NBSIR 74-359

SEMI-ANNUAL REPORT ON MATERIALS RESEARCH

W SUPPORT OF SUPERCONDUCTING MACHINERY

L.L. Sparks, F.R. Pickett, J.G. Must, P.J. Giarratano,

H.M. Ledbetter, E.R. Naimon, W.F.Weston, M.B. Kasen,

R.L Tobler, R.P. Mikesell, R.L. Durcholz, C.W. Fowlkes

and R.P. Reed

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

March 1974

Prepared for:

Advanced Research Projects Agency

1400 Wilson Boulevard

Arlington, Virginia 22209

NBSIR 74-359

SEMI-ANNUAL REPORT ON MATERIALS RESEARCH

IN SUPPORT OF SUPERCONDUCTING MACHINERY

L.L. Sparks, F.R. Fickett, J.G. Must, P.J. Giarratano,

H.M. Ledbetter, E.R. Naimon, W.F.Weston, M.B. Kasen,

R.L. Tobler, R.P. Mikesell, R.L. Durcholz, C.W. Fowlkes

and R.P. Reed

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

March 1974

Prepared for:

Advanced Research Projects Agency

1400 Wilson Boulevard

Arlington, Virginia 22209

U.S. DEPARTMENT OF COMMERCE, Frederick B. Dent. Secretary

NATKDNAL BUREAU OF STANDARDS. Richard W Roberts. Director

SEMI-ANNUAL REPORT ON MATERIALS RESEARCH

IN SUPPORT OF SUPERCONDUCTING MACHINERY

Sponsored byAdvanced Research Projects Agency

ARPA Order No. 2569

Program Code 4D10

August 10, 1973 - August 9, 1974

Program Director

Dr. E. C. van Reuth

Materials Sciences

Advanced Research Projects Agency

1400 Wilson Boulevard

Arlington, Virginia 22209

Program Manager

Dr. R. P. Reed

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

The views and conclusions contained in this document are those of

the authors and should not be interpreted as necessarily representing

the official policies, either expressed or implied, of the Advanced

Research Projects Agency or the U. S. Government.
ABSTRACT

Results of six months research are reported to the

sponsor, the Advanced Research Projects Agency of the Department of Defense. Subjects include magneto-thermal conductivity,

composites, elastic properties , fracture toughnes s , fatigue, and

tensile. All measurements include the temperature range 4 to

300K. Materials examined are those either presently being used

in superconducting machinery or considered for use in future

prototypes. Material classes include stainless steels, inconels,

titanium alloys, and composites.

Special results include: the thermal conductivity in a magnetic field is considerably lower than would be predicted;

a comprehensive review of glass -reinforced composite behavior

at low temperatures is included; the elastic moduli of 12 engineering

alloys from 4 to 300 K are reported; and fracture toughness and

fatigue crack growth rate data on AISI 304, AISI 316, A286, Ti5A1-4V and Ti-6Al-2. 5Sn at 4, 76 and 300K have been measured.

At the beginning of each individual report a Summary is

provided to highlight the project results.

n
TABLE OF CONTENTS

Page

Magnetothermal Conductivity

L. L. Sparks and F. R. Flckett 1

Thermal Conductivity

J. G. Hust and P. J. Giarratano 29

Elastic Properties/Engineering Alloys at Cryogenic Temperatures

H. M. Ledbetter, E. F. Naimon, and W. F. Weston A5

Advanced Composites

M. B. Kasen 101

Fatigue and Fracture Toughness Testing at Cryogenic Temperatures

R. L. Tobler, R. P. Mikesell, R. L. Durcholz,

C. W. Fowlkes, R. P. Reed 181

ill
DISCLAIMER

Tradenames of equipment and materials are used in this report

for clarity and in order to conform with standard usage in the literature.

The selection of materials for discussion and examination with regard to

application in superconducting machinery is based on properties reported

in the literature, and must be regarded as preliminary and tentative. In

no case does such selection as the results reported imply recommendation

or endorsement by the National Bureau of Standards, nor does it imply

that the material or equipment is necessarily the best available for the

purpose.
SEMI-ANNUAL REPORT ON MATERIALS RESEARCH

IN SUPPORT OF SUPERCONDUCTING MACHINERY

Magnetothermal Conductivity

by

L. L. Sparks and F. R. Fickett

NBS

Cryogenics Division

Institute for Basic Standards

Boulder, Colorado
SUMMARY

The purpose of this study is to determine the thermal conductivity

of technically important materials in high magnetic fields at cryogenic

temperatures. This information is becoming increasingly important as the

use of machinery operating at cryogenic temperatures develops. There

has been essentially no previous experimental work on the magnetothermal

conductivity of materials which will be used in cryogenic machinery.

Appendix A of this report is a bibliography of work in the area of

magnetothermal conductivity; this literature deals mostly with high

purity materials which are of little use in machine design. It does

indicate, however, that the effect of the magnetic field on the thermal

resistivity can be large. Our program will determine the extent of this

effect on technical alloys and other commercial materials which will be

used in actual structures.

The experimental approach utilizes the axial heat flow method of

determining thermal conductivity. This method basically introduces a

measured, steady heat flow at one end of a cylindrical specimen; the

temperature gradient caused by this energy is measured. The temperature

gradient, the heat applied, and the specimen dimensions are used to

compute the thermal conductivity. Measurements will be made in magnetic

fields to 100 kOe. This magnet produces a one inch diameter sphere of

homogeneous (1%) field. One of the difficulties of the experiment is

the miniaturization needed to keep all critical components within the

homogeneous field. A complete description of the experimental system is

given in the Procedures section of this report.

2
Preliminary magnetothermal conductivity results in this report are

on Inconel 718. These data are primarily for system evaluation and are

very limited in the temperature range examined (~ 4.8 K to 5.8 K) . If,

however, the results prove to be valid, the magnetic field effect is

much larger than expected. The implication is that effects other than

those expected from simple theory are being encountered. Additional

materials being considered for testing (see Procedures, specimen section)

include two coppers and one additional, as yet undetermined material. If

the unexpectedly high effect on Inconel 718 is substantiated by further

testing, an additional Inconel, stainless steel, or some other near

magnetic material will be the fourth specimen.

3
TABLE OF CONTENTS

Page

Summary 2

Introduction 5

Procedures 7

Apparatus

.

7

Probe Design 7

Thermometry 11

Equipment 11

Experimental procedures 14

Specimens , 14

Discussion 18

Results 2 0

References 22

Appendix A 24

LIST OF FIGURES

Figure 1. Magnetothermal conductivity probe and magnet. 8

Figure 2. Photograph of completed probe with specimen in place. 9

Figure 3. Resistance bridge for measuring carbon resistance

thermometers (CRTs). 12

Figure 4. Block diagram of the system electronics. 13

Figure 5. Magnetoresistance of a 100 Q, 1/8 watt, AllenBradley carbon resistor at selected cryogenic

temperatures. 21

Figure 6. Preliminary data on unannealed Inconel 718 showing

^ (H) / X (0) as a function of applied magnetic field. 22

LIST OF TABLES

Table 1. Magnetothermal conductivity test specimens. 17

4
INTRODUCTION

Optimum design of machinery operating at cryogenic temperatures

requires a detailed knowledge of the thermal and electrical characteristics of the materials at low temperatures. Most applications also

involve magnetic fields of 0-10 Tesla, and it is well known that fields

of this intensity alter the physical properties of conducting materials,

sometimes causing changes of several orders of magnitude in a given

property.

The purpose of this study is to determine the behavior of the

thermal conductivity of technically important materials in high magnetic

fields at cryogenic temperatures. The materials are those actually used

for the structural and electrical components of machinery. This class

includes, but is not restricted to, alloys such as stainless steels and

Inconels, metals used as stabilizing materials (Cu, Al and Ni) and,

possibly, composite superconductors. The data generated from this

program is needed for optimum machine design, and also it will help us

to gain a better understanding of heat transfer in metals and alloys

under the influence of high magnetic fields.

The effect of magnetic fields on the thermal resistivity of metals

is not known. The world's literature on magnetothermal conductivity is

essentially given in Appendix A of this report. Most of these papers

deal with very pure metals and single crystals and are of little interest

to this program. No literature at all exists on modem technical alloys.

The literature on pure metals indicates that large increases in thermal

resistivity can occur at high fields. Our program will determine how

large this effect is in the various technical alloys and less pure

metals used in actual structures.

5
It is possible in many cases to predict the approximate zero field

low temperature thermal conductivity of a given material from the electrical

resistivity for the specimen in question and knowledge of the Lorenz

ratio for that class of alloys. Lorenz ratio data currently exists only

for the zero field case, and in general one would expect the ratio to

show a different behavior when a magnetic field is applied. Data from

this project will allow us to evaluate the possibility of a "magnetic

field" Lorenz number for predicting the thermal conductivity of alloys

in a magnetic field from a measurement of the electrical resistivity

inherently a much easier measurement to make than the thermal one.

6
PROCEDURES

Apparatus

The experimental determination of the magneto thermal conductivity,

X (H) , of metals is complicated by the requirement that the specimen be

contained in a homogeneous magnetic field. For all but the very largest

laboratories, this restriction necessitates small specimen lengths and

thus relatively small temperature gradients along the specimens. Furthermore,

the precision thermometry required for the measurement is considerably

complicated by large, nonlinear magnetic field effects on the thermometers.

The probe described here is designed to be used in a superconducting

solenoid with a 1.5 inch bore and a 1 inch homogeneous (1%) field sphere.

The maximum available magnetic field is 100 kOe (7.96 MA/m)

.

Probe Design Figure 1 shows the principal components of the probe

with the specimen in the longitudinal position (H parallel to heat flow,

Q) . The temperature controlled heat sink (TCHS) is designed to allow a

(somewhat shorter) specimen to be installed in a horizontal position for

transverse (H Q) measurements.

At our operating temperatures of 4-20 K, radiative heat transfer is

not important and convective losses from the specimen are prevented by

evacuating the copper SPECIMEN CHAMBER,which is usually immersed in a

liquid helium bath. The sink (TCHS) temperature is selected by adjusting

the power input to a wire heater (Evanohm, 36 AWG, 40 fi ) wound on the

TCHS. Energy is transmitted up to the liquid helium (LHe) bath via the

THERMAL LINKS which are three stainless steel (310 ss, 3/16" diam.)

rods. Provision is also made for connecting a //18 copper wire in parallel

with each of the stainless steel rods in order to give additional heat

7
Vacuum and

Electrical Leads

Support Tubes Liquid Fill Tubes

Thermometer Block (carbon]

Specimen (longitudinal]

(transverse not shown]

Thermometer Block

(carbon]

Specimen Heater

Vacuum and

Electrical Leads

Liquid Helium Temperature Cap

Thermal Links

Temperature Controlled Heat Sink

Tempering Post

Specimen

Chamber (Cu]

100 kOe Superconducting Solenoid -

Homogeneous Field in 2.54 cm

diameter Sphere

Figure 1. Magnetothermal conductivity probe and magnet.

8
9
transfer to the bath. This added heat leak capacity will be necessary

for measurements on high conductivity specimens. Once set, the temperature

of the TCHS is controlled automatically. For purposes of temperature

measurement and control, the sink is instrumented with a 100 Q , 1/8 W

carbon resistance thermometer (CRT) and a germaniiun resistance thermometer

(CRT). The CRT, while very accurate, is useless in a magnetic field,

so high field temperature readings are taken on the field-calibrated CRT.

The temperature gradient, A T, along the specimen is established by

heating the bottom end of the specimen with an electrical heater (Evanohm,

40 AWG wire, 40^). The absolute temperature and A T are determined by

measuring the resistance of CRTs embedded in each of the THERMOMETER

BLOCKS mounted on the specimen. The stainless steel (310 ss) TEMPERING

POST thermally anchors the wires to the specimen temperature before they '

actually contact the specimen. This tempering eliminates energy transfer

via the wires. The TEMPERING POST is attached to the TCHS at one end

and is adjusted to match the specimen temperature by an electrical

heater (Evanohm, 40 AWG wire, 40 Q ) on the lower end (not shown in Fig.

1) . A differential thermocouple (KP vs Au 0.07 at % Fe) is connected

between the lower end of the tempering post and the lower themometer

block on the specimen in order to monitor and maintain the necessary

zero temperature difference.

A photograph of the completed probe with a specimen mounted in the

longitudinal configuration is shown in Fig. 2.

The probe is also instrumented for measurements of electrical

resistivity and thermopower ~ information which is sometimes helpful in

interpreting the thermal conductivity data.

10
Thermometry Carbon resistors (Allen-Bradley, 1/8 watt, 100^) are

used to measure both the absolute temperature of the specimen and the

temperature difference between the two THERMOMETER BLOCKS. A calibrated

CRT is situated in the TCHS for purposes of zero field calibration of

the CRTs.

Neuringer and Shapira (1969) found that the relative effect of

magnetic fields, AR/R^ = (R^ - R^)/R^, is very nearly the same for all

Allen-Bradley resistors with the same power rating and nominal room

temperature resistance. These authors have generously supplied detailed

data for AR/Rq ^.s a function of temperature for 100 Q , 1/8 watt, AllenBradley CRTs.

Measurements of CRT resistances are made on an a.c. Wheatstone

bridge. Low frequency a.c. power eliminates the need for current reversal

when determining the resistance of the CRTs. A lockin amplifier serves

as both the detector and the bridge power supply. The bridge system

(Fig. 3) is designed to allow direct measurement of the resistance of

all of the CRTs as well as the resistance difference measurement between

the thermometers in the THERMOMETER BLOCKS.

Equipment The electronics associated with the experiment are shown

diagrammatically in Fig. 4. The 100 kOe superconducting magnet system

and the electronics associated with it are not shown in the figure. The

precision voltage measurement system is indicated by a single box. This

system, capable of measurements in the .1-1 nV range, is described in

detail by Clark and Fickett (1969).

11
(1) ABSOLUTE TEMPERATURE

MEASUREMENT RaT

(2) DIFFERENTIAL

TEMPERATURE

MEASUREMENT RaAT

Figure 3. Resistance bridge for measuring carbon resistance thermometers

(CRTs). Drive voltage and detection are provided by a lockin

amplifier, A and B are lOK Q, 0. 01% resistors. R is a series

of precision decade resistors. (1) X is any one of the three CRTs

(2) Xi and Xg are the CRTs used to determine the specimen AT.

12
THERMi,OMETRY n

LOCKIN AMPLIFIER

WHEATSTONE

BRIDGE

SPECIMEN

HEATER

POWER s

TCHS CONTROL

POTENTIOMETER

TO DVM ^

j SPECIMEN

TO DVM

GRT POWER

GALVANOMETER

AMPLIFIER

I I TCHS HEATER CONTROLLER

JCRT TCHS

GRT

"1

1

L

CRYOSTAT _JL—^N^/t>

THERMOCOUPLE

OUTPUT AMPLIFIER

POST HEATER

CONTROL

r RESISTIVITY

~^

MEASUREMENT

GALVANOMETER

AMPLIFIER

I

POTENTIOMETER

I TE_M_PER I NG ^ST QNTROL

J

PRECISION

VOLTAGE DETECTION

SYSTEM

SAMPLE CURRENT

SUPPLY lT I

Figure 4. Block diagram of the system electronics

13
Experimental Procedures

Two basic procedures are needed In order to acquire the experimental

data. The first Is calibration of the CRTs and the second Is the actual

measurement of the magnetothermal conductivity,

X (H) , and the electrical resistivity, p (H)

.

As mentioned previously, work on CRTs has shown that A R/R^ =

R )/R holds for all similar resistors. Thus, to find T(R,H) all we

o o need is R versus T at zero field. These data are obtained by a series

of isothermal measurements — isothermal in the sense that the TCHS, the

specimen, and the THERMOMETER BLOCKS are all at the same temperature.

This is done as follows: 1) the evacuated SPECIMEN CHAMBER is immersed in

LHe and allowed to reach a stable temperature; 2) the system reaches an

equilibrium in less than 5 minutes below 20 K; 3) the equilibrium temperature

is determined by the CRT, and the resistances of the three CRTs at this

temperature are recorded; 4) a higher equilibrium temperature is achieved

by powering the TCHS heater. In this manner the zero field R-T profile

is found for each CRT.

A check of the Neurlnger and Shapira formula for R(H,T) was made

for our CRTs. For this procedure helium gas is Introduced into the

evacuated SPECIMEN CHAMBER which then serves as a constant volume gas

thermometer. The system is cooled by a transfer gas arrangement between

the SPECIMEN CHAMBER and the liquid helium surrounding the superconducting

magnet. Isothermal data is taken as before at zero field, then at

successively higher fields. Constant temperature is maintained by

14
adjusting the TCHS heater power so as to maintain constant pressure in

the SPECIMEN CHAMBER.

The second procedure is the experimental determination of X (H) and

p (H) . Refrigeration of the SPECIMEN CHAMBER during these tests is

either by transfer gas to the magnet LHe or by immersion of the chamber

in LHe. In either case the SPECIMEN CHAMBER is evacuated. The thermal

conductivity of the specimen at T is calculated from (H,T) = ^

.

The parameters are: bl {= distance between THERMOMETER BLOCKS), A {=

(cross sectional area of specimen) , Q ( = power used to establish the

temperature gradient) and A T ( = temperature difference between the

THERMOMETER BLOCKS)

.

In an actual measurement the temperature of the TCHS is set at some

selected temperature. Power to the SPECIMEN HEATER is applied until the

desired A T is obtained. The heater power is measured by the system

shown in Fig. A. The absolute temperature of the two specimen thermometers

is measured, as is the resistance difference between them. A series of

these measurements are made as the field is progressively increased.

Data are taken over the range from 4-20 K and from 0-100 kOe.

At each gradient and field, the electrical resistivity of the

specimen is also measured at T = 1/2 (T„^_ + T ) by a conventional

nUi LULD

four probe method. A potentiometer is used to measure the voltage drop

for high resistivity specimens and the precision detection system for

lower resistivity ones.

Specimens

Table 1 lists the initial set of specimens to be tested in the

magneto thermal conductivity program. Inconel 718 was the first to be

installed and preliminary data taken on this specimen is being used in

the system evaluation. The two copper specimens will be tested following

15
the Inconel. A final decision has not been made as to what material

will be used as the fourth specimen. Iron, as listed in table 1, and a

different Inconel are being considered.

16
Table 1

Magnetothermal Conductivity Test Specimens

Material

Inconel 718

Copper (OFHC)

Copper (STOCK 7)

Iron (OSRM)

Specimen Characterization

Produced by International Nickel Company,

HT4675E. Specimen not annealed,

hardness B39.

Swaged to ~ 0.1 inch diameter from 3/16

inch rod stock. Etched (50-50 H2O-HNO2)

and annealed at BOO^C for 1 hour in vacuum

several times during swaging process.

Final anneal at 850°C for 1 hr. in vacuum.

< 1 ppm of any impurity with possible

exception of Fe. Specimen was centerless

ground to ~ 0.1 inch diameter. Annealed

at 850°C for 1 hr. in vacuum.

Specimen was centerless ground to ~ 0.1

inch diameter. To be annealed at 400*C

for 46 hours in vacuum.

17
DISCUSSION

In this section we present a very brief explanation of the mathematical

basis of our experiment. We also define the terminology commonly used

in the literature.

Thermal conductivity, in the absence of a magnetic field, is probably

one of the best understood of the transport effects observed in solids.

It is however a relatively complex effect because many separate mechanisms

are operating simultaneously to transmit energy through the material.

The basic quantity of interest is the thermal conductivity, \ , as

defined by the heat transfer equation, which, in its simplest one dimensional

form, is »

Q = X(T) a£ ,

where, conventionally, Q is the heat current in watts, A is the cross2 sectional area in cm , T is in Kelvin and thus X has the units of

watt /cm K. In general the thermal energy in a material is transported

both by the lattice vibrations (phonons) and by the conduction electrons.

These two modes operate as thermal resistors in parallel, such that the

total thermal conductivity is

X = X + X , e g

where the subscript g is the conventional designation for the lattice

component. Each of these two contributions are in turn made up of a

number of separate components, reflecting the modes by which energy may

be lost in transit. These are series additive such that the thermal

resistivity, W, of, say, the electronic branch may be written,

W = = W(electron-lattice) + W (electron-impurity) + W(electron-defect)

.

e

18
The lattice contribution, X , is small in pure metals but may be

significant in alloys, and in nonconductors (where X ^ = 0) it represents

the total contribution. Application of a magnetic field to a thermal

conductor results in a change primarily in the electronic component of

the conductivity. Theory predicts little effect on the lattice component

although no experimental data exist to verify this result. Our experiment will be the first to look at this effect in detail.

As with the electrical (galvanomagnetic) effects three common

orientations are possible between the temperature gradient, the magnetic

field, and the heat flow:

(1) All three are parallel fH , Q , dT/dzl. This is the longiz z tudinal magnetothermal resistance.

(2) The temperature gradient and heat flow are parallel and the

field is normal to that direction [H , Q , dT/dx] . This is

z x

the transverse magnetothermal resistance also known as the

Maggi-Righi-Leduc effect.

(3) All three are at right angles to each other [H^, Q^, dT/dy]

.

This is the thermal Hall effect or the Righi-Leduc effect.

Other effects, such as the behavior of the thermopower under the influence

of a magnetic field, can be included in the list, and in fact some have

been referenced in the bibliography of Appendix A; but this project is

primarily concerned with (1) and (2) above.

As is the case with the electrical magnetoresistance, the thermal

resistance is observed to always increase in an applied field. The

effect in pure metals can be quite large, the resistivity increasing by

more than an order of magnitude at 4 K. One would expect a smaller

effect for most alloys, but even a relatively small change can be critical

in some applications.

19
Another parameter of interest is the Lorenz ratio. Because the

conduction electrons carry both thermal and electrical energy, it is

reasonable to assume that the electronic contribution to the thermal

conductivity and the electrical conductivity should be related. The

equation expressing this relationship is the Wiedemann-Franz-Lorenz law,

where, by simple theory, the Lorenz ratio, L, should be a constant.

Sommerfeld calculated the free electron theory value to be = 2.443 x

_g 2 2

10 V /K . Needless to say, L is not altogether constant, but it is the

most constant of the trio, A.

, p , L. In any event, if one knows the L

vs T curve for a given alloy group, and the electrical resistivity vs T

for a particular alloy, it is quite possible to predict the thermal

conductivity at zero field with sufficient accuracy for many applications.

We wish to determine whether or not a similar relationship holds in the

presence of a magnetic field. Indications are that L is field dependent,

at least in pure metals.

RESULTS

Data given below must be considered preliminary and is being used

for purposes of system evaluation. The computer programs necessary to

compute temperatures for each of the CRTs as a function of resistance

and magnetic field are especially needed for accurate analysis of data.

Figure 5 indicates the field sensitivity of one of the three CRTs;

note that the field effect is not large even at 4 K. Figure 6 gives the

ratio of the thermal conductivity of Inconel 718 in the presence of a

magnetic field to the thermal conductivity at the same temperature with

zero field. If this preliminary data remains substantially unchanged

upon further analysis, the magnetic effect is larger than expected and

almost certainly indicates some interaction beyond the electronic effect

considered in the simple theory.

20
MAGNETORES I STANCE

CRT„

Preliminary Data

(4.22K)

(6.20K)

(8.31K)

(10.72K)

(15.38K)

0 20 40 60 80 100

H, kOe

are 5. Magnetoresistance of a lOOQ, 1 / 8 watt, AIIen-Bradiey

carbon resistor at selected cryogenic temperatures.

21
o

o

CO

00 o , •—

o

CO

>-

1

>

1— <

(_> 1- ZD <

oo 1=1 .—1 o >

<_> _i <

1 LUZ

o

en u

_iLU 1— q; TOUJ

<i:

o

CO

in o

sl-J 03

S-01Q.

(U I—

!-> c:

0) n

E CM

01

IT)

10

ID

01

:^<Ao

o

cr. o

00o

(OK

(H)Y

o

Figure 6. Preliminary data on unannealed InconeL 718 showing

X (H) / X (0) as a function of applied magnetic field.

22
REFERENCES

A. F. Clark and F. R. Fickett, RSI 40, 465 (1969).

L. J. Neuringer and Y. Shapira, RSI 40, 1314 (1969).

23
"Bibliography of Magnetothermal Measurements

at Low Temperatures"

1973

T. Amundsen and 0. Kverndalen, "A Note on the Anisotropy of the Thermal

Magnetoresistance of Aluminum," Phys. Lett. 43A, 9 (1973).

R. S. Averback, C. H. Stephan and J. Bass, "Magnetic Field Dependence of

the Thermopower of Dilute Aluminum Alloys," J. Low Temp. Phys. 12^, 319

(1973).

A. Fevrier and D. Morize, "The Effect of Magnetic Field on the Thermal

Conductivity and Electrical Resistivity of Different Materials,"

Cryogenics 13, 603 (1973).

M. V. Glushkov and A. I. Markin, "Longitudinal Magnetoresistance of

n-Si in Strong Magnetic Fields," Sov. Phys. Solid State 14, 2967

(1973).

1972

R. S. Averback and D. K. Wagner, "Low-Temperature Magnetothermoelectric

Power of Aluminum," Solid State Commun. 11, 1109 (1972).

J. M. L. Engels, F. W. Gorter and A. R. Miedema, "Magnetoresistance of

Gallium — A Practical Heat Switch at Liquid Helium Temperatures,"

Cryogenics 12 , 141 (1972)

.

R. S. Newrock and B. W. Maxfield, "Thermal and Magnetothermal Resistance

of Potassium," Bull. Am. Phys. Soc. 17, 256 (1972).

N, N. Sirota, A. A. Drozd and V. I. Gostishchev, "Measurement of Electrical

and Thermal Conductivity of Metals in Strong Magnetic Fields,"

Thermophysical Properties of Substances at Low Temperatures (Proc.

1st all-union meeting) M. P. Orlova ed., Published by All-union

Scientific-Research Inst, of Physicotechnical and Radiotechnical

Studies (Moscow, 1972) In Russian.

N. N. Sirota, V. I. Gostishchev and A. A. Drozd, "Thermal Conductivity of

Aluminum in Strong Magnetic Fields at Low Temperatures," JETP Letters

16, 170 (1972).

N. N. Sirota, V. I. Gostishchev and A. A. Drozd, "Thermoelectric Power

of Aluminum in Strong Magnetic Fields at Low Temperatures,"

JETP Lett. 16, 409 (1972).

H. Van Kempen, H. N. DeLang, J. S. Lass and P. Wyder, "Lattice Conductivity

and Thermal Linear Magnetoresistance of Indium Using the Corbino

Geometry," Phys. Lett. A. 42, 277 (1972).

C. H. Stephan and B. W. Maxfield, "Righi-Leduc Effect in Indium," Phys.

Rev. B 6, 2893 (1972)

.

D. K. Wagner, "Lattice Thermal Conductivity and High-Field Electrical and

Thermal Magnetoconductivities of Tungsten," Phys. Rev. B _5, 336 (1972).

24
1971

J. R. Long, "Thermal and Electrical Transport in a Tungsten Crystal for

Strong Magnetic Fields and Low Temperatures," Phys. Rev. B 3^, 1197 (1971).

1970

T. Amundsen and R. P. Sovik, "Measurements of Thermal Magnetoresistance

of Aluminum," J. Low Temp. Phys. 2^, 121 (1970).

L. J. Challis and R. G. Ward, "Magnetic Field Dependent Phonon Scattering

3+

by MN ion in Al^O-j," 9th Conf. on Thermal Cond. (1970), H. R.

Shanks ed. [USAEC Conf-69-1002] pp 3-8.

L. Halbo and R. J. Sladek, "Magneto-Thermal Conductivity of n-type

Germanium at Low Temperatures," 9th Conf. on Thermal Cond. (1970)

H. R. Shanks, ed. fUSAEC Conf. 69-1002] p 9. Abstract only.

N. S. Natarajan and M. S. R. Chari, "Thermal Conductivity of Rhodium

at Helium Temperatures and in Magnetic Fields," 9th Conf. on Thermal

Cond. (1970) H. R. Shanks, ed. [USAEC Conf - 691002] pp. 208-216.

N. S. Natarajan and M. S. R. Chari, "Thermal Conductivity of Dilute

Copper-Manganese Alloys at Helium Temperatures and in Magnetic

Fields," 9th Conf. on Thermal Cond. (1970) H. R. Shanks, ed.

f USAEC Conf - 691002] pp. 169-177.

1969

T. Amundsen, "The Righi-Leduc Effect of Aluminum and Indium," Phil. Mag.

20, 687 (1969).

S. F. Grishin, E. Ya. Grishina and G. A. Milner, "Feasibility of Using

a Cadmium Thermal Switch in a Magnetic Refrigerator," Ukranian

Physics Journal 13, 969 (1969).

C. H. Stephan and B. W. Maxfield, "High Field Thermal Transport Properties

of Indium," Solid State Coram. 7_, 1039 (1969).

1968

T. Amundsen, "Size Effects in the Thermal Hall Effect of Aluminum

Films," Phil. Mag. 17» (1968).

T. Amundsen, "The Righi-Leduc Effect of Pure Aluminum in High Magnetic

Fields," Phil. Mag. 17 » 1303 (1968).

F. W. Gorter and A. R. Miedema, "Thermal Magnetoresistance of Gallium

and its Use as a Heat Switch," Cryogenics 8^, 86 (1968).

25
1966

Laudy and A. Knol, "A Cadmium Heat Switch," Cryogenics 6^, 370 (1966).

Lerner and J. G. Daunt, "Thermal and Electrical Conductivities of

Mo-Re Alloys in the Superconducting and Normal States," Phys. Rev.

142, 251 (1966).

G. Lipson, "The Righi-Leduc Effect in Copper at Low Temperatures,"

Proc. Roy. Soc. A 93, 275 (1966).

1965

Amundsen and T. Olsen, "Size-dependent Thermal Conductivity in

Aluminum Films," Philosophical Magazine 11, 561 (1965).

G. Lipson, "The Thermal Hall Effect in Copper," Proc. 9th Int. Conf. on Low Temp. Phys., Daunt et al., eds. (Plenum Press, 1965) pp. 814-817.

R. Long, C. G, Grenier and J. M. Reynolds, "Electron and Lattice Transport

4 Phenomena in an Antimony Crystal at Liquid-He Temperatures," Phys.

Rev. 140, A187 (1965).

1963

Amundsen and T. Olsen, "Oscillatory Thermal Magnetoresistance in

Aluminum," Physics Letters 4_, 304 (1963).

G. Grenier, J. M. Reynolds and J. R. Sybert, "Electron Transport

Phenomena in Bismuth at Liquid-Helium Temperatures," Phys. Rev.

132, 58 (1963).

G. Grenier, J. M. Reynolds, N. H. Zebouni, "Electron Transport Phenomena

in Zinc at Liquid-Helium Temperatures," Phys. Rev. 129, 1088 (1963).

1962

Berger and D. Rivier, "Au Resistivite Electrique et Thermique du

Nickel pur et d'un Alliage Fer-Nickel en Champmagnetique aux Basses

Temperatures," Helv. Phys. Acta 15, 715 (1962).

L. Olsen, Electron Transport in Metals , (Interscience Publishers,

1962) p. 77.

1961

Douthett and S, A. Friedberg, "Effects of a Magnetic Field on Heat

Conduction in Some Ferrimagnetic Crystals," Phys. Rev. 121, 1662 (1961).
1959

M. S. R. Chari and J. de Nobel, "Thermal Conductivity of Some Dilute

Silver Alloys at Low Temperatures and in Magnetic Fields," Physica

25, 60 (1959).

M. S. R. Chari and J. de Nobel, "Thermal Conductivity of Some non- Superconducting Alloys at Low Temperatures," Physica 25_, 84 (1959).

1957

J. de Nobel, "Thermal and Electrical Resistivity of Some Tungsten Single

Crystals at Low Temperatures and in Strong Magnetic Fields," Physica

23, 261 (1957).

1956

P. B. Alers, "Thermal Magnetoresistance of Zinc at Low Temperatures,"

Physical Review, 101, 41 (1956).

1953

K. Mendelssohn and H. M. Rosenberg, "The Thermal Conductivity of Metals

in High Magnetic Fields at Low Temperatures," Proc. Roy. Soc.

A218 , 190 (1953).

J. L. Olsen and H. M. Rosenberg, "Thermal Conductivity of Metals at

Low Temperatures," Adv. in Physics, 2y 28 (1953).

1951

K. Mendelsohn and H. M. Rosenberg, "Thermal Conductivity of Cadmium in

Magnetic Field at Low Temperatures," Proc. Roy. Soc. A64 , 1057 (1951).

1950

E. Gruneisen, K. Rausch and K. Weiss, "Conduction of Electricity and

Heat in Bismuth Single Crystals in a Transverse Magnetic Field,"

Ann. Phys. Lpz. 7_, 1 (1950).

J. K. Hulm, "Thermal Conductivity of Tin, Mercury, Indium and Tantalum

at Liquid Helium Temperatures," Roy. Soc. London (Proc.) 204A , 98 (1950).

27
1949

J. de Nobel, "The Thermal and Electrical Resistance of Tungsten at

Low Temperatures and High Magnetic Fields," Physica 15, 532 (1949).

1947

K. Rausch, "Investigations of Antimony Single Crystals in a Transverse

Magnetic Field," Ann. Phys. Lpz. 1, 190 (1947).

1944

S. Shalyt, "Thermal Conductivity of Bismuth at Low Temperatures,"

J. Physics (USSR) 8, 315 (1944).

1940

E. Gruneisen and H. D. Erfling, "The Electrical and Thermal Resistances

of Beryllium Crystals in a Magnetic Field," Ann. Phys. Lpz. 38 , 399 (1940).

1938

E. Gruneisen and H. Adenstedt, "The Influence of Transverse Magnetic

Fields on the Electrical and Thermal Conductivities of Metals

at Low Temperatures," Ann. Phys. Lpz. 31, 714 (1938).

W. J. de Haas and J. de Nobel, "Thermal and Electrical Resistance of

a Tungsten Single-Crystal at Low Temperatures and in a Magnetic

Field," Physica _5, 449 (1938).

1936

W. J. de Haas, A. N. Gerritsen and W. H. Capel, "Thermal Resistance of

Bismuth Single-Crystals at Low Temperatures and in a Magnetic Field,"

Physica 3^, 1143 (1936).

1934

A. Reddemann, "Anderung der thermischen und electrischen Leitfahigkeit

eines Bi-Einkristalls im Magnetfeld," Ann. Phys. 20, 441 (1934).

1926

N. C. Little, "Thermagnetic and Galvanothermomagnetic Effects in

Arsenic," Phys. Rev. 2^, 418(1926).

28
SEMI-ANNUAL REPORT ON MATERIALS RESEARCH

IN SUPPORT OF SUPERCONDUCTING MACHINERY

Thermal Conductivity

by

J. G. Hust and P. J. Giarrantano

Cryogenics Division

NBS - Institute for Basic Standards

Boulder, Colorado
SUMMARY

Thermal conductivity measurement philosophy Is outlined and two

apparatus are described. The need for precision measurements, lower

accuracy measurements, and predictive methods Is outlined. Candidate

materials for superconducting machinery are listed and existing data on

these materials have been surveyed. Based on this survey, materials

have been selected for measurement and specimens In various metallurgical

conditions have been obtained. Measurements on OFHC copper have been

completed, but data analysis Is still in progress; therefore, experimental

values of thermal conductivity are presented. An example Is presented

of a predictive procedure, based on Lorenz ratio and electrical resistivity

data. Upon completion of the OFHC copper data analysis, comparisons

will be made to Indicate the degree of accuracy of the predictive method.

A schedule of measurements for this contract year Is Included.

30
TABLE OF CONTENTS

Page

SUMMARY 30

INTRODUCTION 32

PROCEDURES 33

Precision apparatus 34

Fixed point apparatus 36

Material Selection and Measurement Criteria 38

RESULTS 39

DISCUSSION 40

REFERENCES 43

LIST OF FIGURES

Figure 1. Precision Thermal C onductivity Apparatus 35

Figure 2. Fixed-Point Thermal C onductivity;- Apparatus 37

LIST OF TABLES

Table 1. Calculation of thermal conductivity of OFHC copper

based upon electrical resistivity and Lorenz ratio

data 42

31
INTRODUCTION

Optimum design of superconducting machinery requires a knowledge of

thermal properties of technically important materials at low temperatures.

But, often, components extend from ambient temperatures to low temperatures,

and, therefore, data are needed from 4 K to 300 K. Thermal conductivity

of metals is a function of the host metal and its constituent elements

and temperature; but it also depends on the microstructure of the material,

strain, fatigue, heat treat, and in some cases environmental effects are

important (e.g., in the case of superconducting machinery, magnetic

field intensities may produce significant changes). Specific thermal

conductivity data for particular components are required to limit heat

losses and thermal stresses in machines.

With such a myriad of parameters it is difficult (at best) to

determine accurately, by experiment, the thermal conductivity of each

component. Instead, we often rely on a limited number of measurements

on specific types of materials, characterized according to the above

parameters, and indicate the degree of variability of these data. This

uncertainty is usually limited primarily by material variability, i.e.,

variations in the material parameters listed above.

It should also be noted that accurate thermal conductivity measurements

are much more expensive and time consuming relative to many other properties

measurements. The ease with which high accuracy can be obtained is

primarily a function of temperature and conductivity itself. The extremes

of conductivity (both very high, such as in pure metals, and very low,

such as in insulating materials) promote inaccuracies in measurement.

Generally, low temperature data (below about 100 K) are more readily

obtained with accuracy because of the absence of radiation effects.

32
For the above reasons, various methods of obtaining thermal conductivity data are used. The method used depends primarily on the accuracy

dictated by the application. Usually accuracies of 10% are sufficient;

in other cases values accurate to 50% suffice. The present state-ofthe-art in thermal conductivity measurements is about 1% accuracy, but

seldom is this required except in theoretical or standardization work.

PROCEDURES

In this work, three methods of data determination will be used:

(a) A precision apparatus, capable of measuring at any temperature

from 4 to 300 K, with an accuracy of about 2%. This apparatus

also simultaneously measures electrical resistivity and thermopower for additional specimen characterization and subsequent

predictions of similar materials. This apparatus will be used

to measure materials that have not been sufficiently measured

and characterized previously and for accurate temperature

dependence determinations.

(b) A fixed point apparatus, capable of measuring only near the

fixed temperatures of various boiling fluids (such as liquid

helium and liquid nitrogen) and melting or subliming solids

(such as ice and CO^) . The accuracy of this apparatus is at

best about 5%, depending mainly on the conductivity of the

specimen.

(c) Predictive methods such as those described by Hust and Clark

(1) and Hust and Sparks (2) . The latter paper also contains

data for many metals and alloys, necessary to carry out such

predictions.

33
These predictive methods work best for pure metals (5% accuracy

or better) , not as well for intermediate alloys (10-20%)

,

worse for structural materials (20-60%) , and are not applicable

to non-metals. The figures in parenthesis are to be considered

as rough estimates of the validity of these techniques and are

strongly dependent on temperature, as explained in the above

references. There is evidence that the above techniques will

also be reasonably valid for predicting the effects of magnetic

field. (Magnetic field effects is a separate area of investigation under this contract and is not discussed here)

.

The above techniques are listed in order of decreasing cost per

measurement and likewise increasing inaccuracies, with occasional exceptions. They will be applied to obtain the most cost effective data

consistent with system requirements.

Precision Apparatus

The precision apparatus has been described in detail by Hust et al.

(3) . It is a longitudinal heat flow apparatus requiring a 23 cm long

specimen. The diameter of the specimen is dictated by its approximate

thermal conductivity. The largest diameter which can be accommodated is

about 3 cm and the smallest is limited only by one's dexterity. Specimens

as small as 2 mm diameter have been measured. Figure 1 depicts the

cryostat of this apparatus.

The precision apparatus has been in operation for many years; the

first model was built in the late 50's. Recently, to correct some

faults which have developed during the past ten years and to incorporate

a few improvements in thermometry and temperature control, the cryostat

34
Top Plate

Floating Sink

Heat Exchanger

Specimen Clamp

^ Differential

Thermocouples

(at positions 2^5j8

and specimen heater)

(Front half of shield

removed)

Vacuum and Wiring Ducts

Reference Posts

Floating Sink, Heater

and Temperature Sensors

(carbon and alumel)

8 Tempering Rings and

Shield Trim Heaters

3 Thermocouples

and Clamps

Specimen Heater

Shield and Heater

Figure 1. Precision Thermal Conductivity Apparatus

'•"•-MS I

35
was rebuilt and new wiring was installed. For this reason, the first

specimen to be measured was chosen to fulfill several criteria:

(1) It is to be a material of interest to this program which was

not previously measured over at least a portion of the range 4

to 300 K,

(2) It is to be a good conductor to facilitate in-place calibration

of the new thermocouples, and

(3) It should be a material whose conductivity is predictable over

most of the temperature range to check the operation of the

system.

OFHC copper was chosen to satisfy these requirements. It is of obvious

interest to the program, it was never previously measured in the range 4-20

K, is a good conducting material, and is most predictable based on the

previously mentioned Lorenz ratio method.

Fixed Point Apparatus

This apparatus is inherently less accurate than the precision

apparatus; it can measure thermal conductivity at only a few temperatures

below ambient but requires less time to obtain data and requires only a

2 inch long specimen. It is a new apparatus in this laboratory and its

accuracy limitations are not well defined. Its design accuracy for

midrange conductors is 10%. A schematic drawing of the cryostat is

shown in figure 2.

The principal of operation is as follows. For each bath temperature

a series of empty chamber runs are conducted to establish = f(AT)

for relatively small AT. Is the heat input required to establish a

given temperature difference AT. A specimen is installed and the heat

input, Q, to establish a small A T is measured. The increased heat input

36
Figure 2. Fixed-Point Thermal Conductivity Apparatus

37
at this AT, Q - Q^, is determined and X, thermal conductivity of the

specimen, is calculated from

(Q - Q R

where A is the cross-sectional area of the specimen and t is the distance

between the copper clamps. The principal source of error with this

technique is due to the temperature discontinuity at the interface

between the specimen and the copper clamps. Indium is used at this

junction to improve the thermal contact. Other sources of error are

thermocouple calibration, temperature determination, loss of heat by

radiation, non-steady state, and strain caused by differential contraction.

For specimens which are small compared to the specimen chamber, radiation

losses should be nearly equal in the determination of and Q and,

therefore, are eliminated by subtraction. Temperature drift and radiation

errors are most important when measuring poor conducting, large specimens.

The temperature discontinuity errors are largest with small, good conducting

specimens. This apparatus is checked, using NBS standard reference

materials of thermal conductivity , to determine it's accuracy.

Material Selection and Measurement Criteria

This project is to be conducted as follows:

(1) Select candidate materials of interest to the ARPA superconducting machinery program.

(2) Search existing compilations to determine the candidate materials

whose thermal conductivities are not sufficiently defined

(including effects of heat treatment, microstructure

etc).

38
(3) Measure these materials in the appropriate apparatus or

predict their thermal conductivities based on the Lorenz

ratio method and their measured electrical resistivities.

RESULTS

Materials selected for their potential application to the ARPA

superconducting machinery program are listed below.

Iron Alloys

AISI-301*, 303*, 304*, 30AL*, 310, 316*, 321, and 347*

ARMCO 31-6-9 and 21-13-5*

CARPENTER 20cb-3

KROMARC 58

A-286*

W-545

PYROMET 860

DISCALOY

MARAGING STEELS

Nickel Alloys

Inconel 625, 718*, and X750

Hastelloy C-276

PYROMET 680 and 102

UNITEMP-HX

HAYNES 188

MULTIPHASE MP35N

INCOLOY 903

INCO (low expansion, unnamed)

INVAR

Titanium Alloys

Ti-6A1-4V (ELI)*

Ti-5Al-2.5 Sn (ELI)*

Aluminum Alloys

2014

X2048

2219

5083*

39
Copper Alloys

OFHC copper*

Those alloys marked with an * have been measured over most of the

range 4 to 300 K and would be of interest only from the standpoint of

heat treat, strain, and microstructure variation. Any of these of interest

would be measured in the fixed-point apparatus or predicted on the

basis of electrical resistivity [See also Clark et al (3)], with the exception of OFHC copper. Of the remaining alloys, the iron base alloys

generally have well-behaved thermal conductivities, as a function of temperature, and will usually be measured in the fixed point apparatus.

Aluminum alloys will be treated similarly. The thermal conductivity

curves of nickel alloys tends to be less systematic and the precision

apparatus will be used most frequently for these.

The following materials have been selected and several specimens

procured for additional study:

AISI 310S

Inconel X750

Kromarc 58

Inco Low expansion alloy (unnamed)

OFHC copper.

Most of the materials have been obtained in various heat treat and strain

conditions

.

The following measurements are scheduled for this contract year:

Precision Apparatus

OFHC copper

Inco Low Expansion Alloy (unnamed)

Inconel X750

Fixed Point Apparatus

Iron and Steel, SRM's for apparatus accuracy check

AISI 310S

Kromarc 58.

40
DISCUSSION

Measurements on OFHC copper are nearly complete and data analysis

has been initiated. No measured thermal conductivity data are availably

however, this is an excellent opportunity to illustrate the value of the

Lorenz ratio prediction method mentioned earlier. Based upon preliminary

data from the precision apparatus, the residual electrical resistivity

of this specimen of OFHC copper is 7.4 nQ -cm. Lorenz ratio values,

L =pX/T, for a copper of similar residual resistivity are taken from

Hust and Sparks (2) and listed in table 1. Electrical resistivity

values are obtained from the intrinsic resistivity of copper, the measured

residual resistivity and Mathiessen's rule. These along with calculated

values (predicted) of thermal conductivity, ^ = TL/p, are tabulated in

table 1. Upon completion of the thermal conductivity measurements on

this OFHC copper specimen, comparisons will be made to illustrate the

degree of agreement.

41
Table 1. Calculation of thermal conductivity

of OFHC copper based upon electrical

resistivity and Lorenz ratio data

2 2 +8 LT

T,K L, V /K X 10 p., nQ cm p, nQcm X

=

, W/cmK

1 D

5 2.4 0.0 7.4 16.2

10 2.0 0.0 7.4 27.0

15 1.8 0.3 7.7 35.1

20 1.5 0.8 8.2 36.6

25 1.2 2.5 9.9 30.3

30 1.2 6.3 13.7 26.3

40 1.2 22 29 16.6

50 1.2 50 57 10.5

70 1.4 153 160 6.12

80 1.5 215 222 5.40

100 1.7 350 357 4.76

150 2.00 700 707 4.24

200 2 . 16 1060 1070 4.04

250 2.26 1400 1410 4.01

300

^ 2.31 1750 1760 3.94

42
REFERENCES

1. Hust, J. G. and Clark, A. F., The Lorenz Ratio as a Tool for Predicting

the Thermal Conductivity of Metals and Alloys, Materials Research

and Standards, 11, No. 8, 22-24 (Aug. 1971).

2. Hust, J. G. and Sparks, L. L., Lorenz Ratios of Technically Important

Metals and Alloys, Nat. Bur. Stand. Tech. Note 634, 133 pp. (Feb.

1973).

3. Clark, A. F., Childs, G. E. and Wallace, G. H., Electrical Resistivity

of Some Engineering Alloys at Low Temperatures, Cryogenics 10, 295-

305, (Aug. 1970).

43
44
SEMI-ANNUAL REPORT ON MATERIALS RESEARCH

IN SUPPORT OF SUPERCONDUCTING MACmNERY

Elastic Properties of Engineering Alloys

at Cryogenic Temperatures

by

H. M. Ledbetter, E. R. Naimon, and W . F. Weston

Cryogenics Division

NBS - Institute for Basic Standards

Boulder, Colorado
SUMMARY

The problem was to determine the elastic properties of engineering

alloys at cryogenic temperatures, 4 - 300 K. These properties include:

longitudinal modulus , Young's modulus, shear modulus, bulk modulus, and

Poisson's ratio. Such data are vital to areas such as structural design,

fracture analysis, and properties of composites. Data have been acquired

for thirteen materials: Ti - 6 Al - 4V and Ti - 5 Al - 2.5 Sn; Inconels 600 and

X-750; stainless steels A286, 304, 310, and 316; aluminum alloys 1100,

5083, 7005, and 7075: and invar. The hardness, mass density, and (when

appropriate) grain size of these materials were measured. Chemical analysis

and thermal-mechanical history are also reported. Elastic properties were

determined by measuring ultrasonic velocities using pulse-echo and pulse-'

superposition techniques. In final manuscripts, the few existing data are

compared with present results. Originally, fixed-temperature measurements

were planned at 300, 195, 77, and 4K. However, an experimental system

was developed capable of semi- continuous measurements between 300 and 4K,

both cooling and warming. Recommendations for further research in the

cryogenic elastic properties of engineering alloys include: (1) similar studies

on additional materials; (2) effects of mechanical deformation and heat treatment; (3) effects of magnetic fields on ferromagnetic and antiferromagnetic

materials; (4) effects of manufacturing and processing variables; and (5)

effects of low-temperature phase transformations. The primary problems

46
in extending this work are: (a) constructing a helium cryostat for measurements in a magnetic field (presently we are limited to room-temperature

magnetic measurements); (b) developing a new measurement technique for

copper alloys, which highly attenuate ultrasonic waves in the 1-10 MHz

region - a composite oscillator technique (10 - 100 kHz) is being studied.

RESULTS

Results from this program are given below and include:

(1) A manuscript "Elastic properties of two titanium alloys at low

temperatures," by E. R. Naimon, W. F. Weston, and H. M. Ledbetter.

This manuscript has been accepted for publication in Cryogenics

,

(2) A manuscript "Dynamic low-temperature elastic properties of

two austenitic nickel-chromium-iron alloys," by E. R. Naimon, H. M.

Ledbetter, and W . F. Weston. This is to be submitted for publication.

(3) A set of ten figures showing the elastic properties of four stainless steels and four aluminum alloys.

Details concerning experimental procedures are given in items (1) and

(2) or in the references cited therein.

47
Elastic Properties of Two Titanium Alloys at Low Temperaturest

E, R. Naimon-ls W. F. Weston^ls and H. M. Ledbetter

Cryogenics Division, Institute for Basic Standards,

National Bureau of Standards , Boulder, Colorado 80302

ABSTRACT

Sound velocities and elastic constants were determined

semi -continuously for two annealed polycrystalline titanium

alloys between 4 and 3 00K. Results are given for: longitudinal

sound velocity, transverse sound velocity, Young's modulus,

shear modulus, bulk modulus, Poisson's ratio, and elastic

Debye temperature. A pulse -superposition technique was used.

Keywords: Bulk modulus; compressibility; Debye temperature;

elastic constant; Poisson's ratio; shear modulus; sound velocity;

titanium alloys; Young's modulus.

"nRC-NBS Postdoctoral Research Associate, 1973-4.

tContribution of NBS, not subject to copyright.

48
Elastic Properties of Two Titanium Alloys at Low Temperatures

E. R. Naimon'-, W. F, Weston-', and H. M, Ledbetter

Cryogenics Division, Institute for Basic Standards

National Bareau of Standards, Boulder, Colorado 8 03 02

Introduction

Despite the fact that titanium alloys Ti-6A1-4V and Ti-5Al-2. 5Sn are

intended mainly for high- strength high-temperature applications, they also

have cryogenic uses. Their important properties include: ease of fabrication,

corrosion resistcince, high strength-to-density ratios (especially at cryogenic

temperatures), and high strength-to-thermal-conductivity ratios.

The low-temperature elastic properties of these alloys are reported

here, A pulse -superposition method was used to determine the ultrasonic

wave velocities in annealed specimens. From longitudinal and transverse

wave velocities, together with the mass density, the elastic constants were

calculated: Young's modulus, shear modulus, bulk modulus (reciprocal compressibility), Poisson's ratio, and the elastic Debye temperature. These

constants are useful in a wide variety of applications, from engineering design to equations of state for solids.

Experimental Procedures

Alloys were obtained from commercial sources in the form of 3/4 -in.

(1.9-cm) diameter rods. Cylindrical specimens 5/8-in. (1,6-cm) in diameter

NRC-NBS Postdoctoral Research Associate, 1973-4,

49
and 5/8-in. (1. 6-cm) long were prepared by grinding. Opposite faces were

flat and parallel within 100 X 10 ^ in. (2, 5 m). Specimens were annealed

at a pressure of 5 X 10 ^ torr, or less, and cooled in the furnace. Annealing

times and temperatures are given in Table 1, together with chemical compositions (obtained from mill analyses), hardness numbers, microstructures, grain

sizes, and mass densities. Hardness, microstructure , and grain size were determined by standard metallurgical methods. Mass density was determined by

Archimedes' method using distilled water as a standard.

The specimen holder is shown in Fig. 1. The holder was placed in the

ullage of a helium dewar and lowered stepwise to achieve cooling. Temperatures were monitored with a chromel-constantan thermocouple contacting the

specimen.

Quartz transducers (10 MHz) were bonded to the specimens with phenyl

salicylate for room-temperature measurements and with a stopcock

grease for lower temperatures. In a few cases, failure of these bonds at very

low temperatures required using a silicone fluid (viscosity = 200,000 cP

at 2 5 °C) for bonding.

1,2 A pulse-superposition method was used to determine the sound-wave

velocities over the temperature range 4-300K. No thermal-contraction corrections were made; for titanium alloys this introduces a maximum error (over a

300 K range) of about 0.2 percent. No bond corrections were made; this error

is insignificant for the purposes of the present study. Maximum uncertainties

50
in the velocity measurements are estimated to be about one percent.

Results

Longitudinal and transverse sound-wave velocity data are shown in

Figs, 2 and 3, where the longitudinal modulus is given by

= pv^^ = B +1 G (1)

and the transverse modulus is given by

= pv^^ = G. (2)

Here v^ and v are the longitudinal and transverse sound-wave velocities,

p is the mass density, B is the bulk modulus, and G is the shear modulus.

Temperature dependences of both C, and C were fit to a semi-theo-

'L t 3

retical relationship suggested by Varshni:

C = C° - s/(e*^'^ - 1), (3)

where C° , s, and t are adjustable parameters and T is temperature. The

value of C at T = 0 K is C° , and -s/t is the high-temperature limit of the

temperature coefficient dC/dT. By invoking an Einstein oscillator model of

solids, it can be shown (in the absence of electronic effects) that t is the

T-l- . 1 . . o . . Einstein characteristic temperature. Parameters C , s, and t are given

in Table 2. Room-temperature values of the temperature coefficients of the

elastic moduli are given in Table 3,

Curves in Figs, 2 and 3 are plots of Eq, (3) determined by an

unweighted least- squares fit of the data. Average percentage differences

between measured and curve values are 0,03% and 0,07% for the longitudinal and transverse moduli, respectively.

51
While polycrystalline aggregates (quasi-isotropic solids) have onlytwo independent elastic constants, several constants are commonly used for

various applications. The four most common are the bulk modulus B,

Young's modulus E, the shear modulus G, and Pois son's ratio v • The

relationships among these are

_]_ 1 1 E ~ 3 G 9 B

and

E

These elastic constants were calculated from the moduli shown in Figs.

2 and 3 by the relationships:

= S - I

, (7)

and

V . i(C^ - 2C^)/(C^ - C^), (8)

The constants E, B, and v are shown in Figs. 4-6.

It is of interest to calculate the elastic Debye temperature 0 for the

two alloys. This fundamental parameter is important in the lattice properties

4

of solids and is related to the elastic wave velocities by

52
e = K <v>, (9)

where

k \4ttA/

Here h is Planck's constant, k is Boltzmann's constant, N is Avogadro's

constant, p is the mass density, and A is the atomic weight. The average

velocity is given by

<v) = [ 3

J

. (11)

The Debye temperatures for the two alloys at T = OK , and also for pure

titanium, are given in Table 4.

Discus sion

The elastic properties of both Ti-6A1-4V and Ti-5Al-2, 5Sn behave regularly with respect to temperature. All of the elastic moduli (C^, = ^»

B, E) decrease with increasing temperature, show a relative flatness at low

temperatures, achieve zero slope at T = 0, and approach linear behavior at

high temperatures, Poisson's ratio also behaves regularly, having a positive

temperature coefficient.

Assuming that the specimens studied are representative of the two

alloys, then conclusions concerning their relative elastic behavior can be

drawn. Not surprisingly, as shown in Figs, 2-6 and Tables 2 and 3, for most

practical purposes the two alloys are elastically identical. Thus, whether the

second alloying element is vanadium or tin, the elastic properties are essentially

the same.

53
As is well known, titanium undergoes a c.p.h. -to-b.c.c. (a-to-P)

5 crystal- structure transition when heated to 1155 K. Fisher and Renken

found a large temperature dependence of the c . . shear modulus in titanium:

66

1 ^^66 4 1 —r= =-11.93 X 10 K . A small value of c., is expected

c,, dT 66

66

from the shear mechanism of a c.p.h. -to-b.c.c. transition proposed by

Burgers^. For hexagonal symmetry, the shear modulus G is an approximate average of the two single-crystal shear constants c and c . Thus,

66 44

1 dG from the smaller values of — —— in Table 3 one can infer that both G dT

alloys have higher c.p.h. -to-b.c.c. transition temperatures than pure

titanium. That is, the cr-titanium phase is stabilized in the alloys. As

7 shown in Hansen and Anderko , aluminum is a strong stabilizer of a titanium,

while both vanadium and tin are moderate stabilizers of 3 titanium.

Despite the extensive use of these alloys, very little elastic data

exist for them. Most information has appeared in engineering reports and is

summarized in references 8-12. For comparison, the room-temperature

values of E, G, B, and v are given in Table 5, Generally good agreement

is observed between previous and present results.

The elastic moduli of both alloys are several percent higher than those

12

reported for commercicil-grade titanium . But they are lower by several

percent than the quasi-isotropic moduli obtained from a Voigt-Reus s-Hill

5

average of the data obtained from zone-refined single crystals . The elastic

Debye temperatures of the alloys are cilso about 2% lower than that calculated

54
from the single -crystal elastic data of titanium. The relatively low elastic

anisotropy of titanium does not allow for a large error in the Voigt-Reus s -Hill

averages. Thus, existing data suggest a strong impurity effect on the elastic

constants of titanium, perhaps due to interstitial impurities.

Finally, it is emphasized that the data reported here are dynamic

(adiabatic) rather than static (isothermal) and apply to rapid, rather than slow,

loading. In most cases the differences between adiabatic and isothermal elastic

13

constants are small. Conversion formulas are given in Landau and Lifshitz ,

for example. For titanium at room temperature:

(Eg - E^)/Eg = 0,001, (Bg - B^)/Bg ^ 0. 009,

(Vg - v^)/Vg = 0.005, and (Gg - G^)/Gg = 0,

where subscripts S and T denote adiabatic and isothermal, respectively.

Acknowledgment

This work was supported in part by the Advanced Research Projects

Agency.

55
References

1. McSkimin, H. J., J. Acoust. Soc. Amer. 33 , 12(1961).

2. McSkimin, H. J. and Andreatch, P., J. Acoust. Soc. Amer. 34 ,

609 (1962).

3. Varshni, Y. P., Phys. Rev. B2, 3952 (1970).

4. Debye, P., Ann. Phys. (Leipz. ) 39, 789 (1912).

5. Fisher, E. S. and Renken, C. J., Phys. Rev. 135, A482 (1964).

6. Burgers, W. G. , Physica J_, 561 (1934).

7. Hansen, M. and Anderko, K. , Constitution of Binary Alloys (McGrawHill, New York, 1958).

8. Metals Handbook, 8th Edition, Vol, 1, Properties and Selection of

Metals (Amer. Soc. Metals, Metals Park, Ohio, 1961).

9. Cryogenic Materials Data Handbook, AFML Rep. No. ML-TDR-64-280

(PB 171809, revised) (Aug. 1964).

10. NERVA Program Materials Properties Data Book, Vol. 1, Introduction

and Light Metal Alloys (Aerojet Nuclear Systems, Sacramento, Calif. , 1970).

11. Hanlein, S. L. , Hinckley, W. M. , and Stecher, F. P., NOLTR Report

70-141 (1970).

12. Fahey, N. H. , WAL Tech. Rep. No. TR 118.1/1 (I960).

13. Landau, L, D. and Lifshitz, E. M. , Theory of Elasticity (Pergamon,

London, 1959), p. 17.

56
List of Tables

1. Compositions and properties of alloys.

2. Parameters in equation 3.

3. Temperature derivatives of elastic constants at room temperature

-4 -1

(10 K ).

4. Elastic Debye temperatures at T = 0 K.

5. Comparison of present and previously reported results.

List of Figures

1. Specimen holder,

2

2. Longitudinal modulus C^ = pv^ of two titanium alloys.

2

3. Transverse or shear modulus C^ = G = pv^ of two titanium alloys.

4. Young's modulus of two titanium alloys.

5. Bulk modulus (reciprocal compressibility) of two titanium alloys,

6. Pois son's ratio of two titanium alloys.

57
OU

C

<:

o

OU

x:U

(UN

><

^2 o

g

^ n °

u

>

o

1—

I

0)

ft

U

oo

o

o

o

o

o

o

o

o

o

o

(NJ

>

I

to -cJ

2 ^ S

-

d a r-i

Oh nJ a d) CD o

rW -l-> " "

^ 4J ft D O

^ "J ft H .S C. ft n)

5s

O

o

O

o

tsj o
Table 2. Parameters in equation 3

C° s t Alloy Mode 11 2

(10 N/ra ) do" 2 N/m ) (K)

i 1 - oAi -4 V

2 1 A Q Q 0

.

r\ o c 085 213,7

2

Pv^ t 0. 467 0. 039 183, 6

Ti -5A1-2. 5Sn

2

1,716 0, 085 208, 8

2

0. 470 0. 051 210. 4

59
Table 3. Temperature derivatives of elastic constants at

room temperature (10"'^K~-'-)

1 dB 1 dE 1 dG 1 d y

B dT E dT G dT v dT

Ti-6A1-4V -1.01 -4.44 -4.90 1.89

Ti-5Al-2.5Sn -0.75 -4.99 -5.5 5 2.26

Pure Ti" -0.93 -6.63 -7.15 2.80

* Calculated from data in ref. 5 using a Voigt-Reuss -Hill average.

60
Table 4. Elastic Debye temperatures at T = 0 K

Alloy e (K)

Ti-6A1-4V 418. 1

Ti-5Al-2. 5Sn 416.9

Pure Ti 425. l''

Ref, 5

61
Table 5, Comparison of present and previously reported results; room-temperature

11 2

values; units of 10 N/m except v (dimensionle s s)

Ti-5Al-2.5Sn Ti-6A1 -4V

Source E G B V E G B V

Ref. 12 1.11 0.421 1.06 0. 325

Ref. 8 1.14

Ref. 9 0.96-1.17 - - - 1. 07-1.21 0,421

Ref. 10 1.07 - 1.07 1. 10 1, 13

Ref. 11 1.10 1.03

Pre sent 1.11 0,420 1.07 0.327 1.11 0,420 1.05 0. 3 23

62
To coaxial leads

Bakelite cylinder

i 4- n'jfh ~ 'inductor l^l'S'^^y "l''""^^'^

tuning circuit

Brass nut

Thermocouple

Brass plunger

Bakelite cylinder

Beryllium -copper spring

Quartz transducer

Specimen —Thin-wa!led stainless tubing

(electrical ground)

Brass base

Heat vane {optional)

Fig. 1 Specimnen holder.

63
o

1—

I

r-Hci

ziC

rj

o

>

CL

u

in

i-i

,01) sninaoi/'j ivNianiioNoi

64
CD CO o C\J CO o CvJ CO

CD CO IT) vr ro rO CJ o v-f- d d d d d d d d d d d

(jii^/NiiOi) smnaoiAj hvbhs

65
66
(gUj/N „oi) smnaoiN Mine

67
68
Dynamic Low-temperature Elastic Properties of

Two Austenitic Nickel-Chromium-Iron Alloys*

E. R. Naimon"^* H. M. Ledbetter, and W. F. Weston**

Cryogenics Division, Institute for Basic Standards,

National Bureau of Standards, Boulder, Colorado 80302

Abstract

The zero-magnetic-field low-temperature elastic properties of two

polycrystalline nickel-chromium-iron alloys were determined ultrasonically

between 4 and 300 K. Results are given for: longitudinal and transverse

sound velocities. Young's modulus, shear modulus, bulk modulus, Poisson's

ratio, and elastic Debye temperature. Effects of alloying are discussed.

Key words: Bulk modulus; compressibility; Debye temperature;

elastic constant; nickel-chromium-iron alloys;

Poisson's ratio; shear modulus; sound velocity; Young's

modulus

.

* Contribution of NBS, not subject to copyright.

NRC-NBS Postdoctoral Research Associate, 1973-4.

69
Introduction

Elastic properties of metals at low temperatures have twofold

interest. First, the absence of thermal vibrations permits a determination of intrinsic lattice properties; such information is essential for

understanding the basic aspects of mechanical deformation, for example.

Second, low-temperature elastic properties are essential design parameters

for cryogenic structures.

Austenitic nic*kel-base alloys containing chromium and iron are standard

engineering materials. They have high heat resistance, high corrosion resistance, good high-temperature strength, and can be readily fabricated

into structures. Many of these materials also have low-temperature applications. In particular, Inconel* 600 has exceptional cryogenic mechanical

properties. Its strength increases with decreasing temperatures, and

its ductility and toughness are preserved. Inconel X-750 (formerly

Inconel X) is similar to Inconel 600, but it contains small amounts of

aluminum, titanium, and niobium. Aluminum and titanium make precipitation

hardening possible by forming Ni^ (Al, Ti) by suitable thermal treatment.

(Inconel 600 is usually hardened by low-temperature mechanical deformation.)

Niobium further stiffens the matrix and stabilizes the carbides.

The zero-magnetic-field dynamic low-temperature elastic properties

of Inconel 600 and Inconel X-750 are reported here. A pulse-superposition

method was used to determine ultrasonic wave velocities in specimens prepared from as-received commercial bar stock.

* Tradenames are used to characterize materials; they are not NBS

endorsements of particular products.

70
Experimental

Inconel 600 and X-750 alloys were obtained from commercial sources

in the form of about 3-in (7. 6 -cm) and 4-in (10.2-cm) diameter rods,

respectively. Specimens 3/8-in (1-cm) long were prepared by grinding.

—6

Opposite faces were flat and parallel within 100 x 10 in. (2.5 t-im).

Chemical compositions (obtained from mill analyses), hardness numbers,

microstructures, and mass densities are given in Table 1. Hardnesses,

microstructures, and grain sizes were determined by standard metallurgical

methods. Mass density was determined by Archimedes' method using distilled

water as a standard.

1 2 A pulse-superposition method ' was used to determine all ultrasonic

velocities except the shear mode of Inconel X-750. This particular mode

was highly attenuated, and for its measurement the standard pulse-echo

technique was used.

Quartz transducers (10 mHz) were bonded to the specimens with phenyl

salicylate for room-temperature measurements and with a stopcock grease

for lower temperatures. Occasional failure of the bond at very low temperatures resulted in using a silicone fluid (viscosity = 200,000 cP at

25°C) for bonding.

The specimen holder is shown schematically in Fig. 1. The holder

was placed in the ullage of a helium dewar and lowered stepwise to achieve

cooling. Temperatures were monitored with a chromel-constantan thermocouple

contacting the specimen.

No thermal contraction corrections were made for the sound velocities

or for the elastic constants; for nickel alloys this introduces a maximum

error in the velocity of about 0.2 percent. Bond corrections were not

71
made; this error is insignificant for the purposes of the present study.

Maximum uncertainties in the velocity measurements are estimated to be

about one percent.

For Inconel X-750, velocity measurements were made over the range

40-300 K. Below 40 K, the attenuation was too high for both longitudinal

and transverse modes. However, little change is expected in the velocities

from 40-0 K and theoretical curves fit to the data points (as explained

in the next section) should give accurate low-temperature values . For

Inconel 600, velocity measurements were made from 4-300 K.

Results

Longitudinal and transverse modulus are shown in Figs. 2 and 3.

The longitudinal modulus is given by

and the transverse modulus is given by

r

C^ = pv^ = G. (2)

Here v and v are the longitudinal and transverse sound-wave velocities,

"t/ t

P is the mass density, B is the bulk modulus, and G is the shear modulus.

Temperature dependences of both C and C were fit to a semi-theoretical

-L t

3 relationship suggested by Varshni :

C = C° - s/(e^/^ - 1), (3)

where C*^, s, and t are adjustable parameters and T is temperature. The

value of C at T = 0 K is C°, and -s/t is the high-temperature limit of

dC/dT. By invoking an Einstein oscillator model of solids, it can be

72
shovm (in the absence of electronic effects) that t is the Einstein

characteristic temperature. Parameters C°, s, and t for Inconels 600 and

X-750 are given in Table 2. Room- temperature values of the temperature

coefficients of the elastic moduli are given in Table 3.

Curves in Figs. 2 and 3 are plots of eqn. (3) determined by an

unweighted least-squares fit of the data. For Inconel 600, average percentage differences between measured and curve values are 0.05% for both

the longitudinal and transverse moduli. For Inconel X-750, the average

percentage differences between measured and curve values are 0.04% and

0.18% for the longitudinal and transverse moduli, respectively. The comparatively large error for the transverse modulus of Inconel X-750 was due

to a relatively poor echo pattern.

While polycrystalline aggregates (quasi-isotropic solids) have only

two independent elastic constants, several constants are commonly used

for various applications. The four most common are the bulk modulus B,

Young's modulus E, the shear modulus G, and Poisson's ratio v . The relationships among these are:

+ (4) E 3G 9B

and

V =^ - 1. (5)

These elastic constants were calculated from the moduli shown in Figs.

2 and 3 by the relationships:

B = - fC^, (7)

73
and

V = I

(C^ - 2C^)/(C^ - C^). (8)

The constants E, B, and v are shown in Figs. 4-6.

It is of interest to calculate the elastic Debye temperature 9 for

the two alloys. This fundamental parameter is important in the lattice

4 properties of solids and is related to the elastic wave velocities by

e=K<v), - (9)

where

1/3

K 4nA

Here h is Planck's constant, k is Boltzmann's constant, N is Avogadro's

constant, p is the mass density, and A is the atomic weight. The average

velocity is given by

-3 -3 -1/3

V + Z V

(v)= (-^-^ — ) . (11)

The Debye temperature of each alloy at T = OK, and also of nickel, is

given in Table 4.

Discussion

The elastic properties of both Inconel 600 and Inconel X-750 behave

regularly with respect to temperature. The elastic moduli (C , C = G,

'L t

B, E) decrease with increasing temperature, show a relative flatness at

low temperatures, achieve zero slope at T = 0 K, and approach linear behavior at high temperatures. Poisson's ratio also behaves regularly,

having a positive temperature coefficient.

74
Assuming that the specimens studied are representative of the two

alloys, then conclusions concerning their relative elastic behavior can

be drawn. Not surprisingly, as shown in Figs. 2-6 and in Tables 2 and 3,

for most practical purposes the two alloys are elastically identical. Thus,

additions of small amounts of aluminum, titanium, and niobium have little

effect on the elastic properties of nickel alloys containing about 15

percent chromium and about 7 percent iron.

Very little elastic data exist for these alloys^ especially

Inconel 600. Most information has appeared in engineering reports and is

summarized in refs. 7-11. For comparison the room-temperature values of

E, G, B, and v are given in Table 5. Generally good agreement is observed

between previous and present results.

The elastic moduli of both alloys range from several percent lower

12

(E, B) than polycrystalline nickel to about the same (G, v ) as polycrystalline nickel. This also holds if the elastic moduli of the two

alloys are compared with the quasi-isotropic moduli obtained from a 13 Voigt-Reuss-Hill average of the single-crystal data , although there is

14 some discrepancy for the value of G obtained by this method . The elastic

Debye temperatures of the alloys are also within 2% of the value calculated

from the single-crystal elastic data of nickel.

It is emphasized that the data reported here are dynamic (adiabatic)

rather than static (isothermal) and apply to rapid, rather than slow,

loading. In most cases the differences between adiabatic and isothermal

elastic constants are small. Conversion formulas are given in Landau

and Lif shitz'''^ , for example. For nickel at room temperature:

(E sis - E)/E = 0.003, (B sis B_)/B = 0.022,

(v sis -v^)/v =0.012, and (G sis - G^)/G = 0, (12)

where subscripts S and T denote adiabatic and isothermal, respectively.

75
Acknowledgment

This work was supported in part by the Advanced Research Projects

Agency.

76
References

1. H. J. McSkimin, J. Acoust. Soc. Amer. , 33 (1961) 12.

2. H. J. McSkimin, and P. Andreatch, J. Acoust. Soc. Amer. , 34, (1962) 609.

3. Y. P. Varshni, Phys . Rev

.

, B2 (1970) 3952.

4. P. Debye, Ann. Phys. (Leipz.) , 39 (1912) 789.

5. H. C. Burnett, unpublished results (1956).

6. W. H. Hill, K. 0. Shimmin, and B. A. Wilcox, Pro-! . ASTM , 61 (1961) 890.

7. International Nickel Co., Report on Inconel All' 600 (1973).

8. International Nickel Co., Report on Inconel Alio X-750 (1970).

9. NERVA Program Materials Properties Data Book, Vol. lA, Nickel Base

Alloys (Aerojet Nuclear Systems, Sacramento, Calif., 1970).

10. Cryogenic Materials Data Handbook, AFML Rep. No. ML-TDR-64-280

(PB 171809, revised) (Aug. 1964).

11. Aerospace Structural Metals Handbook, Vol. II, AFML Rep. No.

ASD-TDR-63-741 (Mar. 1966).

12. H. M. Ledbetter, and R. P. Reed, J. Phys. Chem. Ref. Data , Vol. 2,

No. 3, (1973).

13. S. G. Epstein, and 0. N. Carlson, Acta Met. , 13 (1965) 487.

14. G. Simmons, and H. Wang, Single Crystal Elastic Constants and Calculated

Aggregate Properties: A Handbook , (M.I.T. Press, Cambridge, 1971)

pp. 223-227.

15. L. D. Landau, and E. M. Lifshitz, Theory of Elasticity (Pergamon,

London, 1959) p. 17.

16. M. Dixon, F. E. Hoare, T. M. Holden, and D. E. Moody, Proc. Roy. Soc.

285 (1965) 561.

77
List of Tables

1. Compositions and properties of alloys.

2. Parameters in equation 3.

-4 -1

3. Temperature derivatives of elastic constants at room temperature (10 K ).

4. Elastic Debye temperatures at T = 0 K.

5. Room-temperature elastic constants of Inconel 600 and Inconel X-750,

11 2 units of 10 N/m except V (dimensionless)

.

List of Figures

1. Specimen holder.

2 2. Longitudinal modulus C = pv of two nickel-chromium-iron alloys.

L 't 2 3 . Transverse or shear modulus = p v^ = G of two nickel-chromium-iron

alloys.

4. Young's modulus of two nickel-chromium-iron alloys.

5. Bulk modulus (reciprocal compressibility) of two nickel-chromium-iron

alloys.

6. Poisson's ratio of two nickel-chromium-iron alloys.

78
13 0 O to

cd o n3

*r 5i •"in c

c

QJ c QJ vD c > CO > .H o •H •rl fcil o (U X) (1) 13 O QJ O (U 00

QJ .-1 (1) rH m >-( .-( M t-H .H O O .H

CO M m tj --^ <:

T3

nJo

iH

(0

(0 A!

01C H

•oM

0) O

!3Q

o

<:

a o

o o

c

79
Table 2. Parameters in Equation 3

Alloy Mode

(lO-^'-'-N/m^)

C°

(10-^-^N/m^)

St

(K)

2 Inconel 600 pv 2.909 0.107 203.1

pv^ 0.828 0.063 246.0

Inconel X-750 pv^ 2.920 0.105 165.9

Pv^ 0.846 0.120 315.8

1

80
Table 3. Logarithmic temperature derivatives of elastic

-4 -1

constants at room temperature (10 K )

Material i^B l^E l^G 1 dv

B dT E dT G dT V dT

Inconel 600 -1.07 -2.82 -3.08 1.10

Inconel X-750 -0.89 -3.94 -4.40 1.97

Nickel* -1.56 -4.16 -4.53 1.72

* Ref. 12

81
Table 4. Elastic Debye temperatures at T = 0 K

Material 9 (K)

Inconel 600 464.9

Inconel X-750 473.5

Nickel* 476.2

* Ref. 16

9x
Table 5. Room-temperature elastic constants of Inconel 600 and Inconel X-750;

11 2 units of 10 N/m except v (dimensionless)

Source Inconel 600 Inconel X-750

E G B V E G B V

Present 2.036 0.780 1.760 0.307 2.040 0.784 1 .736 0.304

Ref. 5 0.789

Ref. 6 - - - 2.144 - - -

Ref. 7-8 2.11 0.75 2.11 0.75 — 0.29

Ref. 9 ; 2.075 0. 714 n ?Q

Ref. 10 2.041 0.766-

0.813

Ref. 11 2.109 2.109 0.749

83
To coaxial leads

Bokalita cylinder

2-6 inductor

4-1! jsh inductor Ipigidly mounted

tuning circuit

Brass nut

Thermocouple

Brass plunger

Bokelite cylinder

Beryllium -copper spring

Quartz transducer

Specimen

Thin-v/alled stainless tubing

(electrical ground)

Brass base

Heat vcne (optionc!)

S I)

•"' c I'l o n. n o 1. cl c L"

,

84
03

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en ID <n <—

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85

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(^lu/N-j^oi) sninaow s.sNnoA

87
88
T

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I—

1

f—

1

o o o o o en cri cr> ro ro m CO CO CO CM CM CM CM O O o o o o o CD o o O

oim s.NOssiod

89
Elastic Properties of Four Stainless Steel Alloys

Elastic Properties of Four Aluminum Alloys

The following ten figures are unaccompanied by text and give

the elastic constants

Longitudinal modulus

Shear modulus

Young's modulus

Bulk Modulus

Poisson's ratio

for

and

Stainless steels 304, 310, 316, A286

Aluminum alloys 1100, 5083, 7005, and 7006.

The preceding two manuscripts for titanium alloys and nickel - chromiumiron alloys give experimental procedures, etc. that largely apply to the

stainless steels and the aluminum alloys.

90
91
oocococooocacooooooooo LO CM cTi ro'O «^ t-i co in cnooooooo o\ 00 00 00 '00 oooo«x> vo

(^uj/N^^oT) sninaow «v3hs

92
T

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1

o cn CO CO CM T-H O o 00 00

• • • • • csi CVJ CNJ V C\j CM f—

1

I—

1

I—

1

1—

(^uj/N^^OTj snmaow s.SNnoA

93
oom

o

in

ooCVJ LU

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o

o

in

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00

00 00

00

00

CM U3 o

in

10

o

o

in

in

CM

I—

I

in

in

oCM

94
Oim S.NOSSIOd

T

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• • • • • • • • • o o o ' o o o O O o o

(^ui/N^^OT) smnaow W3hs

97
r

/ o

/ CD

/ 1^

/ O

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o

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ooCM

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r>. CD O CD CD CD CD CD o

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98

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ro

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CM

o ^

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100
SEMI-ANNUAL REPORT ON MATERIALS RESEARCH

IN SUPPORT OF SUPERCONDUCTING MACHINERY

Advanced Composites

by

M. B. Kasen

Cryogenics Division

NBS - Institute for Basic Standards

Boulder, Colorado
SUMMARY

A comprehensive review of the literature on the mechanical and

thermal properties of glas s -filament reinforced structural composites at

cryogenic temperatures has been completed. This is Part I of a two-part

series, Part II reviewing the literature on advanced composites. The latter

review is presently 75% complete.

The review includes tensile, flexural and compressive strengths

and moduli, interlaminar shear, ultimate tensile strain, static and dynamic

fatigue, fracture properties, thermal contraction, thermal conductivity and

specific heat. Uniaxial, crossply and cloth-reinforced epoxies, polyurethanes,

phenolics, polyimides, polyesters, silicones, phenyl silanes, teflons and

Imidites are included in the review.

The literature indicates that glass -reinforced composites are most

useful in cryogenic applications requiring high strength combined with high

toughness and low thermal conductivity, but where stiffness is not a major

requirement and where cyclic fatigue is not a major problem. There are

indications that static fatigue may not be a problem at cryogenic temperatures; however, both static and dynamic fatigue must be studied more

thoroughly before glas s -reinforced composites are utilized in critical areas

of superconducting machinery.

No literature data were available on the thermal fatigue resistance

of glas s - reinforced composites within the cryogenic range. As elevated

temperature thermal fatigue is known to seriously degrade composite

properties, this parameter should also be investigated for those composites

most likely to be used in superconducting machinery, particularly for any

component likely to be subjected to thermal cycling.

At the present time, the composite most widely used for cryogenic

applications is S-901 glass in an epoxy matrix composed of Epon 828/DSA/

Empol 1040/BDMA in proportions 100/115. 9/20/1 by weight. This resin

probably represents the current state-of-the-art for cryogenic applications

in glas s -reinforced systems.

102
The literature data suggests that, while epoxy matrices probably

yield superior overall cryogenic properties, the glass -polybenzimidazole

(PBI or Imidite) composites appear to have almost as good static strength

properties with superior resistance to cyclic fatigue. The latter type of

composite should be more thoroughly investigated.

Recent studies have indicated that the mechanical properties of

weakly cross -linked polymers at cryogenic temperature are altered by the

activity of the gas or liquid in contact with the specimen surface. The

possibility therefore exists that data obtained with polymer- matrix

composites at 77K in liquid nitrogen might differ from data obtained at 77K

in helium vapor. This phenomenon should be further investigated.

Construction of a tensile cryostat capable of testing composite

specimens at temperatures from 295K to 4K has been completed. Initial

tests shall be run on state-of-the-art commercial boron-epoxy, boronaluminum and graphite -epoxy materials. All but the boron-aluminum

materials have been received. The testing program is scheduled to begin

upon completion of the literature survey (approximately May 1).

103
TABLE OF CONTENTS

Page

SUMMARY 102

LIST OF FIGURES 106

INTRODUCTION 107

MECHANICAL AND THERMAL PROPERTIES OF FILAMENTARYREINFORCED STRUCTURAL COMPOSITES AT CRYOGENIC

TEMPERATURES- I: Glass -Reinforced Composites 112

SCOPE OF THE LITERATURE SURVEY 112

ORGANIZATION OF THE BIBLIOGRAPHY , , 112

GLASS-REINFORCED COMPOSITES AT CRYOGENIC

TEMPERATURES 113

STATIC MECHANICAL PROPERTIES 114

Tensile Strength and Modulus 114

Flexural Strength and Modulus 121

Compressive Strength and Modulus 127

Interlaminar Shear Strength 132

Ultimate Tensile Strain 135

Static Fatigue 137

Bearing Yield Strength 138

DYNAMIC MECHANICAL PROPERTIES 140

Cyclic Fatigue 140

Impact Strength and Fracture Toughness 144

THERMAL PROPERTIES 145

Thermal Contraction 145

Specific Heat 149

Thermal Conductivity 151

COMMENTS ON GLASS-REINFORCED COMPOSITES AT

CRYOGENIC TEMPERATURES .............. 155

BIBLIOGRAPHY 158

BIBLIOGRAPHIC CROSS-REFERENCE 174

104
TABLE OF CONTENTS (cont)

Page

COMPOSITE TESTING PROGRAM 176

PROCEDURES 176

CRYOSTAT CONSTRUCTION 177

SELECTION OF MATERIALS 179

105
LIST OF FIGURES

Page

Figure 1, Ultimate Tensile Strengths of Glas s -Reinforced

Composites 115

Figure 2. Initial Tensile Modulus of Glas s -Reinforced

Composites 119

Figure 3, Ultimate Flexural Strength of Glas s -Reinforced

Composites ' 122

Figure 4. Initial Flexural Modulus of Glas s -Reinforced

Composites 125

Figure 5. Histogram Illustrating Changes in Tensile and

Flexural Properties of Glass -Epoxy Composites

During Cooling 126

Figure 6, Ultimate Compressive Strength of Glass -Reinforced

Composites 128

Figure 7, Compressive Modulus of Glas s -Reinforced

Composites 131

Figure 8. Interlaminar Shear Strength of Glass -Epoxy

Composites 133

Figure 9, Ultimate Tensile Strain of Glas s -Epoxy Composites 136

Figure 10. Bearing Yield Strength of Glas s -Reinforced

Composites 139

Figure 11. Cyclic Tensile Fatigue Strength of Glas s -Reinforced

Composites After 10^ Cycles 142

Figure 12, Thermal Contraction of Glas s -Reinforced Composites 146

Figure 13. Thermal Contraction of Glass-Cloth Reinforced

Composites 147

Figure 14, Specific Heat of Glas s -Reinforced Composites 150

Figure 15. Thermal Conductivity of Glass-Fiber Reinforced

Composites 152

Figure 16. Thermal Conductivity of Glass-Cloth Reinforced

Composites 153

Figure 17, Schematic Cross -Section of Tensile Cryostat 178

Figure 18. Installed Tensile Cryostat Without Dewars 180

Figure 19. Tensile Cryostat With Dewars Installed 180

106
INTRODUCTION

The primary impetus for structural composite development has

arisen from the need to obtain improved long-term mechanical properties

at elevated temperatures or to reduce the cost of structures designed for

ambient temperature use. Comparatively little effort has been expended on

development of composites for use at cryogenic temperatures, A notable

exception has been the rather extensive body of work sponsored by NASA

during the Apollo Program wherein a series of glas s -reinforced plastics

were characterized to 20K. The other major field of cryogenic development

has been concerned with composite reinforcement of pressure vessels for

aerospace use, largely exploiting the continuous -filament method of fabrication. To a large extent, the remaining published data on composite

properties at cryogenic temperatures reflects work on which the generation

of cryogenic property data was peripheral to the main work objective.

. This relative lack of emphasis on cryogenic structural composites is

perhaps understandable, as the majority of such structural applications are

presently satisfied by readily available and well-characterized metals and

alloys. In view of the extensive data base available on metals, it is probable

that metals will continue to constitute the main body of structural materials

at low temperatures.

Why, then, should one consider composites? The answer lies in the

increasingly stringent demands made on materials in advancing cryogenic

technology, of which superconducting machinery may serve as an example.

Undoubtedly, the first generation of superconducting motors and generators

will be dependent almost entirely on metals technology. However, it is

highly probable that succeeding generations of such equipment will capitalize

on advanced composite technology for reasons of increased reliability,

reduced weight and increased efficiency, reflecting the higher specific

strengths and moduli of advanced composites coupled with a wider range of

thermal and electrical properties than are obtainable with any conventional

metal.

107
The technological problems associated with integration of connposites

into superconducting machinery are threefold: a) most designers lack a

feel for the properties available with composites, b) an adequate data base

does not exist for composites at cryogenic temperatures, particularly at

4K, and c) most existing composites are optimized for service at room

temperature and above--not for cryogenic service. The current program at

NBS has taken aim at these three problem areas.

Our first effort has been to initiate a comprehensive review of what is

known about the mechanical and thermal properties of composites at cryogenic temperatures. The objectives of the review are threefold: a) we

wish to provide the designer with a feel for the general magnitude of property

values which may reasonably be expected from a given category* and class of

composites within the cryogenic range, b) we wish to provide him with a

feel for the ranking of specific composite classes with regard to a specific

property, c) we wish to impart a feel for whether the property of interest is

likely to increase, remain unaffected, or to decrease with lowering of

temperature, and d) we wish to define those areas in which further data are

needed and to define the direction that future work should take in optimizing

composites for cryogenic service and for implimenting their use in the

construction of superconducting machinery.

The literature review covers I960 to the present time, as it is within

this time span that almost all of the significant work was undertaken. We

include only filamentary-reinforced structural composites, and among these,

only such composites as are amenable to a general characterization. The

review, therefore, excludes cryogenic insulations, superconductor composites,

*We define a composite category by the general reinforcement type, e, g, , glass -fiber or advanced fiber (graphite, boron, etc. ), We define a composite

clas

s

by the general matrix type, e. g, , glas s -polyester or graphite -epoxy.

We define a composite type by a specific reinforcement/ matrix combination,

e. g. , S-901 glass/E787 epoxy or HT-S graphite/X-904 epoxy.

108
thin films, honeycomb structures and composite -overwrapped metal. Filled

composites (as distinct from filamentary reinforcement) are also excluded.

There remains the very large field of composites reinforced by a variety of

fibers in a variety of layups in a variety of matrix materials, and it is with

this body of data that this review is concerned.

The wide variety of composite formulations and layups are further

complicated by lack of standard test procedures. Furthermore, as the field

is relatively new, much of the earlier work was performed on relatively

poorly characterized composites. We have attempted to cope with this

complexity by dividing the review into two major sections: the first treating

glass -reinforced composites and the second treating the so-called advanced

composites. The rationale for this separation is the distinctly different use

of these two composite categories in engineering practice, i, e. , glassreinforced composites are used in applications where stiffness is not a design limitation, while the advanced composites are used where a high

modulus material is essential.

Within each of these categories, we present the reader with a series

of graphs on which appear the average literature values of each property for

each composite class from room temperature into the deep cryogenic range.

Admittedly, presenting average data is in danger of being misleading, as each curve has associated with it a considerable scatter band. For this

reason, we discuss the range of values associated with each curve, emphasizing those specific composite types for which the highest values were

reported. It is of utmost importance, however, that the reader understand

that the graphed data appearing in this review are class averages and are not

to be used for engineering purposes.

The second basic objective of the present NBS program is to initiate a

test program to establish a viable data base for advanced composites potentially useful in superconducting machinery. This program will emphasize

advanced composites, as glass -reinforced composites have been rather

widely used for cryogenic applications over the past decade, primarily for

support structures where the low modulus may be tolerated. Such composites

109
have also found applications in fiber -overwrapped cryogenic pressure

vessels. As a result, the technological base for their application is

comparatively well established. Conversely, the cryogenic application of

advanced composite technology is relatively new with the potential for

exploitation of their unique properties hardly investigated. In terms of the

overall scope of the present ARPA program, the present NBS effort interfaces between the basic composite screening studies being undertaken by

the General Electric Company and the eventual development of design

allowable data required by the design engineer. The screening studies

utilize the simple and inexpensive flexural strength test which is valuable

in establishing a relative ranking among a large number of experimental

composite formulations. However, flexural tests do not generate data which

can be directly applied to design allowables, primarily because the state of

strain is continually changing throughout the test. The tensile test is more

difficult and time consuming; however, it is the least complex test which

yields data of use to the design engineer. Furthermore, the tensile test

may be used to evaluate the properties of a uniaxial lamella, the "fundamental

building block" of composite laminates. Modern composite theory predicts

that, having valid data on the ultimate values in tension, compression and

shear for such lamella along with the pertinent poissons ratios, the mechanical properties obtainable in the complex crossply layups required in

engineering structures may be calculated.

We have chosen to use the tensile test as our primary experimental

method, working with unidirectional lamella for the above reasons. This

decision was arrived at after extensive survey and consultation in person

with a large number of individuals prominent in the composite field. Furthermore, this approach allows us to capitalize on the extensive experience of

our laboratory in conducting tensile tests at cryogenic temperatures.

Tensile testing of composite materials is not as easy as with metals,

even at room temperature. An extensive survey was also made of those

individuals who had worked in this area in order to determine what, if any.

110
problems remain to be solved. It does appear that some development

work will be necessary. The basic problem stems from the extreme anisotropy of composites and of the unidirectional composite in particular. The

very high strength of the fibers coupled with a relatively low interlaminar

shear strength in the composite frequently results in invalid test data due to

specimen failure in shear within the grip area, A relatively high fraction of

such failures may be tolerated during room temperature testing; however,

the high cost of testing at cryogenic temperatures dictates that every effort

must be made to obtain valid data from each specimen under test. Our

laboratory can make a significant contribution in this area. Tensile testing of composites requires the use of a much longer

specimen than is normally used for metals; consequently, it has been

necessary to construct a tensile cryostat capable of taking specimens up to

11 inches in length. The expense of this unit is being equally shared with

another project.

The materials selected for our initial test program represent a selection of the current state-of-the-art commercial production of boronepoxy, graphite -epoxy, and boron-aluminum composites. The reason for

this approach is twofold. First, cost is reduced by working with materials

donated by the manufacturer who finds it an advantage to have his materials

evaluated for potential cryogenic use. Secondly, the data obtained in this

initial phase of the program will provide a reference against which experimental composite formulations optimized for cryogenic service may be

compared.

The present report contains that portion of the literature review

concerned with glass -reinforced composites and that portion of the experimental program concerned with construction of the tensile cryostat and the

acquisition of the test materials.

Ill
The following symbol nomenclature is used in this report:

tu

CT - tensile ultimate strength EC - compressive modulus

E*

1 - initial tensile modulus cysi - interlaminar shear strength

Et

2

- secondary tensile modulus aby - bearing yield strength

tensile ultimate strain - im.pact strength

tu

n tensile fatigue - therm.al conductivity

^fu flexural ultimate strength AL/L - thermal contraction

^1 - initial flexural modulus - specific heat

- secondary flexural modulus

„cu - compressive ultimate strength

MECHANICAL AND THERMAL PROPERTIES OF FILAMENTARYREINFORCED STRUCTURAL COMPOSITES AT CRYOGENIC

TEMPERATURES- I: Glas s -Reinforced Composites

SCOPE OF THE LITERATURE SURVEY

We initially conducted a subject search using the data bases of the

NBS Cryogenic Data Center, DDC and NTIS, Additionally, a subject search

was conducted through the volumes of NASA STAR, and the ASM - AIME

Metals Review. As the search progressed, a series of contract numbers

were identified as being associated with studies of the cryogenic properties

of composites. The DDC and NASA data bases were then searched for all

reports issued under such contracts. Finally, the DDC data base and that of

the Smithsonian Science Information Exchange were searched for current

work in progress.

ORGANIZATION OF THE BIBLIOGRAPHY

The appended Bibliography contains 148 references. As the work

progressed, it became apparent that a large part of the relevant data had

been produced under a relatively few contracts sponsored either by NASA or USAF, Eleven such contracts have been listed at the start of the Bibliography

with references to the most pertinent publications issued under each contract.

112
Only final reports are listed, as they adequately summarize the data which

also appeared in numerous interim reports on each project. Journal

publications are listed as they often provide a convenient review of the

subject matter contained in the comprehensive reports and are generally

more readily available to the reader.

The Bibliography also contains a general section, alphabetically

arranged by author, listing relevant publications sponsored by other

contracts or by corporate in-house funding, A separate part of the Bibliography itemizes handbooks or reviews which will be found useful references

but which do not contain original data. Finally, a miscellaneous reference

section lists publications which are referenced in the text which do not

contain relevant mechanical or thermal property data.

Wherever possible, the pertinent NASA or DDC code number is

included to facilitate retrieval of specific publications. Corporate reports

not so identified must be obtained from the corporate source.

An extensive cross-reference relating mechanical and physical

properties of specific types of composites to specific literature references is

included so as to simplify literature retrieval by the reader. The latter

includes separate references to filament-wound pressure vessels in recognition of the importance of such applications to cryogenic technology, A

separate listing is also provided of reports containing information on the

effect of combined cryogenic temperature and nuclear radiation,

GLASS -REINFORCED COMPOSITES

The mechanical and thermal properties of glas s -polymeric composites

are summarized on Figures 1-16, Where available, data are presented for

295K, 200K, 77K, and 20K {4K data are almost nonexistent). Straight lines

connect average values at each temperature. Absence of a data point for a given temperature implies no significant data. An asterisk adjacent to the

number identifying a curve indicates that the data for that particular composite

type was minimal relative to that available for the other composite types

included on the Figure.

113
In considering the mechanical property data, the reader should be

aware that there exists no universally accepted method of determining these

properties for composites, although committees of the ASTM are working

diligently on the problem of standardization. Furthermore, testing of

composites, particularly at cryogenic temperatures, introduces numerous

complexities. The problems of tensile testing are primarily those of

obtaining fractures within gage lengths in uniaxial longitudinal layups and

dealing with misalignment in off-axis layups. The problems in compression

testing are column buckling in the high-aspect-ratio specimens normally

used for composite testing. The data discussed in the present review were

for the most part obtained in the course of comprehensive research programs

by reliable investigators who were concerned with obtaining the most valid

results possible. Nevertheless, it remains a possibility that some of the

scatter in the data reported in the literature, particularly for compression

and interlaminar shear, reflects differing test procedures. Where this has

become apparent, the data have been separated by test method,

STATIC MECHANICAL PROPERTIES

Tensile Strength and Modulus - Figures 1 and 2

The reason for the widespread use of glas s -reinforced composites

is evident from the tensile strength data presented on Figure 1. No other

type of composite can match the uniaxial tensile strengths of 250-300 KSI

provided by the glass -epoxy formulations. Even in the 0°/90° crossply

layup, the glass -epoxy strength is almost equal to that of the advanced

composites in the uniaxial longitudinal configuration. Unfortunately, the

moduli of glas s -reinforced composites are quite low, as may be seen in

Figure 2, It is this low modulus of glass that has given impetus for

development of the advanced composites.

From Figure 1 we see that the tensile strength is reduced about 50%

at all temperatures in a 0° / 90° configuration as compared to uniaxial, e. g. , from about 300 KSI to about 150 KSI for glass -epoxies. This is expected

from a 50% decrease in longitudinal fiber content, the crossply fibers contributing nothing to the overall strength when tested parallel to one fiber

direction.

114
Minimal data

0 50 100 150 200 250 300

TEMPERATURE, K

Fig. 1, Ultimate Tensile Strength of Glas s -Reinforced Composites.

(1) Epoxy (4) Polyimide (7) Phenyl Silane

(2) Polyurethane (5) Polyester (8) Teflon

(3) Phenolic (6) Silicone (9) Imidite

115
Approximately another 25% decrease in strength is found upon going to

woven cloth reinforcement of an epoxy matrix, reflecting the decrease in load

carrying capacity of fibers which are slightly bent in the weaving process.

Epoxy resins are most widely used as matrices for structural

applications where maximum strength is required. This appears justified on

the basis of the data of Figure 1 wherein the ultimate tensile strengths of the

cloth-epoxy composites are overall higher than those of other cloth-polymeric

composites at all temperatures. The polyurethane-, teflon-, phenyl silane-,

polyimide-, and silicone-matrix composites appear to have the poorest

strength properties at 295K, with phenolic-, Imidite--'', and polyester-matrix

composites being intermediate in strength. At 77K, the polyimide-, silicone-,

and phenyl silane-matrix composites continue their relatively poor performance, while the polyurethane-, Imidite-, and teflon-matrix composites

have about equalled the phenolics and polyesters.

Among the uniaxial glass -epoxy composites, the highest reported

strengths were 360 KSI^^^^ and 375 KSI^'^'^^ during NOL ring test of S-HTS/

660 FW and S-901/ERL 2256 glass composites. This work was directed

toward filament -wound pressure vessels, which has been the incentive for

much of the cryogenic composite development work. Other authors^^' have

also reported strengths of about 330 KSI at 77K for S-901/E-787 and S-HTS/

Epon 828 filament -wound uniaxial composites respectively. The overall data

range was 213-375 KSI.

'i' The use in this paper of tradenames of specific products is essential to the

proper understanding of the work presented. Their use in no way implies

approval, endorsement, or recommendations by NBS. Imidite is the Narmco

Division, Whittaker Corporation tradename for a polybenzmidazole resin of the

polyaromatic family. The generic name rarely appears in the literature, although the abbreviation PBI is sometimes used. The present review follows the

literature in using the tradename. Polyimide resins are of the same family,

sometimes referred to as PI. In the latter case, the generic name is used in

this report, again following the literature convention.

For present purposes, tensile data obtained from Naval Ordnance Laboratory

ring tests has been combined with data obtained from testing of flat tensile

specimens in the uniaxial longitudinal direction, even though it is recognized

that the NOL test is not strictly a uniaxial test.

116
Among the 0°/90° data, the best results have been reported for

S-901 /E -787^^^ and S-901 /modified Epon 828^^^ with tensile strengths ranging

from 170-200 KSI. The overall range was 115-200 KSI.

Among the woven cloth composites, the highest strength epoxy-matrix

(2) (3) composites were 181/modified Epon 828 and 1581/E-787 having strengths

on the order of 125-145 KSI at 77K tested parallel to the woof or warp. The

phenolic -matrix composite data shows quite consistent strengths of 60-70 KSI

(71)

at 77K except for the work of Levin, et al. , who has reported 100 KSI

at 77K in a composite based on a butvar -phenolic adhesive. The polyestermatrix data scatters from 50-80 KSI at 77K except for one report of

(2) 100-105 KSI for 181 glass in Hetron 31 or Narmco 527 resin. ^ The teflonmatrix composites ranged from 50-80 KSI at 77K, the highest value being

reported for type 116 glass in TFE or FEP. ^^^^ The silicone -matrix data

showed relatively large scatter from 25-70 KSI at 77K with the highest

values being reported for 181/Trevarno F-131^^^ and for 181/Narmco 513.

Few data were found for polyurethane-, phenyl silane- and polyimide -matrix

composites. About 73 KSI was reported at 77K for 181 glass reinforced with

the flexible polyurethane Adiprene L-lOO and 60 KSI for the same reinforce-

(2) ment in the phenyl silane Narmco 5 34. Polyimides are relatively new

matrix materials, having been developed primarily for elevated-temperature

use, particularly for stability. The one reference reports a value of 43 KSI

at 77K for a glas s -polyimide composite, a value which is hardly impressive,

On the other hand, the Imidite -matrix composites which are of the same

family as the polyimides, developed excellent strength at 77K and at 20K,

being second only to the epoxies.

All of the above comparisons were made at 77K because, as seen on

Figure 1, while the tensile strength in all cases increases between 295K and

77K, further cooling to 20K produces erratic results. This is more clearly

illustrated in Figure 5(a) which shows that cooling of glass-epoxy composites

from 295K to 77K can produce a strength increase of from 10-140 KSI with

117
a high probability of an increase on the order of 30-60 KSI , essentiallyindependent of the type of layup. However, on cooling further to 20K, the

probability is for a slight decrease in strength for cloth and 0°/90° crossply

reinforced epoxies and a reasonably high probability that uniaxial glass

epoxies will suffer a strength degradation which may be as high as 80 KSI.

The phenolic-, polyester-, phenyl silane- and polyurethane-matrix composites

all showed a similar erratic behavior. An exception appeared to be the

silicone-matrix composites which showed consistent moderate increases in

strength at 20K.

The behavior at 200K offers few surprises except for the polyurethanematrix data which indicates an exceptionally large increase from relatively

(2)

poor room temperature strength in 181/Adiprene L-100.

As with the tensile strength, the initial tensile modulus. Figure 2,

shows the expected dependence on fiber orientation. Values range from

7 6

about 10 psi for the uniaxial longitudinal layups to 5-6 x 10 psi for the

0°/90° crossply to 2-5 x 10 psi for the woven cloth composites. The Imiditematrix composites developed much higher moduli than any of the other clothreinforced materials at cryogenic temperatures. Also, the glass -cloth

phenolic composites are found to have, on the average, slightly superior

moduli than glass cloth-epoxies, while glass cloth-polyesters appear almost

as good as the epoxies. The silicone-, polyurethane-, and teflon- matrix

composites displayed the lowest moduli with an indication that phenyl silanematrix composites are of intermediate modulus.

Again taking 77K as a criterion temperature, the uniaxial glass

epoxies showed a modulus range of about 8-11 x 10^ psi with the higher

(66) values reflecting on variants of S-HTS/Epon 828 and on variants of

Hi-Stren/Epon 828. The 0°/90° crossply data ranged from 3-7 x 10^ psi

with values of 5. 5-7 x 10^ psi being reported for S-901 glass with a series

(3) of epoxy resin formulations. The cloth-reinforced epoxies yielded moduli

from 2-5 x 10 psi with the highest values being reported for 181 glass/

(2) Epon 828 formulations.

118
12

10

-

Qfl

UNIAXIAL —

ou

- (1)

70

60 - (Ref: 2, 3, 3. 1,3. 2, 3. 4, 3. 5, 4, 5, 5,1, 66)

50

0 /90

-

40 ——

( 1 ^

\ KGT

.

o,J.i,o.£,o.4,j.D,4,0,D.i) "in

Q.

LU

CLOTH

(Ref. 1,1.2,1.3,2,2.1,3,3.1,3.2,3.4,3.5,

28,34,35,54,55,57,66,80,97)

*Minimal data

I I ! l_

50 100 150 200

TEMPERATURE, K

250

30

25

20

— 15

10

300

Fig. 2. Initial Tensile Modulus of Glass -Reinforced Composites.

(1) Epoxy (4) Polyimide (7) Phenyl Silane

(2) Polyurethane (5) Polyester (8) Teflon

(3) Phenolic (6) Silicone (9) Imidite

119
Although the average modulus of the glass cloth-phenolic composites

was higher than that of the epoxies, literature values ranged from 3-4. 7 x 10^

psi which suggests that no significant difference in moduli should be expected

for good composites made with either epoxy or phenolic matrices. A

detailed look at the glass -cloth-polyester data, however, shows a relatively

narrow modulus range of about 3. 5-4 x 10 psi suggesting that the best

composites made with this polymeric matrix are probably inferior in modulus

to the best epoxies or phenolics by about 10 psi. In a similar way, the

glass cloth-silicone matrix composites are still more inferior having moduli

which vary from 2, 5-2. 9x10 psi at 77K. Somewhat fewer moduli data are

available for Imidite-, silicone-, polyurethane- and phenyl silane-matrix

composites, but that available suggests that the teflon matrix produces

6 6 6 moduli of only 1.6-3x10 psi with 3. 3 x 10 psi and 3. 6 x 10 psi for the

polyurethane and phenyl silane matrices, respectively. The Imidite data

(2)

was obtained with a 181 /Imidite composite and is impressive not only for

the high average value of 4. 38 x 10^ psi developed at 77K, but also for the

indication of a substantial increase to 4. 9 x 10^ psi at 20K.

As with the tensile data, Figure 2 indicates erratic moduli below 77K,

Reference to Figure 5(b) shows that with the exception of a few crossply data,

the literature suggests that cooling from 295K to 77K will produce a modulus

6 6 increase of about 0. 3-1. 2x10 psi with about 0. 7-0, 8x10 psi being most

likely. There does not appear to be a strong dependence on layup. Results

of further cooling to 20K are more difficult to interpret. In general, the data

seem to cluster around a small increase in modulus up to 0, 6 x 10 psi for

cloth-reinforcement and suggests that a somewhat larger increase on the order

of 0. 6-1, 3 X 10^ psi could be expected for crossply and uniaxial composites.

Nevertheless, the scatter from -0. 8 x 10^ psi to +3, 2 x 10^ psi modulus change

is indicative of something erratic occurring below 77K,

Again, the only surprises in the 200K data of Figure 2 is the high

value of the modulus of the flexible polyurethane Adiprene L-lOO compared to

the room temperature modulus and the almost equally large increase in the

Imidite data.

120
Flexure Strength and Modulus - Figures 3 and 4

Flexural tests are frequently used for screening a large number

of composites during development studies, as such tests are simple and

relatively inexpensive compared to tensile testing. In this application,

flexure tests have the added advantage of testing the matrix as well as the

reinforcement fiber. Unfortunately, the state of stress is continuously

changing throughout the flexure specimen as the test proceeds, which makes

engineering interpretation of the data difficult. Consequently, data on

flexural strength and moduli are generally considered valid only for

establishing relative performance ranking.

Figure 3 shows the flexural strengths of the glas s -epoxies to be higher

than that observed in tension by approximately 100 KSI in the uniaxial specimens

and by about 50 KSI in the 0°/90° or cloth layups,

As in the tensile results, the data show the epoxy-matrix composites

to be superior in flexural strength at all temperatures, although the Imidite- • matrix composites are almost as good. Among the other matrix types, the

agreement with the tensile data is less clear. For those composites for which

there is a reasonable amount of data, e, g, , the polyester-, phenolic-, and

teflon-matrix types, the strength order is the same at 77K as it is in tension;

however, the relative strength differences bear little relationship to the

tensile data. The polymide and phenyl silane-matrix composites rank near

the top in flexural strength, while appearing near the bottom in tension.

Conversely, the polyur ethane -matrix composite appears good in flexure but

poor in tension. However, as the data on the latter two composites are based

on only one or two references and on a comparison between composites of

different authors, caution is necessary in interpreting the results. The

polyurethane -matrix data does reflect the same composite tested both in

tension and in flexure.

Verified by comparison of a and a data from the same authors

testing the same composites.

121
400

300

200

100

UNIAXIAL

NOL Ring

100

200

0°/90'^

(Ref: 3,3.1,3.4,3.5,4,66)

150

100

(Ref: 1,1.1,1.2,1.3,2,3,3.1,3.2,3.4,3.5,

12.1,71,80,89,93,97)

CLOTH

Minimal data

50 100 150 200

TEMPERATURE, K

250

- 2

300

Fig. 3. Ultimate Flexural Strength of Glas s -Reinforced Composites.

(1) Epoxy (4) Polyimide (7) Phenyl Silane

(2) Polyurethane (5) Polyester (8) Teflon

(3) Pbtnolic (6) Silicone (9) Imidite

122
Within the unicixial data, a separation has been made between data

generated from flat flexural specimens and that obtained from curved

segments of NOL rings because flexural properties obtained from each type

of specimen are distinctly different.

Examining the available flexural strength data in more detail, we

find that the flat-specimen uniaxial strength of the glas s -epoxies ranged

from 325-470 KSI at 77K with the highest values reported for S-901/E-787, '

The NOL specimen data was significantly lower, ranging 200-270 psi at

(4) 77K. Flexural strengths varied among the 0°/90° epoxy data from 145-

260 KSI at 77K with the highest values again reported in S-901 glass using

(3) either E-787 or an experimental epoxy formulation. Data for the clothreinforced epoxies showed a spread of 95-175 KSI. The highest values

(3) were obtained with 1581/E-787, almost as high values reported for 181

/ (2) glass in a variation of Epon 826 resin. These were the same composites

which excelled in tensile testing.

The glass -phenolic composites ranged from 70-110 KSI at 77K,

The highest values were reported for 1 8 1 / C TL- 91 -LD, ^''"^ Glas s -Imidite

(2) data reflects only the average data with 181 glass reinforcement. The

glas s -polyester s showed a slightly higher range, 80-127 KSI with the best

(2) value reported for 181/Hetron 31. Reported flexural strengths of clothreinforced teflon-matrix composites varied from 30-70 KSI, the highest

(2) values being developed with 181 /FEP,

Following a pattern which is found to repeat itself in all strength

properties of glas s -polymeric composites, the flexural strengths all

initially increased upon cooling from 295K to 77K; however, they then

changed in erratic ways upon additional cooling to 20K. Changes in flexural

strength during cooling as reported in the literature are summarized on

Figure 5(c), which shows that the expected strength increase from 295K

to 77K is about 50-80 KSI for crossply and woven-cloth layups. However,

strength increases of up to 250 KSI have been reported for uniaxial composites, suggesting that the magnitude of the increase is layup-dependent

123
in flexural strength testing. A comparison of Figure 5(c) with that of 5(a)

shows a much greater scatter in the flexural data as compared to the

tensile data on cooling to 77K.

Upon cooling further to 20K, Figure 5(c) indicates that one mayobtain strength changes varying from -50 KSI to +150 KSI with a higher

probability of a decrease than an increase. The data does not appear to

be layup sensitive at 20K.

The flexural modulus data, Figure 4, shows a value of about 8. 5 x

10^ psi for uniaxial glass -epoxy. This is lower than the average moduli in

tension; however, this data reflects only data for a flat specimen of S-901/

(3) E-787. A check of the data shows that this specific composite had an

initial tensile modulus of only about 8. 8 x 10 psi, which suggests that the

two methods are giving about the same answers for the uniaxial case. The

same is true for the 0°/90° case, about 5 x 10 psi being obtained in both

the flexural and tensile modes of testing. The cloth-reinforced polymer

6 6

flexural modulus data range of 2-3 x 10 psi to about 5x10 psi is also

similar to that of tensile modulus.

Except for the teflon-matrix showing poorest performance, the

silicone-matrix showing next poorest, and the Imidite matrix being one of the

best, in both tests there appears to be little correlation between the relative

modulus ranking in flexure and in tension for the same series of composite

types. This is not a blanket condemnation of the flexural test - it may

equally well indicate that average data from the literature cannot be used to

predict tensile behavior from flexural data with any degree of reliability.

Examining the flexural modulus data in more detail, we find that

the 0° / 90° cros sply glas s - epoxy data ranged from 4. 6-5.8x10 psi at 77K

(3) with the highest values for S-901 /E-787. Among the cloth-reinforced

composites, the epoxy-matrix data varied from 2. 6-5 x 10^ psi with maximum

values reported in 181/Epon 828^^^ and in 181 glass with modified Epon 828

. (2)

resm.

124
7

6

- (Ref: 3)

5 —

(1)

(Ref: 3,3.1,3.4,3,5)

0°/90°

a.

4-1 rH _ LU 5

1 —

50 100 150 200

TEMPERATURE, K

250

50

40

30

40 CN

I

- 35

- 30

- 25

20

- 15

- 10

5

0

300

Fig. 4 Initial Tensile Modulus of Glass -Reinforced Composites.

(1) Epoxy (4) Polyimide (7) Phenyl Silane

(2) Polyurethane (5) Polyester (8) Teflon

(3) Phenolic (6) Silicone (9) Imidite

125
• +

0 UNIAXIAL 0 0°/90°

0 CLOTH

295K — 77K

Ao-^"

77K — 20K

0 2 4 6 10 12 14

X lO'^psi

AE^

295K — 77K

-6 -4 -2 0 2 4

X 10 pSi

(a) Ae^

77K — 20K

FR R

-0.8 0.8 1.6 2.4 -0.8

X lO^'psi

0.8 1.6 2.4 3.2

(b)

+ + +

Ao-f"

295K— 77K

Aa^"

77K — 20K

0 2 4 6 8 10 12 14 16 18 20 22 24 26

X lO^psi

(c)

-4 -2 0 2 4 6 8 10

X 10 psi

Ae^

295K-" 77K

R f 11 . n .. 11

-0.8 0 0.8 1.6 2.

Ae^

77K— 20K

.

. fI

+

• HI 0 . n

-0.8 () 0.8 1.6

X lO^psi X lO^'psi

Fig, 5 Histogram Illustrating Changes in Tensile and

Flexural Properties of Glass -Epoxy Composites

Upon Cooling.

126
The phenolic -matrix data emcompassed 1. 2-4. 3x10 psi at 77K,

(41 97) the highest values being developed in Conolon 516, ' The siliconematrix type composites showed a relatively small spread of 2. 6-3. 2x10 psi

(2)

at 77K the highest value reported for 181 glass/Trevarno F-131. Among

the composites for which less data were available, the excellent showing of

the phenyl silane-matrix composites in flexural modulus is somewhat of a

(2) surprise, but the data, coming from the work of Chamberlain, et al. , on

Narmco 534, appear to be valid. The data again shows excellent moduli at

(89, 2)

(2) all temperatures for 181/Imidite composites. The polyurethane- and

polyimide -matrix flexural modulus data reflect only one reference each.

Again, cooling below 77K causes the data to become erratic. Figure

5(d) summarizes the glass -epoxy flexural modulus data and shows that,

while on the average one would expect an increase of about 0. 4 x 10 psi on

cooling to 77K and a like increase in further cooling to 20K, one may find

changes ranging from -0. 8 to +2 x 10 psi. The data suggest that the

crossply might be more adversely affected by cooling to 77K than the clothreinforced specimens; however, there are not sufficient data to verify this

indication.

Compressive Strength and Modulus - Figure 6 and 7

The compressive strength data discussed herein was obtained by

compressing in the fiber direction or, in the case of cloth-reinforced

materials, in the plane of the cloth. As most composites are used in fairly

thin sheet form, the major problem is one of avoiding failure by column

buckling during the test. Problems are further accentuated in the uniaxial

longitudinal case, where slight mis orientation of the fibers can substantially

reduce the compressive strength.

One observes from Figure 6 that the average of the data reported in

the literature for the compressive strength of uniaxial glass -epoxy composites

is less than half of the ultimate tensile strength average for the same type

composites. Yet, the 0°/90° compressive strength is but slightly lower than

its tensile strength, while the cloth-reinforced data span about the same

127
10^

140

—

UNIAXIAL —

120

100 - (Ref: 1,2,3,3.1,3.4,3.5,66)

—

80

110

100

90

80

70

60

50

40

30

20

10

CLOTH

(Ref: 1,1.1,1.2,1.3,2,3,3.1,3.2,3.4,3.5,

80,82,97)

Minimal data

(1)

140 0°/90°

120

100

80

(Ref: 3,3.1,3.4,3.5)

50 100 150 200

TEMPERATURE, K

250 300

Fig. 6 Ultimate Compressive Strength of Glas s -Reinforced

Composites,

(1) Epoxy (4) Polyimide (7) Phenyl Silane

(2) Polyurethane (5) Polyester (8) Teflon

(3) Phenolic (6) Silicone (9) Imidite

128
range in compression and in tension. Consideration should thus be given

to the possibility that the uniaxial data, and to a lesser extent the 0°/90°

data, are lower than the true values due to testing problems.

Among the cloth-reinforced polymers, the epoxy-matrix composites

continue to show superiority over all others. Once again, as in the tensile

case, the polyur ethane -matrix materials indicate a remarkable transformation from an extremely low strength at 295K to one of the strongest of

the group at 77K. The glas s -phenolics continue their reasonably good

performance previously noted in the tensile results. The glas s -Imidite

appears to rank about average in compressive strength, similar to its

performance in tension. The glass -phenyl silanes appear to rank somewhat

better in compression than in tension, although few data are available. The

glass -polyesters appear to have relatively poor compressive strength,

although they were ranked among the top in tension. Finally, the teflon- and

silicone -matrix composites display consistently the lowest compressive

strengths of all materials surveyed.

Considering the uniaxial compressive strength data in more detail,

we found the reported data at 77K to vary widely from 100-240 KSI with

about equal scatter at the other temperatures. This is almost twice the

percentage variation found in the uniaxial tensile data even though the latter

data was much more extensive. This large scatter very likely reflects the

aforementioned problems inherent in compression testing. The highest value

(3) reported at 77K for uniaxial compression was 237 KSI in S-901/E -787.

The 0°/90° data showed much less scatter, ranging from 106-130 KSI

at 77K, the highest value reported in biaxially filament-wound S-901 glass in

DER 332 epoxy.

Among the glas s -reinforced composites, the epoxy-matrix data varied

from 90-138 KSI at 77K with the highest value reported for a modified

(2) 181/Epon 828 composite. Glass -phenolic properties covered a range of

60-100 KSI at 77K, distinctly inferior to the epoxies. The highest value was

reported for 181/Narmco 506. The polyester -matrix composites were

another notch down in strength with a 35-68 KSI showing at 77K, the best

129
(2) results being obtained in 181 glass with Hetron 31 or Narmco 527 resin.

The 19-40 KSI range of the glass -silicones and the 25-40 KSI range of the

glass -teflons at 77K leave little doubt of the inferiority of the latter composite types in compression. The data for polyurethane-, Imidite-, and

(2) phenyl silane-matrix composites reflects the work of Chamberlain et al.

As with tensile and flexural properties, cooling below 77K produces

somewhat erratic results in compression. The 200K data appear to be in

line except for the glass -teflon composite for which cooling to 200K had no

apparent effect.

No uniaxial or 0°/90° data were available for compressive moduli.

Average values among the cloth-reinforced polymers which appear in Figure

7 bear a striking resemblance to the tensile modulus data. Again, the

phenolic -matrix composites were superior with the epoxy-matrix second.

Teflon-matrices were clearly the poorest with the polyurethanes starting out

equally low in modulus but again rapidly increasing its value to the middle of

the group. The silicone-matrices were again on the low modulus side. The

only significant change from the tensile modulus ranking is the somewhat

poorer relative performance of the polyester and Imidite -matrix composites

in compression.

The phenolic composites ranged from 4-8, 4 x 10^ psi in compression

modulus at 77K, maximum value being reported for 181 /CTL-91 -LD. The

glass-epoxy data showed much less scatter at 77K ranging from 4-5, 15 x 10^

(2)

psi. The highest reported value for 181 glas s /modified Epon 826 composite.

The phenyl silane and polyurethane data reflect only the values for Narmco

(2) 534 and 181/Adiprene L-lOO, respectively. Somewhat more compressive

modulus data were available for glass -polyester composites, values ranging

from 2. 5-4, 3 x 10^ psi at 77K the highest value obtained with 181/

Polyester C, The silicones not only averaged about like the polyesters

but had about the same spread in 77K values at 2. 3-4, 6x10^ psi. The data

on glass -teflon and glas s -Imidite composites were provided by Chamberlain

(2)

et al. It is of particular interest to note that the 181 glas s /Imidite

130
X 10

6

COQ.

2 —

CLOTH

Minimal data

(Ref: 1,1.1,1.2,1.3,2,80,82,97)

I I J 50 100 150 200

TEMPERATURE, K

250

X 10

40

35

30

- 25

- 20

15

- 10

- 5

300

Fig. 7 Compressive Modulus of Glas s -Reinforced Composites.

(1) Cpoxy (4) Polyimide (7) Phenyl Silane

(2) Polyurethane (5) Polyester (8) Teflon

(3) Phenolic (6) Silicone (9) Imidite

131
composite shows a sharp decline in both compressive strength and modulus

on cooling from 295K to 77K, the only composite to show this behavior,

Interlaminar Shear Strength - Figure 8

Interlaminar shear strength is a property unique to composites. It

is the resistance to shearing in the plane of the lamella defined by the fiber

reinforcement. It is believed to strongly affect structural integrity, particularly in compression loading. Like the flexural test, the results of interlaminar shear tests evaluate several composite parameters, including resin

strength, resin-fiber bond strength, filament distribution, and matrix

porosity. As such, interlaminar shear joins the flexure test in being a valuable indication of overall composite quality but providing few data

directly useful in engineering calculations. Its most valuable application

may well be in quality control during composite manufacture.

Nevertheless, it is of interest to consider the published values for

interlaminar shear, if for no other reason than to compare the results obtained

with the different methods. Interlaminar shear is usually measured either by

the guillotine method in which interlaminar shear is forced by the imposition

of opposing but offset cuts in a flat tensile specimen or by a dimensional

modification of the flat-specimen flexural specimen so as to force failure by

shear on the central layers of the composite (short-beam shear). Of the two

methods, the latter is most widely used. Unfortunately, the results obtained

by the two methods are not comparable, the short-beam test usually yielding

values higher than that of the guillotine. The overall situation is made still

more complex by the understandable desire of some investigators to obtain

interlaminar shear data from filament -wound composites prepared by the

NOL ring method in which case the flexural method must be used with a short

section of the ring. Because the specimen is not flat, the results are not

comparable to either of the above methods. It is for this reason that the interlaminar data appearing on Figure 8 have been separated according to the

various test methods.

132
0 50 100 150 200 250 300

TEMPERATURE, K

Fig. 8 Interlaminar Shear Strength of Glass -Epoxy Composites.

133
The largest discrepancy is observed for the uniaxial composites, the

values obtained by the NOL short-beam method being much higher than those

in conventional short-beam or guillotine methods. It is obvious from Figure

8 that one must be very cautious in comparing interlaminar shear values

published in the literature. In interpreting these kind of data, it is also

necessary to take into consideration that the very high values developed in

the NOL short-beam test may also reflect the generally lower void content

in filament -wound composites as compared to vacuum- bagged or autoclavecured flat layups.

In the case of crossply or cloth laminates, only the conventional

short-beam flexure test or the guillotine test may be used. Figure 8 shows

that the results of these two test methods are in reasonably close agreement

for the 0°/90° crossply layups, while the cloth-reinforced composites show

lower values for the guillotine as compared to the short-beam test mode.

The same is true for the uniaxial layups tested by these two methods.

The variety of test methods used and the variety of different epoxy

matrices which have been evaluated make it almost an exercise in futility to

attempt to identify composite types which are clearly superior in interlaminar

(3)

shear. The S-901/E-787 composite reported by Toth et al. at 14. 7 KSI

(77K) in short-beam flexure looks good as this value is almost as high as those reported for the NOL' short-beam test. The same composite also

looked comparatively good when tested by the same author in a 0°/90° biaxiallyfilament-wound layup. Other composites which appeared to be relatively

superior among their group were S-901 glass in an experimental resin

(5) Epon 826/Empol 1 040/Z-6077/DSA/BDMA in NOL short-beam and a 0°/90° layup of Hi-Stren glass in Epon 828/LP-3 / Cure agent D resin in

conventional short-beam shear.

^'^^

Ignoring the NOL flexure data, one observes on Figure 8 that the

interlaminar shear values are the highest for the cloth-reinforced composites

and lowest for the 0°/90° crossplies with the uniaxial layups in between. The

134
relatively high values for the cloth composites probably reflect the added

shear resistance provided by the convolutions in the woven glass cloth. By

similar reasoning, the uniaxial composites may be superior to the 0°/90°

crossplies simply because the former provides no distinct lamella along

which shear can propagate.

Again, in a repetition bordering on monotonous, the interlaminar

shear strength is found to become erratic upon cooling from 77K to 20K,

while the 200K properties appear to be as expected.

Ultimate Tensile Strain - Figure 9

In view of the high strength of the glas s -polymeric composites coupled

with their relatively low modulus and absence of significant plastic flow at

rupture, it might be expected that the strains accommodated at fracture

would be somewhat larger than those of the higher -modulus advanced

composites. This is substantiated by Figure 9 on which it is seen that

_2 ultimate tensile strains are in the 10 range for glass reinforcement while, as we shall see later in this review, similar data for the advanced composites

-6

are in the 10 range.

This relatively high fracture strain is, of course, a direct reflection

of the high fracture strain of the glass reinforcement. This has some interesting consequences because while the fracture strain of some polymeric

matrix materials may approach that of glass at 295K, the ductility of almost

all polymers falls drastically upon cooling to cryogenic temperatures with

the result that at low temperatures, fracturing of the matrix occurs well before

fracture of the glass fibers. A dramatic example of this phenomenon is

found during proof testing of filament-wound pressure vessels where the

cracking of the polymeric matrix is very audible. It is surprising that, at

least in the case of hydrostatically loaded pressure vessels, such rupturing

of the matrix does not decrease the overall rupture strength of the tanks and

may even produce an increase. One must emphasize, however, that this does

not imply that fracturing of the matrix is always acceptable or that it may be

tolerated in other structures.

135
(Ref: 3,3.4,4,5,5.1,35)

100 150 200

TEMPERATURE, K

300

Fig. 9 Ultimate Tensile Strain of Glass-Epoxy Composites,

136
Figure 9 indicates that at 295K, the average fracture strain will be

similar for uniaxial, 0°/90° and woven-cloth glass -epoxy composites, the

-2

strains being on the order of 3. 2 x 10 . On cooling, however, marked

differences develop between the layup types. Ultimate tensile strains in the

uniaxial composites (tested in the longitudinal direction) decreased upon

cooling to 77K while cooling the 0°/90° and cloth layups resulted in increased

fracture strain. Detailed examination of the data shows this to be a true

behavior, as five of the six composites forming the uniaxial data showed the

(3) decline (the exception being S-901/E-787 ) while all eight of the 0°/90° and

all three of the cloth composite types showed substantial increases upon

cooling to 77K,

Again, considering 77K as a reference temperature, the uniaxial data

-2

were found to range from 1, 9-5. 3x10 , the highest being the previously

(3) noted work of Toth, et al. Among the 0°/90° crossply data, values showed

_2

a relatively narrow spread of 3. 2-4. 5x10 , the S-901/E-787 composite

(3) -2

again topping the list. Cloth-reinforced data ranged from 3. 5-5 x 10 , (3) the highest ultimate tensile strain being reported for 1581 /E -787.

Data on strain to fracture at cryogenic temperatures are rare for

(64 65) other than the epoxy-matrix types. Kerlin, et al, ' have reported a

-2

fracture strain of 2. 75 x 10 for a Mobaloy 81 AH7 glass -cloth phenolic

-2

composite at 77K and a value of 5, 3 x 10 for a Selectron 5003 glass-cloth

(3) polyester composite at the same temperature. Also, Toth, et al. , have

-2

reported 5. 3 x 10 for a crossply glass -polyester composite Selectron 5003.

Static Fatigue

Under sustained room temperature loading, glass filaments have been

found to deteriorate and fracture under applied stresses substantially below

that of their normal ultimate tensile strength. As such, the failure is analogous to stress corrosion in metals. Some published data indicate that glass

(99) filament-wound pressure vessels may undergo similar deterioration,

although some indication has also been obtained suggesting that static fatigue

137
of glas s -reinforced composites may pose less of a problem at cryogenic

(66

)

temperatures than at 295K. Static fatigue will be an important factor in

any composite used in the rotating components of superconducting motors

and generators, and in view of the minimal data presently available , further

studies are needed to clarify the magnitude of the problem and to select

formulations providing utmost resistance to such failure at cryogenic temperatures, T, T. Chiao of the Lawrence Livermore Laboratories is

currently directing an extensive static fatigue program at room temperature

on dead-weight loaded filaments coated with resin. As this latter method

tests a basic structural element of the composite and permits many specimens

to be run concomitantly, it should provide a useful and relatively inexpensive

method of evaluating static fatigue at cryogenic temperatures.

Bearing Yield Strength - Figure 10

Bearing yield strength is a test designed to determine bearing stress

as a function of the deformation of a hole through the composite. The load is

applied by a pin inserted into the hole. The intent of the test is to provide

information on the stress that may be sustained across riveted or bolted joints

without loosening the joint. Bearing yield strength is defined as that stress

on the stress -strain curve which corresponds to a distance of 4% of the bearing

hole diameter when measured from the intersection of a line tangent to the

stress -strain curve at this point and the zero load axis. (See insert, Figure 10. ) Bearing yield strengths for a series of epoxy-matrix composites have

(3) (2) been reported by Toth et al. , while Chamberlain et al. , have provided

data on silicone-, polyur ethane-, and phenyl silane -matrix composites.

These data are summarized on Figure 10,

The glass -epoxy composites developed bearing yield strengths which

increased from about 50 KSI at 295K to 70-75 KSI at 77-20K, The uniaxial

and 0°/90'' data were very similar, while the cloth-reinforced epoxies were

lower at 295K and at 200K but increased their strength to equal that of the

others at 77K,

138
Fig. 10 Bearing Yield Strength of Glass -Reinforced Composites.

139
Data on the other matrix types were only available for cloth

reinforcements. The phenyl silane composite developed a bearing strength

equal to that of the epoxies, while the silicone- and polyurethane -matrix

materials were distinctly inferior.

Again, there is indication of erratic behavior on cooling from 77K to

20K,

DYNAMIC MECHANICAL PROPERTIES

Cyclic Fatigue - Figure 11

Cyclic fatigue of glas s -reinforced composites at cryogenic temperatures

has been studied by Brink, et al, , and by Chamberlain, et al. , in tension,

by Fontana, et al, , ^^^^ by reciprocating beam, while Lavengood and Anderson^"^^^

have contributed data on torsional fatigue. The data are not extensive, reflecting

the very high cost of generating complete S-N curves at cryogenic temperatures.

Yet, such testing is mandatory to provide assurance that composite components

will fulfill their life expectance in such applications as rotating cryogenic

machinery. Every effort must, therefore, be made to restrict such testing

to composites which have the best chance of developing superior fatigue

properties. The data reviewed herein provide some sense of direction;

however, there remains an urgent need for some type of relatively inexpensive

screening test which will permit relative ranking of composite cyclic fatigue

performance at cryogenic temperatures.

In comparison to that of the advanced composites or to most

conventional metals, the dynamic fatigue properties of glass -reinforced

composites are not good. This is primarily due to the fatigue behavior of

gla s s -polymeric materials being controlled by the properties of the matrix.

Even at room temperature, the strain accommodation of the most epoxy resins

is less than that of the glass reinforcement, and as it is the latter which

controls the ultimate strength of the composite, cracking of the matrix will

occur at ultimate loads far below the composite ultimate strength, allowing

corrodents to attack the glass. In crossply layups, such cracking may occur

at stress levels as low as 20% of the ultimate composite strength.

140
Lavengood has pointed out that since fatigue life of a glass -reinforced

composite is related to the strain capability of the matrix, embrittling of the

matrix due to lowering of the temperature should lower the fatigue life.

However, this is not always experimentally verified. Based on an analysis of

experimental torsional fatigue data, Lavengood concluded that fatigue life at

cryogenic temperatures is strongly affected by the composite interfacial stress

which arises due to differential thermal contractions of the filament and the

matrix. Where fatigue strains serve to increase compressive interfacial

stresses, the fatigue life may increase with lowering of temperature^ the

reverse being true when strains decrease the interacial stress.

The published tensile-fatigue data of Brink^^^ and Chamberlain^^^ and

their coworkers were, with one exception, obtained on 181 cloth-reinforced

composites. As an initial criteria, composites were screened by testing at

30% of their ultimate tensile strength developed at 295K and 77K. Those

composites failing to achieve 10^ cycles were dropped while those which

were successful were further tested at 200K and 20K. The conventional

polyester-matrix composites 1 81 / Hetron 92, 181 /Hetron H-31, 181/paraplex

P-43, and 181/Narmco 527 were unsuccessful as was the silicone-matrix

composite 181/Trevarno F-131.

It is beyond the scope of the present task to comprehensively review

this data for which the reader may refer to the original references. However,

in order to provide the reader with a sense of the fatigue performance of glasspolymeric composites, we have prepared Figure 11 from the data of Brink^'^^

(2) and Chamberlain. This Figure presents the fatigue strength of the various

successful materials as a function of temperature after 10^ fatigue cycles,

the maximum studied in this work. It is also instructive to know the percent

of the relevant ultimate strength retained by each composite type at each temperature after 10^ cycles; consequently, these data also appear on Figure 11.

Figure 11 shows that the absolute magnitude of the stress required to

induce fatigue failure at 10 cycles generally increases with decreasing

141
10-

60 I-

(9)

% cr^" at Temperature

295K 200K 77K 20K

Epoxies (1) 30

Phenol ics (2) 41

Polyesters (3) 33

Silicones (4) 35

Phenyl Silanes(5) 32

Polyurethanes (6) 60

FEP Teflon (7) 50

Imidite (8) 48

Scotchply (9)^

32

42

35

42

45

64

37

44

36

31

57

40

35

52

52

High Temperature Types

^Uniaxial. All others

181 Glass Reinforced

Minimal Data

(Ref. 1,2)

37

41

40

66

36

28

57

50 100 150 200

TEMPERATURE, K

250 300

Fig. 11 Cyclic Tensile Fatigue Strength of Glass -Reinforced

Composites After 10^ Cycles.

142
temperature for all materials studied, only the high-temperature polyesters

(Laminae 4232 and Vibrin 135) and perhaps the silicones (Narmco 513 and

Trevarno F-130) not showing an appreciable increase below 200K, At all

temperatures, the Imidite composite demonstrated superior fatigue performance both in terms of absolute retained tensile strength after 10^ cycles

and in the percentage of the original strength retained. The epoxies also

looked good, but primarily because of the somewhat higher initial strength-

their percentage retention of strength was among the lowest of the group.

(2) The best of the epoxies appeared to be Epon 828/DDS, The polyur ethane-,

phenyl silane-, and phenolic -matrix composites appeared to group into an

intermediate performance class while the polyester-, silicone-, and teflonmatrix composites showed distinctly inferior fatigue properties. Data at

' (2) 77K only was available for a Scothply 1002 uniaxial glass -epoxy composite,

which Figure 11 shows to have a much higher strength after 10 fatigue cycles

at 77K than any of the cloth-reinforced types; nevertheless, even this composite showed retention of only about 52% of its original strength at that

temperature.

Do the data indicate the existence of a stress limit below which

fatigue life is essentially infinite? In most cases, testing was not carried

out to a sufficiently large number of cycles to answer this question. However,

judging from the shapes of the S-N curves developed with 181/Epon 1001 by

Brink and his coworkers, it appears that this glass -epoxy formulation may

reach such a limit at about 30 KSI for temperatures below 77K (^30% UTS),

Conversely, the S-N curve for the 181/Epon 828 composite showed no

indication of flattening out at 10 cycles. With the possible exception of the

silicone-resin composites, 181/Narmco 513 and 181/Trevarno F-130, there

was no evidence that cooling the composites had any effect on establishment

of a fatigue limit where none was evident in the room temperature data.

Again, one notices an apparent tendency for the data to become

erratic when cooling from 77K to 20K,

143
Impact Strength and Fracture Toughness

Impact strength and its more sophisticated partner - -fracture

toughness --are measurements of the amount of energy that may be stored in a structure before the energy is released by fracture. As it is only in recent

years that fracture mechanics has been put on solid theoretical grounds and

exploited in homogeneous metals, it is not strange to find that application

of parallel concepts to composites is still in its infancy. Furthermore,

there is a major problem in applying basic principles of fracture mechanics

to composites where, in most cases, multiple cracking occurs so that a unique "crack length" cannot be defined.

Measurement of the energy required to fracture specimens - -the

impact test- -is the most simple method of obtaining data on relative material

toughness. Such tests show the glas s -reinforced composites to have much

greater toughness than do the advanced composites, probably reflecting the

much larger strain-accommodation of glass filaments in comparison to the

advanced fiber reinforcements. Unfortunately, the literature contains few

low temperature data even on this simple parameter, and that which is

available is impossible to systematize due to differing test methods, specimen

(4) design, and filament orientation. Lewis and Bush have evaluated a series of

epoxies and modified epoxies reinforced with Hi-Stren glass using the Izod

impact method. For unidirectional composites, they find the 295K impact

strength to vary from 82-128 ft-lb in of notch. On cooling to 77K, the

-1

measured values ranged from 67-162 ft-lb in , some formulations showing

45% increase in impact strength while others showed as much as a 26%

decline, A 0°/90° crossply test with the same composite formulations yielded

-1

room temperature impact strengths from 49-96 ft-lb in and 77K values of

59-76 ft-lb in ^ with strength changes ranging from +36% to -26%, no

consistent behavior being shown by any specific composite formulation. Data

uT u J u (71,80,103,109,110) „ , . ^ , published by other authors, are equally confusing. Perhaps

(71) part of the answer is provided by Levin who followed the change in impact

strength of glas s -phenolic and glass-epoxy specimens at closely spaced

intervals from 295K to 77K and found that the impact properties peaked

144
sharply at about -30° C followed by a rapid decline at lower temperatures.

If the impact properties are indeed such a rapidly changing function of

temperature, it might account for the lack of systematic change noted in

data taken only at two or three temperatures. Clearly, this is an area which

must be given much more attention in the future if composite behavior at

cryogenic temperatures is to be well understood.

THERMAL PROPERTIES

Thermal Contraction - Figures 12 and 13

All of the glas s -reinforced polymers contract when cooled. As we

wish to review the behavior of such composites on cooling from 295K, the

data have been plotted as thermal contraction, avoiding negative values of

AL/L.

Thermal contraction is very dependent on the type of composite

layup as well as on the orientation of the composite. Figures 12 and 13,

therefore, present data on values obtained normal to the fiber reinforcement

for uniaxial, 0°/90° crossply and cloth layups. Other factors affecting

thermal contraction are the specific resin used and the composite fiber

density.

The thermal contraction is always found to be higher in a direction

normal to the fibers than in the fiber plane, the difference increasing as the

temperature is lowered. This reflects the much higher thermal contraction

of the polymer as compared to the glass reinforcement. This is very evident

in the data for uniaxial epoxy-matrix composites, where at 20K the thermal

contraction normal to the fibers varies from about 2. 5 to 5 times that in the

fiber direction at 20K. An apparent anomaly appears in the 0°/90° crossply

data, which shows a remarkably small difference in contraction between

these two directions. This occurs because, in a filament-wound crossply

composite, each layer of glass is in close proximity to the orthogonal layer

next to it and thus provides resistance to dimensional change in the thickness

direction. This effect is much less evident in the cloth layups because the

145
TEMPERATURE, K

ig. 12 Thermal Contraction of Glass-Fiber Reinforced Composites,

146
A iU

300

250

200

150

140

100

In Plane

Normal

7(a)

, \ :

\

\

\

\7(a)

\

\

\

\

\

\

"--.^ 7(b) "

©-^^ (4)

""--^

Fig. 13

(1)

(2)

(3)

(4)

(5)

150 200

TEMPERATURE, K

Thermal Contraction of Glass -Cloth Reinforced Composites.

Epoxy (Ref. 85, 101)

Polyurethane (Ref. 85)

Phenolic (Ref. 14)

Polyester (Ref, 14)

Silicone (Ref. 14)

(6) Phenyl Silane (Ref. 14)

7(a) Teflon TFE (Ref. 14)

7(b) Teflon FEP (Ref. 80)

(8) Imidite (Ref. 16)

147
latter are prepared from prepreg tape which results in a much lower

filament density than is obtained in filament winding.

As the thermal contraction of the composite is dominated by the

properties of the polymeric matrix on cooling, one might anticipate

considerable variation in contraction even within a given matrix type,

reflecting, for example, the effect of additives to epoxy resins. This is

found to be true; for example, the AL/L data for epoxy matrix composites

-4

varied from 20-75 x 10 in the thickness direction and from 4, 4-11. 5x10

in the fiber direction. In the 0°/90° layups, the epoxy composites showed

-4 -4

variations of 5, 1-13, 8x10 in thickness and 9. 2-13 x 10 in the fiber

direction. The polyester resins gave similar data spread in the 0°/90°

layups.

A much greater variation is observed in the thermal contraction

among cloth-reinforced composites, reflecting both the larger variety of

matrix materials for which data are available and the greater variation in

fabrication method. The glass -teflon composites have the highest thermal

contractions of all the materials examined. Furthermore, the teflon matrix

may be either of the TFE (polytetrafluoroethylene) or FEP (copolymer with

hexafluoropropalene) type which have markedly differing thermal contraction

characteristics. This difference does not appreciably affect contraction in

the fiber plane due to the restraint provided by the fibers. However, the

difference is marked in the direction normal to the fibers. As indicated by

the dashed curves 7(a) and 7(b) on Figure 13, the TFE -matrix composites

have twice the transverse thermal contraction of the FEP-matrix types.

The epoxy- and the polyur ethane -matrix composites also have

relatively high thermal contractions. Unfortunately, no data were available

in the fiber normal directions for these composites. The data indicate that

the polyester-matrix composites are next in decreasing order of thermal

contraction, with the fiber normal data being approximately 2. 8 times

greater than in the fiber plane. Showing the least thermal expansion in the

148
fiber plane are the phenolic-, silicone-, phenyl silane-, and Innidite -matrix

composites. Of these, the phenolic composite data appear to be abnormally

high for the fiber normal case, since, if all else were equal, the order in the

fiber normal direction should follow that in the fiber plane.

All of the polymeric matrices create compressive stresses in the glass

reinforcement upon cooling to cryogenic temperatures, such stresses being

added to those already existing from cooling from the cure temperature to

room temperature. This is primarily manifested as shear stress at the

resin-fiber interface, so that it might be expected that an effect would be

seen on those strength properties which are sensitive to debonding of the

matrix-glass interface. One such property is fatigue, and from Figure 11,

it is noted that the fatigue properties of the teflon-matrix composites are poor,

while that for the Imidite is good, consistent with the respectively high and

low thermal expension of these two composite types. A similar relationship

holds for the flexural strength. Figure 3, Interfacial residual stress is but

one of the factors influencing fiber debonding, of course, others being the

ability of the interface to sustain a shear load and the strength of the matrix

itself.

Specific Heat - Figure 14

The specific heats of glas s -reinforced composites show an almost

linear dependence on temperature from 295K to 77K, and relatively small

difference between matrix types. The specific heats are relatively high

compared to most metals, being roughly comparable to that of aluminum but

substantially higher than that of titanium, iron, or copper. Teflon- and

polyester -matrix composites have the highest specific heat. The epoxy-,

Imidite-, silicone-, and phenolic -matrix composites are bunched on the

lower specific heat side of the group. The phenyl silane -matrix composite

seems to have a slightly larger temperature dependence, starting out in the

middle of the group at 295K, but showing the lowest specific heat at 77K.

149
X 10'

12

11

10

90 bIi.io\4" j^sYis as > 8

i.rbnodsb c-i ©viiiarit ^

ijjd S£J3 2

4

(1

(2

(3

(4

(5

(6

(7

(8

GLASSES Refs:

) Epoxies (15,16)

) Phenol ICS (^15)

) Polyesters (15)

) Silicones (15)

) Phenyl Silanes (15)

) Teflon (15)

) Imidite (16)

) Phenyl Formaldehyde (72)

ooq asix 50qn-roD xxiiBai'.aonsj s.dj io- asx

qidaaoftsls!

(2,4,7)-'

''^(4)

, B 9 q-(J 9 j" i s oqni o

(2)

-i-sraisfi^ wo I

T.9dil §iiic>iis.arix"2x 3to^osi siij' "Ic »;rt'

-G r;isj-8;ja Oj sosl'tsirxi 9ri:t lo y-"^ f-f -jci'-

±

ha., -ncPiisT .:£s§0u^ -xo ,.A00 150 200 _ , 3..;.?50,...,,,,,,^300

aoqiifOD xlrrtisrrf- -fsjasyioq

9x1:

.HTT isaxi ojllosqa isswol sAi ^ah^od^ tod ,Hc9S qxjotrq ad .

150-'-
The data reflects the work of Campbell, et al. ,

' with the sole

(72) exception of the phenyl formaldehyde contribution of Luikov. Except for

the epoxies, data reflect only one composite for each type of matrix. The

epoxy data reflect' ^the average of three compositions: a unidirectional

YM-31 -A/DER-332 and a unidirectional S-994/E-787 composite. The

spread in specific heat values for these three formulations at 295K, 200K,

-2 -1 -1

and 77K was 8. 16-9. 21, 6. 07-6. 7 and 2. 09-2. 72 x 10 J Kg K , respectively, with the highest value associated with the S-994/E-787 composite.

Thermal Conductivity - Figures 15 and 16

As with thermal contraction, the conductivity is dependent on the type

of layup and on the orientation of the composite. Figures 15 and 16, therefore,

present data on values obtained normal to the fiber reinforcement (thickness

direction) and in the plane of the reinforcement for uniaxial, 0°/90° crossply

and for woven-glass cloth layups. Also, as with contraction, the thermal

conductivities are affected by the matrix resin and by the density of the reinforcement fiber. Continuity of the fiber is also a factor affecting in-plane

conductivity.

Both the difference between the fiber normal and the in-plane

conductivities for a given composite and the absolute spread of conductivity

values among the various composite types is the widest at room temperature.

However, these differences rapidly diminish as the temperature is lowered.

Figure 15 contains data for uniaxial and 0°/90° epoxy layups and for a 0°/90° silicone -matrix layup. The highest thermal conductivity was for the

filament -wound uniaDcial epoxy composite in the plane of the fiber. This is to

be expected from the high density of continuous glass fibers in that direction.

The conductivity in the fiber normal direction is 25-50% less, reflecting the

lower thermal conductivity of the matrix. Thermal conductivity in the fiber

plane of the 0°/90° epoxy layup is only about 80% of that for the uniaxial

composite, reflecting the lower effective fiber density in the direction of the

heat flow.

151
0.8

0.7

0.6

0.5

I 0.4

0.3

0.2

0.1

Parallel to Reinforcement

Normal to Reinforcement

EP0XIES,(Ref; 42, 38, 16, 14, 56)

SILICONES(Ref: 14)

Uniaxial

Si 1 icone,

0O/90O

50 100 150 200

TEMPERATURE, K

250 son

Fig. 15 Thermal Conductivity of Glass-Fiber Reinforced Composites.

152
Cloth, in Plane (Ref. 14,16,21,25,56) ^«

I L 1 ! I I L_J

0 50 100 150 200 250 300

TEMPERATURE, K

Fig. 16 Thermal Conductivity of Glass-Cloth Reinforced Composites.

153
Figure 15 shows the 0°/90° silicone-matrix composite to have a substantiallylower in-plane conductivity than the 0°/90° epoxy layup. This does not imply

that the conductivity of the silicone matrix is less than that of the epoxy;

indeed, the marked similarity between the fiber normal data for both polymeric matrices indicates that the matrix conductivity is very similar in both

materials'. The difference reflects the relatively inferior heat transfer

properties of glass roving which was used for the silicone-matrix composite

as compared to the continuous filament used with the epoxy.

Among the cloth-reinforced composites, the highest thermal

conductivity was found for the silicone-matrix composites. Only fiber

normal data were available; however, even in this least conducting orientation,

the conductivity exceeded that of all the other materials in their more favorable in-plane direction. The Imidite composite was next lowest, and in

decreasing order were the epoxies, phenolics and phenyl silanes, all with

approximately equal conductivities, both in-plane and plane normal. The

lowest conductivities were evidenced by the polyesters and the teflons.

An interesting facet of these data is that, unlike the other thermal

properties of composites, the thermal conductivity is affected by the

ambient atmosphere in which the measurement is taken, i, e. , it differs in

helium, nitrogen or in vacuum. The values plotted in Figure 15 and 16 are

averages of data obtained in the various atmospheres. However, a more

detailed analysis of the literature data shows that, compared to values obtained

in helium gas, data taken in nitrogen will average 7% lower, while in a vacuum

the data will, on the average, be lowered by 20%.

The only composites for which there were more than one literature

reference were the epoxies. For the latter, the data spread was found to be

much greater for the in-plane conductivity than for the plane -normal, as would be expected in view of the relatively high conductivity of the fibers

compared to the matrix. As an example, the data which averaged

-1 -1

0. 50 w- m . K for the 295K cloth-reinforced epoxy ranged from

154
0, 30 to 0. 65 w- m " K , while the plane normal data for the identical

composites average 0. 35 w. m ^- K ^ but ranged from 0. 30 to 0, 40 w- m ^' K \

The scatter decreased significantly at lower temperatures.

COMMENTS ON GLASS -REINFORCED COMPOSITES

We have reviewed the properties of filamentary-glass reinforced

composites at cryogenic temperatures in order to provide the reader with an

overall feel for their behavior and the magnitude of the properties which may

be expected. Having this information, for what applications should glass

reinforced composites be used? What composite formulation should be

selected?

It would appear that glas s -reinforced composites are most useful in

applications requiring high tensile strength combined with high toughness and

low thermal conductivity, but where stiffness is not required and where cyclic

fatigue is not a major problem. In such applications, glass -reinforced composites have given and will continue to give excellent performance at relatively

low cost. Filament-winding techniques should be used whenever possible in

order to obtain the highest quality composites; in particularly the lowest

void content.

The choice of composite will be dictated by the particular application,

considering not only the final properties but considering equally the problems

involved in working with a specific composite under production conditions.

At the present time, the composite most widely used for cryogenic applications

(5)

is S-901 glass filament-wound in an epoxy matrix known as Resin 2, This

resin, composed of Epon 828 /DSA/Empol 1040/BDMA in proportions of

100/1 15. 9/20/ 1 by weight, probably represents the current state-of-the-art

for cryogenic applications.

An unexpected result of the present review is the surprisingly good

overall performance of the glas s -Imidite (polybenzimidazole) type. Presently

available data indicate that at 77K, such a composite ranks second only to the

epoxies in ultimate tensile strength and flexural strength while developing

tensile and flexural moduli superior to the epoxies. Compressive properties

155
appear only average; however, in cyclic fatigue, the Innidite -matrix

composite appears superior to all other matrix materials, particularly

in percent of retained ultimate tensile strength, A key to its excellent

fatigue properties may be the comparatively low thermal contraction of

Imidite compared to the other polymers, resulting in comparatively low

residual interfacial stress between the fiber and the matrix. It would seem

that a second look at glas s -Imidite composites is in order,

A pervasive characteristic of glas s -polymeric composites is the

erratic behavior of the mechanical properties on cooling from 77K to 20K,

With few exceptions, the strength properties increase on cooling from 295K

to 77K; however, on further cooling to 20K, the data indicate that such

properties may either increase, decrease, or remain unchanged, the behavior

being quite unpredictable, even among composites of the same matrix type. It

is difficult to attribute this simply to the matrix becoming suddenly much more

brittle between 77K and 20K, as cooling to 77K has already decreased the

strain capability of the matrices to a level far below that of the glass reinforcement. The present task does not permit more than a cursory consideration of

the possible factors involved in this phenomena; however, the author believes

it relevant to call attention to recent studies on the low temperature properties

of polymers which have provided convincing evidence that, in at least the

linear polymers of the polycarbonate (PC) and polyethelene terepthalate (PET)

types, the media in which the low temperature test is conducted can strongly

affect the fracture mode and the resultant mechanical properties measured

(123 126)

for the polymer. ' The experimental evidence indicates that the

failure in such polymers in the cryogenic range is controlled by a crazing

phenomena which, in turn, is related to the activity of the gas or liquid in

contact with the polymer surface. Such studies have not been extended to the

strongly cross -linked polymers; however, until proven otherwise, it must be

considered a possibility that the mechanical properties of glas s -polymeric

composites may be influenced by the ambient media such that data obtained at

20K or 77K in liquid hydrogen or nitrogen may not be the same as those which

would be obtained in a helium atmosphere.

156
This review suggests that the following work would be valuable in

preparing glass -reinforced composites for use in critical components of

superconducting machinery:

a) There are indications that static fatigue of glass -reinforced

composites may not be a problem at cryogenic temperatures; however, data

are minimal and should be expanded,

b) Data on cyclic fatigue of glass -reinforced composites at cryogenic

temperatures, particularly high cycle fatigue, is minimal. That which is

available does not clearly indicate the existence of a fatigue limit, although

the available data does show the fatigue life for a given composite to be

higher at cryogenic temperatures than at room temperature. It is evident

that additional fatigue data will have to be obtained before glas s -reinforced

composites can be used with confidence under cyclic loading in superconducting

machinery,

c) In view of the extremely high cost of conducting fatigue tests at

4K, an effort should be made to devise a screening test to correlate

incipient damage with fatigue life so that full cycle fatigue testing will be

restricted to those composite formulations showing the most promise.

d) No literature data were available on the resistance of glass

reinforced composites to thermal fatigue. As cycling through elevated

temperatures has been shown to seriously degrade many types of composites,

this parameter should be investigated for any composite likely to be subjected

to thermal cycling in its application in superconducting machinery.

157
CONTRACT BIBLIOGRAPHY

I - MECHANICAL PROPERTIES

Contract AF-33(616) -8289

Contractor; Directorate of Materials and Processes, Aeronautical

Systems Division, Air Force Systems Command, WrightPatterson Air Force Base, Ohio.

Research Facility; Narmco Research and" Development, San Diego,

California

.

1,0 Brink, N. O, , "Determination of the Performance of Plastic Laminates

Under Cryogenic Temperatures", ASD-TDR-62-794, August 1962 (AD

288 944).

1. 1 Ibid. , "Mechanical Behavior of Reinforced Plastics at Cryogenic

Temperatures", Technical Papers, 20th Annual Technical Conference, Society of Plastics Engineers , Vol. 10, Section XV-2,

1964, pp. 1-19.

1.2 Ibid. , "Mechanical Behavior of Reinforced Plastics at Cryogenic

Temperatures", Society of Plastics Engineers Journal , Vol. 20,

1964, pp. 1123-1131.

1. 3 Ibid, , "Mechanical Behavior of Reinforced Plastics at Cryogenic

Temperatures", Narmco Research and Development Report, Code

No. 105-4, 1964.

2. 0 Chamberlain, D. W. , Lloyd, B. R. and Tennant, R. L. , "Determination

of the Performance of Plastic Laminates at Cryogenic Temperatures",

ASD-TDR-62-794, Part II, March 1964 {N64-24212).

2. 1 Chamberlain, D. W. , "Tensile Fatigue Testing at Temperatures

Down to 20 K" , Advances in Cryogenic Engineering, Vol. 9, 1964,

pp. 131-138.

2. 2 Ibid, , "Mechanical Properties Testing of Plastic Laminate

Materials Down to 20 K", Advances in Cryogenic Engineering , Vol. 10, 1965, pp. 117-125.

Contract NAS 8-11070

Contractor; National Aeronautics and Space Administration, George C.

Marshall Space Flight Center, Huntsville, Alabama.

Research Facility; Goodyear Aerospace Corporation, Akron, Ohio.

3.0 Toth, L. W. , Boiler, T. J,, Butcher, I. R. , Kariotis, A. H. andYoder,

F. D. , "Program for the Evaluation of Structural Reinforced Plastic

Materials at Cryogenic Temperatures", NASA CR-80061 (Final), August

1966 (N67- 12051).

158
3. 1 Toth, L. W. , "Properties Testing of Reinforced Plastic Laminates

Through the 20 degree K Range", Technical Papers, 20th Annual

Technical Conference, Society of the Plastics Industry , Section

7-C, 1965, pp. 1-10.

3.2 Toth, L. W. andKariotis, A. H. , "'An Assessment of Test Specimens and Test Techniques Useful to the Evaluation of Structural

Reinforced Plastic Materials at Cryogenic Temperatures", Advances in Cryogenic Engineering , Vol. 10, 1965, pp. 126-133.

3. 3 Toth, L. W. , "Properties of Glass-Reinforced Epoxy Through the

20 K Range", Modern Plastics, Vol. 42, August 1965, pp. 123- 130.

3.4 Toth, L. W. and Burkley, R, A. , "Mechanical Response at Cryogenic Temperatures of Selected Reinforced Plastic Composite

Systems", Goodyear Aerospace Report GER-13169, Paper No. 16,

Seventieth Annual Meeting of the American Society for Testing and

Materials , June 1967.

3.5 Toth, L. W. , Boiler, T, J., Kariotis, A. H, andYoder, F. D. , "Program for the Evaluation of Structural Reinforced Plastic

Materials at Cryogenic Temperatures", NASA CR-64005, June

1963 through June 1964 (N65-29724).

Contract NAS 3-6297

Contractor; National Aeronautics and Space Administration, Lewis

Research Center, Cleveland, Ohio.

Research Facility ; Aerojet General Corporation, Azuza, California.

4.0 Lewis, A. and Bush, G, E. , "Improved Cryogenic Resin-Glass Filament

Wound Composites", NASA CR-72163 (Final), March 1967 (N67-31856).

4. 1 Lewis, A. , Bush, G. E, and Creedon, J, , "Improved Cryogenic

Resin/Glass Filament -Wound Composites", NASA Interim Report

CR-54867, May 1966 (N66-28040).

Contract NAS 6-6287

Contractor ; National Aeronautics and Space Administration, Lewis

Research Center, Cleveland, Ohio.

Research Facility: Aerojet General Corporation, Azuza, California.

5.0 Soffer, L. M. and Molho , R. , "Cryogenic Resins for Glass FilamentWound Composites", NASA CR-72114 (Final), January 1967 (N67-25076).

5. 1 Ibid. , "Mechanical Properties of Epoxy Resins and Glass/Epoxy

Composites at Cryogenic Temperatures", Cryogenic Properties of

Polymers , J.L.Koenig, Ed., Marcel Dekker, New York, 1968,

pp. 87-117. (Identical to NASA CR-84451, 1967, (N67-27217).

159
Contract F0470 1 - 69 -C -00 59

Contractor: Space and Missiles Systems Organization, Air Force

Systems Command, Los Angeles Air Force Station,

Los Angeles, California,

Research Facility; The Aerospace Corporation, El Segundo, California.

6.0 Pepper, R. T, , Rossi, R. E. , Upp, U. W. and Riley, W, E. , "Development of an Aluminum- Graphite Composite", SAMSO-TR-70-301 , August

1970 (AD 718 409).

6, 1 Pepper, R, T. , Upp, J. W. , Rossi, R, C. and Kendall, E. G. , "Aluminum-Graphite Composites", SAMSO -TR-70 - 1 74 , April

1970 (AD 706 883). (Identical to Metallurgical Transactions , Vol. 2, January 1971, pp. 117-120.)

6.2 Rossi, R. C. , Pepper, R. T. , Upp, J, W. and Riley, W, C. , "Development of Aluminum-Graphite Composites", Ceramic

Bulletin , Vol. 50, May 1971, pp. 484-487.

Contract NAS 8-11508

Contractor: National Aeronautics and Space Administration, George

C. Marshall Space Flight Center, Huntsville, Alabama.

Research Facility : Harvey Engineering Laboratories, Torrance,

California.

7.0 Sumner, E. V. and Davis, L, W, , "Development of Ultrahigh Strength,

Low Density Aluminum Sheet and Plate Composites", NASA CR-85863

(Final), July 1966 (N67-31181).

7. 1 Davis, L. W. , "Composites at Low Temperature", Paper No. 15 , Seventieth Annual Meeting of the American Society for Testing

Materials

,

June 1967,

Contract NASA DPR C 10360-B

Contracto r: National Aeronautics and Space Administration, Lewis Research Center, Cleveland, Ohio,

Research Facility : Naval Ordnance Laboratory, Silver Springs,

Maryland.

8.0 Simon, R, A. and Alfring, R, , "Properties of Graphite Fiber Composites

at Cryogenic Temperatures", NASA CR-72652 (NOLTR 69-183), May

1970, Tasks I and II (AD 746 885).

9. 0 Larsen, J. V, , "Properties of Graphite Fiber Composites at Cryogenic

Temperatures -- Effect of Elastomeric Additions to Resin Systems",

NASA CR-72804 (NOLTR 70-195), March 1971, Task III (AD 882 972).

160
10,0 Larsen, J. V. and Simon, R, A., "Carbon Fiber Composites for

Cryogenic Filament-Wound Ves sels" , NASA CR-120899, (NOLTR

71-201), May 1972, Tasks IV, V and VI, (N73-1 1553).

10. 1 Simon, R. A. , ^'Graphite Fiber Composites at Cryogenic Temperatures", Technical Papers, 26th Annual Technical Conference,

Society of the Plastics Industry , Section 10-D, 1971, pp. 1-4.

10.2 Larsen, J. W. , "Fracture Energy of CBTN/Epoxy-Carbon Fiber

Composites", Technical Papers, 26th Annual Technical Conference,

Society of the Plastics Industry , Section 10-D, 1971, pp. 1-4.

Contract NAS 8-26198

Contractor : George C. Marshall Space Flight Center, Huntsville,

Alabama.

Research Facility : General Dynamic s / Convair , San Diego, California.

11,0 Scheck, W. G. , "Development of Design Data for Graphite Reinforced

Epoxy and Polyimide Composites" , Report No. GDC-DBG70-005 , General

Dynamics Quarterly Report No. 1, 1970.

12. 0 Maximovich, M. and Scheck, W. G. , "Development of Design Data for

Graphite Reinforced Epoxy and Polyimide Composites", Report No. GDCDBG70-005, General Dynamics Quarterly Report No. 2, 2970.

12. 1 Stuckey, J. M. and Scheck, W. G. , "Development of Graphite/

Polyimide Composites", National Technical Conference, Society

of Aerospace Material and Process Engineers , Vol. 3, 1971, pp.

717-723.

Contract F3 3 6 1 5 - 70 - 1442

Contractor ; Air Force Materials Laboratory, Wright- Patterson Air

Force Base, Ohio.

Research Facility : General Dynamic s / Convair , San Diego, California.

13.0 Hertz, J. , Christian , J. L. and Varlas, M. , "Advanced Composite Applications for Spacecraft and Missiles, Phase I Final Report, Volume II: Material Development" , AFML-TR 71-186, Vol. 2, March 1972 (AD

893 715L).

13, 1 Forest, J, D. , Fujimoto, A. F. and Foelsch, G. F. , "Advanced

Composite Applications for Spacecraft and Missiles, Phase I Final

Report, Volume I: Structural Development", AFML -TR- 7 1 - 1 86 , Vol. 1, March 1972,

13,2 Forest, J. D. and Varlas, M. , "Advanced Composite Applications

for Spacecraft and Missiles , Final Report" , AFML- TR - 72 - 2 78

,

March 1973.

161
13, 3 Christian, J. L, and Campbell, M. D, , "Mechanical and Physical

Properties of Several Advanced Metal-Matrix Composite Materials",

Advances in Cryogenic Engineering , Vol. 18, 1973, pp. 175- 183.

II - THERMOPHYSICAL PROPERTIES

Contract AF- 33(657) -9 160

Contractor : Air Force Materials Laboratory, Wright-Patterson Air

Force Base, Ohio,

Research Facility ; General Dynamic s? Astronautic s , San Diego,

California

The following reports are in a series entitled "Thermophysical Properties of Plastic Materials and Composites to Liquid Hydrogen Temperature (-423°F)",

14, 0 Haskins, J, F. , Campbell, M. C. , Hertz, J. and Percy, J. L. , MLTDR-64-33, Part I, June 1964 (AD 601 337).

15.0 Campbell, M, D, , Hertz, J., O'Barr, H. L. and Haskins, J. F. , MLTDR-64-33, Part II, March 1965 (X65-18921).

16.0 Campbell, M. D. , O'Barr, G. L. , Haskins, J, F. and Hertz, J., MLTDR-64-33, Part III, August 1965 (AD 468 155).

16. 1 Hertz, J. and Haskins, J. F. , "Thermal Conductivity of Reinforced Plastics at Cryogenic Temperatures", Advances in Cryogenic Engineering , Vol. 10, 1965, pp. 163-169.

16.2 Campbell, M. D. , "Thermal Expansion Characteristics of Some

Plastic Materials and Composites from Room Temperature to

-253°C", Advances in Cryogenic Engineering , Vol. 10, 1965,

pp. 154-162.

16.3 Campbell, M, D. , Haskins, J. F. , O'Barr, G, L. and Hertz, J.

" Thermophysical Properties of Reinforced Plastics at Cryogenic

Temperatures", Journal of Spacecraft , Vol. 3, April 1966, pp.

596-599.

(See also Ref, 13.0)

Contract F33 6 1 5 - 72 - C 1 388 (Work currently in progress)

Contractor : Air Force Materials Laboratory, Wright-Patterson Air

Force Base, Ohio.

Research Facility: General Dynamic s / Convair , San Diego, California.

162
17. 0 Forest, J, D, and Schaeffer, W, H. , "Advanced Composite Missile and

Space Design Data", General Dynamics Report GDCA- CHB 72 - 0 0 1 - 1

,

Progress Report No. 1, October 1972.

18,0 Forest, J. D, , "Advanced Composite Missile and Space Design Data",

General Dynamics Report GDCA - CHB 72 - 00 1 - 2 , Progress Report No. 2,

January 1973,

III - GENERAL BIBLIOGRAPHY

19.0 Aleck, B. , "Fiberglass-Overwrapped 2219 -T87 Aluminum Alloy LowPressure Cryogenic Tankage", Society of Aerospace Material and

Process Engineers National Technical Conference, Space Shuttle

Materials , Vol, 3, 1^71, pp, 131-134.

20. 0 Alfring, R. J. , Morris, E. E, and Landes, R. E. , "Cycle - Testing

of Boron Filament- Wo und Tanks", NASA CR-72899, National Aeronautics and Space Administration, Lewis Research Center, August

1971 (N71-38023).

21. 0 Barber, J. R. , "Design and Fabrication of Shadow Shield Systems for

Thermal Protection of Cryogenic Propellants " , NASA CR-72595, National

Aeronautics and Space Administration, Lewis Research Center, Cleveland,

Ohio, November 1969 (N70-25098).

22. 0 Baucom, R. M. , "Tensile Behavior of Boron Filament-Reinforced Epoxy

Rings and Belts", NASA TN D-5053, Langley Research Center, Hampton,

Virginia, March 1969 (N69-19918).

23. 0 Benton, W, , Carr, R, , Cohen, A, , Gustafson, G, , Lankton, C, and

Zeldin, B, , " Propellant Storability in Spac e" , RPL - TDR -64 - 75 (Final),'

Air Force Systems Command, Edwards Air Force Base, California,

June 1964 (AD 603 215).

24. 0 Brechna, H. , "Superconducting Magnets for High Energy Physics Applications", Proceedings 1 st International Cryogenic Conference, Tokyo

and Kyoto, Japan, April 1967 , Heywood Temple Industrial Publishers,

Ltd., London, 1968, pp. 119-129 ( CFSCI N67 - 36009)

.

25.0 Bullard, B. R. , "Cryogenic Tank Support Evaluation", NASA CR - 72546

,

NASA Lewis Research Center, Cleveland, Ohio, December 1969 (N70-13085).

26. 0 Campbell, M. D. , "Development of Thermal Expansion Capabilities and

the Investigation of Exp3,nsion Characteristics of Space Vehicle Materials",

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163
27.0 Campbell, M, D, , "Development of the Thermal Expansion Capabilities

and the Investigation of the Thermal Expansion Characteristics of Space

Vehicle Materials (II)", General Dynamic s / Astronautic s Report ERRAN-450, 26 December 1963.

28.0 Caren, R. P. , Coston, R. M. , Holmes, A. M. C. and Dubus, F. , "LowTemperature Tensile, Thermal Contraction, and Gaseous Hydrogen

Permeability Data on Hydrogen -Vapor Barrier Materials"", Advances in

Cryogenic Engineering , Vol. 10, 1965, pp. 171-180,

29.0 Chiao, T. T. and Moore, R. L. , "Tensile Properties of PRD-49 Fiber

in Epoxy Matrix" , Journal of Composite Materials , Vol. 6, October 1972,

pp. 547-551.

30.0 Cooper, G. A. and Sill wood, J. M. , "Multiple Fracture in a Steel Reinforced Epoxy Resin Composite", Journal of Materials Science , Vol. 7,

1972, pp. 325-333.

31,0 Darwish, F. and Tetelman, A. S. , "Mechanical Behavior of Si02-Epoxy

Composite", Conference Proceedings No. 63, Advisory Group for Aerospace Research and Development, Symposium on Composite Materials , Paper No. 9, Paris, April 1970 (Hartford House, London).

32. 0 Davis, J, G. and Zender, G. W. , "Mechanical Behavior of Carbon Fiber

Reinforced-Epoxy Composites", 12th National Symposium, Society of

Aerospace Material and Process Engineers , Vol, 12, 1967, Section AC- 10.

33. 0 Dervy, A. J. , "Reinforced Plastics of High Strength/Weight Ratio for

Space Applications", Technical Papers, 17th Annual Technical Conference,

Society of the Plastics Industry , Section 7-D, 1962, pp. 1-10.

34. 0 Fontana, M, G. , Bishop, S, M, and Spretnak, J. W, , "Investigation of

Mechanical Properties and Physical Metallurgy of Aircraft Alloys at Very

Low Temperatures, Part 5 - Mechanical Properties of Metals and a Plastic Laminate at Low Temperatures", AF Technical Report 5662,

Part 5, Materials Laboratory, Wright-Patterson Air Force Base, Ohio,

December 1953 (AD 27726).

35.0 Freeman, S, M, , "Properties of Vapor Barriers, Adhesives and Foams

at Cryogenic and Elevated Temperatures", Lockheed Aircraft Corporation

Report ER-5687, April 1962.

36,0 Freeman, W, T. and Campbell, M. D. , "Thermal Expansion Character

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for Testing and Materials, 1972, pp. 121-142.

164
37.0 Funk, C. W, and Dixon, C. E, , "Cryogenic Radiation Damage in

Structural Polymers", Transactions of the American Nuclear Society , Vol. 9, 1966, pp. 406.

38.0 Gille, J. P. , "Development of Advanced Materials for Integrated Tank

Insulation System for the Long Term Storage of Cryogens in Space",

NASA CR- 102570 (Final), National Aeronautics and Space Administration,

Huntsville, Alabama, September 1969 (N70-23348),

39. 0 Gleich, D. , "Development of a Filament-Overwrapped Cryoformed Metal

Pressure Vessel", NASA CR-72753, National Aeronautics and Space

Administration, Lewis Research Center, Cleveland, Ohio, January 1971

(N71-22401).

40.0 Gray, P. D. , Cornelius, G. K. , O'Donnell, J. D. , and Howard, W. W. , "Rockets in Space Environment, Volume 1: The Experimental Program",

RTD-TDR-63-1050, Aerojet General Corporation, February 1963 (N63-

20999).

41.0 Greer, F. , "Flexural Properties of Conolon 506 at Room Temperature,

-320°Fand -423°F", Convair/ Astronautics Report 55E 522, June 1961

(AD 677 565).

42.0 Hale, D. V., "Study of Thermal Conductivity Requirements: MSEC 20- ' Inch and 105-Inch Cryogenic Tank Analyses", NASA CR-61288, National

Aeronautics and Space Administration, Marshall Space Flight Center,

Alabama, June 1969 (N69-3581 1).

43,0 Hall, J, , "Cryogenic Tensile Tests- Epoxy Fiberglas", Douglas Aircraft

Company Report MP 1348, September 1961.

44.0 Hanson, M. P. , "Effects of Temperature and Creep Characteristics of

PRD-49 Fiber/Epoxy Composites", NASA TN D-7120, National Aeronautics and Space Administration, Lewis Research Center, Cleveland,

Ohio, November 1972 (N73-12607).

45.0 Hanson, M. P. , "Tensile and Cyclic Fatigue Properties of Graphite

Filament-Wound Pressure Vessels at Ambient and Cryogenic Temperatures", NASA TN D-5354, National Aeronautics and Space Administration,

Lewis Research Center, Cleveland, Ohio, July 1969 (N69-31300). (Identical to SAMPE Vol. 15, pp. 249-256).

46.0 Hanson, M. P., Richards, H. T. and Hickel, R. O. , "Preliminary Investigation of Filament-Wound Glas s -Reinforced Plastics and Liners for

Cryogenic Pressure Vessels", NASA TND-2741, National Aeronautic

s

and Space Administration, Lewis Research Center, Cleveland, Ohio,

March 1965,

165
47.0 Hanson, M. P., "Glass-, Boron-, and Graphite -Filament-Wound

Resin Composites and Liners for Cryogenic Pressure Vessels", NASA

TN D-4412, National Aeronautics and Space Administration, Lewis

Research Center , Cleveland, Ohio, February 1968. (Identical to NASA

TM X-52350, 1967. )

48, 0 Hanson, M. P. , "Static and Dynamic Fatigue Behavior of Glass FilamentWound Pressure Vessels at Ambient and Cryogenic Temperatures", NASA

TN D-5807, National Aeronautics and Space Administration, Lewis Research

Center, Cleveland, Ohio, May 1970 {CFSTI-CSCL-20K).

49.0 Haskins, J. F. and Hertz, J. , "Thermal Conductivity Testing of Coast

F-224-6 Phenolic -Fiberglas s Laminate", General Dynamics/Convair

Report No. AR- 592 - 1 -482 , July 1963,

50.0 Haskins, J. F. and Hurlich, A. , "Measured Values for the Coefficients

of Linear Expansion of Plycel 420 and Conolon 506. at Low Temperatures",

Convair/Astronautics Report MRG-154, May I960.

51.0 Haylett, J. W. , Rottmayer, E. and Butcher, I. , "Advanced Composite

Material Study for Millimeter Wavelength Antennas", Technical Report

AFML-TR-71 -205, Vol, 1, Air Force Materials Laboratory, WrightPatterson Air Force Base, Ohio, October 1971 (AD 893 368).

52,0 Haylett, C. E, , "Advanced Composite Material Study for Millimeter

Wavelength Antenna s , Volume II, Environmental Tests", AFML-TR-71-

205-Vol, 2, Air Force Materials Laboratory, Wright-Patterson Air

Force Base, Ohio, October 1971 (AD 893 358 L).

53.0 Herring, H. W. , Baucom, R. M. and Pride, R. A. , "Research on Boron

Filaments and Boron Reinforced Composites", 10th National Symposium , Society of Aerospace Material and Process Engineers , Vol, 10, 1966,

pp. B-21 to B-34.

54. 0 Hertz, J, , "Tensile Testing of Conolon 506 at Room and Sub-Zero Temperatures", Convair/Astronautics Repo rt MRF - 1 20 , December 1959.

55. 0 Hertz, J. , "Tensile Testing of Adlock 851 , Adlock PG-LA and Adlock

EG-11A-81A from -423°F to 78°F", Convair/Astronautics Report MRG

237, June 1961.

56.0 Hertz, J. , "Investigation of Potential Low Temperature Insulations",

General Dynamics/Astronautics Report GS/A-ERR-AN-668, December

1964.

57. 0 Hertz, J, , "The Effect of Cryogenic Temperatures on the Mechanical

Properties of Reinforced Plastic Laminates", General Dynamics Report

AR-592-1 -415, March 1963 (AD 40 5 170).

166
58,0 Hertz, J. , "Investigation into the High- Tempe ratur e Strength Degradation of Fiber -Reinforced Resin Composite During Ambient Aging",

General Dynamic s / Convair Report No. GDCA - DBG7 3 - 00 5 (Final Contract

NAS 8-27435), June 1973.

59.0 Hoggatt, J, T. , "Development of Cryogenic PRD -49 - 1 Filament- Wound

Tanks", NASA CR- 120835, National Aeronautics and Space Administration,

Lewis Research Center, Cleveland, Ohio, December 1971 (N72-24941).

60,0 Hoggatt, J. T, , "High Performance Filament Wound Composites for

Pressure Vessel Applications", Society of Aerospace Material and

Process Engineers National Technical Conference, Space Shuttle

Materials, Vol. 3, 1971, pp, 157- 167.

61,0 Hust, J, G. , "Thermal Conductivity of an Epoxy- Fiber gla s s Laminate",

Unpublished data (1973).

62.0 Johnston, H, L, and Brooks, H, E, , ^''Impact Strength of Various Metals

at Temperatures down to 20° Absolute", Ohio State University Cryogenic

Laboratory Report TR 264-17, May 1952.

63.0 Keller, C, W. , "Fiberglass Supports for Cryogenic Tanks^^ NASA CR120937 (Final), National Aeronautics and Space Administration, Lewis

Research Center , Cleveland, Ohio, October 1972 (N72-33564),

64.0 Kerlin, E, E, and Smith, E, T, , "Measured Effects of the Various Combinations of Nuclear Radiation, Vacuum and C ryotempteratures on Engineering Materials: Biennial Report" , NASA CR - 77 772 , National

Aeronautics and Space Administration, George C, Marshall Space Flight

Center, Huntsville, Alabama, July 1966 (N66-35963).

65, 0 Kerlin, E, E, and Smith, E. T. , "Measured Effects of the Various Combinations of Nuclear Radiation, Vacuum and C r yotemperatures on Engineering Materials: Annual Report", NASA CR-58830, National Aeronautics and Space Administration, George C. Marshall Space Flight

Center, Huntsville, Alabama, May 1964 (N64-33043).

66,0 Keys, R, D, , Kiefer, T, F. and Schwartzberg , F. R, , "Cryogenic

Properties of High-Strength Glas s -Reinforced Plastics", Advances in

Cryogenic Engineering , Vol, 11, 1966, pp. 470-477,

67,0 Krause, D. R. , "'Development of Lightweight Material Composites to

Insulate Cryogenic Tanks for 30 -Day Storage in Outer Space", NASA CR123797, National Aeronautics and Space Administration, George C.

Marshall Space Flight Center, Huntsville, Alabama, June 1972 (N72-

30495).

167
68.0 Krause, D. R. , Fredrickson, G. O. and Klevatt, P. L. , "Effects of

Cyclical Environments on High-Performance Multi-Layer Insulation

Materials", Society of Aerospace Material and Process Engineers

National Technical Conference, Space Shuttle Materials , Vol, 3, 1971,

pp. 639-643.

69. 0 Lantz, R. B. , "Materials for Filament Wound Cryogenic Pressure

Vessels", 6th National Symposium, Society of Aerospace Materials

and Process Engineers, Vol, 2, Engineering Paper No. 1750,

70, 0 Lavengood, R. E, and Anderson, R. M. , '^Matrix Properties Controlling

Torsional Fatigue Life of Fiber Reinforced Composites", Technical

Papers, 24th Annual Technical Conference, Society of the Plastics

Industry, Section 11-E, 1969, pp. 1-7.

71.0 Levin, V, A. , Naumenkov, P. G. and Shchitov, M. V, , "Some Properties

of Plastics at Low Temperatures", Plasticheskia Massy , Vol, 11, 1966,

pp. 64-65.

72.0 Luikov, A. V. , Vasiliev, L. L. and Shashkov, A. G, , "A Method for the

Simultaneous Determination of all Thermal Properties of Poor Heat

Conductors Over the Temperature Range 80 to 500 K", Proceedings , Third American Society of Mechanical Engineers Symposium, Purdue

University, March 1965, pp. 314-319,

73.0 Lyon, D. N. and Parrish, W. R. , "Low Temperature Thermal Conductivities of Two High Compressive Strength Materials", Cr yogenic s

,

Vol. 7, No. 1, 1967.

74. 0 Maher, L. E. , "Some Problems Arising from the Use of HydrogenFuelled Propulsion Systems", Journal of the Royal Aeronautical Society , Vol. 68, November 1964, pp. 765 -771.

75.0 McKannon, E. C. and Gause, R. L. , "Effects of Nuclear Radiation

and Cryogenic Temperatures on Non-Metallic Engineering Materials",

Journal of Spacecraft , Vol. 2, August 1965, pp. 558-564.

76.0 Morris, E. E. , "Glas s -Fiber -Reinforced Metallic Tanks for Cryogenic

Service", 12th National Symposium, Society of Aerospace Materials and

Process Engineers , Vol. 12, October 1967, Section AS -4. (Also NASA

CR-72224. )

77,0 Morris, E, E. , "The Performance of Glas s -Filament- Wound Pressure

Vessels with Metal Liners at Cryogenic Temperatures", Journal of

Materials , Vol. 4, December 1969, pp. 970-1004,

168
78.0 Morris, E, E. anclAlfring, R. J,, "Cryogenic Boron - Filament -Wound

Pressure Vessels", Composite Materials: Testing and Design, ASTM

STP 460 , American Society for Testing and Materials, 1969, pp. 430-443.

79,0 Morris, E. E. and Landes, R, E. , "Cryogenic Glas s -Filament-Wound

Tank Evaluation'^ NASA CR-72948 (Final), National Aeronautics and

Space Administration, NASA Lewis Research Center, Cleveland, Ohio,

July 1971 (N72-14696).

80. 0 Mowers, R. E, , Leib, J, H. and Sherman, S. , ''Program of Testing

Nonmetallic Materials at Cryogenic Temperatures'', Rocketdyne Corporation Report R-3498, Rocket Propulsion Laboratories, Edwards,

California, December 1962 (AD 294 772).

81.0 Nadler, M. A., Yoshino, S. Y, andDarms, F. J,, ''Boron/ Epoxy

Support Strut for Non-Integral Cryogenic Tankage", North American

Rockwell Space Division Report SD 68-9950 1 , February 25, 1969. (See

also 15th National Symposium SAMPE April 1969 and North American

Rockv-ell Report SD 995 2, 1968.)

82. 0 Nelson, L. F. , "Compressive Strength of Conolon 506 at -f 75°F and

-320 °F", Convair/Astronautics Report No. 27E 1336, January 1962.

83. 0 Nelson, L. R. , "Mechanical Properties of Adlock 851 at Room Temperature, 1000°, -320° and -423°F", Convair/Astronautics Report No. 55E

812, July 1961.

84.0 Nelson, P. T. and Archer, J. S. , "Graphite Reinforced Plastic EHF

Antenna", TRW Systems Group, Redondo Beach, California Report No.

99900-7128-RO-ll, December 1969.

85.0 Patten, P. M. , "Internal Insulation Liner Alteration", Douglas Aircraft

Company Report No. SM 45975, August 1964.

86.0 Perkins - Elmer Optical Group, Norwalk, Connecticut, Work -In - Pr ogres s on Contract No. F3 3 6 1 5 - 72 - C - 20 3 3 , Air Force Systems Command, WrightPatterson Air Force Base, Ohio,

87. 0 Pink, E. and Campbell, J, D. , "The Effect of Strain Rate and Temperature on the Deformation Behavior of Reinforced and Unreinforced Epoxy

Resin", Oxford University Department of Engineering Report No. 1040/72,

Oxford, England, July 1972 (N73- 10568).

88.0 Pirgon, O. , Wostenholm, G. H. and Yates, B,, "Thermal Expansion at

Elevated Temperatures, IV, Carbon-Fibre Composites", Journal of

Physics D: Applied Physics . Vol. 6, 1973, pp. 30^-321,

169
89.0 Pride, R. A., Stein, B. A. and Schmidt, F. W. , "Mechanical Properties of Polyimide -Resin/Glas s -Fiber Laminates for Various Time,

Temperature and Pressure Exposures", Technical Papers, 23rd Annual

Reinforced Plastics Technical and Management Conference, Washington,

DC, 1968, Section 17-c, pp. 1-8.

90.0 Ratcliffe, E. H. , "Thermal Conductivities of Plastics with Glass,

Asbestos and Cellulosic Fiber Reinforcements", Applied Material Re - search , Vol. 5, 1966, pp. 200-201,

91.0 Roseland, L. M. , "Materials for Cryogenic Usage", Technical Papers,

21st Annual Technical Conference, Society of the Plastics Industry , 1966,

Section 4-C, pp. 1-6.

92.0 Roseland, L, M. , "Investigation of Structural Properties at Cryogenic

Temperatures of Filament-Wound Pressure Vessels Containing Both

Organic and Flass Filaments", Douglas Aircraft Corporation Report

No. SM-48409, January 1966. -

93.0 Ross, J. E. , "Fiberglass Laminate- Ultimate Tensile and Flexural

Strength Tests at Room Temperature, -100°F and -320°F", Convair

Astronautics Report No. 7E 1687, June 1959 (AD 830 230),

94. 0 Sanders, R. H. and Weleff, W, , "Final Report on GTR-17 Effects of

Radiation on Organic Materials Irradiated in Liquid Hydrogen", AerojetGeneral Corporation Report No. RN-S-0327, March 1967.

95.0 Sanger, M. J. , Molho, R. and Howard, W, W. , "Exploratory Evaluation

of Filament- Wound Composites for Tankage of Rocket Oxidizers and

Fuels", AFML-TR-65 -381 , Air Force Materials Laboratory, WrightPatterson Air Force Base, Ohio, January 1966 (AD 477 455).

96.0 Sanger, M. J, and Reinhart, T. J. , "Development of Filament-Wound

Tankage for Rocket Oxidizers and Fuels", Technical Papers, 12th National

Symposium, Society of Aerospace Material and Process Engineers , 1967,

Section AS - 7.

97. 0 Sewell, J, J. and Kuno, J. K. , "Aerospace Use of Plastic Hardware and

Thermal Insulation", Technical Papers, 17th Annual Technical Conference , Society of the Plastics Industry , 1962, Section 7-A, pp. 1-14.

98. 0 Shriver, C. B, , "Design and Fabrication of an Internally Insulated Filament Wound Liquid Hydrogen Propellant Tank", NASA CR-127, National

Aeronautics and Space Administration, Washington, DC, November 1964

(N65-10775).

170
99. 0 Solty-siak, D. J. and Toth, J. M. , "Static Fatigue of Fiber Glass

Pressure Vessels from Arabient to Cryogenic Temperatures", Tech - nical Papers, 22nd Annual Technical Conference, Society of the

Plastics Industry, 1967, Section 14-E, pp. 1-14.

100. 0 Speare, J. C. , "Preliminary Sizing of Filament-Wound RNS Tanks",

Report No. TOR-0066 (5759 -07) - 13 , Space and Missile Systems Organization, Air Force Systems Command, Los Angeles Air Force

Station, Los Angeles, California, June 1970 (AD 872 626).

101.0 Steinhauer, R. A. , "Linear Thermal Expansion of 828CL 181 Cloth

Laminate", Douglas Aircraft Company Report No. MP 11,979, August

1961.

102. 0 Stinnett, W. D, , "Cryogenic Tensile Properties of Selected Materials",

NASA CR-71751, AEC-NASA Space Nuclear Propulsion Office, Report

No. 2712, January 1964 (N66-22816).

103. 0 Suezawa, Y, , Hojo, H. and Nakamura, K. , "Impact Characteristics

of Fiberglass Reinforced Plastics at Low Temperatures", Kagaku

Kogaku (Chemical Engineering, Japan), Vol. 33, 1969, pp. 1051-1059.

104. 0 Toth, J. M. , "Barrier Films for Filament -Wound Fiberglass Cryogenic

Vessels", Advances in Cryogenic Engineering , Vol. 1, 1964, pp. 537-

544.

105.0 Toth, J. M. and Soltysiak, D, J, , "Investigation of Smooth-Bonded

Metal Liners for Flass Fiber Filament -Wound Pressure Vessels",

NASA CR-72165 (Final), National Aeronautics and Space Administration,

Lewis Research Center, Cleveland, Ohio, May 1967 (N67-25070).

106. 0 Toth, J. M. , Sherman, W. C. and Soltysiak, D. J. , "Investigation of

Smooth-Bonded Metal Liners for Glass Fiber Filament-Wound Pressure

Vessels", Douglas Missile and Space Systems Division Report No. SM49384, Quarterly Report No. 3, Contract No. NAS 3 6293, NASA Lewis

Research Center, Cleveland, Ohio, April 1966.

107.0 Toth, J. M. , Sherman, W. C. and Soltysiak, D. J. , "Investigation of

Structural Properties of Fiber-Glass Filament- Wound Pressure Vessels

at Cryogenic Temperatures", NASA CR-54393, National Aeronautics

and Space Administration, Lewis Research Center, Cleveland, Ohio,

1965 (N65-35392).

108.0 Toth, J. M. and Barber, J. R. , "Structural Properties of Glas s -Fiber

Filament -Wound Cryogenic Pressure Vessels", Advances in Cryogenic

Engineering , Vol. 10, 1965, pp. 134-144.

171
109.0 Voloshenko-Klimovitskii, Yu. Ya. , Belyaev, Yu. A., L'vof, B. S.

and Schpakovskaya , E, I, , "Strength of Cold -Hardening GRPs Based

on PN-1 Resin Under Impact Tension at Normal (20°C) and Low (-196°C)

Temperatures", Plasticheski Massy , Vol, 6, 1964, pp. 39-40.

110.0 Voloshenko-Klimovitskii, Yu. Ya. , Belyaev, Yu. A. and Korenkov,

Yu. A. , "Impact Tensile Tests on Glass-Fibre Reinforced Plastics

at Normal and Low Temperatures", Plasticheski Massy , No, 5, 1963,

pp. 51-54.

111, 0 Watson, J. F, , Christian, J, L. and Hertz, J. , "Selection of Materials

for Cryogenic Applications in Missiles and Aerospace Vehicles"^,

Convair/ Astronautics Report No. MRG 132-1, February, I960.

112. 0 Weleff, W. , "Effect of Nuclear Radiation and Liquid Hydrogen on

Mechanical Properties of Three Phenolic Materials", Advances in

Cryogenic Engineering , Vol. 11, 1966, pp. 486-491,

113. 0 Weleff, W, , "Final Report, GTR-16 Radiation Effects Test on Structural

Materials at -423°F", Aerojet-General Corporation Report No, RN-S0290, November 1966.

IV - HANDBOOKS AND REVIEWS

114, 0 Coston, R, M. , "Handbook of Thermal Design Data for Multilayer

Insulation Systems", LMSC - A847882 , Vol. II (Final), George C.

Marshall Space Flight Center, Huntsville, Alabama, June 1967 (N67-

34910). . .

115. 0 Hertz, J, , "The Effect of Cryogenic Temperatures on the Mechanical

Properties of Reinforced Plastic Laminates", Society of Plastics

Engineers Journal , Vol, , February 1965, pp, 181-189.

116, 0 Hertz, J, and Knowles, D, , "Survey of Thermal Properties of Selected

Materials", General Dynamics/Convair Report AAL-65-008 (AR-504-

1-553), February 1965 (N65-31775).

117,0 Jurevic, W, G, and Rittenhouse, J, B. , "Structural Applications

Handbook", AFML TR-67-332, Air Force Materials Laboratory,

Wright-Patterson Air Force Base, Ohio, August 1968 (AD 804 585).

118. 0 Lackman, L, M. , Arvin, G, H. , et al, , "Advanced Composite Design

Guide, Third Edition, Volume IV: Materials", Air Force Materials

Laboratory, Wright-Patterson Air Force Base, Ohio, January 1973,

172
119.0 Landrock, A. H, , "Properties of Plastics and Related Materials at

Cryogenic Temperatures", Plastec Report No. 20, Plastics Technical

Evaluation Center, Picatinny Arsenal, Dover, New Jersey, July 1965

(AD 469 126).

120, 0 Maximovich, M. and Scheck, W. G, , "Data Summary and Reference

File for Graphite and Boron Reinforced Composite Materials", General

Dynamics/Convair Report No, GDCA-DBG71 -006 , 1971 (Contract NAS

8-26198, George C. Marshall Space Flight Center, Huntsville, Alabama).

121,0 Nored, D. L, , Hennings, G. , Sinclair, D, H. , Smith, G, T. , Smolak,

G, R. and Stofan, A, J, , "Storage and Handling of Cryogenic Fluids",

NASA Special Publication SP-5053, Proceedings of Conference on

Selected Technology for the Petroleum Industry, Lewis Research Center,

Cleveland, Ohio, December 1965 (N66-33674).

122,0 "Plastics for Aerospace Vehicles, Part 1, Reinforced Plastics", MILHDBK-17A, Department of Defense , Washington, DC, January 1971,

123. 0 Rittenhouse, J. B. and Singletary, J. B. , "Space Materials Handbook",

Third Edition, NASA Special Publication SP-3051, National Aeronautics

and Space Administration, Washington, DC, 1969 (Limited publication as AFML-TR-68-205).

124. 0 Schwartzberg, F. R. , Hertzog, R. G, , Osgood, S, H. , et al, , "Cryogenic Materials Data Handbook (Revised), Volume H" , AFML-TDR-64-

280-Vol. II (Revised), Air Force Materials Laboratory, Wright-Patterson Air Force Base, Ohio, July 1970 (AD 713 620).

V. MISCELLANEOUS REFERENCES

125. 0 Kastelic, J. R. , Hiltner, A. and Baer, E. , "Crazing, Yielding and

Fracture in Polycarbonate and Polyethelene Terephthalate at LowTemperatures", Journal of Macromolecular Science-Physics , B7(4),

1 973. pp. 679-703"^

126. 0 Relationships Between Structure and Mechanical Behavior in Polymeric

Solids . ASM Materials Science Seminar. Chicago, 111. , 1973 (in

preparation).

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COMPOSITE TESTING PROGRAM

PROCEDURES

The objective of the present effort in the composite testing program is

to make an initial evaluation of the tensile properties of current state-of-theart (commercial production) boron/epoxy, boron/aluminum and graphite/

epoxy composites over the 295K to 4K temperature range. The procedure is as follows:

a) design and construct a tensile cryostat capable of testing

composite specimens of the required length,

b) obtain the required test materials,

c) conduct initial tests at 295K to check out specimen

instrumentation,

d) conduct initial tests at 77K to work out any problems related

to specimen gripping or any other problems adversely affecting

test results at cryogenic temperatures, and

e) evaluate the tensile properties of the commercial composites

at temperatures of 295K, 77K and 4K. Test data shall include

tensile properties in the uniaxial longitudinal and uniaxial

transverse modes including poissons ratio values.

These procedures will lay the groundwork for future work, the

direction of which shall be determined by the results of the work at NBS (both

the literature survey and the experimental work) and by the results of the

experimental composite screening program currently being conducted by

General Electric under ARPA sponsorship. It is anticipated that future

efforts will involve fabrication and testing of experimental composites

optimized for cryogenic service, following the same procedures as for the

commercial products. A comparison of results will dictate the choice of

materials for inclusion in a test program aimed at producing design allowable

data.

176
CRYOSTAT CONSTRUCTION

The tensile cryostat has been completed. Figure 17 shows the unit

in cros s -section. The design is similar to that previously developed in our

laboratory and successfully used for many years. It differs only in the

incorporation of a thermal short designed to minimize helium loss by cooling

the aluminum face plate to which the cryostat barrel is attached. The use of

a fiberglass adapter plate between the face plate and the crosshead of the

tensile machine also provides a thermal barrier.

In operation, the instrumented specimen is carefully aligned in the

grips and hung from the upper pull rod. Instrumentation leads are then

passed out of the instrumentation port. The cryostat barrel is then hung onto

its pin supports, the end cap is attached, and the nut loosely attached to the

lower pull rod. The inner dewar is then attached and filled with liquid

nitrogen to precool the fixture. The nitrogen is removed from the inner dewar

both dewars are attached and the outer dewar is filled with liquid nitrogen.

Liquid helium is then transferred into the inner dewar through the vacuumjacketed transfer tube. Liquid height is monitored through clear sections of

the dewars. The test is then begun.

The present system allows for self-alignment of the specimen in the

cryostat. Where needed, the specimen can be forced into alignment by means

of a self-centering, tapered fixture in the bottom cap of the cryostat.

The cryostat is designed to take the full 20, 000 pound capability of

the tensile frame with which it is to be used. However, the present composite

study will utilize only a small amount of this capability.

Figure 18 shows the unit attached to the tensile frame before

installation of the dewars. Figure 19 shows the unit with both dewars

installed.

177
Fig. 17 C ros s -Section of Tensile Cryostat

1. Cryostat Barrel 5. 2. Heliuni Dewar 6.

3. Nitrogen Dewar 7. 4. Liquid Helium 8.

Liquid Nitrogen

Thernial Short

Fiberglass Baseplate

'^elium Transfer Line

9. 10.

11.

12.

13.

Instrumentation Port

Radiation Shields

Titanium Pull Rod

Specimen Grips

Specimen

178
SELECTION OF MATERIALS

The following materials are being donated by two cooperating

commercial composite fabricators. All materials are unidirectional.

Supplier Plies

5.6 mil Boron-Epoxy A 4 longitudinal on hand

5. 6 mil Boron-Epoxy A 11 transverse on hand

Graphite -Epoxy A 6 longitudinal on hand

Graphite -Epoxy A 16 transverse on hand

5.6 mil Boron-Epoxy B 6 longitudinal on hand

5.6 mil Boron-Epoxy B (12) transverse (committed)

5.6 mil Boron-Aluminum B ( 6)

longitudinal (committed)

Longitudinal and transverse tests will include ultimate tensile

strength, tensile modulus, and poissons ratio data. Svifficient material is

being obtained to prepare a minimum of 15 specimens of each composite

type and thickness, permitting five tests to be performed with each material

at 295K, 7K, and 4K.

A sufficient amount of each of the above materials shall be sent to the

General Electric Research Laboratory for inclusion in their flexural test

program, enabling a comparison to be made between materials representative

of current commercial production and the experimental composites being

evaluated at that laboratory in their current year's research effort.

179
180
SEMI-ANNUAL REPORT ON MATERIALS

RESEARCH IN SUPPORT OF SUPERCONDUCTING MACHINERY

FATIGUE AND FRACTURE TOUGHNESS

TESTING AT CRYOGENIC TEMPERATURES

Semi-Annual Report

March 1974

by

R. L. Tobler, R. P. Mikesell, R. L. Durcholz,

C. W. Fowlkes, and R. P. Reed

Cryogenic Division

NBS - Institute for Basic Standards

Boulder, Colorado
SUMMARY

I. The fatigue crack growth rates for Ti-6A1-4V, AISI 304 and AISI 316

alloys were essentially independent of temperature over the interval

298-76-4 K.

II. The fatigue crack growth rate of A-286 alloy was significantly

lower at 4 K as compared to room temperature.

III. At room temperature, the plane strain fracture toughness of

Ti-5Al-2.5Sn (K =66.9 ksi/in) exceeded that of Ti-6A1-4V.

The value of K for Ti-6A1-4V decreased from 44 ksi^in at

298 K to 35 ksi/in" at 4 K.

IV. The fracture toughness of alloys AISI 304 and 316 was higher at

'

4 K than at room temperature. These stainless steels exhibited

very tough fracture behavior and preliminary K^^ values were

obtained by J integral tests.

V. Valid J-integral tests on A-286 alloy revealed that K^^

decreased from 117 ksi/in at 298 K to a value of 95 ksi yin

at 4 K.

182
TABLE OF CONTENTS

Page

SUMMARY 182

LIST OF FIGURES 184

LIST OF TABLES 188

INTRODUCTION i 189

Fracture Toughness 189

Linear Elastic Fracture Mechanics 190

J-integral Fracture Criterion 192

MATERIALS 194

EXPERIMENTAL PROCEDURES 198

Tensile Tests 198

Low Temperature Fatigue 200

Fracture Toughness 211

RESULTS 215

Tensile Tests 215

Fatigue Crack Growth Rates 233

Fracture Toughness 249

DISCUSSION • 296

REFERENCES 305

183

/
LIST OF FIGURES

Page

Figure 1. Tensile Sample 199

Figure 2. Helium Fatigue and Fracture Cryostat 201

Figure 3. Cryogenic Fatigue Testing Apparatus 2 02

Figure 4. Helium Fatigue Cryostat 203

Figure 5. Specimen and Clip Gage Positioned in

Fatigue Cryostat 205

Figure 6. Compact Tensile Fracture Specimen used

for Ti-6A1-4V Alloy 207

Figure 7, Compact Tensile Specimen used for

AISI 304, AISI 316, and A-286 Alloys 2 08

Figure 8. Fracture Cryostat and Sample with 60,000 lb

(27,200 kg) Machine 212

Figure 9. Strength of Ti-6A1-4V as a Function of

Temperature 218

Figure 10. Strength of Ti-5Al-2.5Sn as a Function

of Temperature 219

Figure 11. Strength of AISI 304 and AISI 316 as a Function of Temperature 223

Figure 12. Stress-Strain Curves of Ti-6A1-4V 22 5

Figure 13. Stress-Strain Curves of Ti-5Al-2.5Sn 226

Figure 14. . Stress-Strain Curves of AISI 304 227

Figure 15. Stress-Strain Curves of AISI 310 228

Figure 16. Stress-Strain Curves of AISI 316 229

Figure 17. Work Hardening Coefficient of Ti-6A1-4V as

a Function of Temperature 231

184
LIST OF FIGURES (continued)

Page

Figure 18. Work Hardening Coefficient of AISI 30A and

AISI 316 as a Function of Temperature 232

Figure 19. Crack Growth Rate Data of T1-6A1-AV at 298 K ... . 234

Figure 20. Crack Growth Rate Data of T1-6A1-4V at

76 and 4 K 23 5

Figure 21. Crack Growth Rate Data of Tl-5Al-2.5Sn

at 298 K 237

Figure 22. Crack Growth Rate Data of AISI 304 at

298 K, 76 K, and 4 K 239

Figure 23. Crack Growth Rate Data of AISI 316 at

298 K, 76 K, and 4 K 240

Figure 24. Crack Growth Rate Data of AISI 304 at 298 K .... 241

Figure 25. Crack Growth Rate Data of AISI 304 at 76 K 242

Figure 26. Crack Growth Rate Data of AISI 304 at 4 K 243

Figure 27. Crack Growth Rate Data of AISI 316 at 298 K .... 244

Figure 28. Crack Growth Rate Data of AISI 316 at 76 K .... 245

Figure 29. Crack Growth Rate Data of AISI 316 at 4 K 246

Figure 30. Temperature Independence of Crack Growth Rate

of AISI 304 247

Figure 31. Crack Growth Rate Data of A-286 at

298 K, 76 K, and 4 K 248

Figure 32. Typical Load-Displacement Record at

4 K for T1-6A1-4V 251

Figure 33. Temperature Dependence of K^^ for T1-6A1-4V .... 254

Figure 34. Fracture Surface Appearance of T1-6A1-4V

Specimens 255

Figure 35. Fracture Surface Appearance of Tl-5Al-2.5Sn at

298 K 257

185
LIST OF FIGURES (continued)

Page

Figure 36. Load-Displacement Curve of AISI 304 at 298 K . . . . 259

Figure 37. Load-Displacement Curve of AISI 304 at 76 K .... 260

Figure 38. Load-Displacement Curve of AISI 304 at 4 K 261

Figure 39. Load-Displacement Curve of AISI 316 at 298 K ... . 262

Figure 40. Load-Displacement Curve of AISI 316 at 76 K .... 263

Figure 41. Load-Displacement Curve of AISI 316 at 4 K 264

Figure 42. The J-Integral as a Function of Crack

Extension for AISI 304 at 298, 76, and 4 K 265

Figure 43. Fracture Surfaces of AISI 304 at 300 and 76 K ... 267

Figure 44. Fracture Surfaces of AISI 304 at 4 K 268

Figure 45. Load Parameters of AISI 304 as a Function

of Temperature 273

Figure 46. Fracture Toughness Parameters of AISI 304 as a Function of Temperature 274

Figure 47. The J-Integral as a Function of Crack

Extension of AISI 316 at 300 and 76 K 276

Figure 48. Fracture Surfaces of AISI 316 at 298 and 77 K ... 277

Figure 49. Fracture Surfaces of AISI 316 at 4 K 278

Figure 50. The J-Integral as a Function of Crack

Extension of AISI 316 at 4 K 279

Figure 51. Load Parameters of AISI 316 as a Function

of Temperature 284

Figure 52. Fracture Toughness Parameters of AISI 316

as a Function of Temperature 285

Figure 53. Load-Displacement Curves of A-286 at

298, 76, and 4 K 286

186
LIST OF FIGURES (continued)

Page

Figure 54. The J-Integral as a Function of Crack

Extension of A-286 at 298 and 4K 288

Figure 55. The J-Integral as a Function of Crack

Extension of A-286 at 76 K 289

Figure 56. Fracture Surfaces of A-286 at 298 K 290

Figure 57. Fracture Surfaces of A-286 at 76 K 291

Figure 58. Fracture Surfaces of A-286 at4K 292

Figure 59. Fracture Toughness Parameters of A-286 as a Function of Temperature 294

Figure 60. Critical Stress Intensity Factor as a Function

of the Work Hardening Coefficient of

AISI 304 and AISI 316 299

Figure 61. Critical Stress Intensity Factor as a Function

of the Work Hardening Coefficient of

T1-6A1-4V 300

187
LIST OF TABLES

Page

1. Material Stock Description 195

2. Chemical Composition of Materials 196

3. Hardnesses of Materials 197

4. Strengths of Ti-6A1-4V and Ti-5Al-2.5Sn 216

5. Ductility of the Titanium Alloys 217

6. Strengths of AISI 304, 310, and 316 220

7. Strengths of A-286 221

8. Ductility of AISI 304, 310, and 316 and A-286 .... 222

9.. Work Hardening Coefficients of Ti-6A1-4V,

AISI 304, and AISI 316 230

10. Dimensional Criteria for Valid K of Titanium Alloys . . 250

11. Fracture Toughness of Ti-6A1-4V 253

12. The J-Integral of AISI 304 269

13. Dimensional Criteria for Valid J : AISI 304,

AISI 316, and A-286 270

14. Toughness Parameters of AISI 304 272

15. The J-Integral of AISI 316 281

16. Intermediate J-values of AISI 316 282

17. Toughness Parameters of AISI 316 283

18. Test Results of A-286 293

19. Dimensional Criteria for Valid K^^ of A-286 295

188
INTRODUCTION

The phenomenon of superconductivity is currently being exploited in

the design and construction of motors and generators. Superconducting

components are normally cooled with liquid helium, and since some

structural parts must operate at 4K, there is a need for property data in this

environment. Existing mechanical property data on structural alloys at 4K

are limited to tensile properties. Few fracture data have been reported at

this temperature. Therefore, it was decided to characterize the fracture

behavior of various structural materials immersed in liquid helium.

This report presents the fatigue crack growth rate and fracture

toughness data of the alloys AISI 304, AISI 316, A-286, Ti-6A1-4V, and

Ti-5Al-2. 5Sn. Both the growth rate and fracture toughness were determined

at 298, 76, and 4K. The stainless steels were chosen because they are

austenitic and, hence, very tough at low temperatures; the titanium alloys

and the A-286 were selected because of their high yield strengths.

Tensile tests were also conducted on the same material used with

the fracture tests in order to describe the material in terms of flow stress,

ductility, and work hardening behavior.

Work on the Ti-6A1~4V, the AISI 304, and the AISI 316 was previously

reported to the Naval Ship Research and Development Center. ^ The present

report includes a more involved study of these alloys as well as presentation

of data on the Ti"5Al"2. 5Sn, the AISI 310, and the A-286 alloy.

Fracture Toughness

Fracture tests are conducted to characterize the strength of materials

in the presence of sharp flaws. Fatigue cracks are a very common type of

189
flaw induced in structural components subjected to fluctuating loads.

A fatigue crack may reduce the strength of a structure to a value far

below the strength predicted by normal continuum failure criteria. The

unexpected and catastrophic fracture of a number of ships, storage tanks,

steam turbines, missiles and aircraft have underlined the importance of

relatively small cracks on strength. The problem is usually more severe

at low temperatures. The fact that the yield strength increases and

ductility decreases for most structural materials as temperature is

lowered increases the likelihood of brittle fracture.

Research into the mechanics of crack propagation or fracture has

resulted in the identification of parameters which describe the fracture

toughness and growth of fatigue cracks in materials. In this report,

two techniques have been applied to characterize fracture behavior:

1) linear elastic fracture mechanics, and

2) J-integral analysis.

Linear Elastic Fracture Mechanics

In cases where a material exhibits insignificant plastic deformation

prior to fracture, the techniques of linear elastic fracture mechanics

apply. The theory of linear elastic fracture mechanics is discussed in

references 2-4. The approach utilizes the stress intensity factor "K"

to describe fracture behavior. The parameter K characterizes the elastic

stress field near the crack tip which governs fracture behavior. For

a specimen of a given geometry, K is calculated from load, crack length,

and specimen dimensions. The stress intensity factor is the fundamental

parameter of interest because it relates all of these variables with a

single term.

190
Many normally ductile materials may fail in a relatively brittle

manner in the presence of a crack and under conditions of plane strain.

Consequently, a wide range of structural materials may be characterized

by linear elastic fracture mechanics. Plane strain conditions prevail

for "thick" specimens where the plastic zone at the crack tip is constrained

by the surrounding bulk of elastically loaded material. Under these conditions, materials which are loaded in the mode of crack opening exhibit

a critical stress intensity, K , at which the crack propagates catastroXL* phically without significant plastic deformation.

The fracture toughness parameter K is a material property and a

useful design criterion. Provided that size requirements are met, K^^

may easily be determined from fracture tests on specimens of standard

geometry. However, specimens which are too thin to establish a state

of plane strain do not render valid K data. Thin specimens are subject

XU

to conditions of plane stress or a mixed mode of plane strain and plane

stress. The fracture toughness (K^) measured on undersized specimens

is therefore less meaningful, varying with thickness and usually higher

than KjQ«

The ASTM guideline^ for assuring valid K^^ is that specimen thickness,

B, (and the crack length, a,) should be greater than or equal to 2.5 (K ^7 Q

iL y

It is apparent from this equation that the specimen size required to achieve

valid tests is prohibitively large for low yield strength, tough materials.

7 -2

For a stainless steel with a yield stress of 40,000 psi (2.8 x 10 Nm )

and estimated at 200 ksi in (220 x 10^ Nm~^ x m'''''^) , the required

specimen would be about 62 inches (158 cm) thick. A specimen this size

191
is obviously impractical, which means that direct measurement of K

for this category of materials is impossible within the framework of

linear elastic fracture mechanics.

J-integral Fracture Criterion

The concept of stress intensity loses its significance in the case

of "thin" specimens which do not provide sufficient constraint at the

crack tip. A different fracture criterion is needed to account for the

effects of large scale plasticity, i.e., the case where linear elastic

fracture mechanics is not applicable.

The recently developed J-integral technique is a valuable tool in

the treatment of thin, ductile materials. The theory of the J-integral

was originated by Rice^. Experimental verification of J as a fracture

criterion for "elastic-plastic" materials has been established primarily

7—8 9—11

by Begley and Landes , and others . The approach is based on the

premise that fracture behavior in elastic-plastic materials is governed

locally by conditions at the crack tip. J is the parameter characterizing

crack tip conditions; its role is analogous in some respects to the role

of K in linear elastic fracture mechanics.

J may be understood as the potential energy difference between two

identically loaded specimens of slightly different crack sizes. J is

represented as^

J = ^(1) da

where "u" is potential energy and "a" is the crack length. The parameter

J is also derived from the area under the load versus deflection curve

of a pre-cracked sample pulled in tension. The value of J just prior

to crack initiation is the critical value J-j-^* assuming that measurements

are made on samples of sufficient size.

192
The J-integral criterion has been successfully applied to cases of

large scale yielding, but only with plane strain conditions predominating,

The applicability of J to cases of plane stress has not been demonstrated.

Nevertheless, size requirements for J^^ tests are not the same as those

required by ASTM for K testing.

The rationale for using the J-integral as a fracture criterion is

8 9 that J may be converted to K by means of the relation '

:

where E is Young's modulus and v is Poissons ratio. J^^ yields the design

parameter K if the test specimen size and design are adequate. Therefore,

K may be determined for non-linear elastic materials from specimens

which are significantly smaller than those required for direct K^^ measurements .

193
MATERIALS

The materials tested in this study were Ti-6A1-4V, Ti-5Al-2.5Sn,

AISI types 304, 310 and 316 stainless steels, and the iron base alloy

ASTM A-286. These materials were obtained in various forms as given by

Table 1. With the exception of A-286, the materials were tested in the

as-received condition. Specimens of A-286 were machined from stock and

subsequently heat treated in two steps as follows:

(1) solution treat at 1650°F for 2 hrs, oil quench

(2) age at 1350°F for 16 hrs, air cool.

The chemical analyses of all materials are given in Table 2. Note

that two different heats of AISI 316 were used. Except for three samples

all of the 316 specimens were machined from the same plate, designated

as "heat 1" . Three samples were machined from the material designated

"heat 2", which was of slightly higher hardness. The hardnesses of

all materials tested are listed in Table 3.

194
Table 1. Original Condition

Material

Ti-6A1-4V

Ti-5Al-2.5Sn

AISI 304

AISI 310

AISI 316, heat 1

AISI 316, heat 2

A-286

Form

1" X 2" bar

4-3/4" square bar

1.5" plate

1.5" plate

1.5" X 4" bar

1.5" plate

4-1/2" square bar

forging

Note: 1 in = 2.54 cm

Condition as Received

Commercially annealed (1300°? 2 hrs)

,

cooled, descaled

Commercially annealed

Commercially annealed, descaled

Commercially annealed, descaled

Commercially annealed, descaled

Commercially annealed, descaled

Forged and annealed

195

Table

Material

Ti-6A1-4V

Ti-5Al-2.5Sn

AISI 304

AISI 310

AISI 316, Heat 1

AISI 316, Heat 2

A-286

3. Hardness of Materials Tested

Hardness

R 37.0

c R 33.8

c h

86.2

Rg 80.1

Rg 84.1

\ 85.7

R 30.8

197
EXPERIMENTAL PROCEDURES

Tensile Tests

Uniaxial tensile tests were performed with a 10,000 lb (4,500 kg)

Instron Universal testing machine. Tests were made with a stainless

12 steel cryostat, as previously described by Reed . Testing in liquid

nitrogen (76 K) was accomplished by immersion of specimen and cryostat

in a single metal dewar containing the cryogen. Testing in liquid helium

(4 K) was accomplished with a double glass dewar arrangement, the outer

dewar containing liquid nitrogen.

Load was monitored with a 10,000 lb (4,500 kg) commercial load

cell. An X-Y recorder was used to obtain load versus displacement curves

in the conventional manner.

Specimen extension was measured with a commercial clip-on extensometer

calibrated with a dial micrometer at each temperature. There was little

decrease in sensitivity with temperature: at 298 K, a strain increment

-4

of 3.2 x 10 was recorded for every small division of chart paper; at

-4

76 K, the sensitivity was 3.4 x 10 strain per division; and at 4 K,

-4

the sensitivity was 3.65 x 10 strain change per division. The extensometer

monitored strain to about 25%. Strain to fracture was monitored according

to a constant chart and crosshead speed. With Ti-6A1-4V, strain at 298 K

was measured with a clip-on Baldwin extensometer. This extensometer

-4

monitored a strain change of 1.2 x 10 per chart division.

With the exception of the alloy Ti-6A1-4V, all tensile samples were

13 machined according to ASTM specifications as shown in figure 1. Ti-6A1-4V

specimens had a reduced section that was 1-1/8 inch long, due to limitations in the width of stock. The Ti-6A1-4V specimens were tested in

198
I ,.500" ] .498

Figure 1. Tensile Sample (1 in. =2. 54 cm. )

199
both the longitudinal and transverse orientations. All other materials

were tested in the transverse orientation so that the plane of fracture

would be in the rolling direction. This orientation of the fracture

plane was the same as for a fracture toughness sample.

Low Teraperature Fatigue

A 20,000 lb (9,000 kg) cryostat was constructed primarily for fatigue

testing and some fracture tests in liquid helium. To minimize the consumption of costly liquid helium over relatively long term (4 hr.) fatigue

test periods, the steady state heat gain into the system had to be minimized.

This was accomplished by efficient use of low thermal conductivity, nonmetallic composite materials.

Figures 2, 3 and 4 illustrate the 4 K fatigue cryostat and associated

apparatus. The load carrying frame consists of two tubular stand-off

compression members which were designed for low thermal conductivity and

buckling resistance. The lower sections of these columns are AISI 304,

0.125 inch (0.3 cm) in wall thickness; the upper sections are fiberglass

reinforced plastic (FRP) , 0.250 inch (0.6 cm) in wall thickness. The

bridge linking the lower ends of the stand-off members is a maraging

steel. The lower specimen grip is inserted into the center of the bridge

and pinned in place. The upper specimen grip, also of maraging steel,

is threaded and attached directly to a titanium alloy pull rod.

A double dewar arrangement is used for fatigue and fracture tests

in liquid helium. Conventional glass dewars were rejected for fracture

testing applications on the basis of fragility. Instead, an inner helium

200
Pull Rod

Stand Offs

0 2 4

1 I I \—l

SCALE, inches

0 5 10

I \ I SCALE, cm

FRP Insulators

Grip ^

ASTM

Clip Gage^

Bridge

CT Fracture Specimen

FRP Helium Dewar, vacuum insulated

FRP Nitrogen Dewar,

foam insulated

Figure 2. Helium Fatigue and Fracture Cryostat

201
202

dewar was fabricated from fiberglass reinforced plastic. The two concentric

cylindrical walls of this dewar are each 0.1 inch (0.25 cm) thick. A

vacuum is maintained in the space between the walls by periodic mechanical

pumping, and five layers of aluminized mylar are located in the vacuum

space for added insulation. During 4 K tests, an outer dewar of liquid

nitrogen surrounds the inner dewar of liquid helium. This outer nitrogen

dewar was fabricated from fiberglass and epoxy. The space between the

walls of the outer dewar contains polyurethane foam and several layers of

aluminized mylar.

Thin copper baffles may be attached with hose clamps at locations

along the stand-off columns of the cryostat frame as shown in figure 5.

These baffles are intended to direct the flow of cold gas resulting from

boil off during transfer of liquid helium. They also serve as radiation

shields. Other features of the cryostat include two vents, a carbon

resistor liquid level indicator, a connection for gas pressurization,

and a lead-through for clip gage wiring.

Liquid nitrogen is used to precool the cryostat. Liquid helium is

transferred into the precooled inner dewar through a vacuum jacketed fill

tube. Under static load conditions, there was only a loss of 0.05-0.10

liter of liquid helium per hour. - The loss of helium is higher under

dynamic conditions. For fatigue loads up to 4,000 lbs (1,800 kg) the

rate of helium loss increases to 0.5 liter/hr (five times the rate for

static conditions) . For fatigue tests at loads greater than 10,000 lbs

(4,500 kg), cryogen loss is such that liquid helium is continuously

transferred into the dewar at a slow rate. The increased loss of cryogen

204
205
during fatigue is attributed primarily to heat generated by the work done

on the cryostat frame. The total helium consumption for this cryostat

has varied between 11 and 35 liters per test, depending on the amount

of reserve fluid added above the specimen, the efficiency of helium

transfer, and the applied fatigue load.

Compact tensile specimens were employed in all fatigue and fracture

tests. The geometry and dimensional tolerances for this specimen design

are described in ASTM standard E-399^. As shown in figure 6, the Ti-6A1-4V

specimens were 1 inch (2.54 cm) thick. Ti-5Al-2.5Sn alloy specimens

were 1.5 inch (3.81 cm) thick. For both titanium alloys, deflections

were measured between attachable knife edges located on opposite sides

of the crack mouth at the specimen edge.

AISI 304, AISI 316, and A-286 alloy specimens were 1.5 inch (3.81 cm)

thick. J-integral fracture tests were to be performed on these materials

subsequent to fatigue crack growth tests. A modification in notch

configuration was therefore introduced to enable measurement of loadline

deflection. Integral knife edges were machined at the loadline as shown

in figure 7.

All alloys were oriented with the machined notch in the rolling direction.

The depth of the machined notch was such that a^/W =0.3 for the majority

of specimens. Some specimens were machined to notch depths of a^/y = 0.46 - 0.

to facilitate precracking. The notch root radius was 0.005 inches (0.013 cm).

Fatigue crack growth tests were conducted at 298 K, 76 K, and 4 K,

using a 20,000 lb (9,000 kg) MTS servo-hydraulic test machine in the mode

of load control. The load was varied sinusoidally and the ratio of minimum

206
p

B = 1.00 inch

W = 2.00 inch

Figure 6. Compact Tensile Fracture Specimen used for

Ti-6A1-4V (1 in = 2.54 cm).

207
Figure 7. Compact Tensile Specimen Used for AISI 304, AISI 316,

and A-286 Alloys (1 in. = 2. 54 cm).

208
to maximum load (stress ratio) was maintained at R = 0.1. The loads

were measured to within 1% during dynamic testing by means of a peak

recording oscilloscope.

A conventional ASTM clip gage (figure 5) with a useful operating

range of 0.17 - 0.37 inches (0.4 - 0.9 cm) was used to measure deflection

in all fatigue tests. An extensometer calibrator was used to ascertain

that clip gage linearity met ASTM requirements^ for K^^ testing.

Changes in crack length were detected by means of compliance mea14

surements. The compliance technique is based on the fact that, for a

given load, specimen deflection increases with increasing crack length.

Experimental crack length-compliance correlations were determined for

each material and temperature. The experimental correlations were not

always in agreement with the theoretical solution as given by Roberts^^.

Discrepancies were attributed to differences in degree of crack front

uniformity. Crack fronts in actual test specimens always exhibit some

non-uniformity (crack front curvature) . The experimentally determined

compliance was correlated to an averaged crack length (a) , measured

according to ASTM standard E-399^.

Testing consisted of cooling the specimen and cryostat to the desired

temperature, initiating a crack, and propagating the crack at various

levels of stress intensity. For titanium alloy and A-286 alloy specimens,

cracks were always initiated at the temperature of testing. For stainless

steels, cracks were always initiated at room temperature. It is known

that crack growth rates may be temporarily retarded in going from a higher

to a lower temperature, or from a higher to a lower applied load. Therefore,

209
whenever such a change was made, the crack growth rate was allowed to

stabilize before data was accepted as valid.

The procedure for determining crack growth rates involved plotting

compliance on an X-Y recorder at intervals during a fatigue test. The

total number (N) of fatigue cycles sustained by the specimen was also noted.

The crack growth rate, da/dN, was obtained by graphical differentiation

of the a versus N curve.

The loads applied in fatigue varied from 3000 lbs (1350 kg) to

12,000 lbs (5400 kg), depending on the material and starting notch depth.

Fatigue cycling was usually conducted at rates of 20-26 Hz and no

variation in crack growth rates was detected as a result of frequency

variations in this range. A few high stress intensity fatigue tests were

conducted at frequencies no greater than 10 Hz.

Three to four specimens were used to generate data at each temperature. The majority of specimens were intended for subsequent fracture

tests, so that fatigue stress intensity was limited to 0.6 K^. A few

specimens were tested solely for fatigue crack growth data; these were

subjected to stress intensities approaching K^, and cracks were extended

to maximum depth of a/W = 0.7.

For specimens employed in subsequent J-integral tests, the stress

intensity factor was no greater than 35 ksi-^in during the final 3%

of crack growth. Fatigue loads were reduced to increase the sharpness of

terminated cracks. When this was done on AISI 304 and 316 stainless steel

alloys, the degree of crack front curvature increased significantly.

Pronounced crack front curvature invalidates the crack length-compliance

210
relationship; however, the fatigue crack growth rate data were not

significantly affected since data were not recorded during the terminal

stages of fatigue cracking. The degree of crack front curvature did not

increase for high crack growth rate tests where stress intensity was not

reduced

.

Fracture Toughness

Fracture toughness tests were carried out on titanium alloys using

the 20,000 lb (9,000 kg) MTS servo hydraulic test machine in the mode of

stroke control. K tests were conducted according to ASTM standards^.

Fracture tests were performed with the helium fatigue apparatus, subsequent to crack growth tests.

J-integral fracture toughness tests on precracked specimens of

AISI 304, AISI 316, and A-286 required loads in excess of 20,000 lbs

(9,000 kg). Fracture tests were performed on these alloys using a

60,000 lb (27,000 kg) controlled hydraulic tensile machine. A 60,000 lb

(27,000 kg) cryostat was designed similarly to the helium fatigue cryostat,

except that composites were not employed in the stand-off columns. The

frame is shown in figure 8 (left view). The stand-off columns are

AISI 321 stainless steel tubes, 1.5 inch CD. by 0.875 inch I.D. (3.7 x 2.2 cm).

This design provides rigidity and greater load capacity. The higher rate

of steady state heat gain is not a vital consideration since the cryostat

is intended for short term tests. About 11-15 liters of liquid helium

are consumed per test.

Load was monitored with a commercial load cell. A clip gage with

a useful operating range of 0.17-0.75 inches (O.A-2.5 cm) was used to

211
Figure 8. Fracture Cryostat and Sample with

60, 000 lb. (27, 200 kg. ) Machine

212
to measure deflection in all J-integral tests. The clip gage is shown

in figure 8 (right view) . This clip gage was evaluated using an extensometer

calibrator; precision corresponds to a maximum deviation of + 0.003 inches

(0.007 cm) from linearity over the useful operating range. The calibration

was performed at each testing temperature. Other equipment such as

specimen grips and dewars were identical to the apparatus described previously for fatigue testing.

J-integral tests were performed similarly to a procedure outlined

by Landes and Begley^^. According to this method, three or four nearly

identical specimens (having equivalent average crack lengths) are tested

at each temperature. The specimens are loaded to various increments of

stable crack extension and the tests interrupted. Specimens were pulled

apart for analysis after a heat tinting treatment oxidized the exposed

surfaces. Each value of J, as derived from the area under the load-displacement curve, is plotted as a function of the measured crack extension, Aa.

The critical J integral, J^^ is obtained by extrapolation of the J versus Aa

plot to the vertical "stretch zone" line. The stretch zone is the transition

region ahead of the fatigue precrack zone where the material deforms

under load prior to crack extension. The measured crack extension, A a, includes an increment contributed by the stretch zone.

For the compact tensile specimens described in this report, J was

calculated from the relation^ '

^ Bb '

where A is the area under the load-displacement curve to a particular value

of extension, B is specimen thickness, and b is the length of ligament or

the uncracked portion of the precracked sample. The area A was measured

using a planimeter.

213
Crack extension was measured in several different ways: the value

taken at the specimen centerline (A a, cent), the value taken as an

average of measurements at the center and edges or midpoints (A a, ave) , and

a "true" average determined by measuring the extension area with a planimeter

and dividing by specimen thickness. Results are presented for the AISI 304

and AISI 316 using the true average. The A-286 results are reported in

terms of " A a, center" and A a, average (of center and the midpoints)

Values of K_ and K were obtained from the secant value of the

Q max

sample load and the ultimate load, respectively. The secant value was

9 determined according to ASTM designation E 399-70T : the level in question

is the intersection with the load-displacement curve of a line that has

5% less slope than the tangent. Both K. and K are obtained from the ^ Q max

relation

K = -~-pr f (a/W)

,

BW '

where P is the load, B is the thickness, W is the length from the hole

centerline to the sample back, and "a" is the fracture crack length measured from the hole centerline. The "a" value used was an average of

three measurements : one taken at the center and the other two taken midway

between the center and outside.

214
RESULTS

Tensile Tests

The tensile strengths and yield strengths measured at 0.2 percent

offset for the titanium alloys are shown In Table 4. Ductility of the

titanium alloys Is shown In Table 5. Only a limited number of tests

were conducted since both the alloys had been thoroughly tested to 4 K,

Some testing was performed In both orientations at 298 K In order to

predict the mechanical behavior at low temperatures for samples oriented

In the transverse direction.

Figures 9 and 10 show the temperature dependence of strength for

1 7-1 Q 1 7 1 Q 20

T1-6A1-4V and annealed Tl-5Al-2.5Sn ' » , respectively. Tests done

by other Investigators are Included. With the exception of Tl-5Al-2.5Sn

at 76 K, the yield strengths are close to the values of the tensile

strengths. The data of this report can be extrapolated to the lower temperatures on noting the temperature trend of the other tests.

Table 6 shows the yield and tensile strengths of AISI 304, 310 and

AISI 316. The strengths of A-286 are shown In Table 7. Ductility of

AISI 304, 310 and 316, and A-286 Is shown In Table 8. The A-286 samples

reported by Warren are essentially the same hardness as the samples of

this report; thus, using Warren's data, the low temperature yield and

tensile strengths of A-286 for the transverse orientation can be approxl21-25

mated. Figure 11 shows the strength of AISI 304 and AISI 316.

22

Note that there is a maximum in the AISI 304 of Guntner and Reed ,

the strength dropping between 20 and 4 K.

215
Table 4. Strengths of Titanium Alloys

Alloy

Ti 6A1-4V

Ti 5A1-2.5 Sn

Temp,

(K)

Orientation 0.2% Yield*

Strength

Tensile Strength*

(psi)

(psi)

298 longitudinal 134,900 145,900

134.300 143,600

Av = 134,600 144,750

298 transverse 145,300 152,700

145,100 152.200

Av = 145,200 152.450

298 transverse 124,600 132,000

118,650 126.850

Av = 121,625 129,425

* 1 ksi = 0.689 X 10 Nm

216
Table 5. Ductility of Titanium Alloys (Annealed)

Alloy Author Temp.

(K)

Orientation Elongation

(%)

Reduction of

(%)

6A1-4V

20

Warren (1963) 1 o 298

76

20

Longitudinal

Longitudinal

Longitudinal

16.6

10.2

6.7

(4D)

(4D)

(4D)

47.4

40.6

31.0

Nachtigall-'-^(1974) 298

76

4

Longitudinal

Longitudinal

Longitudinal —

—

45.

36.

13.5

5Al-2.5Sn This report 298 Transverse 15.2 (4D) 32.4

Warren^^(1963) 298

76

20

Longitudinal

Longitudinal

Longitudinal

19.0

13.8

11.5

(4D)

(4D)

(4D)

43.6

29.5

18.3

Nachtigall-'-^(197A) 298

76

4

Longitudinal

Longitudinal

Longitudinal

45.

34.

30.

217
00 150 200

TEMPERATURE. K

300

Figure 9. Strength of Ti-6A1-4V as a Function of Temperature (1 psi. =6. 895 X 10 N/m )

218
300

X 10-

250

O Warren (1963), Longitudinal

A Nachtigall (1971), Longitudinal

This Report, Transverse

2.0

X 10

Tensile

VVV

Yield

1.5

200 \

\

Q.

ct: I—

aso

100

50

Tensi 1 e, 1.0

Yield

0,5

50 100 150

TEMPERATURE, K

200 250 300

Figure 10. Strength of Ti-5Al-2. 5Sn as a Function

of Temperature

219
Table 6. Strengths of AISI 304, AISI 310, and AISI 316

Alloy Temp.

(K)

Orientation 0.2% Yield Strength*

(psi)

Tensile Strength*

(psi)

AISI 304 298

298

76

Longitudinal

Av =

Transverse

Av =

Transverse

Av =

39,800

36,600

38,200

36,900

36,900

36,900

62,600

62.600

62,600

90,500

86.400

88,450

86,100

86.900

86,500

207,000

207,000

AISI 310 298 Transverse

Av =

35,200

34,200

34,700

78,950

78.700

78,800

AISI 316 298

76

Transverse

Av =

Transverse

Av =

Transverse

Av =

35,800

35,100

35,450

75,300

68,500

71,900

78,600

79.600

79,100

77,400

78.700

78,050

172,200

170.500

171,350

193,000

193,000

* 1 ksi = 0.689 X 10^ Nra ^

220
Table 7. Strengths of A-286

Author Condition Temp. Orientation 0.2% Yield*

Strength

(psi)

Tensile*

Strength

(psi)

This ReDort 1650°F

2 h-O.Q.,

Aged 1325 °F16 h-AC

298 K X L CLLLO V Wi. O C

Av =

93 300

83,100

88,200

156 J_ _y \J % 100\J \J 144,850

150,475

Warren^°(1963)

(Average

Values)

1800°F1-1/2 h-O.Q,

Aged 1350°F16 h-AC

298 Longitudinal 111,100 160,000

195 Longitudinal 120,300 175,800

76 Longitudinal 135,400 209,400

20 Longitudinal 150,200 235,300

4 Longitudinal 155,000(est.) 245,000(est.)

1 ksi = 0.689 X 10^ Nm

^

221
Table 8. Ductility of Stainless Steels

Alloy

AISI 304

Author

Guntner^^(1962)

Annealed

Temp,

(K)

298

76

4

Orientation Elongation

Longitudinal

Longitudinal

Longitudinal

84 (1"GL)

47 (1"GL)

44 (1"GL)

Reduction of Area

84

69

57

AISI 310 This report,

Annealed

Guntner^^(1962)

,

Annealed

298

298

76

4

Transverse

Longitudinal

Longitudinal

Longitudinal

46.9 (4D)

59 (4D)

68 (4D)

50 (4D)

70.7

71

50

41

AISI 316 This report.

Annealed

Desisto^^(1960),

Annealed

298

298

76

4

Transverse 75 (4D)

Longitudinal 46.5

Longitudinal 59

Longitudinal 50.5

73.2

76.5

74.7

58

A-286 Warren (1963) , 298 Longitudinal 25.4 (4D) 49.3

Solution Treated 76 Longitudinal 35.8 (4D) 49.2

1800°F, aged 20 Longitudinal 36.2 (4D) 42.8

at 1350°F,

R 30

c

222
300x10

1—I I I

r~v I I I I I I I I I I I r 1 I . T T I I I I t I I > I I I I I I I I I I I I I r r r I 40% Cold rolled, 36 j ^ 304, Annealed.Trans., This Report

4- 607oCold rolled, Rc 40 > A 304, Annealed ) Hoke

a 707oCold rolled, Rc 44 * ^ 316, Annealed i (1949)

O 304 L, Annealed 3l6,CoJd drawn257o,Kropschot(l956)

' O 304, Annealed, Long., Guntner (1962) • 316, Annealed, Desisto (I960)

316, Annealed,Trans., This Report

50 2b6

TEMPERATURE, K

300

Figure 11. Strength of AISI 304 and AISI 316 as a Function

of Temperature (lpsi=6. 895x 10^ N/m^)

22.3
Stress-Strain curves for the titanium alloys is shown in figures 12

and 13. Typically, there is little work hardening of these alloys even

to the lowest temperatures. Note the discontinuous yielding of Ti-5Al-2.5Sn

alloy at 4 K, as reported by Nachtigall.

Stress-strain curves for AISI 304, 310, and 316 are shown in figures

14, 15, and 16. The three stainless steel alloys work harden at a much

higher rate with temperature than do the titanium alloys. In particular,

the AISI 304 work hardens extremely at the lower temperatures. The

discontinuous yielding observed with the stainless steels at 4 K is a

result of localized heating and is dependent on the rate of strain. This

rate dependence is also observed at 20 K with AISI 310 (see figure 15)

Explanation of the discontinuous phenomenon has been presented by

„ . 1 .26 Basxnskx

Table 9 and figures 17 and 18 shows the work hardening coefficients of

Ti-6A1-4V, AISI 304, and AISI 316 as determined from the stress or load-strain

curves. The strain values were taken from the third, or temperature dependent,

stage of work hardening, i.e., the latter part of the curve to the ultimate.

The coefficients are derived from the empirical relation

where a ^ is the true stress, e ^ is the true strain, n is the work

hardening coefficient, and K is the strength coefficient. Taking the

logarithm of both sides of the above equation,

log a = log K + n log s .

224
300x10

250

200

Q.

io 150

UJ

DCcn

20 K

o Warren (1963) , Longitudinal

This Report .Longitudinal

This Report .Transverse

X Indicates Specimen Fractured

120

50

0 004 008 0.12

STRAIN

0.16 0.20 Q24

Figure 12. Stress -Strain Curves of Ti-6A1-

(1 psi. = 6,895 X lo'^N/m^)

4V

225
0 0.04 0.08 0.12 0.16 0.20 0.24

STRAIN

Figure 13. Stress -Strain Curves of Ti-5Al-2. 5Sn

226
300x10

0.3

STRAIN

Figure 14. Stress -Strain Curves of AISI 304

(1 psi. = 6. 895 X 10 ^ N/m ^)

227
300x10'

250

200

Q.

'C^ 150

CO

100

50

O Warren (1963), Longitudinal,

0.02 in/min Crosshead Speec

Except as Shown.

This Report, Transverse

0.2 0.4 0.6

STRAIN

0.8 1.0 1.2

Figure 15. Stress -Strain Curves of AISI 310

(1 psi.=6. 895 X 10^ N/m ) 228
Figure 16. Stress -Strain Curves of AISI 316

(1 psi. =6, 895 X lo'^ N/m ®)

229
Table 9. Work Hardening Coefficients of the Alloys

Alloy Temperature Orientation n Ave. Deviation,

Log True Stress

Ti-6A1-4V 298 K longitudinal 0.071 0.01%

2 0 (Warren ) 195 0.062 0.01

1 D U . UoU n no

U . lib U . Uj

Ti-6A1-4V 298 longitudinal 0.068 0.09

(this report) transverse n A C "7 A A/.

OA/ 304 o r\ o 298 longitudinal 0. 333 A on 0. 39

22

(Guntner, Reed ) 76 0.875 0.27

20 < 1(1.04) 0.06

4 0.654 0.08

304 298 longitudinal 0.392 0.53

(this report) 298 transverse 0.403 0.44

76 transverse 0.897 0.10

316 298 transverse 0.383 0.34

(this report) 76 0.649 0.04

4 0.585 0.12

230
0.12

Z 0.06

UJ

o

<

X

o

^ 0.04

0.02

0 50 100 150 200 250 300

TEMPERATURE, K •

Figure 17. Work Hardening Coefficient of Ti-6iAl-4V as a Function of Temperature

231
O Guntner(l962),304, Longitudinal

This Report ,304, Longitudinal

This Report ,304, Transverse

A This Report ,316, Transverse

0.3

02

0.1

0 50 100 150 200

TEMPERATURE, K

250 300

Figure 18. Work Hardening Coefficient of AISI 3 04 and AISI 316 as a Function of Temperature

232
Thus, if the points are plotted on log-log paper, "n" is the linear slope

of a straight line. A computer program was developed that enables the

coefficient to be determined through these steps. The straight line plot

is established by a linear least squares fit. The straight line data

fit is indicated by the numbers in the right hand column of Table 9.

These values represent the difference between a single number and the

straight line. The coefficients were also determined as a function of

orientation; there was a possibility that hardening due to dislocation

pile-up would vary with inclusion direction. For the AISI 304 at 20 K

(Guntner) , the coefficient calculated to be 1.04. Since obviously this

is wrong, the coefficient is reported to be a number less than 1. A

discussion relating the rate of work hardening to fracture toughness will

be presented in the last section.

Fatigue Crack Growth Rate

Ti-6A1-4V

The results obtained at room temperature are presented in figure 19,

and data obtained at 76 K and 4 K are shown together in figure 20. The

data are described in terms of the equation

da/dN = C( AK)^

where "a" is crack length, N is number of cycles and C and n are empirical

27

constants. This equation was originated by Paris and applies to all

materials tested in this program.

Superimposed on the data is a band which represents the room tempera28

ture data appearing in the Damage Tolerant Design Handbook . The room

temperature results are in good agreement with this reference; they fit

within the band of currently accepted data.

233
Figure 19. Crack Growth Rate Data of Ti-6A1-4V at

Z98 K. (1 ksi. JiK=l. 093 X 10^ INj/m ^ x m^^ j

. = 2. 54 cm. )

234
Figure 2 0. Crack Growth Rate Data of Ti-6A1-4V at

76 and 4 K. (1 ksi. yin. =1. 093 x 10° N/m^ x ^ and

1 in. =2. 54 cm. ) 235
Figure 20 shows that there is not a significant temperature dependence

of the crack growth rate for Ti-6A1-4V alloy at cryogenic temperatures.

The 76 K and A K data are plotted on the same graph, since any variation

due to temperature was of the same magnitude as the degree of scatter

among test specimens. In relation to the apparent temperature independence

29 of this property, it is significant to note that Wei, et al , found the crack

growth rate of a Ti-6A1-4V alloy to be temperature independent over the

interval from 298 K to 563 K.

Crack growth tests were conducted on two specimens immersed in a

solid-liquid mixture of alcohol and granulated dry ice, 195 K. These

tests were regarded as invalid because compliance curves were non-linear

and the crack fronts exhibited irregularities (crack branching) . These

results imply that fatigue cracking behavior of Ti-6A1-4V is deleteriously

influenced by the alcohol-dry ice environment.

30 Bucci, et al , measured the threshold value of stress intensity (the

value at which the rate of crack growth becomes vanishingly small) for

Ti-6A1-4V and found it to be 6-8 ksi-^in. This experimental result

compares well with the present room temperature data extrapolated one

decade lower in da/dN (figure 19)

Ti-5Al-2.5Sn

Crack growth rate tests of the alloy Ti-5Al-2.5Sn are not yet

completed. The results obtained for two specimens at room temperature

are presented in figure 21. Testing is currently in progress to expand

the crack growth rate data at 298 K, 76 K and 4 K.

236
20 40 60 80

STRESS INTENSITY FACTOR RANGE, AK, (ksi-in'"^)

100

Figure 21. Crack Gr^owth Rate of Ti-5Al-2. 5Sn at 298 K.

(1 ksi ^^=1. 093 X 10^ N/m^ x m^^ and 1 in. =2, 54 cm.)

237
AISI 304 and 316

Crack growth rate results for the stainless steels are summarized

in figures 22 and 23. The bands in these figures were drawn to envelop

all data obtained at a given test temperature. Original data are presented at specific test temperatures in figures 24-29.

The crack growth rates of AISI 304 and 316 are similar. Results

indicate that there is not a large temperature dependence of crack growth

rate for either material. The scatter bands are coincident and independent of temperature over the range of A K from about 20-40 ksi-^ in.

Above 40 ksi-^ in the scatter bands appear to diverge depending on the

temperature but it is uncertain whether this effect is real. The data

above 40 ksi-^in represent tests performed on only one specimen per

temperature. It is likely that additional tests will enlarge the scatter

bands in this region.

Since small temperature effects may be masked by the degree of

scatter among individual test specimens, it is revealing to test a single

sample at more than one temperature. This was investigated in a test

on AISI 304. Figure 30 shows the results. The specimen was precracked

as usual at room temperature and crack growth rate data were obtained at

76 K. The test was halted at a midway point and liquid nitrogen was replaced with liquid helium. Testing was then continued under identical

conditions of dynamic load and test frequency. The resulting a versus N

curve was essentially independent of temperature. This is reflected

in the results presented in Figure 31, where all data fit a single

equation of the form da/dN = C(^ K)^.

One of the interesting aspects of cryogenic tests on austenitic

stainless steels is the phenomenon of martensitic phase transformations.

238
^0-61 1 \ \ \ \ 1 I I I I 10 20 30 ^0 60 80 100

STRESS INTENSITY RANGE.AK. (ksi-in-^)

Figure 22. Crack GroNvth Rate Data of AISI_304 at 298 K, 76 K and 4 K. (1 ksi % in =1. 093 x 10° N/m" x m^^

and 1 in. =2. 54 cm. )

239
Figure 23. Crack Growth Rate Data of AISI 316 at

298 K, 76 K, and 4 K. (1 ksi vm=l. 093 x 10® N/m^x

and 1 in. =2. 54 cm. )

240
10"^

TYPE 30^ STAINLESS STEEL

AT 298 K

(Jo

o

UJ

I—

210"

<t: Q-Oai

D-

<c

CXI

<C

10

-6

10 20 30 ^0 50 80 100

STRESS INTENSITY RANGE. AK. (ksi-in^^)

Figure 24. Crack Growth Rate Data of AISI 304 at Z98 K.

(1 ksi Vin=l. 093 x 10*^ N/m^ x m^^ and 1 in. =2. 54 cm. )

TAX
10

o

>>u

T3

eel

:= 10

^

<cQ_O

D_

CJ3 I—

<C

10'

—

TYPE 304 STAINLESS STEEL —

AT 76 K ^ —

— • —

•

•

• •

• •

• —

m —

— • - •

•

— —

— • • — —

—

• — —

—

\

•

•1* — •• —

- • •

•

1 1 1 1 1 1 1 1

-

10 20 30 i|0 60

STRESS INTENSITY RANGE. AK. (ksi-in'

80 100

Figure 25. Crack Growth Rate Data of AISI 304 at 76 K.

(1 ksi Vki=l. 093 X 10^ N/m^ x m^^ and 1 in. =2. 54 cm

242
10

-14

TYPE 30^ STAINLESS STEEL

AT i| K

o

o

<0 -a

10

-5

<x. <C

Q-OCXl a.

<c

cd

CD

10

-6

10 20 30 ilO 60 80 100

STRESS INTENSITY RANGE. AK. (ksi-in''^)

Figure 26. Crack Growth Rate Data of AISI 3 04 at 4 K.

(1 ksi '/in=L 093 x 10 N/m"

243

and 1 in. =2. 54 cm.
10

TYPE 316 STAINLESS STEEL

AT 298 K

T3

I

10"^

<c

CD

<C

Q-Oor

CL.

<c

<c

10

-6 ±

10 20 30 40 50 80 100

STRESS INTENSITY RANGE. AK. (ksi-in^^)

Figure 27. Crack Growth Rate Data of AISI 316 at 298 K.

(1 ksi Vin=l. 093 x 10® N/m^ x m^^ and 1 in, =2. 54 cm.

)

244
TYPE 316 STAINLESS STEEL

AT 76 K

10 20 30 ^0 60 80 100

STRESS INTENSITY RANGE. AK. (ksi-in*'^)

Figure 28. Crack Growth Rate Data of AISl 3l6 at 76 K.

(1 ksi v''iri^l. 093 x 10*^ N/m"^ x m^^ and 1 in. =2. 54 cm. )

245
10

-4

o

>>> u

UJ

Ocx:

<c

•alio"

9

TYPE 316 STAINLESS STEEL

AT ^ K

10 20 30 ^0 60

STRESS INTENSITY RANGE. AK. ( ksi -i

n'

80 100

Figure Z9. Crack Growth Rate Data of AISI 316 at 4 K.

(1 ksi v'in=l. 093 x 10^ N/m^x m^^ and 1 in. =2, 54 cm.

246
10

-4

o

o

ta -a

<C

cc:

<c

D_OcnQ_

c_> <c

cac:

CD

I—

<c

10

-5

10

-6

• 4 K

A 76 K

A

\

A

A

A

TYPE 30i| STAINLESS STEEL

(all data from a single specimen)

10 20 30 ^0 60

STRESS INTENSITY RANGE.AK. (ksi-in'-

80 100

Figure 3 0. Temperature Independence of Crack Growth

Rate of AISI 3 04 (1 ksi Vin=l. 093 x 10® N/m^x m^'^ and

1 in. =Z. 54 cm. )

247
5x10

-4

1—Mill 1—I I II

10

-5

o

o

<CD

cC

Q.O

o

110-6

A - 786

Solution Treated and Aged

A

A

./

A

A

% J*

%*

r

•

• *

m

A

A

A

A

• 298K - 4 Specimens

A 76K - 3 Specimens

4K - 3 Specimens

10' J I I II I 8 10 20 40 60 80 100

Figure 31,

STRESS INTENSITY FACTOR RANGE, AK, (ksi-in"^)

Crack Growth Rate Data of A-286 at 298, 76, and 4K.

(1 ksi ^/in. = 1. 093 X 10^ N/m^ X m^^ and 1 in. =2. 54 cm)

248
An austenite-to-martensite transformation occurs as a result of plastic

22 deformation at low temperatures . It is clear that martensite formed

during fatigue tests. This was most obvious in the case of AISI 304

stainless steel tested at 4 K. Austenite is paramagnetic while martensite

is ferromagnetic, allowing qualitative detection by a magnetic device.

These magnetic measurements indicate appreciable martensitic transformation occurred when the stress intensity had reached approximately

The crack growth rate results for alloy A-286 are presented in

figure 31. This was the only alloy tested which exhibited a temperature

dependence of crack growth rate.

Figure 31 shows that the crack growth rate is significantly lower

at cryogenic temperatures as compared to room temperature. However,

comparison of data obtained at 76 K and 4 K show no significant difference.

Fracture Toughness

Ti-6A1-4V

One inch thick specimens of Ti-6A1-4V amply satisfied the ASTM

thickness criterion (Table 10) . Valid plane strain fracture toughness

tests have been performed at 298 K, 76 K, and 4 K.

A typical load-displacement record for a specimen at 4 K is shown

in figure 32. There is a slight departure from linearity during the

last 5% of the curve. The specimen failed spontaneously from the

maximum load

.

50 ksivin at 4 K.

A-286

249
Table 10. Dimensional Criteria for Valid K^^: Titanium Alloys

Alloy Temp.

(K) a

(in)

B

(in)

0.2% K

IC

(in)

Yield Strength . n-^

(psi) (ksiv/in)

'Ti-6A1-4V 298

76

4

1.12

1.12

1.12

1.0

1.0

1.0

145,200 43.65

237,325 35.32

278,300 35.03

0.22

0.06

0.04

Ti-5Al-2.5Sn 298 1.62 1.5 121,625 66.9 0.76

Notes: 1 in = 2.54 cm

1 ksi = 0.689 X 10^ Nm ^

1 ksi /in = 1.093 x 10^ — - mm

1/2

250
5,000 lb

Figure 32. Typical Load-Displacement Record at 4 K

for Ti-6A1-4V (1 lb. =0.454 kg)

251
The room temperature tests revealed a more noticeable pop-in behavior

prior to catastrophic crack propagation. The load-displacement records

at 298 K resembled the Type III record described in ASTM E-399^.

Table 11 presents the values of K^^ obtained at each temperature.

The results are presented in figure 33 as a function of temperature, along

with results obtained by other authors.

There is a 20% reduction in fracture toughness in going from 298 K

to 4 K. This reduction in K^^ is monotonic and there is no anomaly at

4 K.

Figure 34 illustrates the fracture surface appearance of four

specimens — one tested at each temperature. The fracture surfaces are

predominantly flat. Room temperature specimens display about 2% oblique

fracture; this decreases to about 1% at 4 K. The flat fracture region

is relatively rough at room temperature and becomes smoother with decreasing temperature.

Ti-5Al-2.5Sn

Cryogenic fracture toughness tests on Ti-5Al-2.5Sn alloy specimens

are currently in progress, but low temperature results are not yet available.

Preliminary results at room temperature show that the fracture toughness

of this alloy exceeds that of Ti-6A1-4V.

The 1.5 inch (3.81 cm) thick specimens are more than sufficient to

assure valid K^^ tests. The ASTM thickness requirement is 0.76 inches

(1.93 cm) as indicated in Table 10.

Two tests at room temperature have yielded K^^ of 62.15 ksi'^^ln

and 71.72 ksi-^In. The average K^^ is 66.9 ksi-^Itu These results

39 are in good agreement with the values reported by Witzell for a similar

Ti-5Al-2.5Sn alloy at 298 K.

252
Table 11. Fracture Toughness of Ti-6Al-AV

Specimen No. Test Temp. K^^, ksiyin +

A 12 298 K 43.09

A 15 " 42.71

A 16 " 45.00 t A 18 '^^ 43.80

A 13 195 K 38.45-

A 14 " 36.21*

A 6 77 K 34.33

A 7

" 34.64

A 8 " 37.13

A 9 " 35.19

A 10 4 K 35.51

A 11 " 34.79

A 21 " 34.80

t Fatigue stress intensity exceeded 0.6 K.

* Invalid, irregular crack front

+ 1 ksiyin = 1.093 X 10^ Nm"^ m"^''^

253
254
255
The load-displacement records of Ti-5Al-2.5Sn fracture specimens

were similar qualitatively to those of Ti-6A1-4V specimens at 298 K.

Slight non-linearity occurred during the final 5% of deflection. A

single pop-in occurred prior to catastrophic crack extension.

The fracture surface appearance of a Ti-5Al-2.5Sn specimen is

presented in figure 35. The region of oblique fracture (shear lip)

amounts to no more than 3% of specimen thickness.

Z56
257
AISI 304 and 316

As was stated in the section on experimental procedure, the J integral

is determined from the load displacement curve. Typical curves for AISI 304

are shown in figures 36-38; curves for AISI 316 are shown in figures 39-41.

As with the tensile tests, discontinuous yielding with both steels was

observed at 4 K. Although the phenomenon is rate dependent, it does exist

even to a crosshead rate of 0.005 inch per minute. A rate of 0.02 inch

was maintained at 4 K for the curves shown. For the most part, the rate

was held steadily throughout each test except at 4 K. At this temperature, the rate increased rapidly with each drop in load but could be

maintained steadily on the rise in load. The arrow marked "P^" on each

curve indicates the load at which the crack begins to extend. This

particular load is derived from the critical value of the J-integral;

derivation of the critical load will be discussed later. The load P shows

c a temperature dependence with both steels: as the temperature is lowered,

the critical load level increases from 298 to 76 K and then decreases

slightly from 76 to 4 K.

Figure 42 is a plot of the J integral for AISI 304 as a function of

the crack extension. Values of J are derived from the intersection of

J- \^ the J versus A a lines with the vertical stretch zone lines. The zone

lines are plotted at an average value of the stretch zones measured from

each specimen. This derivation of J^^ is somewhat of a departure from

16

the method of Landes and Begley who use a "flow" curve instead of a

zone line. They show that J values for very slightly extended cracks

fa.ll on the flow curve. However, data of this report, such as the lower

258
259
sOao . CO c ^ o

<; If) o>udU

rsj

II

-13

O

CO o Q II IO

•rH

260
261
262
263
264
oU3

oooLO o Ln o LT) o

n o CM cvj <—I •—

UL/qL-UL 'f

G 265
point at 298 K, did not fall on such a curve. The J value for this

particular point at the stretch zone line is of interest because it

represents the quantity of energy required to initiate the crack, after

having "stretched" the material a maximum amount. This point is the

critical value of J, J-ro* iC

Fracture surfaces for increasing extensions at 298, 76 and 4 K

are shown in figures 43 and 44. The depicted loads are the values at

which the particular tests were terminated. Notice the nature of the

crack profiles which developed quite differently at low temperatures,

particularly 4 K.

Results of the tests to determine the J integral of AISI 304 are

given in table 12. Three different extensions, as defined under "Experimental Procedures", are tabulated. The plot of figure 42 uses A a, true

average

.

Each of the critical values given can be considered to be a valid

J according to a tentative criterion proposed by Begley and Landes^.

Their criterion, based on limited experimental results of a few materials,

is the following:

ultimate strengths, a is the crack length, B is the sample thickness,

and b is the ligament or uncracked length. In other words, for the

critical value of J to actually be J-^r' above ratio should be equal

to or less than the particular dimensions. According to table 13,

a, B, b ^ a

where a is between 25 and 50, a flow is the average of the yield and

266
Figure 43. Fracture Surfaces of AISI 304 at 298 and 77 K.

267
304 Stainless

4 Kelvin

iiiiiaii^#iiiiiiiirfi^^

Figure 44. Fracture Surfaces of AISI 304 at 4 K

268
Table 12. Results for J Integral of AISI 304

Temp. Spec. Cr hd. Rate, J Aa Aa Aa Stretcl

in/min ,ln-lb , ^

. 2

) m

center

(.in;

out/cent

Cin;

true av

(in;

Zone

Cin;

298 5 0.02 9.37x10^ 0.074 0.129 0.137 0.025

298 4 0.02 9.63 0.080 0.111 0.061 0.032

298 3 0.02 2.74 0.029 0.021 0.026

298 2 0.02 13.38 0.151 0.185 0.109 0.031

298 18 0.02 25.40 0.299 0.207 0.232 0.022

IC

. In-lb . in

2.7x10-

Av = 0.027

76 14 0.02 16.3x10 0.252 0.351 0.326 0.015

76 15 0.02 20.2 0.410 0.479 0.446 0.017

76 12 0.02 14.4 0.251 0.329 0.399 0.012

76 20 0.02 13.1 0.130 0.214 0.142 0.009

76 19 0.02 8.04 0.061 0.077 0.065 0.010

Av = 0.013

6.5x10-

4 17 0.02 12.37x10-^ 0.389 0.460 0.500 0.0028

4 7 0.005 12.25 0.424 0.540 0.570 0.0014

4 13 0.02 13.17 0.733 0.609 0.540 0.0028

4 \ 10 0.02 10.6 0.245 0.252 0.267 0.0016

4 16 0.02 7.8 0.128 0.147 0.140 0.003

Av = 0.002

7.1x10-

Notes1 in = 2.54 cm

. in-lb . -^2 m-N

1 = 1.75 x 10 —r—

in m

269
Table 13. Dimensional Criteria for Valid J

Steel

IC"

Temp.

(K) a B

(in) (in)

b

(in)

Stainless Steel Samples

25

flow

(psi) (in-lb/in ) flow

(in)

50 J,

flow

(in)

AISI 298 1.89 1.50 1.11 61,700 2.7x10-^ 1.094 2.188

304

76 1.93 1.50 1.07 134,800 6.5 1.20 2.41

4 1.94 1.50 1.06 151,000 7.1 1.18 2.35

AISI 298 1.90 1.50 1.10 56,750 3.41x10^ 1.50 3.00

316

76 1.88 1.50 1.12 121,625 7.20 1.48 2.96

4 1.89 1.50 1.11 136,050 5.0 0.92 1.84

298 1.62 1.49 1.38 119,350 4.5x10^ 0.095 0.19

A-286

76 1.65 1.50 1.35 152,225 3.54 0.06 0.12

4 1.65 1.50 1.35 176,725 2.90 0.04 0.08

Notes 1 in = 2.54 cm

1 ksi = 0.689 Nm

, in-lb

-2

. 2

in

= 1.75 X 10

2 m-N

m

270
AISI 304 does meet the Begley-Landes requirement for the most part if

a is chosen to be 25. Thus, considering the critical value of J to

7 8

be the latter can be converted to K^^ through the relation '

E

where E is Youngs modulus and v is Poissons ratio.

Table 14 tabulates the values of K for AISI 304 as well as other

toughness parameters. The load required for crack initiation is derived

from the load displacement curve knowing the value of Jj^q' The parameter

is determined from the relation

J'^Q = ^

Bb '

where A is the area under the load-displacement curve to the ultimate

load. This particular parameter is convenient to measure since only

one test per temperature is required. The value of decreases with

decreasing temperature, as opposed to the manner in which J^^ and K^^

change with temperature. This temperature decrease of is due to the

fact that the load reaches the ultimate value more rapidly the lower

the temperature.

Figure 45 is a temperature plot of the load parameters P , P ,

and Pq for AISI 304. The critical load approaches the ultimate load

level at 76 K, and then at 4 K, P^ has "moved" beyond the ultimate.

Figure 46 is a plot of the toughness parameters as a function of temperature.

271
Table 14. Toughness Parameters of AISI 304

Temp.

(K)

298

298

298

298

298

76

76

76

76

76

Spec,

5

4

3

2 18

14

15

12

20

19

17

7 13

10

16

Load for

Crack

Initiation

(lb)

13,850

K, K

IC Q Q max max

(ksix^) (lb) (ksiN/in) (lb) (ksij±ii)

285

27,200 457.1

28,600 476.4

. in-lb . in

7,900 48.

y , DDV DH . J 8,100 49.5

bU

.

iy , u /

J

iiu

.

10 0 Q-...1 n 13 . JoxiU

£5 , ZUU /. c hD . J

11/. 114

.

on 1 Q ZU. ly

AV — c^Jl . 110 IIZ

.

It 7 Q ID . /o

iJ , UUU ol

.

T7 "7 r>r> Z/ , /UU IT/. 3 o o n«i n*' o . iUxlU

12,500 77.5 27,400 173.

12,100 76. 27,500 174. 8.05

10,200 63. 28,100 174. 9.45

12,000 67.5

Av = 75. 174. 8.60

19,600 122. 31,900 199. 4.9x10^

16,600 112. 28,600 197.

17,200 109. 30,550 190.5

21,550 127. 32,400 194. 4.81

17,500 93. 33,400 195. 5.2

Av = 113. 194. 4.97

Notes: 1 lb = 4.448 Newtons

1 ksiv/in = 1.093 x 10^ ]^/m^-m^^'^

T in- lb --2 m-N

1 z— = 1.75 X 10 —

in' m

272


The graph of the J integral versus crack extension for AISI 316

at 300 and 76 K is presented in Figure 47. Two heats were tested at

300 K. In heat 2, although only 3 data points were obtained, one specimen

represents a point just after crack opening initiated. Since there are

no points for heat 1 near the stretch line and since heat 1 data points

do exhibit considerable scatter, the J^^ value obtained from heat 2

data is considered more reliable.

The fracture surfaces of AISI 316 at 300 and 76 K are shown in

Figure 48; the surface of a 4 K specimen is presented in Figure 49. The

heat tinting of the 4 K specimen left well-defined ridges very apparent.

The number of these ridges exactly correspond to the number of major

load drops obtained from the load-deflection curve (see Figure 41)

Presumably these ridges, defined by differing heat tint coloring, result

from varying martensite concentrations along the fracture surface. These

sudden, repeated crack advances are thought to result from internal,

localized adiabatic heating, leading to reduced flow stresses at the

higher temperatures and sudden shearing with a concomitant drop in load.

Therefore, by measurement of crack extension to each ridge and of J

values to each associated load drop, it is possible to obtain a series

of J- A a points from one specimen.

Figure 50 shows the J plots for three specimens of AISI 316 at

4 K. Each ordinate data point represents the J value for the integrated

area to the bottom of a particular load drop. The corresponding A a is

that extension measured from the end of the fatigue crack at the center

to the particular ridge, as shown in Figure 49. The point corresponding

to the maximum extension (i.e., at the interface between the heat tint

area and the remaining part of the fracture) , falls in line with the points

illustrated.

275
276
277
yure 49. Fracture Surfaces of AISI 316 at 4 K

278
o

o -o tjn o LO o tn o

UL/qL-UL 'r

279
Results for the J-integral of AISI 316 are shown in table 15.

Intermediate J values for AISI 316 at 4 K, each point corresponding to

the particular load drop, are shown in table 16. In spite of the steel

not conforming completely to the dimensional criterion (see table 13)

,

the J^^ values were converted to K^^^* Table 17 tabulates the values

of as well as the other toughness parameters.

A temperature plot of the load parameters of AISI 316 is shown in

figure 51. The toughness parameters as a function of temperature are

plotted in figure 52. The P and P curves are nearly parallel. Notice

that the J^^ and K^^ values decrease between 76 and 4 K, as opposed

to the behavior of AISI 304.

A-286

The alloy A-286 was found to be less tough than AISI 304 and 316.

In fact, the alloy exhibited "pop-in" at 76 and 4 K.

Figure 53 shows the load displacement curves for A-286 at 298, 76,

and 4 K. Since it was desirable to obtain J-values for very small ex*-

tensions, most tests were terminated shortly after the load-displacement

curve would flatten out. At 76 K, the first drop in load was accompanied

by a fairly low sounding "pop"; the load then continued at nearly the same

level until complete fracture occurred with a very loud sounding "pop"

.

The other two tests at 76 K were terminated before pop-in occurred. At

4 K, slight extension occurred with 2 tests before termination; very low

pops were heard. A louder, more definite pop occurred with a third test

at 4 K; a relatively larger extension accompanied the pop-in.

280
Table 15. Results for J Integral of AISI 316

Temp. Spec. Cr hd. Rate, J Aa Aa Aa Stretch

(K) in/min ,in-lb , center out/cent true av Zone

and Cond. in

(in) (in) (in) (in)

298 1 0.02, Heat 1 11.12x10^ 0.263 0.264 0.347 0.013

298 10 II 10.68 0.237 0.240 0.30 0.013

298 3

II 14.6 0.407 0.381 0.386 0.020

298 11 II 8.45 0.140 0.136 0.107 0.015

298 14 0.02, Heat 2 13.1 0.078 0.137 0.10 0.030

298 15 M 9.8 0.068 0.097 0.094 0.027

298 17 II 3.41 0.024 0.024

Av = 0.020

IC

, ln-lb .

4 2''

in

3.41x10-

76 12 0.02, Heat 1 10.2x10 0.184 0.148 0.167 0.010

76 2

II 13.65 0.358 0.240 0.30 0.008

76 4

II 12.32 0.277 0.217 0.30 0.011

76 8

II 8.53 0.155 0.117 0.139 0.014

76 13 It 10.47 0.115 0.106 0.120 0.013

Av = 0.011

4 5 0.02, Heat 1 14.8x10-^ 0.709 0.578 0.004

4 6

M 12.66 0.475 0.309 0.002

4 7 0.005, Heat 1 15.16 0.643 0.492 0.007

Av = 0.004

Notes: 1 in = 2.54 cm

. in-lb , _c irv2 m-N

1 2~ ~ 1*75 X 10 —

2"

in m

281
Table 16. Intermediate J integral values of AISI 316 at 4 K

, 2*

Specimen No. Load Drop No. J, in-lb/ln A a, in*

1 5.10 0.065

2 '6.23 0.120

3 7.62 0.176

4 8.59 0.229

5 9.65 0.336

6 10.8 0.433

7 12.04 0.516

8 12.95 0.581

9 13.6 0.646

10 14.4 0.671

1 5.39 0.032

2 6.09 0.092

3 7.31 0.150

4 8.65 0.214

5 10.10 0.299

6 11.59 0.374

7 12.66 0.454

1 5.07 0.053

2 6.48 0.141

3 7.75 0.168

4 9.44 0.238

5 11.5 0.350

6 12.88 0.452

7 13.69 0.520

8 14.76 0.589

* 1 inch = 2.54 cm, i-lllzi^ ^ ^^^3 ^ ^^2 m; in m'

282
Table 17. Toughness Parameters of AISI 316

Spec,

76

76

76

76

76

1

10

3

11

14

15

17

12

2

4

8

13

Load for

Crack

Initiation

(lb)

14,400

29,200

^IC ^0 P K

lUclA ILiCt A. X

/ (ksiV in) (lb) (ksiv

/M^) (lb)

/ (ksi/in) .in-lb

in^

343 1 9 080 55 5 14 600 90.

3

7 14x10

6,800 40. 15,200 89. 6.95

8 580 48 15 150 89

7 075 40 5 8 45

7 550 46

6 250 36 5

44 5 O 7 • 7 SI Av (Heat 2) = 42.3

515.6 17,050 93. 30,250 164. 7.09x10-^

29,400 166.

15,000 89. 27,800 166. 7.94

17,100 91.5 30,050 159. 6.43

18,150 108.

Av = 95.4 164. 7.15

4 5 31,600 430. 16,700 89.5 32,200 173. 4.41x10

4 6 16,000 89.5 32,550 183. 4.62

4 7 17,600 100. 33,400 190. 4.02

Av = 89.5 178. 4.52

for 0.02 in/min (see table 15)

1 lb = 4.448 Newtons

1 ksi/in = 1.093 x 10^ Vi/m^-ra-^^

1 in-lb/in^ = 1.75 x 10^ ^

m

283

0 50 100 150 200 250

TEMPERATURE, K

Figure 52. Fracture Toughness Parameters of AISI 316 as

a Function of Temperature (1 ksi Vin=l. 093 x 10

N/m^ X and 1 in-lb/in^ =1 . 75 x 10^ m-N/m^)

285
DISPLACEMENT, cm

0 0.02 0.04 0.06 0.08

0 0.005 0.010 0.015 0.020 0.025 0.030 0.035

DISPLACEMENT, in

Figure '3. Load-Displacement Curves of A-286 at 300, 76, and 4K.

286
Figures 54 and 55 are plots of the J-integral of A-286 as a function

of extension. The J^^ values are obtained from the zero value of extension, i.e., there is essentially no stretch zone with this alloy.

Notice that there are extremely small extensions at 76 K.

Fracture surfaces of A-286 tested at 298, 76, and 4 K are shown in

figures 56, 57, and 58, respectively. Observe the very small extensions

with most of the samples.

Test results of A-286 are shown in table 18. The K^^ values are

obtained from J^^ through the relation involving Youngs modulus and Poissons

ratio. What is called "K " is derived directly from the load at pop-in.

pop

It is interesting to note that the K values are higher than the K^^ ° pop ° IC

values. Figure 59 is a plot of the toughness parameters as a function

of temperature. As opposed to the AISI 304 and 316, the J^^ and K^^ decrease

with a decrease in temperature.

Referring back to table 13, it is seen that A-286 easily meets the

requirements for a valid J-rp. However, the alloy does not comply with the

requirements for a valid K (see table 19) according to ASTM specification

E-399 that states

a, B > 2

•'ft)

The K is close to being valid at 4 K. It is possible that the flow stress

is higher than estimated at 4 K, thus giving a valid K at this temperature.

287
288
289
igure 56. Fracture Surfaces of ASTM A-286 at 298 K

290
291
I

29Z
Table 18. Test Results of A-286

Temp.

(K)

298

298

298

Spec.

Q

(lb)

18,370

18,500

(ksiv/in) . in-lb v in

Aa

center

(in)

25,000 105. 5.07x10 0.026

78.

80.

Av = 87.7

76 13

76 9 27,150 122.

76 6 27,900 117.

Av = 119.5

4 7 26,220 114.

4 0 Same as 120.

Ppop

4 10 II 117.

Av = 117.

7.25

4.08

0.124

0.079

Aa

av

(in)

0.022

0.080

0.048

IC

K

IC

P K

pop pop

4.62x10 0.004

3.96 0.0015

5.71x10^ 0.156 0.148

3.51 0.049 0.039

3.73 0.035 0.034

. in-lb > (kst/in) (lb) (ksi^)

in

4.50x10

for

A a/center

4.25x10

for

Aa/av

Av =

3.54x10'

Av =

2.78x10^

for

A a/center

2.90x10^

for

A a/av

Av =

120.9

117.5

119.2

107.2

107.2

29,420 133.2

133.2

95.05 28,280 123.

27,650 120.

97.08 26,980 117_.

96.06 120.

Notes

;

1 lb = 4.448 Newtons

1 ksi^^= 1.093 X 10^ N/m^ - m^^^

T in-lb , --,2 m-N

1 2~ ~ ^'7^ X 10 —

2"

in m

1 in = 2.54 cm

T in-lb _ -„2 m-N

1 2~ 1«75 x 10 —

Jin m

293
TEMPERATURE, K

Figure 59. Fracture Toughness Parameters of A-286 as a Function of

Temperature

294
Temp.

(K)

Table 19. Dimensional Criteria for Valid K : A-286

a B 0.2% 2.5 /K^A K 2.5 /K \

(in) (in)

298 1.62 1.49 88,200 120.9 4.69

76 1.65 1.50 107,500 107.2 2.49 133.2 3.84

4 1.65 1.50 123,050 97.08 1.55 120. 2.38

notes: 1 in = 2.54 cm

1 ksi = 0.689 Nm~^

1 ksi/in = 1.093 x 10^ N/m^ - to'^'^

295
DISCUSSION

The discontinuous yielding observed with AISI 304 and 316 at 4 K in

both tensile and fracture toughness tests has been attributed to a combination of highly localized heating and the temperature dependence of the

22 26 flow stress ' . The most acceptable explanation of this phenomenon may

be described as follows: at 4 K the specific heat and thermal conductivity

of alloys are very low; therefore, any heat generated internally cannot

easily escape and local warming occurs. Temperature rises in tensile

40 specimens as high as 40 K have been measured . In some alloys the

temperature dependence of the flow stress at low temperatures is quite

large, insuring that, if local temperature rises do occur, plastic deformation occurs more easily. In tensile specimens the local temperature

rises and associated plastic deformation surges do occur at intervals

along the reduced specimen length, much in the same manner that Luders

bands propagate along the specimen length (see figures 14-16) . These

40 effects are a function of the strain rate . Observation of figures 38

and 41 indicate that the same phenomena occur in fracture toughness tests

of austenitic stainless steels. Testing a few specimens at reduced load

rates revealed that, again, the magnitude of the load drops (higher load

rates producing higher load drops) was influenced by load rate. The

surprising evidence was that, at least in AISI 316, very distinct crack

front ridges could be directly associated with these load drops. This

indicates that the crack front propagated discontinuously , influenced

considerably by local specimen temperature.

296
Another possible indication of local temperature effects is the

nature of the crack front geometry after partial propagation in J integral

tests (see figures 49 and 50). The unusual tendency of reduced propagation

rate in the specimen center and increased crack propagation tendencies

nearer the surfaces are very apparent. Now, if local heating did occur,

the interior would be expected to be warmer than the exterior (due to

reduced specimen thermal conductivity and lower heat capacity near 4 K)

.

If the propagation rates were temperature dependent (in the temperature

range 4 to about 50 K)^then crack fronts such as observed could be rationalized.

Another explanation, of course, is that loading conditions, after crack

growth to achieve an a/w = 0.8 - 0.9, tend to produce such a crack front.

The da/dN fatigue results do not portray a marked temperature

dependence. More experiments are required to sort out the possible temperature dependence exhibited by fracture toughness tests, opposed to

the apparent temperature independence portrayed by fatigue results.

The martensitic transformation does not seem to significantly influence this discontinuous yielding phenomena. The transformation can

best be considered as an after-effect, i.e., the sudden strains cause

the transformation. In regions of large plastic deformation there will

be more martensite. Toughness testing of AISI 310, which has no martensite

transformation, yet has similar chemical composition (thus similar thermal

conductivity and specific heat) should considerably assist in describing

the role of phase transformations on low temperature discontinuous yielding

of austenitic stainless steels. It is interesting to note that this discontinuous yielding occurs in tensile testing AISI 310 (see figure 15)

.

Of course, the transformation does influence the work hardening rates and,

therefore, should influence the absolute value of the fracture toughness.

297
Divergent suggestions have been made to relate the temperature

dependence of the toughness behavior to the change in the rate of work

41 hardening. The theoretical paper of Rice and Rosengren suggests that

there will be a decrease in toughness of steel with an increase in the work

hardening coefficient due mainly to a rapid rise in stress triaxiality at

42 the crack root. On the other hand, Krafft argues that the plane strain

toughness, K^^, is directly proportional to the work hardening coefficient;

fracture performance is favored by a relatively high rate of work hardening.

The relation of Krafft is supported by his experimental values on selected

steels.

Plots of as a function of the work hardening coefficient for

Ti-6A1-4V and AISI 304 and 316 are shown in figures 60 and 61. Our data

are, at first glance, conflicting. The titanium alloy data support the

41 suggestions of Rice and Rosengren ; i.e., as the work hardening coefficient increases (figure 23) , the fracture toughness decreases (figure 36)

.

42 The stainless steel data support Krafft ; for these alloys, to a rough

approximation, K (calculated from J^.^) is proportional to the work hardening coefficient. It may be significant that in the case of the Ti alloy

plane strain fracture conditions existed, while the stainless steels

exhibited extensive plastic deformation.

The decreased work hardening rates of AISI 304 and 316 at 4 K may

be attributed to two factors: (1) Adiabatic heating insures a higher

specimen temperature than 4 K and (2) below 60 K, the isothermal shear

modulus is lower. Since the interaction force between dislocations is

directly proportional to the shear modulus, which decreases between 76 and

4 K, it is expected that the work hardening rate would decrease between

these two temperatures.

298
600

500

(76K)

400 —

;;; soo

u

H

(298K)

(298K)

200

100 —

1

X 10'

600

500

400

300 u

H

200

100

0.2 0.4 0.6

WORK HARDENING COEFFICIENT

0.8 1.0

Figure 60. Critical Stress Intensity Factor as a Function

of the Work Hardening Coefficient of

AISI 3 04 and AISI 316

299
60

50 —

40 —

^ 30

u

20 —

10 —

0.04 0.08 0.12

WORK HARDENING COEFFICIENT

0.16 0.20

Figure 61. Critical Stress Intensity Factor as a Function

of the Work Hardening Coefficient of Ti-6A1-4V

300
The critical load of AISI 304 and 316, P^, cannot be clearly related to

either P. or P . For example, at room temperature the critical loads for

Q max both steels were nearly the same; however, 304 supported a higher maximum

load than 316. Nor was there a direct correlation between P and P„.

c Q

In 316 stainless steel, P^ increased from 76 K to 4 K, but P^ slightly

decreased in the same temperature interval.

The fracture parameters and J^^ (J integral) were determined

for AISI 304 and 316 at 298, 76, and 4 K. The newly developed J integral

approach was used since plane strain conditions were not possible to

achieve using compact tensile specimens up to 1.5 inches (3.0 cm) thick.

These stainless steels are extremely tough; the estimated K^^ values

from J integral tests is about 300 ksi^in at room temperature and about

450 ksi-^^ at 4 K. The values of at 300 K average 50 ksi-^^iiTand

at 4 K average about 100 ksi-^ in. Therefore, caution must be exercised

in using the absolute K^^ values derived from J^^ tests, as they are

considerably higher than the and K values obtained for each specQ max

imen. Careful consideration and additional testing must be applied

Two conclusions can be reached: (1) AISI 304 and 316 are tougher at

4 K than at 300 K; all fracture toughness parameters are higher at 4 K

than at 300 K; (2) both alloys have about the same toughness.

Additional discussion on these J integral results seem in order.

The parameter has little practical meaning since the concept of

stress intensity loses its significance in the presence of plastic deformation. The parameter retains significance as the value of stress

intensity at which plastic deformation first initiates.

301
Intuitively one expects the critical plane strain value, K^^, to

have a value between K_ and K . That the value is considerably

Q max IC ' higher than both K_. and K is not out of line, if one considers that

Q max

K- and K increase with sample thickness. The larger thickness leads

Q max ^ ^

to a greater constraint which, in turn, raises the flow curve. Thus, at

a thickness considerably larger than the 1-1/2 inch (3.8 cm) dimension

of these experiments, the K^^^ should approach and perhaps exceed the

value of K-j-j^.

It is clear that these calculated values of K could not be used

for design purposes for 1.5 inch (3.8 cm) thick material since failure

would occur at lower stress intensities (K„ - K ) . This result implies

Q max

. that

1) 1.5 TCT specimens were too thin to yield valid

J integral results, and/or that

2) and the load-deflection record in general must

increase considerably with thickness.

If the test specimen is too small, J will not accurately characterize

the crack tip field. Specimens of insufficient thickness yield values

of J which are higher than the true value of J-rp« Thus, for small specimens the value of K is overestimated. The state-of-the-art of J integral

testing is such that rigorous size requirements for valid testing have

not been ascertained. The size criterion suggested by Begley and Landes,

as previously described, is tentative and will undoubtedly be subject to

302
revision in the future as more results are made known. In view of

present uncertainty as to the size requirement, the values of J^.^

and K^^ derived for stainless steels in this study must be regarded

as tentative.

Again referring to the thickness effect, toughness measurements on

invalid specimen sizes show a decrease with increasing thickness until

a limiting minimum value K is reached. However, it is possible for

J. Li invalid toughness parameters to increase with thickness in cases where

the plastic zone size is of the order of specimen thickness. For

stainless steels, calculations indicate that this occurs for thicknesses

of 1.5 inch. The large lateral contractions displayed in fractured

stainless steel specimens is an effect of insufficient constraint. Thicknesses over 1.5 inch would lend greater constraint in the thickness

direction and prevent early yielding. The stress intensity Kp would then

be higher for thicker material, and this was indicated by a single isolated

test performed on 1 inch thick AISI 316 stainless steel at 4 K.

There is some further evidence to support the above ideas. Nelson

43

et al, found a continuous increase in as a function of thickness

in tests on the aluminum alloy 6061-T651. increased as much as

100% before reaching K^^ which was not a "lower limiting value" of

fracture toughness. There is no comparable published literature to

relate the present results obtained for stainless steels.

The alloy A-286 exhibits much less plasticity than does the

AISI 304 or 316. The alloy would approach true plane strain conditions

with a somewhat larger increase in thickness.

303
Results of other investigations of Ti-6A1-4V at cryogenic temperatures were shown in figure 33. Previous authors have not reported data

at A K. The data of this report are in accord only with the results of

37 Vishnevsky and Steigerwald , who obtained valid K data to 76 K.

Their specimens were equivalent in thickness to those tested in this

report (one inch) , and this size largely accounts for the agreement in

results.

37 With the exception of Vishnevsky and Steigerwald , the bulk of

data from other investigators are in rather poor agreement. In general,

their results are too high. Direct comparison is complicated because

of variations in heat treatment, purity and specimen design. Variation

in heat treatment affects the morphology and relative proportion of

phases in the microstructure, while minute differences in interstitial

impurity content (the elements C, 0, N, and H) should influence fracture

toughness of titanium significantly. Nevertheless, the major cause of

high K values must be attributed to specimen size effects. The

"surface flawed" specimens tested by Hall^"^ (t = 0.375") and Tiffany, et al,^"'"

(t = 0.25") nominally satisfy the present ASTM thickness requirement,

but the specimens of other authors referred to in figure 33 clearly do

44

not. In addition, recent results by Golda and Munz indicate that the

ASTM thickness requirement should be doubled in the case of the Ti-6A1-4V

45

alloy. Jones and Brown have also indicated that the present ASTM size

requirement should be doubled for some materials. This possibility casts

33 31 doubt on the results of Hall and Tiffany et al . so that the results

37

of Vishnevsky and Steigerwald and the current NBS data appear to be the

most reliable.

304
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305
13. "Standard Methods of Tension Testing of Metallic Materials," E8.

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in Ambient Air (300 K)," NASA TN D-7532 National Aeronautics and

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18. Watson, J. F., Christian, J. L., and Hertz, J., "Selection of

Materials for Cryogenic Applications in Missiles and Aerospace

Vehicles," Rep. MRG-132-1, Convair Astronautics, San Diego, Calif.

(1960).

19. Espey, G. B., Jones, M. H., and Brown, W. F., Jr., "Sharp-EdgeNotch Tensile Characteristics of Several High Strength Titanium-Sheet

Alloys at Room and Cryogenic Temperatures," Amer. Soc. Testing

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20. Warren, K. A., and Reed, R. P., "Tensile and Impact Properties

of Selected Materials from 20 to 300°K," National Bureau of Standards

Monograph 63 (June 28, 1963).

21. Hanson, M. P., Lewis Research Center, NASA, Cleveland, Ohio, private

communication to T. F. Durham, National Bureau of Standards, Cryogenic Engineering Lab. (Div. 275.03), Boulder, Colorado (1959).

22. Guntner, C. J., and Reed, R. P., "The Effect of Experimental

Variables Including the Martensitic Transformation on the LowTemperature Mechanical Properties of Austenitic Stainless Steels,"

ASM Transactions Quarterly 55, 399-419 (1962).

23. Hoke, J. H., Mabus, P. G., and Goller, G. N., "Mechanical Properties

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of Standards Technical Memorandum No. 39 (1956).

3 06
25. Desisto, T. S., and Carr, F. L., "Low Temperature Mechanical

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.

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463, (1970), p. 97.

3 08
NBS-1 14A (REV. 7-73)

U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET

1. I^UHLICATION OR RIvPOR l' NO.

NBSIR 74-359

2. (lOV 't Accession

No.

3. !<! c ipiiiii s Ac (.•c-.s.sion No.

4. TITLE AND SUBTITLISemi -Annual Report on Materials Research in Support of

Superconducting MachinerySparks, Fickett, Hust, Ledbetter, Kasen, Tobler.

5. l-'ubl it at ion Date- Ma rr h 1 Q74-

6. I'c-rforming Organizaiion ( luii

7. AUTHOR(S) Mikesell, Fowlkes, Durcholz, Reed, Giarratano,

Naimon, and Weston

8. Performing Organ. Report No.

9. PERFORMINC; ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS, Boulder Labs.

DEPARTMENT OF COMMERCE

Boulder, Colorado 80302

10. Projeet/Ta.sk 'Work Unit No.

11. ( 'ontrac t /(irant No. ARPA Order #2569

12. Sponsoring Organization Name and ( omplete Addres.s (Street, City, State, ZIP)

Advanced Research Projects Agency

Departraent of Defense

Washington, D. C.

13. Type of Report tk Period

(-ovcred

14. Sponsoring Agency Code

ARPA

15. SUPPL1-:MENTARY NOTES

16. ABSTRAC'T (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here.)

Results of six months research are reported to the sponsor, the Advanced Research

Projects Agency of the Department of Defense. Subjects include magneto-thermal

conductivity, thermal conductivity, composites, elastic properties, fracture toughness,

fatigue, and tensile. All measurements include the temperature range 4 to 300 K.

Materials examined are those either presently being used in superconducting

machinery or considered for use in future prototypes. Material classes include

stainless steels, inconels, titanium alloys, and composites.

Special results include: the thermal conductivity in a magnetic field is considerably

lower than would be predicted; a comprehensive review of glas s - reinforced composite behavior at low temperatures is included; the elastic moduli of 12 engineering

alloys from 4 to 300 K are reported; and fracture toughness and fatigue crack growth

rate data on AISI 304, AISI 316, A286, Ti-5A1-4V and Ti-6Al-2. 5 Sn at 4, 76 and

300 K have been measured.

At the beginning of each individual report a Summary is provided to highlight the

project results

.

17. KEY WORDS (six to twelve entries; alphabetical order; capitalize only the first letter of the first key word unless a proper name; separated by semicolons

)

Composites; fracture; liquid helium; mechanical properties; structural materials;

superconducting machinery; thermal conductivity.

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