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REFERENCE

IITED STATES IRTMENT OF MMERCE

JLICATION NBS TECHNICAL NOTE 722

*^rts o' ^

175103

QCIOD

,115753

-Oe|

OCT l 4 7975

Tensile Creep

of Boron/Epoxy and

Boron/Epoxy-Reinforced

7075-T6 Aluminum Alloy

U.S.

ARTMENT

OF

DMMERCE

National

100

S7$l

'•• 7*Z

'TO
NATIONAL BUREAU OF STANDARDS

The National Bureau of Standards1 was established by an act of Congress March 3,

1 901 . The Bureau's overall goal is to strengthen and advance the Nation's science and

technology and facilitate their effective application for public benefit. To this end, the Bureau conducts research and provides: (1) a basis for the Nation's physical measure- ment system, (2) scientific and technological services for industry and government, (3) a technical basis for equity in trade, and (4) technical services to promote public safety.

The Bureau consists of the Institute for Basic Standards, the Institute for Materials

Research, the Institute for Applied Technology, the Center for Computer Sciences and

Technology, and the Office for Information Programs.

THE INSTITUTE FOR BASIC STANDARDS provides the central basis within the United States of a complete and consistent system of physical measurement; coordinates

that system with measurement systems of other nations; and furnishes essential services leading to accurate and uniform physical measurements throughout the Nation's scientific community, industry, and commerce. The Institute consists of a Center for Radiation Research, an Office of Measurement Services and the following divisions: Applied Mathematics—Electricity—Heat—Mechanics—Optical Physics—Linac

Radiation2—Nuclear Radiation2—Applied Radiation2—Quantum Electronics3—

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consists of the following offices and divisions: Information Processing Standards—Computer Information—Computer Services —Systems Development—Information Processing Technology.

THE OFFICE FOR INFORMATION PROGRAMS promotes optimum dissemination

and accessibility of scientific information generated within NBS and other agencies of

the Federal Government; promotes the development of the National Standard Reference

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and directs the public information activities of the Bureau. The Office consists of the following organizational units: Office of Standard Reference Data—Office of Technical Information and

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1 Headquarters and Laboratories at Gaithersburg, Maryland, unless otherwise noted; mailing address Washington, D.C. 20234.

* Part of the Center for Radiation Research.

* Located at Boulder, Colorado 80302.
\oo

7lt~

Tensile Creep of Boron/Epoxy and

Boron/Epoxy-Reinforced

7075-T6 Aluminum Alloy

Daniel J. Chwirut

Mechanics Division

Institute for Basic Standards

National Bureau of Standards

Washington, D.C. 20234

*'Te^oJuu*2jO-0 hfrCz^. no, 7X7-

o <^4

o J

U.S. DEPARTMENT OF COMMERCE, Peter G. Peterson, Secretary

NATIONAL BUREAU OF STANDARDS, Lawrence M. Kushner, Acting Director

Issued May 1972
National Bureau of Standards Technical Note 722

Nat. Bur. Stand. (U.S.), Tech. Note 722, 60 pages (May 1972)

CODEN: NBTNAE

Issued May 1972)

For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402 (Order by SD Catalog No. C13.46:722). Price 60 cents.
CONTENTS

Page

1. INTRODUCTION 1

2. SYMBOLS 2

3. TEST SPECIMENS 3 3.1. Boron/Epoxy Specimens 3

3.2. Aluminum Alloy Specimens 4

3.3. Sandwich Specimens 4

4. TEST EQUIPMENT 5 4.1. Creep Testing Machines 5

4.2. Strain Measurement Equipment 6

4.3. Heating Systems 6

5. TEST PROCEDURE 7

6. TESTS AND EXPERIMENTAL RESULTS 7 6.1. Static Tests 7 6.2. Residual Strain in Sandwich Specimens 8

6.3. Creep Tests 8

7. ANALYTICAL PREDICTION OF CREEP CURVES 9

7.1. Theory 9

7.2. Analytical Procedure 10

7.3. Illustration 12

7.4. Analytical Results 13

8. DISCUSSION 14

8.1. Creep of Boron/Epoxy Material 14

8.2. Creep of Aluminum Alloy 14

8.3. Creep of Sandwich Specimens 14

9. CONCLUSIONS 16

10. REFERENCES 18

APPENDIX A 56

APPENDIX B 57

III

TENSILE CREEP OF BORON/EPOXY AND BORON/EPOXY-REINFORCED

7075-T6 ALUMINUM ALLOY

Daniel J. Chwlrut

Tensile creep tests were performed on twelve

specimens of 0° unidirectional boron/epoxy and

on twenty-four specimens of 7075-T6 aluminum alloy

reinforced with 0° unidirectional boron/epoxy.

An analytical procedure for predicting the creep

properties of composite-reinforced metals is presented. Agreement between calculated and experimental creep curves varies with test temperature.

These discrepancies between experiment and analysis are probably not due to an error in the

analysis itself, but rather to uncertainty in

the residual stresses in the specimens.

Key words: Aluminum alloy; boron/epoxy; composite

materials; creep; elastic follow-up technique;

sandwich specimen.

INTRODUCTION

Within the past few years, advanced composites have made dramatic

progress, going from a laboratory-availability status to a production

commitment in some components of the next generation of advanced aircraft systems [1]. There seem to be two different approaches toward

: the application of advanced composites in the United States at present.

One approach advocates the use of composites for complete components

whenever practical, whereas the other approach advocates the use of

composites only to reinforce metals. The investigation reported herein

was undertaken to help answer one of the questions about composites when

applied using the latter approach: What are the long-term strength

properties of metals when reinforced with an organic-matrix composite,

specifically a structural aluminum alloy reinforced with boron/epoxy?

The resistance of boron filaments to creep at temperatures below

1200 °F (650 °C) has been reported to be excellent [2] . At these temperatures no measurable creep was observed at stresses up to the fracture

stress.

figures in brackets indicate the literature references on page 18,
Experimental data on creep of 0° unidirectional organic-matrix

composites are scarce. The data that have been reported are somewhat

conflicting. Shockley et al. [3] report that a creep strain of about

1400 x 10~ 6 was measured after 4 hours at 106,000 lbf/in2 (7.3 x 10 8 N/m 2

) and 270 °F (132 °C) in 0° unidirectional boron/epoxy. Lou and

Schapery [4] claim that negligible creep is found in unidirectional

glass fiber-reinforced epoxy when loaded parallel to the fibers. Most

work on creep of organic-matrix composites reported in the literature

involves either crossplied specimens or off-axis loading of unidirectional specimens, indicating that this is where most of the concern lies. In

a 0° unidirectional boron/epoxy specimen with 50 percent by volume of

fibers, the fibers support approximately 98 percent of the total load on the specimen. Since the fibers do not creep, little or no creep would

be expected in the composite.

Metal-matrix composites, however, specifically boron/aluminum, have

been found to undergo creep and rupture at temperatures below 930 °F

(500 °C) when loaded along the axis of the fibers in tension [5-6] and

in three-point bending [7]. Ruptures occurred in some tests at stresses

less than 50 percent of the tensile strength of the composite. Reference 7 indicates that a considerable increase in rupture life can be

achieved by increasing the fiber volume fraction from 40 to 60 percent.

The creep resistance of a boron-filament composite thus seems to be

directly related to the percentage of the total load on the composite

carried by the filaments as well as the creep properties of the matrix.

It is logical to assume, therefore, that creep would also be possible in

light-alloy structures reinforced with organic-matrix composites, since

the parent metal still carries a significant part of the total load.

The investigation reported here was undertaken to experimentally

evaluate the creep properties of 7075-T6 aluminum alloy reinforced with

0° unidirectional boron/epoxy, and to determine if these data can be

predicted analytically from creep data for each of the components. This

work was performed at the National Bureau of Standards under the sponsorship and with the financial assistance of the National Aeronautics and

Space Administration, Langley Research Center.

2. SYMBOLS

The units used for physical quantities in this paper are given in

both the U. S. Customary Units and in the International System of Units

(SI). Conversion factors pertinent to the present investigation are presented in Appendix A.

Appendix B contains a glossary of terms whose meaning may not be

widely known or whose usage in this paper is different or more restrictive than the generally accepted definition.
A crossectional area

E modulus of elasticity (when referring to the modulus of

boron/epoxy in this report E denotes the modulus along

the axis of the fibers, commonly called E-q)

P load

AP an incremental change in load

S stress

t txme

At an incremental change in time

a coefficient of linear thermal expansion

e strain

Ae an incremental change in strain

subscripts

a aluminum alloy

b boron/epoxy

Iv residual

t total

superscripts

initial, i.e., at the start of the test

3. TEST SPECIMENS

3.1. Boron/Epoxy Specimens

The composite specimens tested during this investigation were unidirectional boron/epoxy, with the filaments alined along the axis of

loading. The specimen geometry is shown in figure 1(A). The boron/epoxy

material was supplied by the 3M Company^ in continuous tape form with

^The identification of the suppliers of materials used in this program is

for information purposes only and is not intended to indicate endorsement by the National Bureau of Standards, nor does it imply that the

material was necessarily the best available.
nominally 212 filaments per inch (83 filaments per cm) of tape width.

The boron filaments were nominally 0.004 in (0.1 mm) in diameter and preimpregnated with PR-279 epoxy resin. Each ply of boron/epoxy had a glass

scrim backing, 0.001 inch (0.03 mm) thick.

The specimens were cured with the platen-press technique using pressures and temperatures recommended by the manufacturer. Five samples of

material were taken from each plate of composite laminate and mounted in

a cast epoxy metallurgical mount. A photomicrograph at 100 x magnification was taken of each sample, and the volume fraction of the sample was determined from the photomicrograph by determining what fraction of the

total intersections of a grid overlay fell on filaments. The filament

volume fraction for each plate was taken as the average of the fractions

of the five samples. All of the specimens tested in this investigation

were taken from laminated plates having filament volume fractions between

47 and 54 percent.

3.2. Aluminum Alloy Specimens

Aluminum alloy specimens were cut from a single sheet of 0.032-in

(0.8-mm) thick 7075-T6 aluminum alloy. All specimens were cut so that

the longitudinal axis of the specimen was in the direction of rolling of

the sheet. The specimen geometry was the standard flat tensile specimen

(ASTM Designation E8-69, Fig. 6), with a reduced section approximately

7 in (18 cm) long (see fig. 1(B)). Unless otherwise indicated, all

specimens were preconditioned by a thermal treatment of one hour at

325 °F (163 °C) and five hours at 350 °F (177 °C) prior to testing. This

simulated the thermal treatment experienced by the aluminum alloy during

the co-cure processing of the sandwich specimens which is discussed below.

3.3. Sandwich Specimens

Sandwich specimens were tested which consisted of 2, 4, 6, 8, or 12

plies of boron/epoxy bonded between two sheets of 0.032-in (0.8-mm) thick

7075-T6 aluminum alloy. The specimen geometry is shown in figure 1(C).

Two methods were used to fabricate these specimens; a two-step process in

which the boron/epoxy laminate was cured first and then bonded to the

aluminum alloy, and a co-curing process in which the boron/epoxy laminate was cured and bonded to the aluminum alloy in one process.
a. Two-Step Process

The initial sandwich specimens were fabricated by bonding precured

boron/epoxy laminates between two sheets of aluminum alloy using

Armstrong's C-4 epoxy and activator W in a 1:1 ratio and cured at 200 °F

(93 °C) . The advantages of using this technique are good dimensional

control of ti e composite, low residual stresses due to fabrication, and

control of composite volume fractions. The disadvantage is that the

epoxy adhesive is ineffective at temperatures above 200 °F (93 °C), as

the specimens tended to delaminate when tested at these temperatures.

Attempts using a higher temperature rated epoxy adhesive were unsuccessful.

b. Co—curing Process

The majority of the sandwich specimens tested in this investigation

were fabricated by curing the composite and bonding it to the aluminum

alloy in the same operation. Good adhesion between composite and

aluminum alloy was achieved for temperatures up to 300 °F (149 °C) . The

disadvantages of this technique are the lack of control of the volume

fractions of the composite, because no excess epoxy is bled from the

prepreg during the cure, and the introduction of significant residual

stresses in the components due to the different coefficients of thermal

expansion of the composite and aluminum alloy. Volume fractions for

these specimens were not measured, but based on specimen thicknesses the

composite was judged to be approximately 40 percent filaments and 60

percent epoxy.

TEST EQUIPMENT

The experimental apparatus used in this investigation is shown in

figure 2.

4.1. Creep Testing Machines

The test program was conducted in two deadweight creep testing

machines. Both machines have lever arms with ratios of 10:1. Specimen

alinement was achieved by using knife edges in the load train of one machine and spherical seats in the other machine.

Prior to initiation of testing, both machines were calibrated in

accordance with ASTM Designation E4-64, Standard Methods of Verification of Testing Machines. Errors in applied loads were less than one

percent. Bending strains induced in the specimen were measured using an aluminum alloy specimen instrumented with four strain gages located to measure bending in two planes. The maximum deviation of any single

strain reading from the average of the four readings was less than two

percent of the average.
4.2. Strain Measurement Equipment

Two systems were used to measure and record creep strains—a high

sensitivity system used in tests on boron/epoxy and sandwich specimens,

and a low sensitivity system used in tests on aluminum alloy specimens.

Both systems employed extension rods to allow the sensing element to

remain outside the test furnaces.

a. High Sensitivity System

Figure 3 shows a sandwich specimen with extension rods and extensometer sensing elements attached. The sensing elements are two dc

linearly variable differential transformers (LVDT's) which measured the

extension over a 4-in (10-cm) gage length in the center of the specimen.

The outputs from the LVDT's were amplified and recorded continuously on

a dual-channel, voltage-time recorder. The strain in the specimen was taken as the average of the two strain measurements. This system was calibrated using a micrometer extensometer comparator. Strains of 4 x 10~° could be resolved, and uncertainties in strain measurement were

less than five percent of the measured value or 8 x 10~ 6

, whichever is

greater.

b. Low Sensitivity System

Figure 4 shows an aluminum alloy specimen with extension rods and a commercial LVDT extensometer attached. This extensometer measured the

the extension of a 3-in (8-cm) gage length in the center of the specimen.

The output from the extensometer was amplified and recorded continuously

on a time-based recorder. This system did not average the strain on the

two sides of the specimen, but it was considered adequate because bending strains measured during testing machine calibration were small. This

system was calibrated using an extensometer comparator and gage blocks.

Strains of 13 x 10-5 could be resolved, and uncertainties in strain

measurement were less than five percent of the measured value or 26 x 10~ 6

, whichever is greater.

4.3. Heating Systems

The test furnaces were each composed of a cylindrical section of

0. 5-in (1-cm) thick asbestos-cement pipe approximately 10 in (25 cm) in

inside diameter and 7 in (18 cm) in length. These were designed so that

the gage section in the center of the specimen could be evenly heated

while the grips could be outside of the furnace. This was done in an effort to eliminate failures in the grip. The top and bottom of the

cylinder were covered with three layers of 0.75-in (2-cm) thick fiberglass insulation. Heat was supplied by four 1000-watt quartz lamps,

wired in pairs, with separate power control for each pair. Temperature

was measured with three or four thermocouples taped to the specimen in
the gage section. Temperature was controlled and recorded by a timeproportioning, recording controller.

TEST PROCEDURE

The same procedures were followed for tests on all three types of

specimens. T le specimen was cleaned with acetone to remove any dirt and

oil. The ext msometer extension rods were attached at the appropriate

gage points, and thermocouples were taped to the specimen. The specimen

was mounted in the testing machine and final adjustments were made to

the strain measurement system.

The specimen was heated to the test temperature in approximately

20 min. Adjustments were made to the temperature controller to reduce

the temperature variations to an acceptable level. A gradient of 3 °F

(2 °C) or less along the gage length of the specimen and a variation of

±2 °F (1 °C) at the center of the specimen were considered acceptable.

The specimen was then maintained at this condition for 30 minutes prior

to application of load. Loading was accomplished smoothly and rapidly

by lowering the weights onto the lever arms using a hand crank on one machine and a screw jack on the other.

Extension and temperature were continuously recorded for the duration of each test. When extensions approached the limit of the recorders,

the extensometers were reset. Tests were usually terminated after 100

hours if no rupture had occurred.

6. TESTS AND EXPERIMENTAL RESULTS

6.1. Static Tests

Tensile tests at four temperatures were performed on boron/epoxy

specimens and aluminum alloy specimens to determine values for the

elastic moduli to be used in the computations. The procedure outlined

in ASTM Designation E231-69, Static Determination of Young's Modulus of

Metals at Low and Elevated Temperatures, was followed, except that re- sistance strain gages were used instead of extensometers. The results

of these tests are given in table 1. The measured moduli for both

materials were approximately constant at temperatures up to 300 °F

(149 °C) , and for the computations the following values were used for

all temperatures:

boron/epoxy: E = 32 x 10 6 lbf/in2 (22 x 10 1 ° N/m2

)

aluminum alloy: E = 9.6 x 10 6 lbf/in2 (6.6 x 10 1 ° N/m2

)
6.2. Residual Strain in Sandwich Specimens

The longitudinal residual strains in the aluminum alloy of sandwich

specimens, resulting from fabrication, were measured at room temperature

on one specimen fabricated by each process.-' This was done by bonding

a resistance strain gage to the aluminum alloy and measuring the strain

relieved when the specimen was carefully delaminated. For the co-cured

specimen the residual strain was 2270 x 10 in tension. For the specimen fabricated by the two-step process the residual strain was 30 x 10-6

in tension.

6.3. Creep Tests

a. Creep of Boron/Epoxy Specimens

A total of twelve creep tests on boron/epoxy specimens was attempted at four temperatures. Four of the specimens failed immediately

upon application of the test load, which was less than the nominal maximum strength of the material as measured in the static tests. Five

other specimens failed after sustaining the test load for a period of

time. Four of these failed in the grip and one in the gage section.

Limited creep data were obtained for these specimens, although the

failures were not considered valid stress-rupture failures for the

material because they occurred at or near the grip. Three specimens

survived 120 hours without failure. The test conditions and results

for the twelve tests are given in table 2.

b. Creep of Aluminum Alloy Specimens

Sixteen tensile creep tests on 7075-T6 aluminum alloy specimens were performed at four temperatures. The test conditions and results are given in table 3. Except as noted these specimens had been preconditioned

as mentioned in section 3.2. The families of creep curves for the four

temperatures are given in figures 5 through 8.

c. Creep of Sandwich Specimens

Tensile creep tests on nineteen sandwich specimens with four plies

of boron/epoxy between two sheets of 7075-T6 aluminum alloy were performed at four temperatures. The test conditions and results are given

in table 4. The families of creep curves for the four temperatures are

3Measurements made on additional specimens are discussed later
given in figures 9 through 12. Note that some tests are replicates of

others in which the specimen failed prematurely in the grip. The

nominal stresses in the aluminum alloy and the boron/epoxy were calculated by assuming that the total test load on each specimen was distributed between the two components so that the axial elastic strain in

both were equal.

Tensile creep tests on five sandwich specimens with 2, 6, 8, 8, and

12 plies of boron/epoxy were performed at 300 °F (149 °C) and one initial

stress level to determine the effect of composite-to-metal volume ratio.

Test conditions and results for these tests are given in table 5. The

creep curves for these specimens are given in figure 13. Figure 14

summarizes the effect of composite-to-metal volume ratio on the creep

behavior and rupture life of the sandwich specimens. Note that specimen

BA-37 seems to be an anomaly, and the results of this test were not con- sidered in constructing the curves in figure 14.

7. ANALYTICAL PREDICTION OF CREEP CURVES

From earlier work done at this laboratory Mordfin [8] presented an analytical procedure for predicting creep curves for creep deflections

of non-uniform beams. Experimental verification of this analysis [9]

showed good correlation between experimental and analytical values. This

analysis used the "elastic follow-up technique" developed by Popov for

stress redistribution in non-uniform structures [10] . A similar analysis

was used by Hayashi to predict creep of unidirectional fibrous composites

[11] . An analysis similar to these is applied to creep of sandwich

specimens in this investigation, using Hayashi 's assumption that the

longitudinal strains in all components are equal.

7.1. Theory

The maximum creep strain in boron/epoxy measured in the investigation was less than 10 percent of the elastic strain in the specimen.

This seems to verify that no significant creep occurs in 0° unidirectional

boron/epoxy at temperatures up to 300 °F (149 °C). However, it was hypothesized that effective creep could occur in composite-reinforced

metals by the following procedure:

a. At time t = 0, the aluminum alloy is under a stress S and the boron/epoxy is under a stress S? . b

a

b. The aluminum alloy tends to creep but, since the strain

in the aluminum alloy cannot exceed the strain in the

boron/epoxy, the aluminum alloy undergoes stress re- laxation instead.
Since the total load on the specimen is constant, the

stress in the boron/epoxy increases as the aluminum

alloy relaxes, thus causing increased elastic strain

which is measured and called "creep" in the sandwich

specimen.

Rupture occurs when enough stress redistribution has

occurred so that the stress in the boron/epoxy exceeds

the tensile strength of the material.

Both the strain-hardening and time-hardening rules for accumulation

of creep streain the the aluminum alloy were used. The assumption of the

strain-hardening rule is that the instantaneous creep rate is a function

of the instantaneous stress and the amount of creep strain that has

taken place. The assumption of the time-hardening rule is that the

instantaneous creep rate is a function of the instantaneous stress and

the time the material has been creeping.

7.2. Analytical Procedure

At time t = 0, the aluminum alloy supports a load P° and is at a pO a stress S° , where S° = 1- S . The boron/epoxy supports a load P° . 9. 3. J\ l D a The total load on the specimen is P^ = P° + P° , and P remains cont a b t stant for the duration of the test. P° and P? are calculated assuming

a b

that the total load P is distributed between the aluminum alloy and

the boron/epoxy so that the axial elastic strain in both are the same.

During an initial time increment, At » the aluminum alloy at stress

S° tends to creep an amount e , determined from the experimental creep

3 3. curves for the aluminum alloy specimens. It is assumed that creep in the

boron/epoxy is negligible. Since the strain in the aluminum alloy is

constrained to remain equal to the strain in the boron/epoxy, the total

load P is redistributed by an amount AP so that at the end of the

time increment At

P = P° - AP

a a

and

P, = P° + AP,

b b

10
Assuming an elastic follow-up, this load redistribution causes a strain

change

Ae =

and

a E A

a a

AP

where the absolute sum of these changes equals e (see fig. 15).

That is,

-Ae aba + Ae, = e . (2)

Substituting eq (1) into eq (2)

AP AP

= e

or

E A E^A, a

a a b d

AP = — a

1

. (3) e a

+ e7aT

a a b b

The creep in the sandwich specimen during time increment At is Ae , which, from eq (1) and (3), is

e a

E A

a a

i +

The aluminum alloy is now at a stress

P

S = T^ + S a A r a

*>—Vf

11
Another time increment is taken, and the procedure is repeated under the

new conditions.

7.3. Illustration

As an illustration, consider Specimen BA-16, tested at 225 °F

(107 °C). The crossectional areas of the boron/epoxy and aluminum were

A^ = 0.0114 in2 (7.35 x 10" 6 m

2

)

A = 0.0315 in2 (2.03 x 10~ 5 m

2

). a Substituting these values and the values of elastic moduli into eq (3)

and (4) gives

e e AP = a_ lbf or Ap = §_ N

6.05 x 10-6 1.36 x 10" 6 and

e

'b 2.2

Ae ' K

= TT ' (4)

The total test load, P , was 2750 lbf (12,232 N)

a' V

The values of P°, P£ , and S were calculated as follows

so

P

a

P

b elastic strain = e = jf- = —and P = P° + P°

a a b d

p° = p° Q b a E A

a a Therefore,

or

P .-. p° (i +-k-b) = 2.2 P°

t a

K E A a a a

P

t P° = -r^r = 1260 lbf (5604 N)

a 2.2

12
and

P° = P -P = 1490 lbf (6628 N) b t a

The value of residual stress in the aluminum alloy, S , was estimated assuming a linear variation from zero at 350 °F (149 °C) where

the composite-to-aluminum bond was cured, to the measured value at 75 °F

(24 °C). At 225 °F (107 °C) , this value is 9700 lbf /in2 (6.69 x 107 N/m2 ). The initial stress in the aluminum alloy, S° , is 49700 lbf /in2

3. (3.43 x 10 8 N/m2

).

Taking, as a first time increment, At - 0.01 h, e = 400 x 10~ 6

a. (interpolated from fig. 7). Therefore, from eq (3) and (4), AP = 66 lbf

(294 N) and Ae, = 181 x 10-8

. The first point on the analytical creep

b curve is thus t = 0.01 h, e = 181 x 10~ 8

. The new conditions at

t = 0.01 h are P = 1194 lbf (5311 N) , S = 47,600 lbf /in2 (3.28 x 10 8

3. 3. N/m2 ), and P^ = 1556 lbf (6921 N)

.

b

During the next time increment, At = 0.04 h, the aluminum alloy is

considered to remain at 47, 600 lbf /in2 (3.28 x 10 8 N/m2 ). The strainhardening rule predicts that the strain rate in the aluminum alloy is a function of stress and previous creep strain, so e is taken from the

3. curve for 47,600 lbf /in2 (3.28 x 10 8 N/m2 ), starting at a creep strain of

400 x 10-8

, over a time increment of 0.04 h. Using this method, e =

9. 80 x 10-6

, AP = 13 lbf (58 N) , and Ae,_ = 36 x 10" 6

. The second point

b on the strain-hardening creep curve is then t = EAt = 0.05 h, e = £Ae, =

217 x 10~ 8. The time-hardening rule predicts that the strain rate in the

aluminum alloy is a function of stress and creep time, so e is taken

from the curve for 47,600 lbf /in2 (3.28 x 10 8 N/m2 ), starting at t = 0.01 h, over a time increment of 0.04 h. Using this method, e = 110 x

9. 10" 6

, AP = 18 lbf (80 N), and Ae,_ = 50 x 10" 6

. The second point on the

b

time-hardening creep curve is then t = EAt = 0.05 h, e = EAe K = 231 x

10"

" ,-6 'b

This incremental process is repeated until the desired creep curve

is completed.

7.4. Analytical Results

Analytical creep curves were calculated for fifteen sandwich specimens using both the strain-hardening and time-hardening rules. Both

analytical curves and the experimental curve for each specimen are shown

in figures 16 through 30. Analytical curves using the strain-hardening

13
theory could not be completed to 100 hours because creep data beyond

100 hours for the aluminum alloy were required for such computations.

8. DISCUSSION

8.1. Creep of Boron/Epoxy Material

The maximum creep strain measured in boron/epoxy in this investigation was less than 10 percent of the elastic strain in the specimen.

This seems to support the view that creep in 0° unidirectional boron/

epoxy is not a serious problem within the temperature range of the epoxy.

The small strains measured as creep at elevated temperatures may be a result of either or both of the following:

1. Straightening of initially bowed or misoriented filaments

which aline themselves as the epoxy softens

2. Fracture of individual defective filaments causing an increased stress in the remaining undamaged filaments.

These random phenomena could also explain the scatter in the creep behavior from specimen to specimen.

8.2 Creep of Aluminum Alloy

In general, the specimens tested in this investigation exhibited

more creep than others reported in the literature [12, 13]. This is to

be expected because of the overaging caused by the thermal preconditioning mentioned earlier.

8.3. Creep of Sandwich Specimens

Tests on sandwich specimens corroborated the supposition that

7075-T6 aluminum alloy reinforced with boron/epoxy can creep at loads

which are less than 75 percent of the static strength of the specimen.

The results of these tests and the analysis are discussed below with

respect to some of the variables in the program.

a. Effect of Fabrication Process

Generally, larger creep strains were observed in the specimens fabricated by the co-curing process than in those fabricated by the twostep process (see table 4) . This can probably be attributed to the

large (21,600 lbf/in 2 at 75 °F, 1.49 x 10 8 N/m2 at 24 °C) residual tensile

stresses induced in the aluminum alloy by the co-curing process. Since

the specimens fabricated by the two-step process have a limited temperature capability and the residual stresses in the co-cured specimens are

14
less at higher temperatures, the co-curing process may be the preferable

process for applications above about 150 °F (66 °C) , while the two-step

process may be preferable for applications below 150 °F (66 °C).

b. Effect of Composite-to-Metal Volume Ratio

The results of the tests on specimens with different composite-tometal volume ratios supported the hypothesis that the resistance to creep

is directly related to the percentage of the load carried by the fibers

(fig. 14).

c. Agreement Between Experiment and Analysis

The agreement between calculated strains and experimentally determined strains seems to vary with test temperature. For tests run at

75 °F (24 °C) and 150 °F (66 °C) the analysis predicts considerably more

creep strain than was observed experimentally. At 225 °F (107 °C) the

agreement was excellent, and at 300 °F (149 °C) the analysis predicts

about 30 percent less strain than was observed experimentally. These discrepancies are probably due to uncertainty in the value of residual stress

used in the calculations, since at 75 °F (24 °C) this constitutes up to

40 percent of the total stress in the aluminum alloy, while at 300 °F

(149 °C) it was negligible.

In an effort to determine the error in the value of residual stress

used in the calculations, measurements on three additional co-cured sandwich specimens were made after the completion of the test program re- ported herein, using the method described previously. The results of

the four tests on co-cured specimens are given in table 6. The average

of the four test results in considerably lower than the value used in the

calculations, but the scatter between specimens is so large that the use

of the average value in the calculations may not yield much better results

for any given specimen. Figure 31 shows two analytical curves, one using

the value of residual stress determined from the one original measurement,

and one using the average value of the four measurements, compared with

the experimental curve for Specimen BA-23. While the agreement between

the experimental curve and calculated curve is better when the average

value of residual stress is used, the discrepancy is still considerable,

and this still may be because the residual stress for that particular

specimen was much less than the average value used. To solve this

problem, some nondestructive method of determining the residual stress

in each specimen to be tested is needed. A more fruitful solution would

result from research leading to the development of specimen preparation

techniques which permit the residual stress to be controlled closely.

15
d. Comparison of Strain-Hardening and Time-Hardening Rules

For the tests at the lower three temperatures, there is little or no difference between the creep curves calculated using the strainhardening rule and those calculated using the time-hardening rule. This was because the creep strains in the aluminum alloy, e , were all in

3. the secondary (steady-state) stage, and it doesn't matter if the translation from one creep curve to another is along a constant time line or

a constant strain line.

For the tests at 300 °F (149 °C) , the creep strains in the aluminum

alloy were larger, entering the tertiary stage (increasing creep rate) of

the curves. According to the strain-hardening rule, the translation from

one curve to another is along a constant strain line. Since the- slopes

of the curves are increasing, this gives a higher value of e than the

3. time-hardening rule, by which the translation is along a constant time

line. Thus, after the aluminum alloy has entered tertiary-stage creep,

the strain-hardening rule predicts more creep strain in the sandwich

specimens than does the time-hardening rule. If creep data for the aluminum alloy specimens were available substantially beyond 100 hours, the

strain-hardening curves for the sandwich specimens could have been calculated to 100 hours, and, judging from the slopes of these curves, the

correlation with the experimental curves would have improved with increased time.

9. CONCLUSIONS

Based on the experiments and analyses covered by this report the

following conclusions are drawn:

1. Creep is not a problem in 0° unidirectional boron/epoxy

at temperatures below 300 °F (149 °C) because the

filaments, which previous work has shown not to creep

at these temperatures, carry approximately 98 percent

of the load. The maximum creep strain observed in

boron/epoxy in the investigation was less than 10 percent

of the elastic strain.

2. Creep can be a problem in 7075-T6 aluminum alloy reinforced with 0° unidirectional boron/epoxy even at

temperatures below 300 °F (149 °C) and loads below 75

percent of the static strength of the specimen. Stress

relaxation in the aluminum alloy causes a continuously

increasing percentage of the total load to be carried

by the composite, resulting in an increased elastic

strain which may be called "creep".

16
The amount of creep in a sandwich specimen is directly

related to the percentage of the total load carried by

the aluminum alloy. In tests run with initial con- ditions of 300 °F (149 °C), 40,000 lbf/in2 (2.76 x

10 8 N/m2

) applied stress in the aluminum alloy and

131,000 lbf/in2 (9.04 x 108 N/m2

) applied stress in

the boron/epoxy, a sandwich specimen with a compositeto-metal volume ratio of 0.18 ruptured after 1.8 hours

with 2410 x 10~ 6 creep strain, while a specimen with

a ratio of 1.17 had not ruptured after 265 hours with

1920 x 10~ 6 creep strain.

Analytical curves calculated using the strain-hardening

rule generally give somewhat better agreement with

experimental results than the curves calculated using

the time-hardening rule.

The analysis described in this report predicts creep curves which are insufficiently accurate to allow

their use in design. However, the shortcoming may

not be in the analysis itself, but rather in the

inaccuracy of the values of residual stress which

are used in the calculations.

The author would like to express his gratitude to Dr. Leonard

Mordfin at the National Bureau of Standards for his encouragement and

many helpful suggestions throughout this program.

17
10. REFERENCES

[1] Fechek, F. J., Advanced Composites Efforts-A Status Report of Air

Force Programs with Graphite Reinforced Composites, Paper 71-DE-13,

ASME, (April 1971).

[2] Ellison, E. G. and Boone, D. H. , Journal of Less Common Metals,

Vol. 13, 103-111 (1967).

[3] Shockey, P. D. , Hofer , K. E. , and Wright, D. W. , Structural Airframe Applications of Advanced Composite Materials, Vol. IVMechanical Properties-Static. AFML TR 69-101, Vol. IV (Oct. 1969).

J4 J Lou, Y. C. and Schapery, R. A. , Viscoelastic Behavior of a Nonlinear Fiber Reinforced Plastic, AFML TR 70-113, (May 1970).

15] Breinan, E. M. and Kreider, K. G. , Axial Creep and Stress Rupture

of B-Al Composites, Metallurgical Transactions, Vol. 1, 93-104,

(Jan. 1970).

16] Toth, I. J. , Creep and Fatigue Behavior of Unidirectional and

Cross-plied Composites, Composite Materials ; Testing and Design , ASTM STP 460, American Society for Testing and Materials, 236-353

(1969).

[7] Menke, G. D. and Toth, I. J. , The Time Dependent Mechanical Behavior of Composite Materials, AFML TR 70-174, (1970).

[8] Mordfin, L. , Analytical Study of Creep Deflection of Structural

Beams, NASA TN D-661, (1960).

[9] Mordfin, L. and Halsey, N. , Experimental Investigation of Creep

Deflection of Extruded and Riveted I-Beams , NASA TN D-662, (1960).

[10] Popov, E. P., Bending of Beams with Creep, Jour, of Applied Physics.

Vol. 20, No. 3, 251-256, (March 1949).

[11] Hayashi, T. , On the Law of Mixture of Creep Strength of Fiber Reinforced Composite Materials, Trans. Japan Society for Aeronautical

and Space Sciences, Vol. 10, No. 17, 54-58, (1967).

[12] Heimerl, G. J. and McEvily, A. J. , Generalized Master Curves for

Creep and Rupture, NACA TN 4112, (1957).

[13] Heimerl, G. J. and Farquhar , J., Compressive and Tensile Creep of

7075-T6 and 2024-T3 Aluminum Alloy Sheet, NASA TN D-160, (1959).

18
Table 1 - Results from Tension Tests on Boron/Epoxy and 7075-T6 Aluminum Alloy

Specimens.

Test Specimen number

Fiber

content

by volume

Elastic

modulus

Maximum

temperature Boron/Epoxy Aluminum alloy strength

°F percent lbf/in2 lbf/inz

75 B-5 - 54 31.3 x 10 6 195,000

75 B-9 - 54 32.7 199,000

150 B-15 - 47 30.3 190,000

150 B-17 - 47 31.5 177,000

225 B-13 - 47 31.9 162,000< a )

225 B--16 - 47 32.7 160,000< a)

300 B-6 - 54 (b) 179,000

300 B-14 - 47 32.0 166,000< a )

75 - A-l - 9.83 x 10 6 74,000

75 - A

- 9.63 (b)

» - A-4 - 9.35 (b)

150 - A8

- 9.94 71,000

225 _ A-9 _ 10.0 62,000

^failed in grip

(b)

measurement not made

19
Table 2 - Results of Tensile Creep Tests on Boron/Epoxy Specimens.

(a)

(b)

(c)

(d)

Failed in gage section.

Failed in gage section upon application of load

Failed partially in grip.

Slipped out of end tabs.

Fiber

content Total Durat ion

Test Specimen by creep of

temperature Stress No. volume strain test Failure

°F lbf/ in2 percent icr 6 b

75 150,000 B-24 52 0.5 (a)

75 150,000 B-25 52 115 160 none

75 175,000 B-26 52 (b)

150 150,000 B-23 52 80 140 none

150 175,000 B-22 52 (b)

150 175,000 B-19 47 40 4.6 (c)

225 125,000 B-30 50 1.5 (d)

225 125,000 B-31 50 60 20.5 (c)

225 150,000 B-18 47 460 120 none 225 150,000 B-20 47 6.2 (c)

300 125,000 B-27 52 (b)

300 150,000 B-21 52 (b)

20
Table 3 - Results of Tensile Creep Tests on 7075-T6 Aluminum Alloy

Specimens.

Total Duration

Test Specimen creep of

temperature Stress No. strain test Failure

°F lbf/in2 io- 6 h

75 50,000 A-ll< a) 430 100 none

75 60,000 A- 13 1,120 150 none

150 30,000 A- 33 110 100 none

150 40,000 A-18 510 137 none

150 50,000 A-17 3,880 100 none

225 30,000 A-25 980 110 none

225 35,000 A-28 1,500 119 none

225 40,000 A- 19 1,840 100 none

225 45,000 A-34 3,290 115 none

225 48,000 A-35 110,000 13. 2 Rupture

225 50,000 A-23 105,000 0. 7 Rupture

300 10,000 A-36 320 100 none

300 20,000 A-27 1,390 140 none

300 30,000 A-24 23,750 110 none

300 35,000 A-29 87,500 16. 6 Rupture

300 40,000 A-26 110,000 2. 1 Rupture

300 44,000 A- 32 116,000 0. 6 Rupture

(a) Specimen A-ll did not undergo thermal treatment prior to testing. All

other^ specimens were exposed to 325 °F (163 °C) for 1 hour and 350 °F

(177 °C) for 5 hours prior to testing.

21
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Table 6 - Residual Strains at 75 °F in Aluminum

Alloy of Co-cured Sandwich Specimens.

Specimen No . Residual strain

in aluminum

BA-30 +2270 x 10" 6

BA-28 +1590

BA-34 + 733

BA-35 + 848

Average +1360 x 10" 6

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M
APPENDIX A

Conversion of II. S. Customary Units to SI Units

For simplicity, only U. S. customary units have been used in the

tables and figures of this report. It should be noted that the U.S.A.

is a signatory to the General Conference of Weights and Measures which

gave official status to the metric SI system of units in 1960. Conversion factors for units used in this paper are given in the following

table:

Physical

quantity

U. S. cust

unit

Dmary

SI unit

Force pound-force (lbf) newton (N)

Length inch (in) meter (m)

Area in2 m

2

Stress lbf/in2 N/m2

Temperature °F °C

Conversion factor

1 lbf = 4.448 N

1 in = 0.0254 m

1 in2 = 6.4516 x 10

-I+ m

2

1 lbf/in2 = 6895 N/m2

°C = 0.556 (°F - 32)

Other conversion factors can be found in ASTM Standard Metric Practice

Guide, ASTM Designation E380-70 (available from American Society for

Testing and Materials, 1916 Race Street, Philadelphia, Pennsylvania

19103).

56
APPENDIX B

Glossary

boron/epoxy - a composite consisting of continuous boron filaments in an epoxy matrix.

co-curing process - a process for fabricating sandwich specimens by which

the boron/epoxy is cured and bonded to the metal in one operation.

composite - a material in which continuous high strength or high modulus

fibers are deliberately oriented in a matrix in such a way as to

increase its structural efficiency.

elastic follow-up technique - a technique for calculating creep strains of

•non-uniform structures which assumes that all load redistributions

between components are elastic.

prepreg - pre-engineered, ready-to-mold combination of resin and reinforcement.

sandwich - a specimen consisting of a sheet of 7075-T6 aluminum alloy on each side of a boron/epoxy laminate.

strain-hardening rule - a theory of cumulative creep deformation which

assumes that instantaneous creep rate is a function of stress and

previous creep strain.

time-hardening rule - a theory of cumulative creep deformation which

assumes that instantaneous creep rate is a function of stress and

the time the material has been creeping.

two-step process - a process for fabricating sandwich specimens by which

the boron/epoxy is first cured then bonded to the metal in a second

operation.

57

ormKiBS-INA (1-71) 2. Gov't Accession No.

U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO. NBS-TN-722

3. Recipient's Accession No.

4. TITLE AND SUBTITLE

Tensile Creep of Boron/ Epoxy and Boron/ Epoxy-Reinforced

7075-T6 Aluminum Alloy.

5. Publication Date May 1972

6. Performing Organization Code

7. AUTHOR(S)

Daniel J. Chwirut

8. Performing Organization

10. Project/Task/Work Unit No. 2130444

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

11. Contract/Grant No. NASA Order L-43,554

12. Sponsoring Organization Name and Address National Aeronautics and Space Administration

Langley Research Center

Hampton, Virginia 23365

13. Type of Report & Period Covered

Final

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

14. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here.)

Tensile creep tests were performed on twelve specimens of 0° unidirectional

boron/epoxy and on twenty-four specimens of 7075-T6 aluminum alloy reinforced

with 0° unidirectional boron/epoxy. An analytical procedure for predicting the

creep properties of composite-reinforced metals is presented. Agreement between

calculated and experimental creep curves varies with test temperature. These

discrepancies between experiment and analysis are probably not due to an error

in the analysis itself, but rather to uncertainty in the residual stresses in

the specimens.

'• KEY WORDS (Alphabetical order, separated by semicolons)

Aluminum alloy; boron/epoxy; composite materials; creep; elastic follow-up technique;

sandwich specimen.

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