(from axisymmetric FEM analysis)

3.27 Tau Max (maximum shearing stress) along failure surface

(from axisymmetric FEM analysis)

3.28 Sigma Theta (circumferential stress) along failure surface

(from axisymmetric FEM analysis)

3.29 Elastic state of stress along idealized failure surface

3.30 Principal stress trajectories for Specimen 1 (from axisymmetric FEM analysis)

3.31 Principal stress trajectories for Specimen 2

(from axisymmetric FEM analysis)

3.32 Elastic deformation of Specimen 1 at P = 50 kips

(from axisymmetric FEM analysis)

3.33 Circumferential stress contour for Specimens 1 and 2

at P = 50 kips. (from axisymmetric FEM analysis)

3.34 Positive strain excursion indicating gage crosses crack surface

3.35 Negative strain excursion (reversal) indicating crack has gone around gage

3.36 Axial gage showing strain reversal — crack has passed around

gage, unloading it

3.37 Typical monotonic non-linear axial strain history

3.38 Discontinuity profiles for Specimens 1 and 2

3.39 Crack propagation sequence for Specimen 1

3.40 Crack propagation sequence for Specimen 2

3.41 Axial strain along failure surface (experimental) at ultimate load

for Specimen 1

xii
LIST OF FIGURES (continued)

3.42 Axial strain along failure surface (experimental) at ultimate load

for Specimen 2

Chapter 4

4.1 Basis for aggregate shear failure calculations

4.2 Basis for calculation of n

4.3 Generalization of aggregate interlock failure

xiii

1

.

INTRODUCTION

The most critical period in the life of a reinforced concrete structure is

during its construction. During this time the concrete is weak and, if unanticipated construction loads are applied, catastrophic collapse may occur. In

the past the development of compressive strength of concrete in such structures

has normally been determined by the testing of field cured specimens. However,

it is now generally recognized that results obtained from such specimens may

vary greatly from the in-place strength of the structure owing to different

casting, compaction, and curing conditions. There exists, therefore, a need

for a method to accurately measure the in-place strength of concrete.

In recent years the pullout test, initially proposed in the U.S.S.R. in 1934,

has attracted much attention as a possible method for measuring in-place

strength and a number of patents have been registered in various countries.

In the U.S.A. Richards [17] has advocated the use of these tests on structural

concrete members and a tentative standard method for such tests has been

adopted by the American Society for Testing and Materials [22]

.

Briefly, a pullout test measures, with a specially constructed center pull tension jack

which reacts against the concrete surface through a reaction ring of specified

geometry, the force required to pull out a specially shaped steel insert (consisting of one or two pieces) whose enlarged end has been cast into the concrete. Because of its shape the steel insert pulls out a cone of concrete.

The precise geometry of this "cone" is determined by a number of factors,

including the diameter of the "disk" (enlarged portion), the depth of embedment,

the diameter of the "stem" (pulling rod), and the diameter of the "counterpressure" (reaction) ring which forms the base of the pulling apparatus (see

figure i.l). The method has a number of advantages over other methods for

in-place strength evaluation. These advantages are:

1. Tne measurements are simple and easy to carry out.

2. It costs considerably less than other in-situ destructive tests, such as drilled cores.

3. The results are available within minutes after the test.

4. It is superior to other in-situ non-destructive methods such as the Rebound

Hammer (ASTM C805) and the Windsor Probe (ASTM C803) because a greater depth

and volume of concrete is tested. The pullout test is a direct measure of

concrete strength, whereas the rebound Hammer and Windsor probe measure other

properties that may be related to strength.

Contrasting these attributes it is generally noted that:

1. The pullout test must be planned in advance of concreting. Other methods,

such as the Windsor Probe can be performed anywhere after the concrete has

hardened

.

2. The pullout test does not measure the strength of the interior of mass concrete because the typical insert usually does not extend further than

1
Figure 1.1 Schematic representation of pull-out test in concrete

2
1.5 inches (38 ram) below the surface. A novel technique for obtaining

pullout measurements at greater depth has recently been proposed by

Richards [8]

.

3. For accurate results it is presently recommended [6,22] that the relationship between the pullout strength and the compressive strength should be

determined for each site and for each type of concrete and aggregate

size. (For such calibrations the in-test variation of the pullout test is

low and on the same order as for standard cylinder tests [6].) This requires

development of a calibration curve before construction begins.

4. The resulting damage to exposed concrete surfaces must be repaired. It is

still a non-destructive test since the structural member need not be

discarded following the test.

The state of stress in the pullout test is complex and difficult to analyze.

Considerable controversy has arisen over just what strength property of con- crete is actually being measured, and what constitutes the physical mechanism

of failure. Prior to 1975 it was generally believed that the concrete was simultaneously in tension and in shear [4] , and it has been suggested that the

pullout force is a measure of the direct shear strength of concrete. Alternatively, some believe [5] that the pullout test measures the punching shear

strength of concrete and that the name should be changed to the Punching-Shear

test. Interestingly, the Danish commercial version of the pullout test is

called the Lok-Test; Lokning is Danish for punching. "Lok strength" was thus

defined as the force required to punch out a small piece of concrete of specified geometry. In a recent survey of pullout tests conducted at a number of

construction sites Bickley [6] reported a high degree of correlation between

the pullout strength and the compressive strength of concrete. He went on to

state that it is, therefore, likely that the pullout test measures a property

of concrete which is either compressive strength, or has a constant relationship

with compressive strength. Whatever the true strength property being measured

is, it is clear that the pullout test, if it is to be widely accepted, must

ultimately present its findings in the form of an equivalent compressive

strength, since this is the basis of design. There are thus two avenues for

improving the method: to determine the strength property being measured and

develop a correlation between this strength property and the compressive

strength; and to experimentally improve correlation with compressive strength

by modifying the test. Using the latter technique Kierkgaard-Hansen [2] performed an extensive experimental investigation in Denmark during the 1960s.

The results of this study indicated that by decreasing the apex angle (see

figure 1.1), one could achieve a higher correlation between the pullout force

and the compressive strength of concrete. This was achieved by varying the

diameter of the counterpressure ring. Lacking a reaction device to control

the upper diameter of the failure surface, a pullout specimen will fail in

tension and exhibit a fracture surface with a trumpet shaped geometry rather

than an approximately straight sided conic frustum. Kierkegaard-Hansen

later adopted a fixed geometry with a 25 mm disk, 25 mm embedment depth, and a

55 mm l.D. reaction ring. Using this device, he established empirical equations for two concretes with differing aggregate sizes. These equations are now

seeing wide commercial use in Europe. The questions of optimum geometry, the

3
reason for different equations for different aggregate sizes, and the ever debatable mechanism of failure gave rise to two important analytical works in

the mid 1970s.

Jensen and Braestrup l 1 ] showed by means of plasticity theory, that the pullout

force is directly proportional to the compressive strength of concrete. Unfortunately, their analysis was based on a series of assumptions having questionable validity. First, there is some doubt that plasticity theory is applicable

to an inherently brittle material, especially when tensile stresses are present

as is the case of the pullout test. Secondly, their derivation assumes that

concrete obeys the Mohr-Coulomb failure theory with a tension cut-off. The

Mohr-Coulomb theory is not the most applicable failure theory for concrete subjected to multiaxial stresses [9,10,11]. Derivation further assumes that the

state of stress on the assumed failure surface would result in "sliding" failure;

the possibility of failure by separation is dismissed. There is no discussion

on the actual stress state within the concrete during the test and, thus, there

is no explanation of why "sliding" failure is the correct failure mode. Finally,

it is assumed that the failure surface defines a conic frustum, whereas tests

indicate that the failure surface is "trumpet shaped" [2]

.

These inconsistencies render their conclusion somewhat less convincing.

In a later study, Ottosen [ 7 J analyzed the pullout test by means of an axisymmetric nonlinear finite element computer program. This analysis followed the

progression of radial and circumferential cracking by means of an iterative

smeared cracking procedure: small load increments were employed and at each

step the state of stress was compared with a specified failure criterion for

concrete. If cracking had occurred the iteration was repeated using the redistributed stress state until a stable crack configuration was achieved. Ottosen

used both the Mohr-Coulomb failure criterion, as had Jensen and Braestrup, and

a more accurate criterion which accounted for strain hardening and softening

in the pre and post failure region, respectively. Failure was determined by

lack of convergence after a specified number of interations. The analysis

showed that circumferential cracks —the ones which form the surface of the

pullout "cone" —begin at the disk edge at about 15 percent of ultimate and

propagate towards the reaction ring with increased load. Furthermore, these

cracks had reached to the reaction ring by 65 percent of ultimate load. Thereafter, Ottosen states that the load is carried by a "compression strut" (see

figure 1.2) —a zone of uncracked concrete which extends from the disk to the

reaction ring between two parallel circumferential cracks —and that ultimate

failure is governed by compressive failure of this strut. From this he con- cluded that the pullout test directly measures the compressive strength of

concrete. The analysis also indicated the formation of radial cracks which

begin at low load levels at the intersection of the top concrete surface and

the disk stem and propagate towards the the circumferential failure surface.

These do not significantly influence the failure load since they form in planes

perpendicular to the principal failure surface (see figure 1.3).

Ottosen' s study is one of the most ambitious analytical attempts to date to

resolve the internal behavior of the pullout test and the results are encouraging. However, a number of inconsistencies cast doubt on the basis for his

primary conclusion: that the pullout test directly measures the compressive

4
REACTION RING

LINE OF SYMMETRY

Figure 1.2a Ottosen's axisymmetric finite element mesh of the pull-out test

d] LOADING = 98% CL

Figure 1.2b Crack development with increasing loadings. The loading is

expressed in relation to the predicted failure load

Figure 1.2 Ottosen's finite element cracking analysis of the pull-out test

5
"P

a) RADIAL (RZ) CRACKS

6r = RADIAL STRAIN

6a = AXIAL STRAIN

6 0 = CIRCUMFERENTIAL

STRAIN

b.) CIRCUMFERENTIAL ( 0 ) CRACKS FORM

PRIMARY FAILURE SURFACE

Figure 1.3 Radial and circumferential cracks in the pull-out test

6
strength of concrete. First, the model assumes perfect bond between the pullout

disk and the surrounding concrete. This is an unlikely boundary condition,

since most contractors coat the pullout inserts with oil prior to casting the

concrete [8]. This serves as a bond breaker which would prohibit vertical load

transmission through the side and bottom faces of the disk. The difference in

the state of stress is significant. Lacking bond on the side and bottom faces,

the load will be transmitted directly to a re-entrant corner in the concrete.

Loading situation will give rise to a stress concentration in the concrete

adjacent to the disk edge. Furthermore, no evidence of a compression strut

failure has been detected in physical test. If such a compression strut did

control failure, one would expect to find an annular frustum of crushed, powdered concrete sufficiently thick to have carried the ultimate load. For a typical commercial pullout test (reaction ring diameter: 2.16" (55 min); disk

diameter: 1" (25 mm); embedment depth: 1" (25 mm)) the thickness of such a strut would have to be on the order of 1/4 inch (6 mm) and it would be highly

noticeable upon completion of the test. The fact that it is not present leads

to the question: if the failure mechanism was incorrectly predicted by

Ottosen's analysis, what assumptions and boundary conditions need to be

modified for further study?

More imposing, however, can we use a continuum theory analysis to model the

failure of the pullout test, or can it only be used up to a certain point

before a discrete failure mechanism governs? An analytical model of a complex

structure must be calibrated so as to reproduce both the actual internal strain

distribution as well as the actual overall load deflection history if it is to

have any validity. This can only be accomplished via calibration with known

experimental data. No such internal strain or load deflection data on the

pullout test existed at the time of Ottosen's study.

In order to provide an experimental basis for future analytical work and to

gain an understanding of the failure mechanism during the pullout test it was necessary to perform a detailed experimental investigation on a level here-tofore unattempted. The NBS study reported herein, sought to provide the following necessary data:

1. An experimental load-deformation history for the pullout test.

2. An experimental record of the internal strain distribution in the

vicinity of the critical failure surface.

3. An experimental record of the internal crack propagation sequence as a function of load.

4. A comparison of the internally measured strains in the pre-cracked state

with an axysymmetric linear elastic finite element analysis.

5. Experimental quantification of the change in the shape of the failure

surface as a function of the apex angle.

7
Together with future experimental and analytical work to be performed in a second study, the ultimate objectives of the NBS investigation of the pullout

are to answer the following questions:

1. What strength property of concrete is measured by the pullout test?

2. What modifications of the current standard test procedure are required to

produce the most reliable results?

8
2 . EXPERIMENTAL PROCEDURE

2.1 General

The primary impediments to gathering precise experimental data from "inside” a pullout test lay in the scale of the test subject. At present, the majority of

the commercially available units have adopted about 1 inch (25 mm) diameter

disk with an embedment depth of the same dimension. Operating within these

geometric restrictions, instrumenting the specimens would prove difficult. To

circumvent this obstacle a reverse-modelling procedure was employed. Scale

modelling of structural systems in the laboratory is often utilized to increase

the efficiency of an experimental investigation. Typically, the technique calls

for a reduction in the size of the structure. For any given model, similitude

requirements must be derived to relate the prototype to the model structure.

If the requirements are accurately met the model can be tested and its results

used to predict the behavior of the prototype. In direct modelling, where the

material properties for the prototype and the model are the same, the prototype

and model are related by only the linear scale factor, Sl. Table 2.1 shows the

relationship generally adopted to relate the prototype properties to the model

properties for the direct modelling case. The reverse (or "magnified") modelling procedure utilizes the same scale factors as shown in table 2.1, but is

now an integer rather than a fraction.

The object then was to "scale up" the prototype pullout test to the maximum

possible degree, as this would facilitate the placement of the internal instrumentation. The "maximum degree" of scaling was limited primarily by cost con- siderations and available loading system components at NBS. The final scale

adopted was 12:1 and the prototype dimensions were taken from a standard commercially available insert (see figure 2.0). This resulted in a pullout insert

with a disk measuring 12 inch (305 mm) in diameter and 4 inch (102 mm) thick.

The counterpressure ring had an inside diameter of 26 inch (660 mm). The depth

of embedment, h, was varied so that the apex angles (see figure 1.1) for the

two specimens tested fell at the upper and lower bounds specified in ASTM C900.

Specimen #1 had h = 10" (254 mm) and specimen #2 had h = 13.5" (343 mm). For

the lower limit of a =27° the pullout force was estimated at 900 kips (4005 kN)

for concrete with a compressive strength of 4000 psi (27.6 MPa). This was considered to be the maximum desirable load while maintaining a sufficient

safety factor on the loading apparatus.

2 .2 INSTRUMENTATION

2.2.1 Micro-Embedment Strain Gages

To measure the internal strain distribution of the pullout test, it was necessary to develop a reliable method by which the strain could be monitored

without disrupting the strain field. It was also desired to have a sufficient

number of measurements along the critical path to completely define a particular

strain profile between the disk edge and the counterpressure ring. This meant

that the measuring device would have to be very small and capable of picking

up the strain in the concrete; i.e. it would require a positive embedment in

the concrete in such a manner as to preclude the possibility of slippage.

9
Table 2.1 Relationship Between Model and Prototype Structures

(Direct Modeling Case)

Variable *1

Modulus of Elasticity 1

Strain, Stress 1

Linear Dimensions SL

Deformations S L

Pressure 1

Concentrated Load (sL

) 2

Moment (sL

) 3

Self-Dead Weight (1/SL)

Steel Reinforcement Area (sL) 2

Sl = Linear Scale Factor

Model Variable = K x prototype variable

10
11

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15

Figure

2.1g

Apply

thin

layer

of

butyl

Figure

2.1h

Wrap

butyl

ruber

with

rubber

waterproofing

compound

thin

teflon

sheet

over

center

portion

of

gage
CO O

a> 05

a) _c

coCO

16

Figure

2.1i

Insert

gage

into

1.25

inch

length

Figur

2.1j

Schematic

drawing

of

insert

of

1/4

inch

diameter

heat

shrink

gage.

All

units

are

in

tubing

and

seal

by

heating.

Attach

inches

nut

and

washer

to

each

end
No such strain measuring devices were commercially available at the time of

this study.

A simple micro-embedment strain gage was developed as shown in figure 2.1. The

gage is composed of a foil strain gage bonded to an aluminum rod and represents

a modified version of a similar gage first developed at the University of Texas

at Austin by W.C. Stone [12]. To meet the requirement of negligible slippage,

an enlarged anchor (consisting of a nut and washer assembly) was attached to

each end of the aluminum rod. The length of the rod between these two anchors

was 1 inch (25.4mm). Over this length the bond between the aluminum rod and

the concrete was broken so that the foil gage measured the average strain in

the concrete between the two end anchors. It should be noted that these gages

measure only axial strain, parallel to the length of the rod. When linked to

the data acquisition system described in section 2.3 the sensitivity of these

gages was +6 microstrain. An extensive search of available foil strain gages

showed that the smallest workable gages had a gage length of .031 inch (.78 mm)

and a grid width of .032 inch (.81 mm). The backing was a flexible polyamide

compound which permitted the foil gage to be glued to a round surface. Rigid

backed foil gages using compounds such as bakelite should be avoided as they

tend to crack when bent.

The foil gage was mounted on a 1.25 inch (32 mm) long, 0.130 inch (3.3 mm)

diameter rod of 70-75 TG aluminum which was prethreaded for 1/8 inch (3.2 mm)

on each end (see figures 2.1a through 2. Id). A two part methyl cyanoacrylate

adhesive was used for the gluing. The foil gages were waterproofed with a liquid epoxy sealant after the lead wires had been soldered in place (see

figures 2.1 e and f). The entire assembly was then encased in a thin shell of

flexible butyl rubber barrier compound and overwrapped with teflon tape

(figures 2.1g and h). A washer and nut (4-40 thread) were then attached on one end (where the lead wires exit) and the central portion of the gage was inserted into a 1-1/8 inch (28.5 mm) piece of 1/4 inch (6.35 mm) diameter heat

shrink tubing which abutted against the washer. A heat gun was used to reduce

this outer barrier so that a durable waterproof shell now encased the gage.

The washer and nut for the opposite end were then screwed on until snug and

the gage was tested to ensure proper electronic functioning.

Before committing the embedment gages to use in the model specimens, a number

of trial runs were performed. Two 3 inch (76 mm) x 6 inch (152 mm) mortar

cylinders were instrumented with a gage positioned parallel to the longitudinal

axis in the center of each cylinder. The cylinders were loaded in a standard

compression testing machine and the strain recorded using a manually balancing

indicator at successive load stages. Likewise, two splitting tension tests were performed with internal gages oriented perpendicular to the splitting

plane. Both types of tests exhibited acceptable linear behavior in the gage

response

.

When placed in the model these embedment strain gages can be classified into

three primary categories: (a) "radial” gages, oriented perpendicular to the

side of the conic frustum defined by the outer disk edge and the inner reaction

ring edge, and parallel to the R-Z plane (figure 1.3); (b) "axial" gages,

oriented tangent to the side of the same conic frustum and parallel to the R-Z

17
plane; and (c) "circumferential" gages, oriented tangent to the conic frustum

but perpendicular to the R-Z plane. Figures 2.2, 2.3, 2.4, and 2.5 illustrate

the different orientations. Additional gages were placed so as to monitor

strains near the side and top faces of the disk, where first cracking was anticipated

.

In order to define a particular strain profile it was necessary to have a string of closely spaced gages suspended along the orientation of interest.

This was accomplished by attaching the gages with soft copper wires to a grid

constructed from .016 inch (.4mm) stainless steel piano wire, as shown in figure

1.6. Each end of the grid was soldered to a brass bar measuring 2 inch (51 mm)

by 1/2 inch (13 mm) by 1/8 inch (3.2 mm), which was machined so that it could

be bolted to attachment points on the disk edge. Figure 2.7 shows the the brass

end bar with additional hardware for achieving circumferential gage orientations.

Typical assembled gage strings, ready for placement in the specimen, are shown

in figure 2.8. An overall view of instrumented specimen #2 is presented in

figure 2.9.

2.2.2 Slip Gages

In order to measure the load-def ormation history of the pullout test, a method

had to be devised to monitor the displacement of the disk with respect to its

original position. Since the concrete surface above the disk would undergo

deformation, the top concrete surface could not be used as an accurate a reference place for displacement measurements. Similarly, the stroke on the hydraulic

ram loading system could not be used as a measure of disk movement since this

includes the elastic deformation of the tendon transmitting the load to the

disk. The method finally adopted was originally proposed by J . 0. Jirsa [13]

and has been used extensively at the University of Texas at Austin for measuring

the absolute displacement of anchor bolts and the relative displacement of

reinforcing bars used in splices. The device is known collectively as a slip

gage, but really consists of about half a dozen commonly available parts.

For this study, the technique involves mounting a 0.045 inch (1.1 mm) diameter

piano wire near the base of the disk as illustrated in figure 2.10. To do this,

a 1/4 inch (6.35 mm) deep hole (approx. 0.046 inch (1.17 mm) diameter) is first

drilled into the side of the disk, a 90° bend is placed in the end of the piano

wire, the wire cleaned with a piece of emery cloth, and finally the bent portion

is bonded into the hole with epoxy cement. This then, provides a positive

mechanical attachment to the disk. The wire is encased in a snug fitting teflon

sheath to prevent bonding to the cast concrete. The encased wire is then guyed

into place so tat it leaves the disk parallel to the vertical axis (the anticipated direction of displacement), forms a smooth curve, and exits through the

side of the specimen perpendicular to the face. As the disk displaces upwards,

the wire will be drawn in from the side face by the same amount. A threadbar

anchor is installed on the side form approximately 4 inch (102 mm) from the

exit point of the slip wire. This is later used to mount the external

instrumentation

.

After casting and form removal, the exposed slip wires are trimmed to 1.25 inch

(32mm) length and placed in tension by means of a spring loaded block which is

18
19

Figure

2.2

General

view

of

instrumented

pull-out

Figure

2.3

Close

up

view

of

RV

(right)

disk

showing,

top

to

bottom:

RV

,

RN

and

RN

(left)

gages

prior

(radial

gages),

TC

(triaxial

compresto

casting

sion,

radial),

and

CW

(circumferential)

gages
20

Figure

2.4

ET

(edge

tension)

gages

measure

Figure

2.5

ANE

(axial,

northeast)

gages

vertical

strain

along

side

face

measure

compressive

strain

of

disk

along

the

failure

surface
21

Figure

2.6

Gages

are

wired

to

piano

wire

Figure

2.7

Typical

end

detail

for

a

grid

using

soft

copper

wire

circumferential

gage

string
22

Figure

2.8

Typical

assembled

gage

strings

ready

for

mounting

on

disk
Figure 2.9 Overall view of instrumented specimen prior to casting

23
24

Figure

2.10

Slip

wire

details

at

the

disk

connection
securely clamped onto the end of the wire. This ensures that there will be no slack in the line and permits both "positive" (IN) and "negative" (OUT) displacement to be measured. The motion of the wire is measured to .0001 inch

(.0025 mm) accuracy using a spring loaded linear voltage displacement transducer (LVDT) . The LVDT is positioned using a clamping block attached to a support rod which has been screwed into the threadbar anchor in the side of the

specimen. Figure 2.11 shows the external details of the slip gage mechanism.

As a backup to the primary slip gages (three were mounted on 120° arcs around

the circumference of the disk on both specimens) two LVDT's measured the displacement of the "stem" of the pullout assembly with respect to the counterpressure ring base. This was accomplished by mounting the LVDT's on long bars

that had beeen welded to the "stem" 4 inch (102 mm) above the concrete surface

(figure 2.12). The two LVDT's were mounted 180° opposite one another to negate

any errors due to lateral bending of the stem.

2.3 DATA ACQUISITION

Each specimen contained a total of 79 physical data channels—71 embedment

strain gages, 5 displacement transducers (LVDTs ) , two 10,000 psi (69 MPa)

pressure transducers and a 6 volt system excitation voltage. Manual monitoring

was deemed prohibitive for all but a few key channels—primarily the pressure

transducer readings from which the load was monitored. Automated data acquisition centered around a NEFF 620 unit driven by a PDP11 minicomputer. The advantage of this system lay in its speed: with a scan rate of 50,000 hertz the

entire data set could be read in less than 2/1000 of a second. The data was stored on a 2.5 million word RK05 high speed removable diskpack and later

transferred to a 9 track 1600 BPI magnetic tape.

Additionally, it was desired to be able to monitor—in real time—the strain

profile along particular gage strings at various load stages. An interactive

graphics/real time data reduction package known as "TEXD," written by K.A.

Woodward of the NBS Structures Division, was utilized in conjunction with the

data acquisition system mentioned above. This program represents a major

advancement over previously available technology and permitted the unique capability of reviewing the test results in reduced form while the test was still

in progress . In the interactive mode, graphics plots could be obtained on a Tektronics 4014 using predefined plot files which had been generated for each

gage string. By obtaining a hard copy of the plot a direct comparison of what

was happening inside the specimen (reduced to units of microstrain) could be

made with an analytically derived (linear elastic) strain distribution. During

the precracking load stages, it was thus possible to isolate anomalies in the

experimental data and identify possible malfunctioning gages or wiring problems

before going past the "point of no return" —internal cracking. This was of

major importance since, due to the high cost of each test, it was desirable to

minimize the number of erratic or malfunctioning gages. A number of bad

channels—caused primarily by malfunctioning plug-in IC cards in the acquisition system—were thus detected, and corrected, prior to loading the specimen

to first cracking. It should be noted that these bad channels would have gone

unnoticed were it not for the graphics plots, since the gages exhibited correct

electronic functioning where the wires exited the specimen.

25
Figure 2.11 External hardware for slip gage

26
LOADING TENDON

SUPPORT ROD

REACTION RING

REACTION FRAME

STEM

LVDT

LABORATORY CLAMP

T

Figure 2.12 Alternate system for monitoring vertical

displacement of disk

27
2.4 SPECIMEN PREPARATION

The central feature in each specimen was the pullout "disk" and "stem" assembly.

The disk was machined from A36 mild steel to the dimensions shown in figure 2.13.

A 3-1/2 inch (89 mm) diameter standard pipe (4.0 inch (102 mm) 0D) was welded

to the top of the disk to achieve the external geometry that was called for in

the 12:1 scaled-up model. It should be noted, however, that there was a significant difference between the inner workings of the model test and actual com- mercial insert. Most commercial pullout inserts utilize a tapered mild steel

stem for positioning of the disk during casting. Prior to testing this stem

is removed—unscrewed—and a high strength (usually with f y

> 120 ksi (828 MPa))

pulling rod inserted for the subsequent load test. To utilize this same technique in the scaled-up model would have required a threaded 120 ksi (828 MPa)

rod 4 inch (102 ra) in diameter and 83 inch (2.1 mm) long. Price quotes ranged

between $10,000 and $15,000, which was deemed prohibitive. An alternative

solution—by machining a tapered hole in the bottom of the disk so that a 270 ksi (1863 MPa) post-tensioning tendon (31-1/2 inch 13 mm) diameter 7 wire

strands) could be threaded through the pipe and up through the loading system

to a standard 31 strand post-tensioning pulling head—allowed for transmission

of the load at a substantial saving of time and funds. The tendon was anchored

to the disk via a standard 31 strand pullling head, with the individual strands

having been pre-seated (with conical chucks) at the factory.

Since it was desirable to maintain a low profile behind the disk, the preset

pulling head was welded directly to the bottom of the pullout disk (figure 2.14).

Due to the sharp curvature imposed on the outer tendons—a commercial 31 strand

anchor normally requires a 34 inch (863 mm) funneling distance, compared with

the 4 inch (102 ram) used here—it was necessary to jack the two pieces together

(figure 2.15) before the weld could be applied. Even though all internal sur- faces had been machined and ground to effect a smooth transition between the

pulling head and the pipe, industry spokesmen cautioned of the possibility of

shearing prematurely the outer strands at high loads. On this basis the mix

design strength of the concrete was lowered from f£ = 4000 psi (27.6 MPa) to

f^ = 2500 psi (17.25 MPa) and the tests were carried out without difficulty.

The assembled disk, stem, and loading tendon are shown in figure 2.16. It was tacitly assumed that the presence of the loading head below the disk would not

adversely affect the state of stress along the critical failure surface. Additionally, the composite thickness (8 inches) of the disk and loading head is

within the current allowable limit of 0.5 times the diameter of the disk as specified in ASTM C-900.

The dimensions of the cast concrete specimen measured 80 inch (2 m) square and

48 inch (1.2 ra) high. This was approximately a 12:1 scale up from specimen

dimensions originally used on a pilot test by Kierkegaard-Hansen [2]. Later

tests by many researchers utilized standard 6 inch (152 mm) x 12 inch (304 mm)

cylinders with the pullout inserts embedded in one or both ends. Adopting a

12:1 scale for this situation would have called for a scaled-up specimen measuring 12 ft (3.7 m) in height! A somewhat more workable figure of 48 inch

(1.2 m) was arbitrarily assigned. In essence the size of the specimen needed

only to be large enough to suppress any perturbations in the vicinity of the

failure surface that might be induced by the specimen’s boundary geometry.

28
314” STANDARD pipe

a = 3.54

Figure 2.13 Pull-out insert specifications (all units inches)

29
Figure 2.14 Assembled loading tendon and pull-out insert (specifications)

(all units inches)

30
31

Figure

2.13

Mating

disk

and

laoding

head

Figure

2.16

Assembled

loading

tendon

and

(with

hydraulic

jacks)

prior

pull-out

disk

(as

built)

to

welding
In his pilot test, Kierkegaard-Hansen noted that some 30 percent of his

specimens failed when flexurally induced tensile stresses split the entire

specimen in two—-thus invalidating the test, as far as pullout results were concerned. Behavior of this type in the present study was to be avoided. Two

supplementary sets of reinforcement were designed to prevent such an occurrence

in the scaled -up pullout tests. These are shown in schematic form in figure

2.17 and as-built in figure 2.18. The primary containment grid consisted of #2

(6.35 mm) hoops (fy - 40 ksi (276 MPa)) 54 inch (1.37 m) in diameter at 3 inch

(76 mm) spacing tied to #3 (9.5 mm) bars (fy = 40 ksi (276 MPa) oriented vertically at a 6 inch (152 mm) spacing around trie hoop. This hoop was 11.25 inch

(286 mm) larger in diameter than the outside diameter of the counterpressure

device and thus was felt to constitute no interference with the pullout test,

while effectively preventing flexural splitting. The remaining reinforcement

consisted of a uniform orthogonal grid of #3 bars (fy = 40 ksi (276 MPa)) at

6.5 inch spacing placed at the sides and bottom of the specimen with a concrete

cover of 1 1/4 inch (31.75 mm). This was to prevent the propagation of any

cracks that might arise due to temperature and shrinkage effects.

After the forms had been assembled the disk and stem were positioned in the

center at the proper elevation (h = 10 inch (254 mm) and h = 13.5 inch (343 mm)

for specimens 1 and 2 respectively). This elevation was established as the

distance from the top face of the disk to a reference string that had been set across the top of the forms. Precision leveling and final positioning were accomplished by clamping the stem to a C-channel cross frame which spanned the

top of the forms (see figure 2.18). A secondary frame, constructed from lightweight perforated 2 inch (51 mm) angle stock (figure 2.19), provided the upper

support for the gage strings which were inserted at this time. The lower

bracket for each gage string was attached to a special lug that had been welded

to the disk edge. Figures 2.4 and 2.5 show the two types of connections used

for mounting circumferential and radial gages respectively.

Each specimen—"pullout 1" (a = 35°) and "pullout 2" (a =27 °) —had eight

gage strings typically consisting of:

1. Ten radial gages (RN) spaced uniformly from the "north" edge of the disk

to the inside edge of the reaction ring (figure 2.3, left hand string).

2. Six gages (RV) oriented parallel to the top face of the disk and uniformly

spaced from the "north” edge of the disk to a point approximately 7 inch

vertically above the north edge of the disk (figure 2.3, right hand string).

3. Ten radial gages (RS) spaced uniformly from the "south" edge of the disk

to the inside of the counterpressure ring. (Specimen 1 (a =35°) had seven radial gages and six axial gages (AS) on the same grid, but this was modified for simplicity in specimen 2). (Figure 2.9, left-most string).

4. Ten circumferential gages (CE) spaced uniformly from the "east" edge of

the disk to the inside of the counterpressure ring (figure 2.9, bottom

string)

.

32
Figure 2.17 Supplemental reinforcement specifications

(all units inches)

33
Figure 2.18 As built supplemental reinforcement

34
Figure 2.19 Perforated angle frame used to support gage strings.

Cross channel frame is used to stabilize and level

disk prior to casting

Figure 2.20 Checking embedment gage resistance

35
5. Ten circumferential gages (CW) spaced uniformly from the "west" edge of

the disk to the inside of the counterpressure ring (figure 2.2, lower left

string)

.

6. Fourteen axial gages (ANE) uniformly spaced in parallel overlapping strings

from the "north east" edge of the disk to the inside of the counterpressure

ring (specimen #1, shown in figure 2.5 had only 11).

7. Six gages (ET) oriented vertically and parallel to the side face of the disk.

These were mounted approximately 1/2 inch (13 ram) away from the disk on the

"south west" edge (figure 2.4).

8. Five gages (TC) mounted 1/2 inch (13 mm) above and parallel to the top face

of the disk on the "north west” side, beginning at the stem and ending at

the disk edge (figure 2.2).

Specific gage locations (numerical coordinates) are given in appendix A. Once

all strings were in place, the slack was taken up via a nut and threaded rod

which formed the upper support. To ensure that each string followed the proper

angle (and the idealized failure surface) a properly trimmed wooden block was

used so that the threaded rod (which exitted through the perforated support,

then through the wooden block) could be tensioned parallel to the desired

angle (see figure 2.19). This tensioning procedure imparted no pre-tension in

any of the gages (due to the method of attachment—flexible copper wires) and

was primarily for ensuring that the gages underwent no undesirable relocation

during casting.

At this point the electronic integrity of the gages was rechecked (figure 2.20)

using a digital ohmmeter. Gages were rejected (and replaced) if readings fell

outside a window of 120 ft to 123 ft. The foil gages normally have a factory

specification of 120 ft ± .5ft. However, most embedment gages fell into the

former window; the higher resistances resulted from the 10 ft (3m) lead wires

and impurities in the soldered connections. This did not adversely affect the

performance of the gages, but occasionally led to difficulties in balancing the

bridge circuits (which used 120 ft precision resistors, instead of temperature

compensating gages) of the data acquisition system due to insufficient range

in the balancing potentionmeter in the event that a particular circuit could

not be balanced additional precision resistors were added in 1 ohm increments

until balance was achieved.

The specimen was now ready for casting. A standard ready mix concrete, with

1/4" (6 mm) pea gravel aggregate, was used.* Total volume per cast, including

* The pea gravel aggregate was chosen because it was the smallest readily

available commercial aggregate. The intent of this choice was to make the

concrete mix as homogeneous as possible from the perspective of the scaled

up pullout insert. This would then allow for better correlation with the

analytically model which assumed the concrete to behave as a perfectly

homogeneous material.

36
Table 2.2

Concrete Mix Proportions

Specimen 1 (a =35°) Specimen 2 (a =27°)

Type 1 cement (lbs) 3172 (1442 Kg) 2961 (1346 Kg)

Silicious Sand (lbs) 12015 (5461 Kg) 12816 (5825 Kg)

Pea Gravel (lbs) 11437 (5198 Kg) 12200 (5545 Kg)

Water (lbs) 2200 (1000 Kg) 2384 (1083 Kg)

W/C (Approximate) .693 .805

test cylinders, was 7.5 yards (7.2 m^). The mix design for both specimens is

given in table 1.2. Concrete was delivered with only a portion of the total

mix water; additional water was added at delivery to achieve a slump of 3 to 5 inches (76 to 127 mm).

2.5 LOADING SYSTEM

The pullout load was applied via the counterpressure system shown in figure 2.21

The bearing ring had an inside diameter of 26 inch (660 mm), an outside diameter

of 42.75 inch (1.09 m) , and was made from 1 inch (25.4 mm) thick plate stock.

Eight W8 x 31 rolled A36 sections, 33.72 inch (856 mm) in length, were welded

to the base as shown in figure 2.22, and tapered inward so as to form a pedestal for a 26 inch (660 mm) O.D. top plate, also made from 1 inch (25.4 mm) plate

The loading system consisted of six 100-ton (892 kn) hydraulic rams, with a maximum operating pressure of 10,000 psi (69 MPa). These were mounted in a hexagonal cluster on top of the counterpressure pedestal. Two steel retainer

rings, made from 2 inch (51 mm) x 0.25 inch (6.35 mm) bar stock, were used to secure the ram cluster. The lower ring (figure 2.21) was welded to the top

plate on the loading pedestal. To transmit the load from the ram cluster to

the post-tensioning tendon (and thus, to the embedded disk) a jacking head 4

inch (102 mm) thick and 22.5 inch (572 mm) in diameter was machined to fit on top of the six hydraulic piston heads (figure 2.23). A 7.5 inch (191 mm)

diameter hole in the center permitted the 31 strand tendon to pass through.

The assembled loading system was placed on top of the cured specimen and leveled

using a base of high strength gypsum plaster (figure 2.24). The strands were

then threaded through a 31 strand pulling head which rested on top of the jacking head. Each strand was pretensioned to 300 lb (1.33 kn) (figure 2.25) and

its conical chucks were firmly seated. When all strands had been set the pistons for the ram cluster were retracted to remove any pre-load. This procedure

was to ensure that all strands would have close to the same load at all times,

thus avoiding a premature "zipper" type failure of the tendon. The six rams were linked to a load maintainer via two-six port manifolds. Two in-line,

10,000 psi (69 MPa) pressure transducers were used as the primary load monitors,

while two additional in-line pressure gauges were used for backup spot-checks.

37
Figure 2.21 Jacking - counter pressure system (schematic) (all units inches)

Figure 2.22 Compression frame specifications (all units inches)

38
Figure 2.23

Figure 2.24

Jacking head and ram cluster specifications (all

Seating j acking-counter pressure system on plaster grout

unit inches)

39
Figure 2.25 Pre-tensioning loading tendon strands

Figure 2.26 Calibrating j acking-counter pressure system

40
The load maintainer was driven by a 40 GPM (0.16 M^/min), 10,000 psi (69 MPa)

portable electric hydraulic pump. The load was advanced manually at a constant

rate in 10 kip (44.6 kn) increments to failure. Before conducting the first

test, the assembled loading unit was calibrated to 1,200,000 lb (5.3 mn) using

the NBS 12,000,000 lb (54 mn) Universal testing machine (figure 2.26).

2.6 INPLACE STRENGTH OF CONCRETE

To aid in interpreting the test results, it is necessary to know the strength

of the concrete in each specimen. Because of specimen size, it was anticipated

that there would be a significant rise in the internal temperature during the

early stages of hydration. It was felt that using ordinary cylindrical test

specimens might not give accurate indications of the inplace strength, due to

differences in thermal history. Therefore, it was decided to use cast-in-place

pushout cylinder (CIPP0C) molds (figure 2.27), as described in standard test

method ASTM C 873 (ref. 18), which would be embedded in the test specimens.

The concrete within the molds would be subjected to approximately the same thermal history as the concrete within the pullout specimens.

For each test, eight 4x6 inch (102 x 152 mm) pushout cylinder molds were

filled with concrete and embedded in the top surface of the test specimens.

Rodding was used to consolidate the concrete within the molds. The top surfaces of the cylinders were kept covered with moist rags. For comparison,

6 x 12 inch (152 x 305 mm) cylindrical specimens were molded and allowed to cure adjacent to the pullout specimens. These cylinders were kept in their

molds, covered with plastic, and water was added periodically to the top

surface to provide moist curing conditions.

Thermocouples were used to measure concrete temperatures. Four pushout and

two 6 x 12 inch (152 x 305 mm) cylinders were instrumented. In addition,

thermocouples were embedded within the pullout test specimens. The

thermocouples were read automatically by a datalogger.

Figure 2.29 shows the temperature histories during the first 10 days for the

first test (pullout #1). The thermocouple within the test specimen was located

near the insert head. It is seen that the 6 x 12 inch (152 x 305 mm) cylinders

did not develop the same temperature as the test specimen. The pushout cylinders, while they were similar for the first 12 hours, did not attain the same maximum temperature as the concrete near the insert head. The discontinuities

in the thermal history of the pushout specimens were due to the removal of the

plastic sheet covering the specimen which resulted in a cooling off of the top

surface of the test specimen.

Because the pushout molds were embedded in the surface of the test specimen,

the concrete within them did not experience the same temperature rise as the

interior concrete. Thus, an alternative procedure was used for the second

test (pulllout #2). An additional set of 6 x 12 inch (152 x 305 mm) cylinders

were made using plastic molds covered with plastic and stored under water. The

temperature of the cylinders in the bath was compared with the temperature near

the insert head, and an electric heater was turned on as necessary to maintain

the cylinder temperature as close as possible to the temperature within the

41
Figure 2.28 Slave cylinders in temperature controlled water bath

Figure 2.27 Placing CIPPOCS model in freshly cast specimen

42
TEMPERATURE

HISTORIES

IN

PULLOUT

SPECIMEN

*1




o

• rJ o >

(S) • o

<x a. Ui • ru X Q. ^

00

-« X

a * Ui o 10

(.ft z u. u.

»-• o o

a • • <Z<JO

Ui => z « «

I

CO

co>

<rQ —

UJ

o

<r

aj

'l l i r LD

ru

t~t~t i i—i—

r





 LO

i r t

~~

t


 LO

t

i—i—

r

LO

ru

I— LUEZGLUJQ£<rt— ZDQ^LlI

43

Figure

2.29

Temperature

history

for

pull-out

specimen

#1
Table 2.3. Cylinder Strength Results

* Standard deviation, psi

1 psi = 0.0069 MPa

large specimen. An electronic circuit was designed and built that turned the

heater on and off automatically (figure 2.28).

Figure (2.30) shows the thermal histories for the second test specimen. It is seen that the "slave" cylinders in the water bath did not heat up as rapidly as the interior concrete. This was due to a failure in the electronic circuitry

and possibly also due to the lack of insulation around the water bath. At an

age of 1 day, the damaged electronic component was fixed and insulation was placed around the tank. The "slave" cylinders then followed closely the interior concrete temperature. The rise in the "slave" temperature at four days

was due to another electronic malfunction that was later corrected. Thus,

except for some "shakedown" problems, the "slave" cylinder concept appears to

have been a good solution to the problem of accounting for the inplace

temperature rise in the massive test specimen.

On the day of the pullout test, the cylindrical specimens were also tested.

Table 1.3 shows the measured compressive and splitting tensile strengths of the

various specimens. The measured compressive strengths of the pushout cylinders

were multiplied by 0.96 to account for their lower length:diameter ratio as specified in test method ASTM C 42 [27].

44
TEMPERATURE

HISTORIES

IN

PULLOUT

SPECIMEN

#2

00

>

<rQ

bJO

ai

nj in nj

h-UJCQ-UiQC<ri-3Q:UJ

45

Figure

2.30

Temperature

history

for

pull-out

specimen

#2

(h

=

13.5

inches,

2a

=

54°)
The compressive strength results are consistent with the known effects of

initial curing temperature on later age strength, namely an increase in curing

temperature lowers the long term strength. This may explain why the "slave"

cylinders were weaker than the pushout cylinders which were in turn weaker

than room cured cylinders. The splitting tensile strengths of the "slave"

cylinders are lower than expected and there is no explanation for this, other

than a "size effect" problem. For specimen #1 the average compressive strength

of the C.I.P.P.O.C. cylinders (f£ = 2660 psi) was used as the basic for material property inputs in the analytical model; for specimen #2 for average of

the slave cylinder compressive strengths (1180 psi) was used, since this was considered to be a more accurate indicator of the concrete strength in the

vicinity of the pullout insert.

2.7 TEST PROCEDURE

Once the electronics for the data acquisition system were checked, a pre-test

data check run was made by loading each specimen to approximately 20 percent

of ultimate load in 10 kip (4545 kg) increments. At each load stage a scan was made of all instruments and the data was stored on magnetic disk. At the peak

load, plots of the observed strain distributions for the key gage strings were produced on site using a Tektronix 4014 terminal and a hard copy device (figure

2.31). If all gages checked out, the load was dropped to zero using the manual

load maintainer and the actual test was begun. Load was monitored using both

a digital voltmeter, which gave a direct reading of the pressure transducer

output, and using the program TEXD. In the latter mode, the transducer channels were automatically converted by the program to give a visual display of

the load in kips on the 4014 unit at the prompting of the operator. The digital

voltmeter served as the primary queue for initiating a load scan. This was then

verified on the 4014 as the load channel values appeared on the screen. The

load was increased in 10 kip (4545 Kg) increments until ultimate failure. In

order to assist in detecting the onset of failure, an acoustic emission receiving transducer (figure 2.32) was affixed to the counter pressure ring and the

output signal amplified through a loud speaker system. The audible cracking

sounds gave sufficient warning so that a large number of scans could be taken

near the failure load. Each test required a full day to perform, including

cylinder tests after the failure of the main specimen. The test data was backed

up to tape from disk and individual strain-history plots for each embedded gage

were generated; similarly, load versus deflection plots were made for each

LVDT. Later, the loading system was disassembled and the pullout "cone" was removed for photographing and measurement. An overall view of the test setup

is presented in figure 2.33.

46
o 3 33 3 > c rb <u 3 «\3 p3

a) a) r-i o 3 -H

•h 3 > y-t oP bO

a.a

O 3 4-1 3

a) 3 g- 3 a

CO *H 3

y-i P o 3

•H bO

4-1 3

i-l 3

Cu P a g 3 £

4-1 r—I O 33 bO P 3

•3 P P O CO P 3

•H P 3O bO S 3

•H

O P P 3 3 CO

CO 3 3 -i-l P H ON 3 P 3 CO -3

•H *H 3 g* * 3P

COo

•H

JZa3po 3 iH

•H3 CO P 3 P

CO 3S

•Ho3a

CO

CO

CM3P3

bO

•H

Cp 47
a) Tektronix 4014 terminal, digital

voltmeter for monitoring load, x-y

recorder for monitoring vertical

displacement of pull-out disk, hard

copy device for computer terminal,

manual load maintainer, thermal data

logger for embedded therocouples

b) Cast specimen, with loading device

in operation

Figure 2.33 Overall view of lab test set-up

48
3. RESULTS

3.1 LOAD DEFLECTION CURVE

The experimental data obtained from each large pullout test are comprised of

three parts: load-deflection data; the failure surface geometry; and internal

strain histories as measured by embedded straingages. Figures 3.1 and 3.2 show

load versus deflection plots for specimens 1 and 2 (2a =70° and 54°, respectively). Also indicated are values of compressive strength from CIPPOCS (for

specimen #1) and slave cylinders (for specimen //2) and computed values of modulus of elasticity using the American Concrete Institute formula 8.5.1 (E

c = wc

* 33 /f’ c

). The leftmost curve in each figure represents the average upward

deflection of the disk as measured by the three slip wire controlled, side

mounted LVDTs. The right hand curve represents the average of the two vertical

LVDT deflections. It can be seen from figures 3.1 and 3.2 that the failure

sequence of the pullout test is marked by three phases signified by pronounced

changes in the slope of the load -deformation curve. These load levels are listed in table 3.1, and their significance in terms of the failure sequence

are discussed in section 3.5.

Table 3.1.

Ranges in Percent of Ultimate Load which Hark Pronounced

Changes in the Slope of the Load Deflection Curve

Specimen #1 (2a = 70°) Specimen #2 (2a =54°)

Phase

I 30- 43 34- 39

II 63- 70 58- 66

III 83-100 84-100

In both specimens the slip wire deflections lagged behind those measured by

the stem mounted gages. This may have been due to a taking up of slack in the

piano wire—despite an effort to reduce this by spring loading the slip wire.

Alternatively, as will be discussed, the load at which the slip gage picked

up a significant deflection in specimen #2 corresonded almost exactly to the

load at which circumferential cracks had completely formed from the outer disk

edge to the inside edge of the counterpressure ring. This would indicate that

large deflection begins only after circumferential cracking is complete.

For comparison of specimens 1 and 2 a normalized load deflection curve is

shown in figure 3.3. Normalization was accomplished by dividing the applied

load by the respective value of the compressive strength for each specimen and

by multiplying the deflection by the calculated modulus of elasticity for each

specimen. Each line represents the average of the vertical LVDT and slip wire

49
LATERAL

MOUNTED

SLIP

GAGE

(LVDT)

50

Figure

3.1

Vertical

disk

deflection

vrs

load

Specimen

#1

2a

=

70°
CO

</> 0> o

CO

51

Figure

3.2

Vertical

disk

deflection

vrs

load

Specimen

#2
225 r

8 - E c

(kips/in)

Figure 3.3 Normalized load deflection curve Specimens 1 & 2

52
deflections. The normalized ultimate load for specimen #2 is twice that of

specimen #1. Thus, a decrease in the apex angle (2a) —or increasing the

depth of embedment, as in this case—leads to a higher pullout force. This was first observed in tests performed by Kierkegaard-Hansen [2] who changed the

apex angle by varying the counterpressure ring diameter, rather than the depth.

3.2 SHAPE OF FAILURE SURFACE

Due to the size of the pullout "cones” (see figure 2.4) it was possible to make

detailed measurements of their geometry. For each specimen several readings of

the variation of radius with respect to height were taken around the circumference of the failure surface from the disk edge to the top edge of the cone which

corresponded to the inside edge of the counterpressure ring. The readings were

taken by dropping a plumb line from the top of the cone and measuring the perpendicular distance to the failure surface by means of a ring stand with an adjustable height horizontal reference rod. These are plotted in figure 3.5.

The dashed line in each figure represents the idealized failure surface that

would be defined by the frustum of a cone, as well as the line of location of

the embedded strain gages. Two important observations that can be drawn from

these figures are that:

a. the failure surface approaches a frustum geometry for low apex angles;

and assumes a trumpet shaped, geometry which deviates further from

the frustum line with increasing apex angles.

b. the failure surface geometry becomes more uniform about the circumference for lower apex angles. Stated differently, there was a greater

scatter in the measurements taken from specimen 1 (figure 2.5a,

2a = 70°) than from specimen 2 (figure 3.5b, 2a = 54°).

3.3 COMPARISON OF EXPERIMENTAL AND ANALYTICAL STRAIN DISTRIBUTIONS

Examination of a typical embedment gage strain history, such as shown in figure

3.6, shows that the load versus strain behavior is nearly linear until roughly

30 percent of ultimate load. Below this load, the concrete can be assumed, for

the sake of analysis, to be linear elastic, homogeneous, and uncracked. After

cracking the state of stress will change continually as radial and circumferential cracks propagate through the concrete. As a precursor to possible future

non-linear, finite element method (FEM) analyses of cracked specimens 1 and 2,

it was felt necessary to first assure that the (FEM) modelling technique (FEM)

would yield reliable reults in the precracked state.

For analyses, a linear elastic, isotropic, two dimensional, axisymmetric (solid)

model was employed. The .element meshes for modelling specimens 1 and 2 are shown in figure 3.7 and figure 3.8 respectively. The mesh generation and post

processing stress contour plotting were done using the UNISTRUC II interactive

graphics package. The problem solution was performed using ANSYS on a CDC

cyber 175 computer. For both specimen models the applied load was 50 kips, or

17 percent and 19 percent of ultimate load for specimens 1 and 2, respectively,

well below the level of first cracking. For each model, input values of the

Elastic Modulus Ec corresponded to values calculated using the American Concrete

53
uoo

I-'-

IIa

CNa

COV

43o

(3 •H 43

4-1

O 3> rH O

CO

II <u 43 43

4J rH S3

=**-

O

J-i c 4-1 <u&

•H <U o 3

<u <D CO >

54

Figure

3.4

Experimental

pullout

cones
oo<rma

CM

03

03 XOa

•Hin• u

<n 03 -I a)

li <13 S? Si ' 4J CM S % OM C 44

03& T3

*H 03 O 5

03 03 a *h

C/3 > aoO

r^-

lla

CM

03

03 J=Oa

•H

•u

O 03 rH <0

03

II 03 Si Si

44

i—i

a

=s= o

14

a 44

033 ra

•H 03 O 5

0) 03 a, -1-1 C/3 > o

03pa

bO

•t4 P>4

55

3.4

Experimental

pullout

cones
•u 4-1 co 3 3 •H o 3 32

1—1 e H ! 1 3 3 3 O i-i (X • 3

"3

B rH

4-1 3 o t4 O 3 4-1 3

4-1 4-1 tH 3 o 3 4-1 33 3 > 3 a o bO 'C to e 3 3 4-1 3 3 3

Sj bO

bO CO 3 U 3 3 3 3 33 32

4-1 33 •H CO

4-4 3 co 3 U 3 3 3

0) 33

O 4-1 CO 5 O 3 O CO J3 cx •u 3

•H >

CM 3 CO •H

=»£: 4J

O 4-1

O

a 3 3 !Z 3 CJ 3 33 t4 e 3?

•H • 3 3 O 3 4-1 •H 3 3 CO CX O 3 CO CO O a- •H % o

LHa

CM

co3

32Os

in ,3 • 4-1 co 3

rH

O

CO

II 3

32 32 w J_i CM 0O

S-t (3 4-1 3B 33

•H 3 O 5 3 a) IX tH CO >

56

Figure

3.4

Experimental

pullout

cones
Figure 3.5a Measured failure surface for Specimen #1

2a = 70°

57
Figure 3.5b Measured failure surface for Specimen #2

2a = 54°

58
59

Figure

3.6

Typical

embedded

strain

gage

history

—

axial

gage

#5

from

Specimen

1
REACTION

RING

(see

Detail

B)

60

Figure

3.7

2D

axy

symmetric

finite

element

mesh

for

Specimen

#1
4

EXTERNAL

APPLIED

^

LOAD

61

Figure

3.8

2D

axysymmetric

finite

element

mesh

for

Specimen

#

2
.16

Institute formula 8.5.1. The value of Poisson's Ratio v c was obtained

experimentally for each specimen. These were E

c = 2.82(10)° psi, u

c = for specimen 1 and E = 1.91(10)^ psi, u = .16 for specimen 2.

Comparisons between all experimental gage strings and the FEM predicted values

are given in figures 3.9 through 3.20. For the most part the agreement between

analytical and experimental values is good. 3.3.1

Axial Strain

Tne axial strain (i.e. the strain tangent to the surface of a conical frustum

defined by the outer disk edge and inner reaction ring edge and parallel to the

R-Z plane (figure 1.3)) is compressive and increases towards the disk edge and

reaction ring. The maximum compressive strain occurs at the disk edge with a value approximately twice that of the compressive strain at the reaction ring.

The minimum axial strain occurs at about 65 percent of the distance from the

disk, edge to the reaction ring (see figures 3.9 and 3.10). These latter two

figures indicate a good fit with the FEM calculated strains, both in the general

trend as well as the absolute numerical values.

3.3.2 Radial Strain

The radial strain (i.e., the strain perpendicular to the side of a conical

frustum defined by the outer disk edge and inner reaction ring edge (figure

1.3)) decreases from a maximum tensile value at the disk edge to near zero at

the reaction ring (see figures 3.11 and 3.12). The middle portion of the data

(from x = 3 to x =10 for specimen #1 and from x =4 to x = 12 for specimen #2)

indicates a relatively constant tensile strain in this region. While the general trends agree, the absolute numerical values predicted by the FEM analysis

are substantially lower than those observed experimentally.

3.3.3 Circumferential Strain

The circumferential strain is tangent to the surface of the conical frustum

defined by the outer disk edge and inner reaction ring edge, and perpendicular

to the R-Z plane. Experimental data from specimen #1 (2a = 70°) indicates that

the circumferential strain is small and varies from slightly compressive near

the disk edge to slightly tensile near the reaction ring. This contrasts with

the analytical solution which indicates a nearly uniform low tensile strain

between the disk and reaction ring. Specimen #2, on the other hand, agrees well

with the analytical solution indicating a near uniform low (less than 10 microstrain) tensile strain from the disk edge to the reaction ring (see figures 3.13

and 3.14). With strains this small the experimental gage sensitivity (+/- 6

microstrain) could account for the majority of the differences between the

analytically predicted and experimentally observed values.

3.3.4 Vertical Strain Near Side Face

The highest recorded tensile strains occurred near the side face of the disk,

with the maximum values near the top edge of the disk (see figures 3.15 and

62
-120

63

Figure

3.9

Axial

strain

along

Frustum

surface

expermental

vrs

FEM

Specimen

#1

2a

=

70°
64

Figure

3.10

Axial

strain

along

Frustum

surface

experimental

vrs

FEM

Specimen

#2

2a

=

54°
I

QC 'U

vu =t

65

Figure

3.11

Radial

strain

along

Frustum

surface

experimental

vrs

FEM

Specimen

#1

2a

=

70°
66

Figure

3.12

Radial

strain

along

Frustum

surface

experimental

vrs

FEM

Specimen

#2

2a

=

54°
•

CW

GAGES

o

CE

GAGES

X

_L

CM

67

Figure

3.13

Circumferential

strain

along

Frustum

surface

experimental

vrs

FEM

Specimen

#1

2a

=

70°
X

VU ^

68

Figure

3.14

Circumferential

strain

along

Frustum

surface

experimental

vrs

FEM

Specimen

#2

2a

=

54°
69

Figure

3.15

Vertial

strain

adjacent

side

of

experimental

vrs

FEM

Specimen

#1

2a

=

54°
3.16). Because there is little bond between the concrete and the side face of

the disk, the load is transmitted to the concrete primarily in compression via

the top face of the disk. At the disk edge the load is applied to a right

angle or "re-entrant” corner, giving rise to a large tensile strain concentration [15] . This strain concentration was monitored in specimen #2 as indicated

by the high point near the top edge of the disk. Due to a faulty gage this was

not seen in specimen #1 . The maximum experimental value of 165 microstrain

(specimen #2) exceeds the nominal limiting tensile strain capacity of concrete

as recommended by Rusch [14] . We would thus expect that the concrete in the

vicinity of the top edge of the disk would be cracked at this load stage, and

indeed, further study showed this to be the case (see figure 3.40 section

3.5.3). The general trend, and absolute numerical values, agree fairly closely

for experimentally observed and FEM predicted strains, with the latter being

slightly higher than the former.

3.3.5 Radial Strain on Vertical Line Above Disk Edge

These data indicated the radial strain (parallel to the r-axis) to be

compressive at the disk edge, but changing to a tensile state a short distance

above the disk (see figures 3.17 and 3.18). The location of this cross over

point varied, occuring at approximately 0.5 inches (12.7mm or 0.037h) above the

top face of the disk for specimen # 2, while occuring at 1.5 inches (38mm or 0.15h) above the top face of the disk for specimen # 1. Good correlation

between analytical and experimental results was obtained for both specimens.

This strain pattern is very typical of concrete structures with post-tensioned'

anchorages and is ’usually referred to as the "bursting" strain/stress—the

transverse tensile strain ahead of the loading plate. In thin web structures

it was believed that this tensile strain caused cracking or "bursting" along

the tendon path ahead of the anchor. For the pullout study these gages were

used to attempt to detect the deviation of the failure surface from the linear

conic frustum.

3.3.6 Radial Strain Near Top Disk Face

These gages were used to measure the strain parallel to the top disk face in

an effort to study the triaxial state of stress which exists in this area. The experimental data indicate a linear increase in lateral compressive strain

towards the outer edge of the disk. The analytical solution, for both specimens

#1 and #2, agrees with the experimental edge compressive strains, but indicates

a slight tension toward the stem (see figures 3.19 and 3.20). This difference

may be due to the inability of the program to precisely model -the complex friction boundary conditions existing at the interface of the concrete and the top

face of the disk. ,

3.4 ELASTIC BEHAVIOR

3.4.1 Stress Contours

Given the reasonably good correlation between the results of the finite element

model and the experimental gage data in the precracked state, we can proceed

with an examination of the overall precracked state of stress by means of the

70
71

Figure

3.16

Vertical

strain

adjacent

to

side

of

disk,

experimental

vrs

FEM

Specimen

#2

2a

=54°
72

Figure

3.17

Radial

strain

on

vertical

line

above

disk

edge

experimental

vrs

FEM

Specimen

It1

2a

*

70°
RV

GAGES

Q£ VL> « a.

73

Figure

3.18

Radial

strain

on

vertical

line

above

disk

edge

experimental

vrs

FEM

Specimen

#2

2a

=

54°
o

to

II Q.

<

>z vu

74

Figure

3.19

Radial

strain

along

disk

face

experimental

vrs

FEM

Specimen

#1

2a

=

54°
75

Figure

3.20

Radial

strain

along

disk

face

experimental

vrs

FEM

Specimen

#2

2a

=

54°
analytical model. One of the most useful devices for identifying critically

stressed regions is a stress contour plot. Figures 3.21 and 3.22 present the

maximum principal stress contours for specimens 1 and 2 respectively. Tensile

stresses are positive. From these plots it is clear that the maximum tensile

stresses occur adjacent to the top edge of the disk. As mentioned previously

this is due to the presence of the loaded re-entrant corner which creates a strain (and therefore a stress) concentration. This tensile stress decreases

rapidly along the failure surface towards the reaction ring and a sign reversal

occurs at approximately 90 percent of the distance to the reaction ring.

9 Contour plots of the minimum principal stress for specimens 1 and 2 are given

in figures 3.23 and 3.24 respectively. From these it is apparent that the peak

compressive stresses occur just ahead of the upper disk edge and beneath the

reaction ring inner edge. The largest compressive stress at the disk is approximately three times the compressive stress at the reaction ring for specimen

1 and 3.75 times that at the reaction ring for specimen 2. There is a lower

stress "saddle" between these two maxima, with the minimum value occuring at

approximately 65 percent of the distance to the reaction ring.

3.4.2 Principal Stress Profiles Along Failure Surface

The variations in stresses along the idealized conic frustum failure surfaces

are presented in figures 3.25 through 3.28. For these plots the analyses were

conducted with E

c^ = ^2 = 2.82(10)° psi and u

c ^ = u

Q 2 = .16, such that the

effect of the geometry change between specimens 1 and 2 could be isolated,

independent of varying material properties. Figure 3.25 (maximum principal

stresses) shows that the peak tensile stress for specimen 2 is, on the average,

about 37 percent lower than for specimen 1 with the greatest difference (45 percent) occurring at the disk edge. Figure 3.26 (minimum principle stresses), on

the other hand, shows a close similarity between specimens 1 and 2. The maximum

difference occurs at approximately x/d - 0.1 where the minimum stress for specimen #2 is 25 percent below that for specimen #1. However, the peak compresive

stresses at the disk edge and reaction ring edge show very little difference.

The maximum shear stresses in the r-z plane (figure 3.27) exhibits the same trends as the minimum principal stress with nearly identical peak values for

specimens 1 and 2. Figure 3.28 shows the circumferential stress variation

along the frustum surface. Notably this stress is tensile, and nearly uniform

over the majority of the surface. Towards the disk edge, and towards the reaction ring the circumferential stress becomes compressive. From figures 3.25,

3.26 and 3.28 it is apparent that the state of stress along the majority of the

frustum surface (see figure 3.29) is biaxial tension-compression, with the

highest tensile stresses occurring near the disk edge.

3.4.3 Principal Stress Trajectories

The principal stress direction (or trajectory) plots for specimens 1 and 2 are presented in figures 3.30 and 3.31 respectively. The solid lines represent

the maximum principal stress directions while the dashed lines represent the

minimum principal stress directions. Interestingly, the geometry of the experimentally measured failure surface coincides closely (more so with specimen #1

than with specimen #2) with the minimum principal stress trajectories between

76
Figure 3.21 Maximum stress contours for Specimen #1 (2a = 70°) for a load of

P - 30 kips ( 17 / of ultimate) (from 2D axisymmetric FEM analysis)

All units are in psi (pounds per square inch)

77
Figure 3.22 Maximum principal stress contours for Specimen #2 (2a = 54°) for

a load of P = 50 kips (19% of ultimate) (from 2D axisymmetric

FEM analysis). All units are in psi (pounds per square inch)

78
-160

Figure 3.23 Minimum principal stress contours for Specimen #1 (2a = 70°)

for a load of P = 50 kips (17% of ultimate) (from 2D

axisymmetric FEM analysis). All units are in psi (pounds per

square inch)

79
-120

Figure 3.24 Minimum principal stress contours of Specimen #2 (2a = 54°)

for a load of P = 50 kips (19% of ultimate) (from 2D

axisymmetric FEM analysis). All units are in psi (pounds

per square inch)

80
9 CM CD 9 8

XTO „E </> O Eb

81

Figure

3.25

Variation

of

maximum

principal

stress

along

Frustum

surface

(from

axisymraetric

FEM

analysis)
82

Figure

3.26

Variation

of

minimum

principal

stress

along

Frustum

surface

(from

2D

axisymmetric

FEM

analysis)
83

Figure

3.27

Variation

of

maximum

shearing

stress

(R-Z

plane)

along

Frustum

surface

(from

2D

axisymmetric

FEM

analysis)
0£=*>2

‘l#

moTM

<3i &

84

Figure

3.28

Variation

of

circumferential

stress

along

Frustum

surface

(from

2D

axisymmetric

FEM

analysis)
Figure 3.29 Elastic state of stress along idealized

pullout failure surface

85
I 1 1 1 I I L

J I 1 CO CM = 2 com com o

r>.

IIu

M

H *-> a> co 0

co 0

•U *H

CO

X

cc! 1—1

CCS

O

O- «M

•HO 0

fl

O

•H S-1 M CM

CU ^ Oro• coVu3

bO

rl

ClH

86

ACTUAL

FAILURE

SURFACE
87

Figure

3.31

Principal

stress

trajectories

for

Specimen

#2

(from

2D

axisymmetric

FEM

analysis)
the disk and reaction ring. This means that the maximum (tensile) principal

stress, for the most part, acts perpendicular to the failure surface while the

minimum principal (compressive) stress acts parallel to the failure surface.

The bold solid line in each of figures 3.30 and 3.31 represents the boundary of

the observed failure surface. It is based on an average of the separate readings taken around the circumference shown in figures 3.4 and 3.5. It should be

noted that the above comparison is for the concrete in the pre-cracked state,

and the situation may change due to cracking.

3.5 CRACK PROPAGATION ANALYSIS

3.5.1 Post Cracking Behavior

As the load is increased towards ultimate failure two crack systems develop in

the pullout test. These are:

1. Circumferential Cracks: Cracks which form the outer surface of the pullout

"cone". As will be shown later these cracks begin at the disk edge and

propagate towards the reaction ring as the load is increased. These cracks

follow the curved paths previously shown in figures 3.4 and 3.5.

2. Radial Cracks: These cracks form parallel to vertical planes extending

radially from the disk stem (r-z plane). Although the physical presence

of these cracks was verified following testing — absorbtion of methyl

alcohol poured on the top surface of the pullout cone clearly revealed

them — the actual path of radial crack propagation was not monitored

experimentally due to a lack of sufficient micro-embedment gages (which

had to be meticulously hand crafted at the lab). However, FEM studies by

Ottosen [7] and the ones by the author in this report indicate cracking

initiates at the annulus formed by the top surface of the concrete and the

disk stem and propagates as a curved wave from this point. Radial cracks

are formed when the circumferential stresses exceed the tensile capacity

of the concrete. The high circumferential stress values at the point of

crack initiation are set up principally due to flexural type upward deformation of the top concrete surface. This as can be visualized in figure 3.23

which shows the peak vertical deformation, and from the circumferential

stress contour plots shown in figure 3.33.

As these two crack systems propagate a discontinuity will occur in the loadstrain history of any embedded gage in the vicinity of the crack. As previously

shown, the actual failure surface did not follow the idealized conic frustum

trajectory along which the gages were located. Because of this those gages

oriented perpendicular to the principal cracking planes (radial gages for circumferential cracks, and circumferential gages for radial cracks) exhibited two

types of load-strain behavior:

a. For those cases where the gage crossed the failure surface a large

change in the slope, or a large positive strain excursion indicated

that cracking had propagated to that location. This type of behavior

is shown in figure 3.34.

88
Figure 2.32a Elastic deformation of Specimen #1 (2a = 70°) at a load

of 30 kips (from 2D axisymmetric FEM analysis)

NOTE: Dashed lines denote undeforraed shape

89
REACTION RINC

DISK

i

i

1

Figure 3.32b Elastic deformation of Specimen #2 (2a = 54°) at a load

of 50 kips (from 2D axisymmetric FEM analysis)

NOTE: Dashed lines denote undeformed shape

90
Figure 2.33a Circumferential stress contours for Specimen #1 (2a = 70°)

at 50 kips load (from 2D axi symmetric FEM analysis)

91
/

/

z_ r = 50 k

Figure 3.33b Circumferential Stress contours for Specimen #2 (2a = 54°)

at 50 kips load (from 2D axisymmetric FEM analysis)

92
Figure 3.34 Gage RS5 , Specimen #2 positive strain excursion indicating

gage crosses crack surface

93

mw
b. For gages not located along the failure surface the formation of a crack is evidenced by a sharp reversal or relaxation of strain. This

type of behavior is shown in figure 3.35.

Although one would anticipate a monotonically increasing, compression strain

history for gages oriented approximately parallel to the minimum principal

stress (ANE and AS axial gages) this was not always the case. As cracking progressed those gages not on the failure surface (such as shown in figure 3.36)

show a strain relaxation as the crack passes adjacent to the gage, unloading it.

Once the radial and circumferential cracks have propagated from the disk to the

reaction ring — as evidenced by large strain excursions having occurred for

all key gages—there still remains a considerable reserve strength. At

approximately 80 percent of ultimate load there begins a pronounced stiffness

softening—a noticible change in slope of the load-strain plot —and finally,

beginning at 97-98 percent of ultimate a large strain excursion is evidenced

by all gages. These changes in the geometry of the load strain curve can be

generally categorized as discontinuities each marking a specific phase of the

failure process.

3.5.2 Discontinuity Histograms

By carefully inspecting the strain histories of each embedded gage a list of

all discontinuities and the load stages at which they occur was compiled. As

previously mentioned, a discontinuity was evidenced in the load-strain plots

by either a pronounced change in slope or by a sharp strain excursion such as shown in figures 3.34 and 3.35. A bar chart of these points as a function of

percent ultimate load is shown in figure 3.38. From this plot it is clear that

there are three distinct phases which constitute the failure sequence of a pullout test. These phases vary slightly for each specimen and are listed below In

table 3.2.

The "peak" values indicate the percent of ultimate load at which the maximum

number of discontinuities were observed within a given phase. Note the close

correlation between the above ranges and the ranges listed in table 3.1, which

represent the discontinuities observed in the external load-deflection plots of

the specimen. This close similarity would suggest that future tests to investigate changes in the failure process due to changes in the geometry of the test

could be more efficiently undertaken by carefully monitoring the external load

deflection behavior, rather than using the tedious embedment gage method. The

latter, of course, is considerly more precise, owing to the superior sensitivity

of the strain gage over a typical displacement LVDT.

94
RW Q?

-25-300 - 15.600 -5.100 5.100 15.500 55.500 33.500 15.000 50.100

Figure 3.35 Gage RN6 , Specimen #1 negative strain excursion (reversal)

indicating crack has passed adjacent to gage

95
Figure 3.36 Gage ANE7 , Specimen #1 axial gage showing strain reversal

crack has passed adjcent to gage, thereby unloading it

96
Figure 3.37 Gage ANE2 , Specimen #2 typical monotonic axial strain history

4

97
PULLOUT

t

1

DISCONTINUITY

PROFILE







to0

•H

3

s; •Hc

O 0

U D) « id -4.P

*« «o *« *o

o (do

O

ja x: e y

3 id c 0

0*>

-p rj

o

C*T3

0 0

^a-P

o

<0 0

1*4*

rd 0

CQ “d

0

.p

o

z

t

1 r o

n

LO

OJ

<&

OJ

LD <S> in ©

A •-» z: t- lu q: z: <r QMt^oozhwrrjMhHunA

98

PERCENT

OF

ULTIMATE

LOAD
PULLOUT

#2

DISCONTINUITY

PROFILE

o

G>

<S> O)

<£> 00

©

r-


 CO


 Lf)

©

'T

©

n

©

nj

©

©

^ HZhUQ:Z<rJ QH^OOZhHZDHf-HUW

99

PERCENT

OF

ULTIMATE

LOAD
Table 3.2. Ranges in Percent of Ultimate Load of Primary Discontinuity Zones

Observed by Embedment Gages

SPECIMEN #1 SPECIMEN #2

PHASE / ULT. LOAD

Range

% ULT. LOAD

Peak

% ULT. LOAD

Range

% ULT. LOAD

Peak

I 32 - 40 36 28 - 38 33

II 63 - 70 66 50 - 70 61

III 80 -100 98 80 -100 83/98

3.5.3 Crack Propagation Sequence

By tabulating the percent of ultimate load at which each "cross crack" gage

records its first discontinuity it is possible to trace the formation and propagation of internal cracking during the pullout test. As previously mentioned,

the key gage strings which most reliably monitored the progression of circumferential cracks were the gages oriented perpendicular to the plane of these

cracks: the radial gages. Similarly, the circumferential gages were the most

accurate indicators of radial crack propagation. Figures 3.39 and 3.40 present,

in schematic form, the progression of circumferential and radial cracks as inferred from discontinuities in the strain histories for specimens 1 and 2.

These figures are broken into eight load stages which mark significant crack

propagation. The right hand diagonal columns for each sequential plot represent the radial gages in the proper relative positions within the specimen.

The short vertical strings above and below the right disk edge represent the

KV (radial) and ET (edge, tension) gages. All of the above gages monitored

the progression of circumferential cracks. The left hand diagonal columns

represents the circumferential gages which recorded radial crack development.

A blackened circle indicates that the gage has undergone a strain excursion at

or before the labelled load stage. From these plots it is clear that:

1 . Circumferential cracking begins at the disk edge at low load levels—25 to

35 percent of ultimate load. Computer studies by Ottosen [7] and those by

the author in the report have shown that the cracks may begin on the side

face of the disk near the top edge at loads below 25 percent of ultimate.

This was verified in pullout specimen #2 as shown in the first schematic

drawing in figure 3.40. A faulty gage precluded monitoring this early

cracking in specimen #1 . Once the cracking threshold is reached the circumferential failure surface begins to propagate from the disk edge towards

the inside edge of the reaction ring. It is interesting to note that for

specimen #1 the RV (radial) gages —the short vertical string above the disk

edge—indicate major cracking in their vicinity at early loads, while the same gages for specimen #2 indicate that no cracking occurs until much

higher loads. This may be inferred to mean that the failure surface near

the disk for specimen #1 (2a = 70°) deviated significantly from the

100
Figure

3.39

Crack

propagation

sequence

for

Specimen

#1

(2a

=

70°)

black

circles

denote

gages

which

have

recorded

cracking

at

a

particular

load

stage.

Left

had

diagonal

columns

represent

circumferential

gages;

right

hand

diagonal

columns

represent

radial

gages

GO »*>

T

T

1

T

101
Figure

3.40

Crack,

propagation

sequence

for

Specimen

#2

(2a

=

54°).

Black

circles

den

gages

which

have

recorded

cracking

at

a

particular

load

stage.

Left

hand

diagonal

columns

represent

circumferential

gages;

right

hand

diagonal

col

represent

radial

gages

102
idealized conic frustum surface. This was born out in subsequent

measurements of the failure cone following the conclusion of the test

(figure 3.5). At approximately 65 percent of ultimate load it can be seen

that in both specimens circumferential cracking has progressed from the

disk edge to the reaction ring. This is an extremely important observation

with the failure surface completely formed there is still a 35 percent

increase in load necessary to remove the cone.

2. Radial cracking, as previously mentioned, begins at the annulus formed

by the intersection of the top concrete surface and the disk stem. This

cracking probably begins at low load levels [7] and rapidly propagates away

from this point. At approximately 35 percent of ultimate load the crack

front reaches the circumferential failure surface. The experimental data

shown in figures 3.39 and 3.40 indicate, that the radial crack front is

curved, since the circumferential gages in the middle of the string are

the first to record cracking in both specimens. As the crack front progresses outwards the end gages near the disk and reaction ring eventually,

with increased load, signal that radial cracks have extended beyond the

middle region of the frustum. At approximately 65 percent of ultimate

load both specimens had developed radial cracking from the stem to some point beyond the failure surface. Although an additional grid of circumferential gages towards the disk stem would have given a more complete

picture of radial crack propagation, it was felt that radial cracking plays

only a minor role in the failure process, and that the extra available

embedment gages, which were in short supply and difficult to fabricate,

should be used to monitor radial strain along a second position on the

frustum.

In the paper summarizing his nonlinear, finite element, cracking analysis of

the Lok-test, Ottosen [7] presented a figure which showed the analytically

predicted cracking pattern. This pattern, reproduced in figure 1.2, shows a remarkably good correlation with the results above. However, two notable differences are evident: because Ottosen' s model assumed perfect bond between the

disk and concrete, large stresses —resulting in circumferential cracking—are developed at low loads beneath the pullout disk. In practice this is not the

case. Most contractors have the disks and stems coated with oil to prevent

rusting which, in effect, serves as a bond breaker along the side and bottom

face of the disk. The stress concentration at the upper disk edge, as shown in

the present study, is not accounted for in Ottosen' s model. We would therefore

not expect to see in the physical specimen the lowermost circumferential cracks

shown in Ottosen's figures 1.2a and b. Furthermore, a comparison of Ottosen's

figures 1.2c and d with figures 3.30 and 3.31 shows that the major circumferential cracks (from the disk edge to the reaction ring) generated by his analysis

follow the minimum principal stress trajectories. Since these trajectories do

not directly connect the disk edge to the inner edge of the reaction ring, the

analytical failure surface will form such that a band of uncracked concrete

between the disk and reaction ring remains after completion of circumferential

cracking. After circumferential cracking is completed at approximately 60-65

percent of ultimate load, the uncracked concrete acts as a "compression strut"

which carries the additional load. Ultimate failure, Ottosen asserts, is

therefore due to compressive failure of the concrete in the uncracked strut.

103
If we ignore the fact that the measured failure surface for specimens 1 and 2 connected the disk edge and inner edge of the reaction ring, and assume that

Ottosen's compression strut does indeed exist, then the compressive strain

along the strut should approach the limiting value of approximately 3000 microstrain as the failure load is reached. Figures 3.41 and 3.42 show the measured

axial strains along the frustum surface at the ultimate load for specimens 1 and 2, respectively. From these it is seen that the peak compressive strain is

1500 microstrain for specimen 1 and 1140 microstrain for specimen 2, both well

below that required to produce a compressive failure in the concrete. Additionally, there was no observed evidence in either of the specimens of a strut having failed in compression. It is, therefore, necessary that the load beyond

that required to complete circumferential cracking be carried by a mechanism

other than the "compression strut".

104
Figure 3.41 Axial strain profile along Frustum surface (experimental)

at ultimate load for Specimen #1 (2a = 70°)

105
MICROS

TRAIN

Figure 3.42 Axial strain profile along Frustum surface (experimental)

at ultimate load for Specimen #2 (2u = 54°)

106
4. DISCUSSION

4.1 FAILURE MECHANISM

The close correlation between the internal discontinuity histograms (presented

in figure 3.38) and the overall load-deflection response (figures 3.1 and 3.2)

indicates that there are three distinct phases in the failure sequence of a pullout test. The first of these phases is marked by a deviation from elastic

behavior. Previous work by Buyukozturk, et al. [16] has demonstrated that,

under unaxial loading, the main cause which precipitates nonlinear response in

concrete is raicrocracking at the aggregate-mortar interface. Under uniaxial

loading nearly elastic behavior was obtained to about 40 percent of ultimate

load in Buyukozturk ' s tests. For the pullout test, where the stress state

along the majority of the failure surface is biaxial tension-compression, this

value marking the beginning of nonlinear response will be less than 40 percent,

as is evidenced in the discontinuity histograms.

As the load is further increased, circumferential cracks begin to propagate

from the disk edge towards the inner edge of the reaction ring. This crack

system appears to be completely formed at approximately 65 percent of ultimate

load. Likewise, radial cracks, beginning at the intersection of the top surface of the concrete at the disk stem, have propagated through the entire failure surface by 65 percent of ultimate load (see figures 3.39 and 3.40). The

above events constitute the second key phase in the failure mechanism. If the

pullout test had been conducted using a perfectly homogeneous, brittle material,

the completion of circumferential cracking would have marked the ultimate load

of the test, since there would be no physical mechanism for load transfer

between the pullout cone and the base material. For concrete it is thus likely

that the entire load past 65 percent of ultimate must be supported by something

other than a continuous stress field. The logical mechanism is one of aggregate

interlock—by which the load is carried via the shear strength of the mortar

in which the aggregates crossing the failure surface are embedded. Beyond

approximately 80 percent of ultimate load, individual aggregates begin to shear

free giving rise to a softening of the system, marked by a change in the slope

of the load deflection curve.

It is, therefore, likely that the failure mechanism proceeds as follows:

Phase 1. Cracking initiates in the critical biaxial tension-compression zone

near the disk at 25-30 percent of ultimate. This phase marks the

initial load departure of the load-deflection curve from linear

behavior

Phase 2. Circumferential cracking, beginning at the disk edge, propagates

toward the reaction ring forming the completed failure surface at

approximately 65 percent of ultimate load.

Phase 3. Load past 65 percent of ultimate is carried entirely by discrete

forces developed via aggregate interlock. Ultimate failure comes about through gradual shear failure of mortar as embedded aggregates

which cross the failure surface are pulled free.

107
Note that the previously discussed phases are the same as those discussed in

table 3.2.

4 .2 AGGREGATE INTERLOCK MODEL

The load carrying mechanism of phase III can be demonstrated in the following

simplified model. First, envision an idealized spherical aggregate with diameter d located on the failure surface such that one half of the aggregate is

embedded in the pullout cone and the remaining half is embedded in the base

material such as shown in figure 4.1a. For the sake of simplification we assume the failure surface to follow the idealized conic frustum previously

discussed. The vertical force P required to pull the aggregate out of either

side is equal to the shear strength of the cement paste times the shear area. The shear area Av , as shown in figure 4.1b is calculated as:

Ay = 2 /qO

x r sin 0 dO (see figure 4.1b)

= -2 x r cos 0 |g0

= +2 x r but since r = d/2

Av

= xd (1)

whe re

0 = base angle for integration

r = radius of the aggregate

d = diameter of the aggregate

x = height of the shear surface = d / ( 2 tan a)

a = half the pullout apex angle

The force, P, to remove one aggregate is thus

P T „ X d

C c 2tana

( 2 )

where x c = shear strength of the paste

We now investigate the more generalized case where the failure surface is

represented as shown in figure 4.3. Here the aggregate spacing from the disk

to the failure surface is arranged such that each aggregate will shear out an area equal to Av as defined by Eq. (1). Assume also that the aggregates touch

one another in the circumferential direction. In this manner, the number of

aggregates around any circumference of the failure surface can be calculated

as

108
a) IDEALIZED AGGREGATE MECHANICALLY

BRIDGING THE FAILURE SURFACE

dA - Zrd 6

r sin 0

Z

= tan a

z

_

r sinfl = tana

dA = xr sin 0 d 6

VERTICAL DISTANCE TO

BE SHEARED

= d/2tana

x sin 6

b.) PASTE SHEAR AREA FOR SINGLE AGGREGATE = A v

Figure 4.1 Basis for aggregate shear failure calculations

109
Aggregate spacing L is chosen so that

there will be no interference between

shear surfaces in adjacent levels of

aggregates

J

L

Z

N

tf

. y

/ sm a

d

f 1 T 1 + T—

2 sm a

n d = < + T

L DOS a = y

Z

d / sin a + 1 sm a

cos a

2 h sin a

d cos a (1 + sin a)

Figure 4.2 Basis for the calculation of n used in eq . (6)

110
Figure 4.3 Generalization of idealized aggregate interlock failure

111
( 3 )

where = the number of aggregates around the circumference at a specified

level

= the diameter of each successive aggregate ring. (D^ is approximately

equal to the diameter of the disk.)

d = the diameter of the aggregate

The number of aggregates in the following successive rings will be:

N = H_?2 = iRl +

7Td(l + sing) 2 d d d

i( D,

N-, = + tt( 1 + sina)

^ d

ttD, = + 2tt ( 1 + sina) (4)

J d

7T 0

Na = —- + 3tt ( L + sina)

4 d

ttDi N = + (n-1) tt ( 1 + sina)

n ^

The total number of aggregates along the failure surface for any given pullout

test is thus:

n = N

nrr D, n + it ( 1 + sina)

(n-l)n

(5)

where a = the number of successive aggregate rings

between the disk and reaction ring

2hsina

dcos a(l+sina)

and h = the depth of embedment of the disk

The force required to overcome aggregate interlock is thus given by:

( 6 )

(7)

112
where A

t = total shear area = T

n

• Av= T

Q 2tana

( 8 )

and x c = shear strength of the cement paste binding the aggregate.

A cursory search of the literature has revealed little definitive

data relating this value to the compressive strength of concrete,

f£. However, it is known that under combined states of stress

the shear strength for concrete can vary from 20 to 90 percent of

f£ or higher [28]. Pending future experimental quantification,

and a detailed literature search an intermediate value of t c = .4

f'c has been assumed arbitrarily for the sake of illustration in

the following calculations.

Substituting eqs (8), (6), and (5) into (7), and assuming = D pullout force

is

T c "d 2

, 2h(sina)D (1+sina) / 2hsina \ ^ 2hsina

|

2 tana d^cosa ( 1+sina) 2

X dcosa( 1+sina) '

' dcosa( 1+sina)

Units are kips, inches and degrees. As an illustrative calculation we may

substitute the following data from specimens 1 and 2 into eq (9) to obtain the

pullout force:

Specimen 1 Specimen 2

h 10 in. 13.5 in.

D 12 in. 12 in.

a 35° 27°

d 1/4 in. 1/4 in

a v = .4 f£. 1.064 Ksi 0.472 Ksi

predicted

ultimate

pullout

force 399 kips 257 kips

experimental

ultimate

force 298 kips 259 kips

The predicted pullout force value for specimen 1 is about 30 percent higher

than the actual value. This is likely due to the large deviation of the actual

failure surface for specimen 1 from the idealized conic frustum assumed in the

113
calculation. The aggregates at the top of the actual trumpet shaped failure

surface wouLd puLl out at considerably lower loads than for aggregates on the

steeper, idealized surface. It should also be noted that the model assumes

that all aggregates pull out at the same time, and that the aggregates are packed as tightly as possible without interfering with adjacent shear surfaces.

Both of these may lead to overestimations of the actual pullout force. The predicted and actual pulLout forces are quite close for specimen #2. However,

this should not be misconstrued as proof of accuracy for eq (9), since the value

of t cis somewhat arbitrary.

What can be drawn from this hypothetical analysis is that a reasonable

approximation of the pullout strength was obtained by assuming that after 65

percent of ultimate load —when circumferential cracks have propagated from the

disk to the reaction ring —the entire load is carried by the mechanism of

mechanical aggregate interlock. Ultimate failure thus appears to be governed

not by tension, not by compression, but by shear failure of the cement paste

which prohibits vertical displacement of those aggregates which cross the failure surface. For this reason analytical solutions, based on a continuum theory,

are not applicable for predicting the ultimate pullout force, since beyond 65

percent of ultimate the load is carried via a noncontinuous , discrete mechanism.

An interesting corollary follows directly from this hypothesis: in an actual

pullout test the spacing of the aggregates, and thus the number of aggregates,

which cross the failure surface will be a random process; no where near as idealized as we have assumed for the above calculations. As was shown above,

for a given concrete the pullout force will be directly proportional to the

number of aggregates crossing the failure surface: for a smaller number of

aggregates mechanically bridging the gap, there will be a lower pullout force.

The "scatter" associated with the pullout strength for a given value of f£ is

thus likely to be a function of the random manner in which the aggregates are located aLong the failure surface. One might anticipate that greater precision

for the test could be achieved by enforcing a homogeneous material state along

the failure surface. This could be accomplished by selectively screening out

random aggregates and essentially testing the cement paste with the pullout test.

This, then, would have the effect of deleting aggregate interlock from the failure mechanism, and ultimate failure would occur upon completion of the circumferential failure surface. As previously mentioned, this latter phase (formerly

phase II) is principally governed by the tensile strength of the concrete, which

is known to have a direct correlation with the compressive strength.
5. SUMMARY AND CONCLUSIONS

At the inception of this study very little fundamental information was available concerning the internal behavior of the pullout test. The two notable

exceptions involved a plasticity analysis by Jensen and Braestrup [1] and a nonlinear, cracking finite element analysis by Ottosen [7]. Both analyses con- cluded that the pullout test directly measures the compressive strength of the

concrete. In an effort to provide experimental data for the verification of

these studies, as well as to calibrate further analytical methods, efforts were begun at the National Bureau of Standards in the summer of 1980 to design a scaled up version of the pullout test. These specimens were fabricated at a 12:1 scale based on commercial pullout test equipment and heavily instrumented

with waterproof embedment gages to measure the internal strain field at critical

locations. Two tests were conducted so as to achieve geometries at the upper

and lower apex angle (2a) bounds currently recommended by ASTM C 900. The

most important results of these tests, and of an auxiliary series of 2 axisymmetric finite element analyses, are as follows:

5.1 FAILURE SEQUENCE

The load-deflection curve and internal gage strain histories mark three distinct

phases in the failure sequence of a pullout test. These are:

a. Phase I: Initiation of circumferential cracking near the upper edge

of the disk between 30 to 40 percent of ultimate load.

b. Phase II: Completion of circumferential cracking between 60 and 70

percent of ultimate load.

c. Phase III: Shear failure of paste (or mortar) and degradation of

aggregate interlock beginning at 80 percent of ultimate.

5.2 INTERNAL STRAINS

Six gage strings were oriented in each specimen so as to record the variation

of internal strain with load along six paths in the vicinity of the pullout

disk and failure surface. The significant findings for each string are as follows

a. Axial strain parallel to the failure surface: Large compressive

strains exist near the outer disk edge and beneath the inner edge of

the reaction ring, the former being approximately twice the magnitude

of the latter. Between these two points the compressive strain

decreases, reaching its lowest point at about 60 percent of the

distance to the reaction ring with a value roughly half that at the

reaction ring. Studies of axial strain data at ultimate load revealed

that the compressive strains adjacent the failure surface were insufficient to initiate a compression fai ure.

115
b. Radial strain perpendicular to the failure surface: Large tensile

radial strains exist near the outer disk edge and decrease rapidly

along the failure surface towards the reaction ring. Prior to the

initiation of cracking the elastic radial strain reaches zero at

approximately 90 percent of the distance to the reaction ring, and is

slightly compressive at the reaction ring. The large tensile strains

at the disk edge are sufficient to initiate circumferential cracking

near the disk edge at approximately 30 percent of ultimate load. As

circumferential cracking progresses towards the reaction ring all

radial gages began to pick up large tensile strains.

c. Circumferential strain along the failure surface: These strains are small (less than 20 raicrostrain) , nearly uniform, and tensile along

the failure surface at a load equal to 20 percent of ultimate. One

specimen, with 2a =70°, indicated small negative (compressive)

strains (< 15 microstrain) in the vicinity of the disk edge. These

values disagreed with a companion finite element analysis which

indicated small tensile strains near the disk edge.

d. Vertical near side face of disk: These strains are tensile and

increase exponentially from the bottom disk edge towards the top disk

edge. Peak values exceeding the average limiting tensile strain of

the concrete (180 microstrain) were measured at the upper disk edge

at only 17 percent of ultimate load. This indicates that there is

little or no bond of the concrete along the side face of the disk.

The force is thus transmitted only via the top face of the disk,

producing a stress concentration at the upper disk edge.

e. Radial strains near top face of disk: These stains increase linearly

from near zero at the disk stem to a maximum corapresive strain at the

disk edge. Peak values were between -40 to -50 raicrostrain at 20 percent of ultimate load, and reached a maximum of -350 microstrain at

approximately 80 percent of ultimate, afterwhich the compressive strain

rapidly decreased and became tensile for the outermost gage at the

disk edge. The latter behavior was likely one to develop bending of

the gage.

f. Radial strain on a vertical line above disk edge: This gage string

indicated compressive radial strains at the disk face changing to

tensile strain at a point approximately 1/10 the diameter of the disk

above the upper disk face. The experimentally measured tensile strains

were generally less than + 10 microstrain at 20 percent of ultimate.

Large tensile strains at ultimate near the disk face for specimen #1

(2a =70°) showed that circumferential cracks had passed through

these gages to a height of 37 percent of the embedment depth above

the top face of the disk for these gages to monitor cracking the

failure surface must deviate towards a trumpet shape (rather than a conic frustum) for higher values of 2a.

116
5.3 PRINCIPAL STRESSES

A two dimensional linear elastic axisymmetric , finite element analysis was performed for each experimental specimen. Following good correlation with the

experimentally observed strains in the precracked state, the analysis was modified so that both specimens had identical material properties. It was thus

possible to isolate the effect of geometric change on the stress field along

the idealized frustum failure surface. These analyses indicate:

a. The maximum principal stress along the failure surface is positive

(tensile) and decreases with a decrease in apex angle. For equal

values of pullout force, the peak principal stress (near the disk

edge) for the specimen with 2a = 54° was 45 percent lower than for

a similar specimen with 2a = 70°.

b. The minimum principal stress along the failure surface is compressive

and, for the two specimens investigated, showed virtually no change in

the maximum values. This would indicate that the minimum principal

stress along the frustum surface is fairly insensitive to changes in

geometry. The peak compressive stresses occur near of the disk edge

and beneath the reaction ring with the latter being approximatley

half the value of the former.

c. The circumferential stress magnitude is not greatly influenced by

changes in the apex angle and is uniform and tensile over the majority

of the failure surface. It is compressive just near of the disk (to

10 percent of the distance to the reaction ring) and just beneath the

reaction ring (beginning at 90 percent of the distance to the reaction

ring)

The variation of these three principle stresses indicate three states of stress

along the idealized frustum failure surface (see figure 3.29):

1. Biaxial compression-tension from x/d = 0 to x/d ~ .1

2. Biaxial tension-compression from x/d ~ . 1 to x/d ~ .9

3. Triaxial compression from x/d ~ .9 to x/d = 1

where x/d represents the fractional distance from the disk edge to the inner

edge of the reaction ring. These values are identical for both specimens.

5.4 FAILURE SURFACE SHAPE

As the apex angle (2a) is increased the shape of the failure surface

changes significantly. For the low apex angle (2a = 54°) the failure surface is

nearly linear, following the idealized shape of a conic frustum defined by the

disk edge and the inner edge of the reaction ring. For increasing values of

2a the surface becomes more curved, assuming a trumpet shape. Correlations

between experimental and analytical studies indicated that the failure surface

very nearly follows the minimum principal stress trajectory from the disk edge

117
to the reaction ring . This means that the principal tensile stresses are acting

almost normal to the failure surface. Experimental data show that the compressive stresses (and strains) along this path are insufficient to cause compressive failure and additionally, that the normal (radial) tensile strains do

substantially exceed the average limiting strain of concrete along the circumferential failure surface. It can, therefore, be concluded that the propagation

of circumferential cracks (which form the failure surface) is controlled by

tensile strength rather than compressive strength.

5.5 FAILURE MECHANISM

It has been shown experimentally that three distinct phases occur prior to

ultimate failure of a pullout test, which are marked by changes in the slope of

the load deflection curve and by discontinuities in the load-strain histories

of embedment gages placed along the idealized frustum failure surface. What

precisely takes place during the first and last of these phases has been hypothesized in this study to be initiation of circumferential cracking, and a progresive ultimate failure by degradation of aggregate interlock across the

failure surface. Both of the first and third phases will require further

experimental work to quantify precisely. However, the second phase—that of

the completion of circumferential cracking between the disk and reaction ring

has clearly been shown experimentally to take place via propagation beginning

at the disk edge and ending at the inside edge of the reaction ring. Completion of circumferential cracking occurs at approximately 65 percent of ultimate

load (regardless of variation in apex angle). At this point, for a homogeneous

material, all continuity between the pullout "cone" and the base material would

have been severed, and indeed, if the material were perfectly homogeneous, it

would fail at this stage. But concrete is not homogeneous. Its fine grained

cement paste binds together a matrix of larger particles —aggregates—which

usually, but not necessarily, have material strengths greater than that of the

binder. Even though complete propagation of circumferential cracking has taken

place at 65 percent of ultimate load the presence of randomly spaced large

particles mechanically bridging the failure surface prohibits failure until all

such particles have pulled out of the retaining cement paste. This assertion

is an extremely important one and means that beyond 65 percent of ultimate load

any analytical model—including a nonlinear cracking finite element analysis-

which bases its failure criteria on material failure in a continuum is not

applicable. Beyond 65 percent of ultimate load in real concrete we are dealing

with a discrete failure mechanism, not a continuous one. A rudimentary idealized, discrete shear failure model, developed in the present study, gave failure

load estimates which were encouraging. Further experimental work is justified

to address the roles which are played by both the aggregate and cement paste,

and to further quantify material constants such as t c—the shear strength

(in pure shear) of the cement paste—and to relate this quantity to f£.

5.6 FURTHER RESEARCH

As stated in the introduction, the ultimate goal of the NBS study is to

determine which strength property of concrete is being measured in the pullout

test, and to optimize the test geometry—perhaps even to modify the test—so as to reduce bandwidth scatter relating the pullout force to a key strength

118
property of concrete, be it compressive strength, tensile strength, or shear

strength of cement paste. In essence the goal is to make the pullout test as reliable as possible. Realistically the strength property of concrete which

best correlates with the pullout force, if it is not compressive strength f^,

must eventually be related in terms of f£. for it to be of practical use. Further experimental work is justified at this time to investigate the

following topics:

1. The effect of aggregate on the pullout force (load-deflection study)

a) shape of aggregate

b) content of aggregate

c) strength of aggregate

d) size of aggregate

2. A study of the aggregate pullout mechanism

3. A study of the effect of apex angeeand aggregate content of the scatter

associated with the pullout test for multiple tests at a given concrete

strength.

119
REFERENCES

1. Jensen, B. C. and Braestrup, M. W. , "Lok-tests Determine the Compressive

Strength of Concrete," Nordisk Betong , V. 19, No. 2, 1975.

2. Kierkegaard -Hansen, P., "Lok-strength" , Saertryk af Nordisk Betong, V. 3,

1975.

3. Malhotra, V. M. and Carette, G., "Comparison of Pullout Strength of

Concrete with Compressive Strength of Cylinders and Cores, Pulse Velocity,

and Rebound Number," ACI Technical Paper #77-20. ACI Jounral May-June

1980, pp. 161-170.

4. Malhotra, V. M., "Evaluation of the Pullout Test to Determine Strength of

In Situ Concrete," Materiaux et Constructions , V. 8, No. 43, 1975.

5. Discussion: "Comparison of Pullout Strength of Concrete with Compressive

Strength of Cylinders and Cores, Pulse Velocity, and Rebound Number,"

by Snell, L. M. and Rutledge, R. B. and by Stamenkovic, H. Disc 77-20,

ACI Journal, March-April 1981, pp. 152-155.

6. Bickley, J. A. and Fasullo, S., "Analysis of Pullout Data from Construction

Sites”, 16 Ann. Meeting Transportation Research Board, Jan. 12, 1981,

Washington, D.C.

7. Ottosen, N. S., Nonlinear Finite Element Analysis of a Pullout Test,

American Society of Civil Engineers, Structures Journal, Vol. 107, No. 4,

April 1981, pp 591-603.

8. Richards, 0., personal communication

9. Newman, K., "Criteria for the Behavior of Plan Concrete Under Complex

States of Stess" , Proc . Inti. Conf: The Structure of Concrete, Cement

and Concrete Assoc., pp. 255-274, London, 1968.

10. Newman, K. and Newman, J. B., "Failure Theories and Design Criteria for

Plain Concrete". Proc. Southhampton 1969, Civil Engineering Materials

Conf., Wiley Interscience, pp. 963-993, London, 1971.

11. Carino, N. J. and Slate, F. 0., "Limiting Tensile Strain Criterion for

Failure of Concrete", Journal of the ACI, V 73, No. 3, pp. 160-163, March

1976.

12. Stone, W. C., Paes-Filhd , W. and Breen, J. E., "Behavior of Post-Tensioned

Girder Anchorage Zones", Center for Transportation Research, #208-2,

Austin, TX, 1981.

13. J. 0. Jirsa, personal communications

120
14. Rusch, H. , Grasser, E., and Rao , P. S., "Fundamentals of Design for

Uniaxial Stress Conditions in Concrete Members", Translation from German,

Oct. 1963 by J. V. McMahon and J. E. Breen, University of Texas at Austin,

Department of Civil Engineering, Austin, Texas, 78712.

15. Peterson, R. E., "Stress Concentration Design Factors", Wiley, New York,

1953.

16. Buyukozturk, 0., Nilson, A. H. , and Slate, F. 0., "Stress-Strain Response

and Fracture of a Concrete Model in Biaxial Loading, Paper No. 68-52,

ACI Journal, August 1971, pp. 590-599.

17. Richards, 0. "Pullout Strength Tests of Concrete," Research Paper,

American Concrete Institute, Annual Meeting, Dallas, Texas (1972).

18. C873 Ref. Tentative Test Method for Compressive Strength of Concrete

Cylinders Cast in Place in Cylindrical Molds (C 873-77T), 1980, Annual

Book of Standards, Pt. 14, American Society for Testing and Materials,

Philadelphia.

19. "Lok-test LTD Brochure", 43 Baywood Rd . , Rexdale, Ontario, M9V 3YB, Canada.

20. "Finite Element Cracking Models", Memo from Dave Darwin to Nick Carino,

March 1980.

21. "Axisymmetric Punching of Plain and Reinforced Concrete" Braestrup, M. W.

et al . , Special Report, Structural Research Laboratory, Technical

University of Denmark ( ).

22. ASTM C 900-78T, "Tentative Test Method for Pullout Strength of Hardened

Concrete," 1980 Annual Book of Standards, Part 14, American Society for

Testing and Materials, Philadelphia (1980).

23. Hsu, T. C., Slate, F. 0., Surman, G. M. and Winter, G., "Microcracking of

Plain Concrete and the Shape of the Stress Strain Curve", ACI Journal,

February 1963, pp. 209-223.

24. Launay, P. and Gachon, H., "Strain and Ultimate Strength of Concrete Under

Triaxial Stress", paper SP34-13, Symposium on Concrete for Nuclear Reactors,

pp. 269-282 ( ).

25. "Determining Concrete Strength for Control of Concrete in Structures",

Skrratajew, B. G., ACI Proceedings, Vol. 34, Jan-Feb 1938, p. 285-303.

26. "Stress-strain Response and Fracture of Concrete in Uniaxial and Biaxial

Compression", Liu, T. C., Nilson, A., and Slate, F. 0., ACI Journal,

May 1972, pp. 291-295.

121
27 . ASTM C42

28. Concrete. Mindess, S. and Young. J. F Cliffs, N. J. 07632, 1981, pp 400-401

Prentice-Hall, Inc. Englewood

122
Gage

ID

Adjusted

distance

"X"

wrt/Disk

(inches

)

RN 1 RN 2 RN 3 RN 4

RN 5 RN b

RN 7 RN 8

RN 9 RN 10

1.873

3.075

4.175

5.375

6.275

7.375

8.375

9.375

10.475

11.475

RN Gages (radial strain) N Specimen #1

Appendix A Table A.l RN Gage locations.

Specimen if 1

A-l
Adjusted

distance

"X"

Gage wrt/Disk

ID (inches)

RS 1 RS 2 RS 3

RS 4 RS 5 RS 6

RS 7

1.625

3.075

4.775

6.475

7.975

9.475

11.275

AS 1 AS 2 AS 3

AS 4 AS 5 AS 6

2.375

3.875

5.575

7.075

8.775

10.375

RS (radial strain) & AS (axial strain) Gages n Specimen #1

Appendix A Table A. 2 RS & AS Gage

locations

,

Specimen // 1

AAdjusted

distance

"X"

Gage wrt/Disk

ID (inches)

CE 1 CE 2 CE 3

CE 4

CE 5 CE 6

CE 7 CE 8

CE 9

CE 10

1.125

2.025

3.025

4.125

5.125

6.325

7.425

8.725

9.825

10.725

CE Gages (circumferential strain) n Specimen #1

Appendix A Table A. 3 CE Gage locations,

Specimen # 1

AGage

ID

Adjusted

distance

"X"

wrt/Disk

( inches

)

CW 1 CW 2 CW 3

CW 4

CW 5 CW 6

CW 7 CW 8

CW 9

CW 10

1.375

2.375

3.375

4.375

5.375

6.425

7.575

8.675

9.775

10.775

CW Gages (circumferential strain) n Specimen #1

Appendix A Table A. 4 CW Gage locations,

Specimen # 1

AGage

ID

Adjusted

distance

"X"

wrt/Disk

(inches)

ANE 1 ANE 2 ANE 3

ANE 4

ANE 5 ANE 6

ANE 7 ANE 8

ANE 9

ANE 10

ANE 11

1.875

2.775

3.675

4.275

5.075

5.975

6.775

7.675

8.475

9.275

9.925

ANE Gages (axial strain) n Specimen #1

Appendix A Table A. 5 ANE Gage locations,

Specimen # 1

AGage

Adjusted

distance

"X"

wrt/Stem

ID ( inches

)

TC 1 .75

TC 2 1.45

TC 3 2.35

TC 4 3.05

TC 5 4.05

TC Gages (horizontal strain) n Specimen #1

TC Gage locations,

Specimen # 1

A

Appendix A Table A.
Gage

ID

Adjusted

distance

"X"

wrt/Disk

( inches

)

RV 1 RV 2 RV 3

RV 4

RV 5 RV 6

2.125

3.125

4.125

5.225

6.225

7.325

RV Gages (horizontal strain) n Specimen #1

RV Gage locations.

Specimen // 1

A

Appendix A Table A.
Gage

ID

Adjusted

distance

"X"

wrt/Disk

( inches

)

ET 1 4.5625

ET 2 3.76

ET 3 2.96

ET 4 2.16

ET 5 1.26

ET 6 .66

ET Gages (vertical strain)

Specimen #1

— w

Appendix A Table A. 8 ET Gage locations.

Specimen # 1

A-8
"X"

Gage wrt/Disk

ID (inches)

RN 1 RN 2 RN 3

RN 4

RN 5 RN 6

RN 7 RN 8

RN 9

RN 10

2.125

3.275

4.575

5.775

7.075

8.375

9.775

11.175

12.475

13.875

INNER EDGE OF

RN Gages (radial strain) n Specimen #2

Appendix A Table A. 9 RN Gage locations,

Specimen # 2

AGage

Adjusted

Diameter

"X"

wrt/Disk

ID ( inches

)

RS 1 1.77

RS 2 3.07

RS 3 4.37

RS 4 5.62

RS 5 6.92

RS 6 8.17

RS 7 9.37

RS 8 10.77

RS 9 12.07

RS 10 13.47

INNER EDGE OF

RS Gages (radial strain] N Specimen #2

Appendix A Table A. 10 RS Gage locations,

Specimen # 2

A- 10
Gage

Adjusted

distance

"X"

wrt/Disk

ID ( inches)

CE 1 1.38

CE 2 2.68

CE 3 3.98

CE 4 5.28

CE 5 6.68

CE 6 8.08

CE 7 9.38

CE 8 10.68

CE 9 11.98

CE 10 13.28

INNER ED6E OF

\ REACTION RING

REACTION RING

CE2

CE Gages (circumferential strain) N Specimen #2

Appendix A Table A. 11 CE Gage locations,

Specimen # 2
Gage

Adjusted

distance

"X"

wrt/Disk

ID (inches)

CW 1 1.2

CW 2 2.6

CW 3 4.0

CW 4 5.3

CW 5 6.7

CW 6 8.0

CW 7 9.3

CW 8 10.7

CW 9 12.0

CW 10 13.4

INNER EDGE OF

CW Gages jeireumfsrentiaS strain) n Specimen #2

Appendix A Table A. 12 CW Gage locations.

Specimen # 2

A-l 2
Gage

ID

Adjusted

distance

"X"

wrt/Disk

(inches)

ANE 1 1.89

ANE 2 2.79

ANE 3 3.59

ANE 4 4.49

ANE 5 5.29

ANE 6 6.29

ANE 7 6.89

ANE 8 7.89

ANE 9 8.69

ANE 10 9.69

ANE 11 10.39

ANE 12 11.29

ANE 13 11.99

ANE 14 12.79

INNER EDGE OF

ANE Gages (axial strain) n Specimen #2

Appendix A Table A. 13 ANE Gage locations,

Specimen it 2

A-l 3
Gage

Adjusted

distance

"X"

wrt /Stem

ID ( inches)

TC 1 1.02

TC 2 1 .82

TC 3 2.82

TC 4 3.52

REACTION RING

E—

INNER EDGE OF

REACTION RING

TC Gages (horizontal strain) N Specimen #2

Appendix A Table A. 14 TC Gage locations.

Specimen // 2

A- 14
Gage

Adjusted

distance

"X"

wrt/Disk

ID (inches

)

RV 1 2.83

RV 2 4.28

RV 3 5.63

RV 4 7.03

RV 5 8.48

RV 6 9.68

INNER EDGE OF

RV Gages (horizontal strain) n Specimen #2

Appendix A Table A. 15 RV Gage locations.

Specimen // 2

A- 1 5
Gage

Adjusted

distance

"X"

wrt/Disk

ID (inches)

ET 1 .63

ET 2 1.33

ET 3 2.13

ET 4 2.93

ET 5 3.73

ET 6 4.63

INNER EDGE OF

ET Gages (vertical strain) n Specimen #2

Appendix A Table A. 16 ET Gage locations,

Specimen # 2

A- 16
RN

(north

radial

gages)

see

Appendix

for

gage

location

and

orientation.

3 b!

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r-H

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B-l

*MSTR

=

microstrain
Appendix B - Table B.2 - Pullout #1

Raw Experimental Data

RS (south radial gages) See Appendix A

for gage location and orientation.

LOAD

KIPS

SCAN

#

RSI

MSTR*

RS2

MSTR

RS3

MSTR

RS4

MSTR

RS5

MSTR

RS6

MSTR

RS7

MSTR

0 1 0 0 0 0 0 0 0

9.54 2 16 -2 4 4 -4 -8 -16

21.48 3 34 8 10 8 4 0 -20

33.37 4 46 14 16 12 -8 -8 -20

42.91 5 56 20 26 18 16 4 -20

52.44 6 73 26 32 20 20 4 -20

66.74 7 77 34 36 26 24 4 -20

72.28 8 85 38 46 30 28 12 -16

85.81 9 103 42 50 34 32 12 -16

95.35 10 113 46 58 38 36 12 -16

104.88 11 111 50 60 42 48 12 -16

114.42 12 125 50 64 46 56 16 -52

128.72 13 133 52 60 44 60 20 -48

138.25 14 153 52 56 42 71 16 -48

147.79 15 196 48 50 40 79 16 -40

157.32 16 204 48 50 38 79 8 -36

166.86 17 220 46 48 36 79 0 -36

176.39 18 244 44 50 36 83 -8 -32

185.93 19 272 46 52 38 83 -8 -32

197.82 20 310 50 56 38 95 -12 -32

197.82 21 316 48 56 36 99 -12 -36

207.35 22 327 54 60 30 103 -16 -64

219.30 23 369 62 65 30 111 -12 -67

226.42 24 411 67 75 28 127 -8 -67

238.37 25 522 93 99 -4 147 64 -60

247.90 26 812 127 95 -20 175 107 -52

257.44 27 1030 151 93 -24 183 79 -52

266.97 28 1175 161 99 -32 179 59 -44

257.44 29 1286 179 101 -38 175 48 -52

266.97 30 1280 196 107 -40 171 52 -71

278.86 31 1355 212 113 -40 175 56 -71

288.40 32 1449 260 127 -40 175 71 -67

293.17 33 1363 323 139 -40 163 119 -71

295.57 34 1411 335 139 -40 159 131 -71

293.17 35 1417 353 141 -40 156 151 -71

293.17 36 1469 383 139 -40 151 175 -71

293.17 37 1574 447 127 -40 147 226 -79

*MSTR = microstrain

B-2
Appendix B - Table B.3 - Pullout #1

Raw Experimental Data

RV (radial) gages. See Appendix A

for gage location and orientation.

LOAD

KIPS

SCAN

// RV1

MSTR*

RV2

MSTR

RV3

MSTR

RV4

MSTR

RV5

MSTR

RV6

MSTR

0 1 0 0 0 0 0 0

9.54 2 -2 -2 0 -2 -2 -2

21 .48 3 0 -2 5 3 3 3

33.37 4 -3 -2 2 3 5 3

42.91 5 -3 0 5 6 6 5

52.44 6 -6 0 3 6 10 5 66.74 7 -6 2 5 11 11 6

72.28 8 -6 2 6 13 13 6

85.81 9 -10 3 5 16 14 6

95.35 10 -11 2 6 17 16 10

104.88 11 -17 2 5 14 16 6

114.42 12 -30 -3 2 16 17 3

128.72 13 -51 -11 -3 11 14 -5

138.25 14 -76 -24 -10 2 10 -21

147.79 15 -106 -38 -21 -6 2 -35

157.32 16 -135 -54 -32 -16 -3 -44

166.86 17 -168 -70 -43 -29 -11 -57

176.39 18 -205 -90 -54 -38 -17 -65

185.93 19 -232 -102 -62 -46 -21 -71

197.82 20 -271 -122 -73 -52 -27 -76

197.82 21 -284 -129 -79 -56 -29 -79

207.35 22 -364 -160 -94 -70 -38 -86

219.30 23 -376 -138 -90 -105 -60 -102

226.42 24 -372 -92 -92 -148 -71 -105

238.37 25 -359 -68 -122 -186 -81 -106

247.90 26 -348 -59 -151 -227 -92 -116

257.44 27 -316 -54 -187 -284 -108 -122

266.97 28 -254 -54 -208 -335 -121 -127

257.44 29 -192 -62 -230 -391 -135 -130

266.97 30 -133 -70 -243 -435 -148 -132

278.86 31 -119 -64 -248 -446 -151 -133

288.40 32 -41 -35 -254 -478 -160 -133

293.17 33 64 27 -248 -486 -160 -127

295.57 34 81 40 -248 -487 -160 -127

293.17 35 103 56 -243 -483 -157 -125

293.17 36 141 94 -236 -478 -154 -124

293.17 37 198 175 -219 -464 -148 -121

*MSTR = microstrain

B-3
Appendix

B

-

Table

B.4

-

Pullout

//I

Raw

Experimental

Data

CE

(east

circumferential)

gages.

See

Appendix

for

gage

location

and

orientation.

<

o o3 o CM NO nO o o CO CO n£> nO nO O'* ON ON CM LO o co o o oo o r—H CO LO CM ON CM CO oo ON C* CM

r—H H «—

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H

r-H r-H r-H »-H ,-H r-H r-H t-H CJ £

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ON H r—H <

i CN CM CO co LO LO nO vO oo oo r>* r^. oo o CM vO O r-H r—

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CN co o ON ON ON oo u CO r-H »—H r-H CN CM CN CN CM CM r—H 1—

1

r—H cm

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r-H r—

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CJ 2 1 l 1 1 1 1 1 1 1 1 1 1 i i i

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uo H l i 1 1 1 t—

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r—H *-H t—

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CM co nO cr\ »-H CO lO CN ON r-H CM CN ro CO CN co CN 00 o ON 00 00 NO

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r-H r-H «-H »—H CNI CM CM CM CM CN CM CN r—H r-H 1 1 i 1 CJ s 1 1 1 1 1 1 i i i i 1 1 1 1 1

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t—

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r-H CM CM CM CM CM CO 00 LO U-l LO NO NO NO

CJ 1 l 1 1 i i i i 1 1 1 i i l 1 i 1 1 1 i i i

2:< T—

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CM CO uo nO 00 on o r-H CM CO <d- LO nO CO ON o r-H CM CO LO vO> oo ON o »-H CM CO U0 NO

CJ r-H r-H r-H r-H r—

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r-H r-H r—

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GO

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uo r—H r—

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i

i

(

B-4

*MSTR

=

mirrostrain
Appendix

B

-

Table

B.5

-

Pullout

#1

Raw

Experimental

Data

CW

(circumferential

west)

gages.

See

Appendix

for

gage

location

and

orientation.

<

s4 o X 00 •<r LO O' CN o vD ON -X) o 00 CO LO O' 00 O' co CO m CM r—

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*—

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od o cn CO o o X X X O' CM o'. ON o> CM o m CM m d* r^. o *-H o X 'D ON LO OO O O'

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CN CM r-H ,—

i—

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o CM o CM CN LO 00 co X r-H o oo O O o vO CT> d" LO On X O' on r». O' CN r—00 oo LO oo O'

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1 1 l 1 1 i i 1 l i i O 1 1 1 i

04 o CM CO CN X 00 X CT\ CO r—

x »-H x> LO co X) O' o o X r\ m CM LO r-H o o LO o CO H i i i 1 1

1 •-H *-H r-H CN CO LO m vO 00 CTn r-H r-H CN co on CO CN t—H »“

»—H o o ON ON O' oo

CO 1 1 1 1 i 1 1 1 1 1 i i i 1 1 •-H r-H »—H <

1 i i 1 1 3 1 1 i i 1 1 1 l 1 1 1 1

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d’ t—H X o O' oo X m *-H CO X on O' CM oo o r-H <3- CN co H i 1 1 i 1 1 CM CO -d" m x oo O' o> oo X m on 1 00 o LO t—H CO X o CO i i i 1 1 i 1 1 1 1 1 1 1 1 l t—

CM CN CM CN co o iZJ

o CM CO co vO r-H X a' r-H oo o r-H CO r^. co CN o <r oo r-- t—H CO "d- 00 oo o 'd- 'd- CO 00

CM H i i 1 i r-H r-H r-H '—1 CM CN co X) c- oo On r-H CN co LO X X X X oo o CM CO oo 00 00 oo 'D 3 CO 1 1 1 1 i i 1 1 1 i i i *—H r—

*-H r-H r-H t—H r-H r-H r-H CN CN CM CN CM CM CM CM CM CN u s 1 1 1 1 1 1 1 1 1 1 i i i i i i i i i i

•KOi o o o CN CO X o o o co "d- CM r-H CM o CO co o- vO CM X O' CO o r—

X t—H o o •—

H i i 1

1—1 r-H r—

r-H *—1 CN CO LO x oo cri r-H CM ^d* oo 1 CM co .—

00 LO CM r-H r-H 3 CO 1 1 1 1 1 i i 1 i 1 i 1 1 t—H r-H r-H 1—1 r-H CM CM CN CN t—H t—H r—

1 1 i u 1 1 1 1 1 i i i i 1 1 1 l i 1

3

< ’—

CN co <3* LO vO oo O' O t——

CM CO LO MD oo O' o t—H CN CO LO v£> oo On o t—H CN co -d" LO vO

C_> r-H *—

r-H r-H t—»-H r-H r~H r-H r-H CM CM CM CM CM CN CM CM CM CM CO CO CO CO CO CO CO co

c/} >r-

<d- oo r—

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p, LO CO o> ^d- CN 00 co 00 *d- CM <n oo CO O' 00 oo CO co 'd* co o> o> oo LO —i1 t—H t—H < cu o O' CO CN CM CN LO LO <3- 00 00 x> v£> LO o> vO 00 r\ N v£) 00 X co LO CO CO CO o M CN CO Mf* LO vO 00 o> o r-H CM CO <3- LO X r^. oo O' On o r-H CM co LO v£) LO MD oo ON O' O' o> O'

t—

r-H —* »-H r-H r-H —< r-H CN CM CM CM CM CM CM CM CM CM CN CM CM CM CN CM

B-5

*MSTR

=

microstrain
Appendix

B

-

Table

B.6

-

Pullout

#1

Raw

Experimental

Data

ANE

(axial,

northeast)

gages.

See

Appendix

for

gage

location

and

orientation.

r—i

,—1 oi o o r—H CM vo r^. ON o eg <p co r-H m O'* in CO MO oo MO co CM >d* eg CO -d- MO o r—H eg i—i Oh r-H Oh H <-H CM CO hP m MO On o <—

H

CO r-H MO r-H CO r-H eg CO oo eg m oo o Ch o r-H CM CM ON NO -d" m

%

ryiZ!

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o-—

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oi o LO *—H V© CM Oh m o vO CM o CO 'd- eg eg NO oo CO On m r^H o NO Oh 00 oo in NO o nO CO r—H O CM o

w H 1 r-H r—H eg CM co m MO ph. eg m 00 r-H Ml- in NO NO NO NO m m m CO o co H^ nd- r—H r-H CM CO NO r^

CO i 1 i i i 1 i 1 1 i i r—i r-H r-H CM eg eg CM CM CM CM CM eg eg eg CM r-H 1 1 1 i 1 1 r-H ' 2 I 1 1 1 i i I 1

i 1 i i i i 1 1 1

Oh oi o NO On CM i-h oo ON o ON ON CM ON ON eo NO <y\ eo r-H O CO ON 00 Oh r-H cn nd- NO r—

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1

r-H

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i 1 a 1 1 1 i i 1 1 l 1

oi o m m co nO m on co oo CO CM i-h r-H M0 in M2 On N CM O m oo eg co co MO oo eg »-H NO CO NO »-H 00 eg <r w H 1 i r—H »—

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eg eg CO co hP m m 00 ON o r-H og co co o NO co i i co m -d* cn r-H r-H eg

CO 1 1 i i 1 1 i i i i 1 I r-H r-H r-H r-H i—H r-H r-H i i a 1 1 1 1 1 1 1

nO oi o co m Oh CM m r—i oo MO 00 ON CM eg co CM CM 00 r-H o m o m NO eg o m CM co m oo 00 CO 'd- r—H w H 1 i »—H *—H eg CO CO <r <P m MO ON o r-H eg co co Mf Oh ON r-H eg 'd- NO r-H NO NO r-H rn r\ r^ r-H NO m

C/1 1 1 i i l i i i i i 1 r-H r-H r-H r-H r-H r-H r-H 1

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r—H r-H r-H r—H »-H i—

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m oi o vD r—H ON m CO oh Os oo r^ o MO o 'd- O NO eg eg m oo <r NO NO <r NO r—H f-H O oo Oh 00 r-H O ON

mj H 1 r—H r—H eg cn <fr m MO i-h oo o r-H co MO r^. oo o o CM -d- oo r-H m *—H o r-H CM rn o CM r-H co m m

JZ CO 1 1 i 1 i i i 1 i 1 r-H r-H r-H *—

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r-H *—

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r—i eg CM CM eg CM co m m MO r-H r-H ON r-H r-H eg co m

< a 1 1 1 1 1 i 1

i i i i 1 1 1 1 i 1 1 i «-H r—H

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<r 00 r-H m f-H oo *-H co r-H m r^ o NO m

p^l CO l i 1 i i i i i 1 1 r-H r-H r—i r-H >-H i—

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CM CM eg CO co CO nd* nd- m m m MO NO r-H r^ ON 2 1 1 1 1 1 1 i i i 1 i 1 i i 1 l 1 1 l i 1 i i 1

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co m o eg co co m 00 O eg m ON nd- o r^ o o m ON co CM r—H

pEp CO 1 i 1 i i I i i 1 r-H r—i r-H r-H *-H r-H eg eg eg CM eg CM co co CO co nd" in m NO r-H 00 oo ON O CM a 1 1 1 1 i 1 i i i i i i i i i I 1 i l 1 i 1 1 1 *—H

1

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r-H co vO oo r—

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)

CM CM eg CO >d- m NO MO oo oo ON ON o r—H eg CO co CO co hT m m m nd- m

CO 1 1 1 i i 1 i i l i i 1 i i 1 l I 1 r-H r-H r-H *—H *-H r—H r-H i—

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CM CM CM CM CM eg CM CM CM CM co CO co CO co co CO co

cn

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i •<r CO 1——

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m oo CM eg m On CM MO CTv CO CM CM m o eg r-H o r-H 'd- r-H vO o pH PH PH pH r-H m <p co ov r-H CM 00 co 00 <p eg CO 00 CO ON 00 oo CO co ^d- co ON *d- Oh nd- Oh oo -d* r—H in r—H r—H »—H Q CO • < 0^ o Oh i

i co CM CM v£> eg m m 00 oo NO MO m Oh M0 00 r-H r^ vO PH MO oo 00 CO in co CO co O M CM co <d* m MD 00 Oh o r-H eg CO m MO oo Oh Oh o r—H CM CO <r in MO in NO l-H co Oh Oh Oh ON Oh J t*i r—i •-H r-H —

r-H r—H r-H r-H -1 -1 CM eg CM CM eg eg eg eg eg CM CM CM eg eg CM CM

B-6
Appendix

B

-

Table

B.7

-

Pullout

//I

Raw

Experimental

Data

I

(axial,

south)

gages.

See

Appendix

for

gage

location

and

orientation.

|

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-d- <r m m

—H r-H

I I

CN

<J\ Oh

CN CN

I I

r—H CN

CN CM

CO CO

I I

on <r

r-~ oo

co cn

I I

—I CO O •—

i

Mf Mf

I I

-d- m m

I I

CM <1-

CTN ON

CM CN

I

~d- Ln

CM CM (O CO

I I

CM 00 ON O

co -dI I

»—H ON

cn <r

-d- <r

I I

N CO ^ O CN CO <r in vO N

<d- \o o

'd' Nt co nJ- ^ UO r-H NONvOCOOOOincOCOO irivoin\DNoo^ON^(^(^

B-7

*MSTR

=

microstrain
Appendix B - Table B.8 - Pullout //I Raw Experimental Data

TC (triaxial, compression) gages. See Appendix A

for gage location and orientation.

LOAD

KIPS

SCAN

#

TCI

MSTR*

TC2

MSTR

TC3

MSTR

TC4

MSTR

TC5

MSTR

0 1 0 0 0 0 0

9.54 2 -4 0 -4 4 0

21.48 3 -4 -8 -8 -8 -11

33.37 4 -4 -8 -12 -16 -24

42.91 5 -4 -8 -12 -24 -40

52.44 6 -8 -16 -20 -40 -52

66.74 7 -12 -20 -28 -44 -64

72.28 8 -8 -24 -32 -48 -79

85.81 9 -16 -28 -32 -52 -103

95.35 10 -20 -36 -32 -60 -111

104.88 11 -20 -36 -48 -67 -115

114.42 12 -24 -40 -48 -83 -139

128.72 13 -28 -52 -60 -99 -159

138.25 14 -32 -56 -67 -119 -191

147.79 15 -40 -71 -83 -143 -214

157.32 16 -48 -83 -91 -159 -238

166.86 17 -52 -91 -103 -183 -270

176.39 18 -56 -103 -115 -210 -294

185.93 19 -64 -115 -123 -222 -318

197.82 20 -68 -119 -139 -250 -325

197.82 21 -64 -123 -143 -258 -329

207.35 22 -71 -143 -163 -302 -333

219.30 23 -79 -147 -163 -322 -345

226.42 24 -87 -167 -183 -349 -345

238.37 25 -91 -175 -194 -385 -341

247.90 26 -95 -191 -206 -405 -329

257.44 27 -99 -206 -226 -452 -310

266.97 28 -103 -222 -238 -488 -286

257.44 29 -99 -226 -250 -528 -250

266.97 30 -99 -234 -266 -556 -222

278.86 31 -107 -246 -278 -579 -226

288.40 32 -111 -262 -290 -623 -210

293.17 33 -119 -282 -306 -671 -206

295.57 34 -119 -282 -310 -687 -214

293.17 35 -119 -286 -318 -695 -222

293.17 36 -127 -294 -322 -714 -246

293.17 37 -131 -314 -345 -766 -345

*MSTR = microstrain

B-8
Appendix B - Table B.9 - Pullout //I Raw Experimental Data

ET (edge tension ) gages. See Appendix A

for gage location and orientation.

LOAD

KIPS

SCAN

#

ET1

MSTR*

ET2

MSTR

ET3

MSTR

ET4

MSTR

ET5

MSTR

ET6

MSTR

0 1 0 0 0 0 0 0

9.54 2 3 0 5 8 3 2 21 .48 3 5 2 5 14 8 3

33.37 4 5 6 14 19 13 3

42.91 5 6 5 14 24 14 3

52.44 6 10 11 21 27 17 5

66.74 7 8 10 21 32 21 5 72.28 8 10 10 22 35 22 5 85.81 9 13 11 24 41 25 6

95.35 10 13 13 27 44 32 6

104.88 11 16 16 33 48 33 6

114.42 12 13 14 30 48 37 8

128.72 13 11 11 30 48 35 6

138.25 14 10 13 32 46 32 6

147.79 15 8 10 30 43 32 6

157.32 16 8 11 29 43 35 6

166 .86 17 10 11 30 44 33 6

176.39 18 8 11 29 41 33 6

185.93 19 8 11 32 40 32 6

197.82 20 3 5 24 41 27 6

197.82 21 3 6 27 40 29 6

207 .35 22 3 8 29 40 21 6

219.30 23 2 6 30 41 19 6

226.42 24 0 6 29 41 19 6

238.37 25 -2 3 25 40 17 6

247 .90 26 3 8 30 37 19 6

257 .44 27 -3 0 24 37 11 6

266.97 28 -3 0 25 37 13 6

257 .44 29 -3 3 29 35 11 6

266.97 30 -8 0 25 35 11 5

278.86 31 -6 2 30 37 11 6

288.40 32 -5 0 29 37 10 5

293.17 33 -11 -5 25 41 11 5 295.57 34 -11 -2 30 43 13 6

293.17 35 -13 -6 25 44 11 6

293.17 36 -14 -8 25 46 11 5 293.17 37 19 25 46 62 14 5

*MSTR = microstrain

B-9
Appendix B - Table B.10 - Pullout #1

Raw Experimental Data

LVDT Displacements in thousandths of an inch

LOAD SCAN S VERT LVDT N VERT LVDT NE HORIZ LVD SE HORIZ LVD W HORIZ LVDT

KIPS # DISP* DISP DISP DISP DISP

0 1 0 0 0 0 0

9.54 2 0 -1 0 0 1 21 .48 3 1 -2 0 0 1 33.37 4 3 -4 0 0 1 42.91 5 4 -5 0 0 1 52.44 6 6 -7 0 0 1 66.74 7 8 -8 0 0 1 72.28 8 8 -9 0 0 1 85.81 9 10 -11 0 0 1 95.35 10 11 -12 0 0 1 104.88 11 12 -14 0 0 1 114.42 12 13 -16 -1 0 0

128.72 13 14 -18 0 0 0

138.25 14 15 -19 0 0 -1

147.79 15 16 -21 0 0 -2

157.32 16 16 -23 -1 -1 -2

166.86 17 17 -25 -1 -1 -3

176.39 18 18 -27 -2 -2 -4

185.93 19 19 -29 -3 -3 -4

197.82 20 19 -31 -4 -4 -5

197.82 21 20 -33 -4 -4 -5

207.35 22 20 -36 -6 -5 -7

219.30 23 20 -39 -7 -6 -7

226.42 24 20 -41 -8 -7 -8

238.37 25 20 -43 -9 -8 -9

247 .90 26 20 -45 -11 -9 -12

257.44 27 19 -49 -13 -11 -14

266.97 28 18 -52 -16 -15 -16

257.44 29 16 -55 -19 -17 -19

266.97 30 13 -58 -21 -20 -21

278.86 31 13 -60 -22 -21 -22

288.40 32 12 -65 -25 -24 -25

293.17 33 8 -70 -31 -28 -30

295.57 34 8 -70 -32 -29 -31

293.17 35 7 -72 -33 -30 -32

293.17 36 4 -74 -35 -32 -33

293.17 37 0 -78 -40 -36 -38

*thousandths of inches

B-10
Appendix

B

Table

B.ll.

Pullout

//2

Raw

Experimental

Data

RN

(North

Radial)

Gages.

See

Appendix

A

for

gage

location

and

orientation.

o

T—l o CO CM co co MO M0 o o o 1—1 in o i—i NO CO CO co CO CO CO CO CO CO CO CO CO Nf •sO CM co

<r CM co i—

1—

<r CNJ Nf Nf 00 MO 00 Nf oh o Oh Oh o> Oh Oh Oh Oh ON Oh Oh Oh CM Oh o co

cd l 1 i 1 i—l T—I r—1 Oh nt Nf <r Nt nt NT NT NT NT NT •NT rH o o CM

i 1 T—

co CO CO CO co CO co CO CO co co CO r—l T—l

1 CO

1 CM

1

Oh o CO nO Nf o 00 nt oh o m CO CM vO ah Nf CNJ NO NO NT un 1—1 co CO Oh o T—l 2 r—

1—

Nf NT m n- 00 oo o CM i—l 00 o co MO o CM MD co i—l Oh 1—1 CM NT NO co T—l m co

1—1 i—

i—l 1—1 CM CM co CO CNJ i—i i—l r—l 1—1 r—l i—i 1—1 T—l 1—1 T—l T—l

00 o co o 00 1—

00 r—

o oo M3 CNJ CO co m CO ON CM CM o cn 00 Oh NT Nf CO Oh hO

r—l nt o> <1- nt co co m CO oo NO o m C- Oh NO 00 ON cn 00 Oh m CM NO eg 1—1 co 3 i—

i—i—i—l i—1—1 T—

nt 00 ON r—

CM co nt m c- NO in CO co cn CO

r—l r—l T—l r—l i—l I—l r—l T—l r—H 1—1 r—l T—l i—l i—l

o oh oo on r—

CM t—l CM oo CO 1“

o NT r—i ON NO CO NT CM Oh ON CM NO m Oh m ON NO

CM co oo 00 r—

oo 00 Oh i—o CO NO T—I oo r\ NO oo o oo o CM CO m Nf NO m o CM 3 i—

1—

7—1 1—1 r—

CM CM 1—

1—1 I—l i—l i—l r—i T—l i—i r—l CM CM CO Nf

vD o nO 00 Nf 00 in o 00 CM NO Oh NO nt CM c- i—i m r—l co CNJ CO Oh Oh CM T—l CO ON

CM CO 00 O co NO no 1^ ON 00 NO CO Nf o NO NO NO e- r—l i—CM co r>. 1 1 t-H 00 NO S t—

i—l 1—1 1—1 T—

1—1 T—

1—1 T—I f—l 1—1 CM CM CM CO

in o LT> m NO CM CO m o 00 CM NO 1^ in in 00 co o NO co NO co CM CM 00 NT 00 00 O' i—l m o

CM CO e- oo 1—in o CNJ in 00 00 o ON NO Oh CM o nt CO 00 CM <r r—l o CM o o ON r^. Nf 3 i—

t—l

CM CM CNJ CM CM CO CO CO CNJ CNJ CM i—i i—

r—l 1 i—i CM CO Nt m 00 00 o NO

i i i i i 1 1 1—1

1

T—l

1

<T o Nf co u~l o 00 MO m nO Oh nt CM i—l oo 1—1 NO CM NO NO o o 00 ON co NT ON Nf NO NT CM in

*2 CM CO 1^ on 1—1 CM 1—

1—1 nt CO CM nt CO CO 1—1 1—1 r—l co NO CO o NO CM CO o 00 CO o

PS T—

i—l 1—1 1—1 t—i i—l i—1 T—

CM CM CM CM CM CM CNJ CO m NO oo o r—l Nf m 00 cn

i—l i—1 i—l i—i 1—1 CO

CO o m o m CM CO 1—1 m nO o m co 1“

Nf o m i—l ON nt 00 NO o 00 NT o Nf i—i 1—1 1—1 co c- co m o CO m 1—1 t—i NT CM o nt 1—1 ON ON i—l NT 00 m Oh NO o m Oh NO o 00 o o o 3 i—

r—

CNJ CM CM CNJ CM CM CM CNJ CM CO CM CM co co CO nt m T—| CM NT 00 o o CM CM CM

1—1 T—l r—l 1—1 CM CO CO CO CO

CNJ o o m LO| c oo o 00 oo NO 00 NO oo oo o 1—1 co nt NO ON <r co 00 ON CM 00 in CM CM 00 NO

*21 r—1—1 T—I T—I i—i T—l 1—1 r—l CM co •<r in oo T—l 00 T—l CO ON ON

Pm T—l 1—1 CM eg CM O1—1

•4- i—

o o CO CM co nO CM CNJ m m o o o 1—1 Oh T—l NO T—l o i—l o NO 00 i—l NO O co CO 1—1 CO

*2* r—

r—

CM co 1 1 co CO CO CM Nf CO Nf in co o m o m NO ON oo c- CNJ ON r^. oo Nf 1—1 o T—l CM CNJ co co in r^. 00 o 1—1 CM m NO 00 co

i—l i—4 i—i r—l 1—1 T—l eg

'c

dJ o CM o ON co CO nt <r on Nf o 00 »—l 00 ON O CM CO nt Oh o m o 00 CM oo o Oh CM NO o Csl CO CO m MO 00 00 Oh o o i—

1—1 CM Nf m VD o co Oh r—l NO ON CO Nf Nf Nf

CO 1—1 T—

t—l I—l i—l 1 1 T—l i—l r—i 1—1 CM CNJ CM CM CO CO CO NT Nf Nf Nf

* o 00 00 o CO co r—

Nf CO NO CG m 00 o co MO 00 1 1 Nt hO Oh CM m 00 vD NO o i—i 00 i—

e- CM Oh 00 o m vn 00 CO ON Nf O' Nf O in o m r—l NO m CO T—l 00 CM CM 00 Nf o r—

m m T—l co m o ON VO co co CNJ CNJ I—l T—l 1—1 o o ON 00 o in O NT C- Nf CO

r—CM co nt vO 00 o o 1—1 CM co nt m vC 00 o o o r—H CM Nf *<r m m m m m Nf CO

1—1 r—1—1—1 1—

i—i i—i r—l T—l r—l r—l CM CM CM CM CM CM CM CM CM eg CNJ eg CM

B-ll

*

kips t

microstrain
Appendix

B

Table

B.12.

Pullout

#2

Raw

Experimental

Data

RS

(South

Radial)

Gages.

See

Appendix

A

for

gage

location

and

orientation.

o

r—

o o rH CO CO i—

O <3- CM rH o CO 00 o rH CM sO rH r-H m CTs o CM o 00 CO 00 sO 0 CO CM

00 I-H i—

r-H 1—

r-H r-H i-H rH rH I rH r-H rH rH i—

rH CN <3- r^> as CO Oh <3* CN sO CN rH rH d I 1 I rH rH CM CN i-H rH <3*

!

rH OS • o 00 sO SO sO as CNJ NT r>. as m CT 04 as co <3- CO m rH CN as CTs CN CO <3- as rH CM

CO rH rH i—

rH r—H CN i-H CN CNJ CM i-H CNJ rH CM r- i-H as CN m 00 <3- 00 o CXD 00 Os 00 d r-H rH rH CN CM CN CN CN i-H r—

rH rH 1 1 rH

00 o CO i-H CNJ m m 00 o sO 00 <3- CO <3- sO CN rH o sO rH O co m O 00 o sO sO CM as

CO 1—i—CN| CN CN CN CO CO <r <3" <3" in as <3* sO o CN <r rH >d- o CO rH rH CN

oi rH CN CO m 00 CM in sO sO SO m rH 0 m CO

rH i—

rH rH rH rH rH rH 1

o <r i-H CN <3- O i-H sD CN CO as sO as as o- 00 as CO rH CN sO rH as o CN sO

CO T—CNI CN CNJ CN CO *<r CO m m •<3 -d- Nt •<f SO rH CO o <3* 00 o CO <3- in df rH SO d CM co co co CO <3- in m in m in m in !

v.O o CO CTs in m CN 00 CM <3" m CTs sO in CN m O 'd' r^. rH sO CM LO rH co CN CM CTs m sO 0 CO ?—

i-H CNI CN| CO co in m n- 00 os co <3* <3* rH 00 00 <r co CO sO 00 in rH 0 m 00

oi rH rH i-H rH 1 I rH CN CO 00 as 0 0

1 i 1 I I 1—

1 rH ,

1

UO o as CNJ 00 CN o co m CTS in o m 00 as rH rH CTS in in <3* r-- m CM rH

CO i-H i-H Cxi CNJ CNJ CN CO CN CO CN CO •vf sO CO o as 00 o o rH as co m i-H 0 CN 00 0

d rH CN r—H rH CN CN CN rH rH rH rH

l rH

1 rHl CNI

o <3- s£> i-H CNI in in O SO rH in 00 00 sO rH sO co 00 sD <3 sD m as in CO CO CO rH

CO i-H r-H CN| CN CN CN co CO NT <r <3- CO co co <3* CM ’d' CO CO <3- o- Os i—H CN <3" m sO m

pi rH rH rH rH rH rH rH rH

CO o i-H as CN CO o n- rH o CN <3* Nt CNJ CM sO m Nt co l"~ CN in co 00 00 in CTs CO

co r-H CN CN| CN CO CO <3* CNJ Nt <3* CO CN CN CM co rH CN i-H rH CN CO m sO so so sD sO <r rH

pi

CN o as C\! <r 00 CO CO i-H rH sO in CNI CTs co <3" n- in <3- sO CN o CN CO 00 in so CO as CN <r CTs CO rH CO <3

- <3- so sO sO <3- m <r CO CN 1 1 rH CNJ as CN sO o CTs sD SO <3- CM rH o CTs in as

pi rH rH CM CN CO in r^. 00 as o o as as rH sO

rH rH rH rH

+-

rH o r- o o i-H CNJ Oh CN o rH rH <3" <3“ CN o CO o o CTs 1—1 o as o CO CN CM sO 00 as CN ! CO r-H CO m in SO m m <3- m sO 00 o in rH n- co so CO co in 00 O <3- sO as d CN CM CO CO -d- in as o rH CO co <3- <3- <3- <3

r-H rH r-H rH rH 1—1 1—1 rH

00co c o CN c Oh CO co -d- as o 00 i-H 00 as O CN CO <3- as o in o 00 CN oo o Oh CN kO o rsi CO CO LO sO 00 00 CTs o o rH rH CN in sO o- o co as i—1 SO Oh CO <3* *<3 <r

CO rH rH rH r-H rH rH rH rH rH i-H CN CN CN CN CO co co <3

•X o CO 00 o CO CO • 1 NT CO so CO in 00 o CO sO co i—1 Nf Oh CN in 00 <3" rH sO 0 rH -o co 1—

CN Oh oo o m SO 00 CO as as vT o m o in i—

vO in CO i—1 00 CM CM 00

• o i-H sO LO IT) «—

CO m o Oh sO CO CO CM CNJ rH rH rH o o as 00 o in o <3- <3- co

hJ i—

CN CO sO oo o o 1 1 CNJ co <3- in kO 00 Oh o o i—

CN *<r n in in m in 'd- co

r-H i-H 1 1 rH rH rH rH rH rH rH rH CN CN CM CN CM CM CN CN CN CN CN CN CN B-12

•x +-

kips

microstrain
Appendix B Table B.13. Pullout #2 Raw Experimental Data RV (radial) Gages.

See Appendix A for gage location and orientation.

Load* Scan# RV It RV 2 RV 3 RV 5 RV 6

0 0 0 0 0 0 0

11.88 4 14 10 10 10 11

26.18 22 22 17 14 11 16

35.70 30 22 16 14 16 14

45.23 39 19 13 10 14 10

61.93 53 27 22 14 24 14

73.81 63 22 22 13 22 13

85.74 74 24 29 21 33 22

100.03 84 -8 14 8 32 5 109.56 89 -10 24 21 43 8

116.68 94 -32 10 5 30 -14

123.85 100 -30 16 16 40 -11

133.38 108 -57 5 8 30 -30

142.90 111 -67 11 13 29 -33

152.43 118 -68 27 11 13 -56

161.96 129 -36 62 10 5 -48

171.48 140 19 62 -10 -22 -78

181.01 152 49 86 0 -14 -62

190.54 163 71 97 -16 -25 -81

200.06 174 71 106 -32 -30 -95

209.59 179 58 127 -40 -24 -98

214.12 200 -40 145 -63 -13 -119

228.65 235 -171 193 -63 10 -117

240.58 270 -574 246 -48 30 -152

245.34 298 -692 274 -2 52 -157

250.11 312 -731 290 59 76 -176

254.87 368 -836 297 162 22 -187

257.26 390 -896 295 211 -19 -188

257 .26 439 -1518 324 463 -190 -195

254.87 442 -1616 339 517 -247 -190

247.70 446 -1906 379 690 -309 -190

233.41 447 -2479 422 926 -368 -187

* kips

t microstrain

B-13
Appendix

B

Table

B.14.

Pullout

#2

Raw

Experimental

Data

CE

(Circumferential

East)

Gages.

See

Appendix

A

for

gage

location

and

orientation.

o

O *—< no co on *-h cy> m c—

00 50 ON vO CM m cn co o NO CM 05 m CO o o 00 CM 00

rH f-H i—i-H CNj CM CM CO Mf <r m m m m vo NO in 50 l m

1 r*H r^. co

1 0

m

1

CMNt

fH

NOm

f—H ONH

MfoCM

O vO 50 H co on <r o <! CO o o F—

o CO CO ON CO CO 50 50 <fr ON o CN cm o ON

r-H f-H CM H H M CO CM CO Mf Mf CO CO <1

* Ml- in m 00 NO f-H Mf —< 00 50 00 H CM CM CM 1

’

1

1 00

1

f-H

i r CMcm

i

8

Oro—

0 (Ti—ir^vooor^OCT\cor^’X)—

HHHHH-iNHcnNcMNNN i i/'iLnoooo<tvOOONOor''.r-~cMaoor-~

I | | *—* >—< —i i—i cm i i <f -<f o

I I I I I I I I I I I I —

I

8

omoo^mooHooNN<oo\0 '£i(MO(nNcn<fNOflONflH(sno^ *—t *—l i—(i-H i-H i—

I HHHHtMCMCMNvt

i—i r.. oom Mf om

8

O'JCliO'nOO'OnN^iCOlNOiiOiOOlSNiONiO'OMOHS —I "-I ^4 — ,~i I HHHCMNniO^f'OOSODOOOCO'OMNiO

I I I I I I I I I I S I —I I I I I

I

mt o m m

m o in o cm oo oo oo

i

I

I 8

OvfH^cON'JNNHinO'OHMOiONffiOlOCMiOinOO^N'O •—I CM CM >—I CM CM CM .—I I HHNNSOOCOi/I^ONON I H N CO

I I I I I —t-—< —i—« —i! I 1 III I I I I I O

o

<3- 00

r-i Mf

Mfm

till

cow

u

o <r id co n cm oo H —I i—l | ci n h o\ co in n n <t co i

I l I l l I

o\ <t h h m o n

05 oi oo n n r> iO

I I I I I I I

CM Mf 00 -H 50 CO Mf

lO <f O H H o 00

I I I •—

I

CM r-< O ON POO# H N CM H

O 1^ o Mf m Mf 00 OO O Mf Mt 50 o CM 05 m CM CM 00 Mf 50 o o o CO 50 CM 05

oo CM Mf

1 CM 1

1 Mf

1 m

l Mf1 in

1 051 501

IN.

i CO1 m

1 CM 1 co

1 CM

1 CM <fr CO

CM of-H

ON

<fr CM omCM

CMm

CM

05CO

CM

8

O CO <T O

I I <r <f to io ill —

i

CM Mf 00 50 oo Mf Mf m

l i l I

cm m co h

N 00 O CO

<3- o o <r

oi h m n —I CM CM CM till

•—

t-H O CO OOCMO eo co >3 in till

H CO N N

m cm m co m id in 05 III!

co o —< in m o o ifi N CO IN ^

c

cOo

C/5 o<-NooicoiO'3

, <tai'C, oco^oooiocMcO'}oioinO(ioN®oaiNioiN cMcocomiosoooooiooH-icvi-jiniorNNOcoiNOi'^iDoicO'3 <f'sf HHHHHHHHHHNN(MCMCO(OCO'J'3 -4'<f

*

-a

cdo OOOOOOCOIOH^IOiOOOmoOOIOOOOH^^CJlNinoO'JHMOONOH 00 ’HrNCM 0i 05 iNOmv005 c0 (7isfai<tOinOLnHi0 in, cOr-icocNcMOON<r

HcDiniOHcoioouiiOcococMCMHHr-lOOoiNj-ooomONfr^rN-vrr-Nco HCMco<f\DNoooo^Nco-3 in'Oscoi3' 00 '-iN'3 '}ininininuw(o hhhhhhhhhhhNNNCMCMNNCMNNCMNN

•K 4-

B-14

kips

microstrain
Appendix

B

Table

B.15.

Pullout

//2

Raw

Experimental

Data

CW

(Circumferential

West)

Gages.

See

Appendix

A

for

gage

location

and

orientation.

0 5 8 2 0 <r co CM <r CO co nO o NO in o no NO NO in r-H O'* o ON o o CM m r-H r-H CM r-H CM r-H CO —

r-H r-H r-H | rH r—

*—CM <1* m <r LO OO or-H

1 CM

in

CM 00

CO

00<r

ON

00

O O no oo —< on

1 '—i r

i-H i—H CM *—*4 U0

cm O N M ON <fco —i CM —i CM h h n oo o o i-H CM t—

»—

UOncT'-^COOnCTnCOCOCMCO NNrH(*1<J-ONtNlCT\vOOO

| rH rH rO N N N

I I I I I I

omocoNOvto^ (MNnH<tfP)socNiH^inoiriOoO'0 (»'0 <f'OooHro

I t HHHHOIMMnnCM fONfvONSXINNfO

I I I I I I I I I I I -I N VO

OcMinininc>ONOocONONOOcoiv.oNCMinao<rcMcoco —ioO\ON<fr'CO'iii^

—I I | | HHMNiOOS 005\OH(*1 (*1 WNHHNOH I I I I I I I I I

*—I *-H <—I I—I »—I 1—I •—I | | —I |

I I I I I I I I

OLO-JOOOO'O'DOO (*10 »tnvCH 0 Hifl 0 NNcM 0 N'i) H(OH<f<tm\0'ffMiriro(ncNi

l l l l I I l I I I l I l

NO U0 CO NO 00 m

I CM NO NO I NO

I I I I I

O in OONOnONO'ONiOOOHNOiflOO^NO"BN>JNHfnOHNHHinN •—< HHncNimuNvON^(j>oorNNONtN(Ni^^<fuNinoooooN

I I I t t I I I <—i I I I I I I I 1 —I —i CM co

3

ONeoifl(YioNninvONCOfOin<fffi'*ONOMnm I ^ HNfONtmNOMOOO'O'OOMNO I I I I I I I I

'—I I I—I—l-H—

I I I I I I I I I

ON CM 00 00 CM ON O' nO CM o ON CM i—

O O ON ON on NO

CM CM CO NO *-H CM m 00

I t

I I I I

ONnonootNiajNjinNonNiriNOssoininoooocio^tnoNNNiorioo

I —1 I I I CSNNfOirnONMvOUNNO'OINNNOONDNMOOONON

I I I I I I I I I I I I I I I —• I I I 1 —l —i —

I

CM

e

cOo

co

OO—icMcOCOOcONOOCMNO<fO<r

—( I HHClNfNflinOOOH

I I I I ! I I -i -i

I I

OnOCOOCOCMOOO fO<tro<frO(Ono r-H *“H *—H »—I r-H *-H *—I *-H

i i i i l l l i

rvcMCMinaNcooococo oONOcscouONomoNLO III co <r cn iv

Ocorv<rmNOcoooo —icncm^v—lOCJNrvONONCMuncMoomNDooooONOOcoirN

—irvcocMOrvCMON|v.<}-mNOoO'<tocM—HCMcooocvicvi | olonocmiv.cmonon

IOCNInT I HHOJHfCI I I |r-tH<f-4(N||H(MI<f <f —I | —I NO V H \D III III I III I I I II I I I —I CV1

I I

O'JNOOW'HCl-JsfON^OCO CMcocomNDrvcoooONOO 00 ON o H <NJ -J

CM CO m no

<j- ON O LO o 00 CM 00 o On CM NO rv o CO ON r—

NO O' CO -tf -3- CM CM CM CM CO CO CO vT vf

*acti o

o oo oo o

00 -H IV COCO—I'JOnOXMOOOOOnDOO—i NO ON n in CO Nt CMONOOrvOLONOOOCOONVl-O'-^OLnOiCN—inOlOOO

rv no no v O —

00 CM N 00 N Nj1 Nomuo^cocnooNNorococMCM^—i—lOOONvtooocnO'd’rvrvvrrvco (NiO'jNOvoooOHNfO'firiNovooo'OO'-iN-jNj'UNiriininin'fo hhhhhhhhhhh!M(MCMCMCMNN(MNN(n|CMN

B-15

kips

microstrain
Appendix

B

Table

B.16.

Pullout

#2

Raw

Experimental

Data

ANE

(axial,

northeast)

Gages

See

Appendix

A

for

gage

location

and

orientation.

t 1 O 0 04 sO Os CO nO <3- ON ON m CO r-H CO r-H ON ON r—H ON m m NO ON 04 04 ON co w 1 r-H t-H r—H 04 04 co <3

- <r <r m m 0- 00 ON O r-H r-H 04 co m m m m <3- 0 04 nD sO ON §

1 1 1 1 1 1 1 1 1 1 ! 1 1 l r-H

I 1

t-H

l

r—H

1

i—

1

r—H

1

r—H

1

<—

1

r—H

1 04

1 CO1 sO

1 nO

1 P^1 P^1

CO

r-H 0 00 04 0 0 04 nO sO CO p^ irH co co r-H r-H CO O <3* 00 r-H m ON Os in NO 04 04 r-H co

1 *—

04 04 CO CO <r in SO nO p^ 00 00 m NO m NO Os r—H CO 0- O CNJ 04 sO r^. 0 nO P^ 0

ANE

1 1 1 1 1 1 1 1 1 1 1 1 1 r-H r—H r-H 04 04 04 CO CO CO co co co 04 t-H I in

l

04

•—

0 in ^H <r ON in 0 NO NO 04 04 r-H r-H <r 00 NO r-H O *-H ON 04 04 nO 00 UO 04 UO p^

1 fH iH 04 co CO in m sO 00 co 1 <3- m NO co 0 04 CO <fr in in 1 00 O P^ p^ w 1 1 1 1 1 1 l 1 1 1 1 1 1 r—

r-H t—r-H t—

ln

l 0^1 ON

1 ON

1

r H

Id 0 04 O CO nO 04 0 00 NO m H Os p^ 00 00 04 00 NO m r-H 00 r-H NO 04 r-H co uo 04 0 CO NO1 t—

t—

04 co CO m sO p^ 1^ 00 NO <3* m <s* <r CO 04 1 04 m r—H m 04 0 00 ON NO Os

l 1 1 1 1 1 1 1 1 1 1 1 ! 1 1 1 1 1 r-H 1—

04 04 co CO CO <3- <3*

OiH 0 CO 0 p^ m 04 0 sO co 0 00 p^ NO ON NO «—

sO NO 0 00 t-H in NO O <3- NO UO p- NO Os

Id 1 *-H i-H 04 CO <3- in NO p^ p^ co 00 <3- <3* <3- m nO NO CO Os 04 CO Nf 04 r-H NO 00 CO NO

1 l 1 1 1 1 1 1 i 1 1 1 1 1 1 1 1 1 1 1 1 r—H

1

r—H

1

»-H

1

r-H

1

r-H

1

r-H

I

r-H

1 041

r-H

1

ON 0 C4 sO r-H <r 04 <r ON in CO 0 t—

0 SO 0 04 CO 04 co n* t—

0 O NO NO co t—

r-H 04

Id 1 iH

1 iHl 041 041 041 co1 CO1 1 1 1

1 r-H

1 041

1 »-H r-H 04 CO <dT NO sO 00 00 p^ CO

1

00 0 0 00 04 NO 00 00 nO <r <r <r 04 0 in CO t-H NO 00 NO 04 NO nO <3- sO 04 <3- Id 1 <—

r—

04 04 CO <3- <r co 04 p** co 0 co NO m CO r—H 04 CO CO -a- n3" co r-H

1 1 1 1 1 1 1 1 1 1 r-H 1 1 1 1 1 1 1 1

p^ 0 CO nO 0 00 04 0 00 00 04 04 0 <r <3- 04 04 04 NO sQ ON co m in Os co 00 UO P^ Os

Id 1 r—H 04 04 CO m «n <3- <3* <3* m m m m m m 00 Os r-H •-H 04 p- NO CO

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 r—H

1

r—H

1

r-H

1

r-H

1

r-H

1

r—H

1

r—H

1

'O 0 00 0 Os co CO t-H 0 04 00 00 00 04 Os r-H ON sO <3- 0 nO sO NO ON CO 04 00 00 uO Os a 1 04 <r <3- m CO 0 m 00 ON «—

r-H 04 sO 04 <3" p*. 0 04 m <-H CO 0 Os 0

- <3- so CO t-H

1 1 1 1 1 •H r-H rH 04 04 04 04 CO CO CO <3- <3* <3- m m NO NO ON ON 0 O

l l l 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 ! 1 r-H

1

r—H

I

O 04 00 00 04 04 ON ON in ON m m 0 00 NO Os co 0- m no 04 CO 04 04

Id iH 1 1 co <r NO 0 ON r-H 04 co m ON CO CO ON 04 CO CO 43 NO 00 r>- 0 uo 00 O- ON UO nO

1 1 1 1 iH

1

1 H

1

r-H

1

r-H

1

r-H

1

r-H

1 04

1 04

1 04

1 CO

1 co

1 CO

1 co1 CO1 co1 co1

<3*

1 <3

-

1 1 in

1 uo1 nO1 NO

1

CO 0 <r ON nO co CO p- ON m m ON t—

p^ m nO CO m p^ <3" 0 04 04 x> ON ON UO 0 nO sO P>» sO uO 04 1 04 0 —l ON <3

* ON m 00 p^ r-H NO 00 Os cO 00 CO in in 00 04 t-H 04 in NO

t—

l

1 04

1

1 co

1 1

r—H

1 04 CO

1

1 co

1 co

1 co1

<3-

1 1

r-H

1

1 r—H

1 co

1 04

1

t—

l

1 04

!

1 1 r—H

1 uo

1

<1*

1 co

1 P^

1

04 0 O UO CO CO ON 1 r-H m 04 04 NO ON m ON 04 m r—H O- 04 m 00 m 0- nO NO <3- 04 CO <p uo W 04 CO NO OD 0 CO p*. 00 0 04 0 nO 04 NO r—H co in m O'

O' NO UO ON NO p-. co

1 1 1 1 r—H H iH »-H 04 04 04 04 co CO <3* <f m m nO O'. 00 0 0 r—H r—H r—H co 00 NO

1 1 1 l 1 1 1 1 1 1 1 1 1 1 1 1 1 •—H

1

r—H

1 1

r-H

1

•-H

1

1 1 1

4—

t—

0 00 00 0 CO CO t—] CO NO 00 in 00 O CO NO 00 r-H sO Os 04 m 00 <r 1—

NO NO p- O r-H 00 r

r-. 04 ON 00 p^ 0 m NO 00 CO ON ON O m 0 m r-H nO m co r-H 00 04 04 00 Ht

Id

t-H nO m m 1 f CO m O ON NO co co 04 04 rH r—

0 O ON <fr co 0 m O <3- CO

<? r—

04 CO sO O'* 00 O 0 r-H 04 CO <3* m NO p>. 00 ON 0 0 r-H 04 <3- in in n in uo CO

*-H 1—

t-H r-H r—H 04 04 CNJ 04 OJ 04 04 04 04 04 04 04 04

1 c

03 O Ol 0 Os CO CO <t Os <r O 00 r-H 00 ON O 04 CO <3- Os O in 0 00 04 00 O ON 04 NO P>. CJ 04 CO co m nO p^ CO 00 ON O 0 r-H r-H 04 <3- m NO 0 co ON r-H sO ON CO <P <P

co

l rH iH r-H r—

r-H - •-H 04 04 04 04 CO CO CO <p

* 0 00 00 0 co CO <3* CO NO CO m 00 O CO NO co r-H <3- sO ON 04 m 00 <3

* r-H NO NO P^ O

1

r—H X)

03O

00 04 ON 00 0 m NO 00 CO ON <f ON O m O m r-H nO m CO r—H 00 04 04 00 P^-

t-H v£> m m CO m 0 ON sO CO CO 04 04 r-H r-H r-H 0 O ON <3

* 00 0 m O <3- P^ co n t-H 04 CO nO p^ 00 0 0 *-H CnJ CO <3

- in NO p» 00 ON 0 0 r—H CNJ <1

- <3- n n n uO n *<3

" CO

t-H r-H r-H r-H r-H r-H r—r-H r—H r—H 04 04 04 04 04 04 04 04 04 04 04 04 04

B-16

kips

micros

train
Appendix B Table B.17. Pullout #2 Raw Experimental Data TC (triaxial compression) Gages.

See Appendix A for gage location and orientation.

Load Scan# TC 1 TC 2 TC 3 TC 4

11.88 4 2 -5 -3 -3

26.18 22 2 -5 -6 -14

35.70 30 -3 -14 -13 -25

45.23 39 -3 -17 -14 -33

61.93 53 -5 -21 -19 -43

73.81 63 -22 -43 -35 -68

85.74 74 -16 -48 -40 -92

100.03 84 -32 -70 -59 -132

109.56 89 -32 -79 -67 -152

116.68 94 -48 -102 -79 -168

123.85 100 -35 -98 -76 -182

133.38 108 -57 -125 -94 -203

142.90 111 -62 -140 -100 -219

152.43 118 -70 -160 -109 -236

161.96 129 -70 -178 -117 -255

171.48 140 -71 -203 -122 -278

181.01 152 -67 -209 -124 -274

190.54 163 -73 -228 -133 -262

200.06 174 -82 -246 -140 -246

209.59 179 -89 -263 -148 -227

214.12 200 -89 -284 -154 -197

228.65 235 -75 -293 -152 -146

240.58 270 -79 -323 -163 -54

245.34 298 -82 -335 -170 2 250.11 312 -73 -336 -175 98

254.87 368 -73 -351 -184 232

257.26 390 -76 -360 -189 289

257.26 439 -84 -378 -208 588

254.87 442 -78 -371 -205 648

247.70 446 -76 -371 -205 828

233.41 447 -79 -370 -200 1072

B-17
Appendix

B

Table

B.18.

Pullout

#

2

Raw

Experimental

Data

ET

(edge

tension)

Gages.

See

Appendix

A

for

gage

location

and

orientation.

nO

<M nO nO 00 nO CM o NO <r CM 00 O O 00 O 00 O' o CM NO CM CO co ON O' UO CM H 5—1 5—1 5—1 5—1 CM rH ?—1 i 7*—l 1 I CM CM 7—1 CO in U0 NO 00 00 O' ON 5—1 CM

CU i 1 I I i I i 1 ! I I 1 f—

1

5—1

1

LO

OsJ 00 o no CM <r NO CM CM NO NO i—

CO LO 00 ON 5—1 CM NO CM 5—1 CM 00 00 H 5—1 CM CM CM 5—1 5—1 7—1 r—

CO CM U0 m 00 00 7—1 CM in CO 5—1 CO CM CM <fr in CM CO O' CM

l 1 i 1 i i i i i i 5—1

1 5—

1

E 1

1

7—1

1

5—1

1

5—1

1

5—1

1

5—1

1

5—1

1

5—1

! CM1 CM1 CM1 <f

1

<f

O') o o nO o 00 NO 00 00 CM o ao vD 00 NO O' 00 <r ON NO m !-» ON m CM 5—1 in I-" 00 H v—

cm CM 1—CM r—

1 i—

r—

CM 5—1 CO CM m in r"- NO 00 00 ON 5—1 CM co co NO O' W i I 1 1 I 1 i i i i I i 1 1 5—1

1

7—1

1

5—1

1

5—1

I

5“1

i

5—1

1

CO

NO CM CM CM no oo CM 00 NO 00 CM o NO NO o NO o oo CM NO NO NO co ON CO 5—1 CO O' m

H r—

CO CO co CO CM CO CM co CO CM CM 5—1 5—

5—1 I 5—1 CM 5—1 CO CO in NO ON CO ON co

CU ! ' 1 ! 1 1 1 i 1 1 I 1 1 5—1

1

CM co U0 ON on m CO UO r—

O' ON r— CO NO NO m o CM m m <1* 5—1 m CM NO o O' co H cm CM CM co co CM CM CM CM T—

5—1 5—1 5—1 1 ! 1 5—1 CM CM CM CO co m in NO w I i 1 I 1 1 1 1 r

5—

uo O' UO O r"- T—

00 NO CM CO CM o uo vf o o 00 00 CM H CM O' CO uo CO ON ON ON oo 00 00 NO NO NO NO NO NO uo m m m m CO CM CM CM CO CO <r in W 5—

5—1 5—1 5—1 5—1 5—

r—

5—

r—

5—1 5—1 1—1 5—1 5“1 5“1 5—1 7—1 5—1 5—1 r—

5—1 5—1 5 1 7—1 5 1 5—1 5—1 5—1 i—

=£=C

CO <r CM o O' co co <r ON o 00 r—

00 ON O CM CO O' O U0 O 00 CM 00 o O' CM NO o CM CO CO in NO 00 00 ON o o i—

5—1 CM in NO r^. o CO O' T—

NO O' CO

CO 5—1 5—1 5 1 1—

T—

5—1 5—1 5—1 5—1 5—1 CM CM CM CM CO CO co <1*

03 00 o co CO r—

CO NO 00 UO 00 o CO NO OO 5—1 <fr NO ON CM m CO <5 5—1 NO NO o 5—1 -a 00 s—

CM ON 00 o UO NO 00 CO O' O' <r O UO O U0 5—1 NO uo CO 5—1 oo CM CM 00 <!

cOo 5—1 no m uo 5—

CO m o ON NO CO CO CM CM 5—1 5—1 5—1 o o ON 00 o m o <3

- CO

-1 5—1 CM co <r NO r-. co o o I—-1 CM CO m NO 00 ON o o 5—1 CM m uo m m m CO

T—

7—1 5—1 5—1 5—1 T—

5—1 5—1 T—5—1 5—1 CM CM CM CM CM CM CM CM CM CM CM CM CM

B-18
Appendix B Table B.19. Pullout #2 Raw Experimental Data LVDT Displacements (averages)

Scan# AVG

KIPS

AVG

VERT.

LVDT

AVG

HOR.

LVDT

4 11.88 0 0

22 26.18 -1.0 0.3

30 35.70 -1.5 0

39 45.23 -2.0 0.3

53 61.93 -2.5 0

63 73.81 -3.0 0

74 85.74 -3.5 0

84 100.03 -5.1 0

89 109.56 -5.6 0.3

94 116.68 -6.1 0

100 123.85 -6.6 0

108 133.38 -8.1 0.3

111 142.90 -8.6 0

118 152.43 -10.1 0

129 161.96 -11.6 -0.7

140 171.48 -15.2 -2.1

152 181.01 -16.7 -3.9

163 190.54 -18.7 -5.6

174 200.06 -20.2 -7.8

179 209.59 -22.2 -9.5

200 214.12 -26.3 -12.7

235 228.65 -30.8 -16.5

270 240.58 -38.4 -23.2

298 245.34 -42.4 -26.8

312 250.11 -48.5 -32.7

368 254.87 -58.1 -40.8

390 257.26 -62 .

-44.3

439 257.26 -83.3 -62.6

442 254.87 -87.9 -66.5

446 247.70 -101.0 -78.1

447 233.41 -124.7 -98.9

B-19
NBS-114A (REV. 2-80

U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

1. PUBLICATION OR

REPORT NO.

2. Performing Organ. Report No. 3. Publication Date

SHEET (See instructions) NBSIR 82-2484 May 1982

4. TITLE AND SUBTITLE

INTERNAL STRAIN, DEFORMATION, AND FAILURE OF LARGE SCALE FULLOUT

TESTS IN CONCRETE

5. AUTHOR(S)

William C. Stone

6. PERFORMING ORGANIZATION (If joint or other than NBS, see in struction s) 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE 8. Type of Report & Period Covered

WASHINGTON, D.C. 20234

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City. State. ZIP)

10. SUPPLEMENTARY NOTES

Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bi bl iography or literature survey, mention it here) A study was performed to obtain detailed experimental data on crack propagation and

internal strain distribution for the pullout test. A 12:1 scaled-up pullout test was designed, using a commercial pullout insert for the prototype dimensions, and was instrumented with small waterproof embedment strain gages so as to obtain

internal strain profiles at critical locations. Two large scale specimens were

tested with apex angles falling at the upper and lower bounds currently recommended

in ASTM C-900. Two dimensional axisymmetric finite element analyses were performed

for the two experimental specimens and the results were compared with measured

strains for load stages below the onset cf internal cracking. The; results showed

good correlation between the analytical and experimentally observed strains. The

experimental data indicate that internal cracking, and the formation of the

failure surface, are principally governed by the tensile strength cf tie concrete.

The failure surface appears to have formed by 65% of ultimate load. Beyond this

point, it is likely that the entire load is carried by the mechanism of aggregate

interlock. Ultimate failure occurs when all aggregates mechanically bridging the

failure surface pullout from the retaining cement paste. It is likely that the

pullout test measures the shear strength of the cement paste cr mortar which

binds the concrete together.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolons)

Concrete; crack propagation; failure surface geometry; failure theory; finite

element method; internal strain; laboratory testing; large scale models; mathematical

model; pullout test; stress contours

13. AVAILABILITY 14. NO. OF

PRINTED PAGES

[X] Unlimited

| 1 For Official Distribution. Do Not Release to NTIS 170

1 1 Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402. 15. Price

fX] Order From National Technical Information Service (NTIS), Springfield, VA. 22161 $15.00

USCOMM-DC 6043-P 80
I