NBSIR 78-884 UC 20 b.c

TECHNICAL REPORTS

MATERIALS STUDIES FOR

MAGNETIC FUSION ENERGY

APPLICATIONS AT LOW TEMPERATURES - I

Edited By

F.R. Pickett and R.P. Reed

Thermophysical Properties Division

National Bureau of Standards

Boulder, CO 80303

April 1978

Sponsored By

Department of Energy

Division of Magnetic Fusion Energy

Washington, D.C. 20545

U.S. DEPARTMENT OF COMMERCE, Juanita M. Kreps, Secretary

Sidney Harman, Under Secretary

Jordan J. Baruch, Assistant Secretary for Science and Technology

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

CONTENTS

Page

SUMMARY V

ORGANIZATIONAL CONTACTS vii

INTRODUCTION 1

PROGRAM DESCRIPTION 3

HIGHLIGHTS OF RESULTS 5

NITROGEN STRENGTHENED STAINLESS STEELS 11 Fatigue Testing of Stainless Steels - Martin Marietta 13

Fracture Testing of 21-6-9 Stainless Steel - Martin Marietta . . 69

Toughness, Fatigue Crack Growth, and Tensile Properties of

Three Nitrogen-Strengthened Stainless Steels at Cryogenic Temperatures-NBS 91

STAINLESS STEEL WELDING 155

Evaluation of Stainless Steel Weld Metals at Cryogenic

Temperatures-NBS 157

Weldability of 21-6-9 Stainless Steel-Armco 167

Ferrite in Type 31 6L Weld Metal -Arcos 179

PHYSICAL PROPERTIES OF STAINLESS STEELS 211

Low Temperature Elastic Properties of 304-Type Stainless

Steels-NBS ^ 213

NONMETALLICS 219

HANDBOOK PREPARATION 223

VAIL WORKSHOP 233

SURVEY OF LOW TEMPERATURE MATERIALS FOR MAGNETIC FUSION ENERGY. ... 243

i i i

SUMMARY

The reports presented here summarize the work of the low temperature

materials research project for the first year of the program. The various

projects are outlined and the research results are presented. The major

thrust of the measurements has been the evaluation of the low temperature

properties of stainless steel base metal and welds, with particular emphasis

on the nitrogen strengthened stainless steels. Some initial work has also

been done on the production and properties of nonmetal 1 ics , primarily industrial laminates, for low temperature applications. A handbook of material

properties is also planned. A survey of low temperature materials needs

and problems related to magnetic fusion energy, performed by NBS as groundwork for the program, is included as is a brief description of the first

workshop held in October 1977.

NOTE

The use of trade names of commercial materials is essential to the

proper understanding of the work presented. This use in no way implies

approval, endorsement or recommendation by the National Bureau of Standards.

V

ORGANIZATIONAL CONTACTS

Listed here are the people responsible for the various aspects of the

program so that those with specific technical questions may contact the

individual directly.

Department of Energy, Division of Magnetic Fusion Energy, Washington, D.C. 20545

Program Monitor E. N. C. Dalder (FTS) 233-4964

(301) 353-4964

National Bureau of Standards, Thermophysical Properties Division, Boulder,

CO 80302

Program Manager R. P. Reed (303) 499-1000 X3870

Program Coordinator F. R. Fickett (FTS) 323-1000 X3785

Welding H. I. McHenry X3268

Elastic Properties H. M. Ledbetter X3443

Nonmetallics M. B. Kasen X3558

Martin Marietta Aerospace, Denver Division, P.O. Box 31, Denver, CO 80201

Project Manager F. R. Schwartzberg (303) 973-3225

Fracture, Fatigue J. A. Shepic

Armco Steel Corp., Research & Technology, Middletown, OH 45043

Project Manager R. H. Espy (513) 425-2490

Arcos Corp., 1500 South 50 Street, Philadelphia, PA 19143

Project Manager R. D. Thomas, Jr. , (215) 727-1500

Mechanical Properties Data Center, 13919 West Bay Shore Drive, Traverse City,

MI 49684

Program Manager R. C. Braden (616) 947-4500

Data Evaluation T. Moore

vil

INTRODUCTION

This report contains results of a research program to produce materials

properties data to facilitate design and development of cryogenic structures,

especially for superconducting magnets, required for magnetic fusion energy

research. The program was conceived and developed jointly by the staffs

of the Cryogenics Division, National Bureau of Standards and the Division

of Magnetic Fusion Energy of the Department of Energy. This program,

sponsored by DOE, is managed by NBS. Research is conducted at NBS and

at various other laboratories, through subcontracts with NBS.
li
PROGRAM DESCRIPTION

The overall objective of the program is to assist the design, construction

and safe operation of low temperature magnetic fusion energy (MFE) systems,

especially superconducting magnets, through effective materials research and

materials technology transfer. The specific steps taken to achieve this

objective are: (1) evaluation of low temperature materials research needs

specific to MFE devices; (2) development and monitoring of a research program to acquire the necessary data; and (3) insuring rapid dissemination of

the data to potential users through personal contacts, publications and

workshops

The first specific objective was accomplished with the publication of the

Survey of Low Temperature Materials for Magnetic Fusion Energy in March 1977.

The second objective is described here in Table 1 in the form of an outline

of the research projects. The results are contained later in this report. The

third objective will be satisfied in part by this report and our handbook

project. It has already been accomplished in large part with the holding of

the first NBS-ERDA Workshop on Materials at Low Temperatures in October 1977.

3
Table 1. Outline of the NBS/DOE Program on Materials Studies for Magnetic

Fusion Energy Applications at Low Temperatures

Program Area Organization

A. Nitrogen Strengthened Stainless Steels

1. Fracture and Martin Marietta

Fatigue

2. Toughness, Crack NBS

Growth and Tensile

Properties

B. Stainless Steel Welding

1 . Armco

Arcos

NBS

C. Physical Properties of Stainless Steels

1. Elastic Properties NBS

D. Nonmetallics NBS

E. Handbook Preparation MPDC/NBS

Program Description

Low cycle, strain controlled, fatigue

testing of AISI 304, 316 and 21/6/9

stainless steel. J-integral fracture toughness testing of 21/6/9.

All tests at 295, 76 and 4 K. Measurement of fracture toughness,

fatigue crack growth rate and tensile

property data for Nitronic 33,

Nitronic 50 and AISI 304N stainless

steel at temperatures of 295, 76

and 4 K.

Development of a satisfactory filler

material for welding of nitrogen

strengthened stainless steel for low

temperature service. Tensile and

impact testing of candidate materials

at room temperature and 77 K. Production of test welds with varying

ferrite content in austenitic stainless steels. Welds will be evaluated

for toughness at NBS later in the

program.

Investigation of metallurgical factors

contributing to the toughness (or

degradation thereof) of stainless

steel weld metals. Later, the program

will evaluate the influence of process

variables on strength and toughness

at 4 K.

Measurement of the elastic properties

of AISI 304, 304L and 304N stainless

steels to determine the effect of

temperature to 4 K and to determine

the variabilities within and among

heats, etc. Acquisition of published formulations

of specific GIO and Gil industrial

fiberglass-epoxy laminates. Low

temperature testing of these materials

in later years of the project to determine mechanical and physical

properties

Production of a handbook containing

low temperature mechanical and physical properties of materials of interest to the MFE community. Presently

designated for inclusion: AISI 304,

316 and 21/6/9 stainless steels; 2219,

5083 and 6061 aluminum alloys.

4
HIGHLIGHTS OF RESULTS

HIGHLIGHTS OF RESULTS

We present a brief description of outstanding results from the

project reports which follow in the next section.

Nitrogen Strengthened Stainless Steels

Tensile, fracture toughness and fatigue crack growth rates have

been measured on nitrogen strengthened austenitic steels containing Fe18Cr-13Mn-3Ni (Nitronic 33), Fe-21Cr-5Mn-12Ni (Nitronic 50), and Fe19Cr-2Mn-9Ni (AISI 304N). AISI 304N is particularly promising for use

at 4 K since the yield strength is doubled (60 to 120 ksi), the toughness

remains acceptably high, and the cost remains near the cost of AISI 304.

Other alloys, including AISI 304LN and AISI 316N, will be measured for

comparison.

Smooth-bar strain-controlled fatigue tests have been conducted on

AISI 304, 316 and Fe-21Cr-6Ni-9Mn stainless steels at 295, 76, and 4 K.

The results indicate reduced fatigue life at 4 K, compared to 76 K.

Alloy AISI 316 appears to have the best fatigue life at 4 K.

Fracture toughness tests at 295, 76, and 4 K on a commercial heat

of Fe-21Cr-6Ni-9Mn (Nitronic 40) indicates lower toughness than a heat

of the same alloy that was electroslag remelted and cross-rolled.

Toughness decreases considerably in this alloy between 76 and 4 K.

Stainless Steel Welding

A literature review has indicated that the toughness of stainless steel

weld metals is marginal for MFE structural applications, particularly in thick

section weldments and in designs where the stresses exceed the ASME allowable,

138 MPa (20 ksi). The best available filler metal appears to be AWS type 330

(18Cr-34Ni); tests on welds prepared by Teledyne-McKay indicate that the Charpy

impact toughness at 76 K is 84 J (62 ft lb) and the yield strength is 860 MPa

(125 ksi) at 4 K. Armco Steel Corp. evaluated the influence of nitrogen on

7
the strength and toughness of a series of filler metals developed for

welding the 21Cr-6Ni-9Mn (Nitronic 40) alloy; nearly matching strengths

can be achieved but the decrease in toughness is unacceptable. Arcos

Corp. welded a series of test plates to evaluate the influence of ferrite

content, ferrite morphology and sensitization on the toughness of 316L

weld metal at cryogenic temperatures; the NBS evaluation is in progress.

Physical Properties of Stainless Steels

Elastic properties of 304-type stainless steels were determined accurately

between room- temperature and liquid-helium temperature using pulse-echosuperposition ultrasonic techniques. Particular elastic constants of interest

were Young's modulus, the shear modulus, the bulk modulus, and Poisson's ratio.

The Young's and shear moduli increase up to eight percent during cooling;

Poisson's ratio decreases about four percent; and the bulk modulus (reciprocal

compressibility) is relatively temperature insensitive. All elastic constants,

except perhaps for the bulk modulus, change anomalously and reversibly at about

45 K due to a Neel (paramagnetic-anti ferromagnetic) transition. Some samples

show anomalies in both the longitudinal modulus and Poisson's ratio at about

16 K, indicating another transition that seems to be reversible. Specimens

studied thus far, which include both bar and plate stock, show very little

elastic anisotropy.

Nonmetal 1 ics

A major problem is posed by the highly proprietary nature of existing NEMAgrade industrial laminates. Discussions with representatives of the industrial

laminating industry (Westinghouse and Spaulding Fiber) and with representatives

of national laboratories capable of performing the required radiation research

(ORNL, ANL) has led to the selection of three laminate systems to be fabricated

from published formulations and included in the mechanical, thermal, electrical,

and radiation characterization work:

8
1. NEMA G-lO/MIL-P-18177, Type GEE

2. NEMA G-n/MIL-P-18177, Type GEB (aromatic amine cure)

3. NEMA G-n/MIL-P-18177, Type GEB (aromatic amine cure, boron-free

E-glass reinforcement)

Laminates 1 and 2 are to be conventional, state-of-the-art materials differing from other NEMA laminates only in being defined material systems as

distinct from proprietary formulations. The laminates will be made by the

Micarta Division of Westinghouse, however, information required to fabricate

the laminates will be made available to all industrial laminators. This will

provide the MFE community with access to G-10 and G-11 type laminates having

a known formulation which will become well characterized as part of the FY 78

NBS-ERDA cryogenic materials program. Laminate formulation data should be

available by 1 January 1978,

9

NITROGEN STRENGTHENED STAINLESS STEELS

11

'mm

Fatigue Testing of Stainless Steels

Martin Marietta

13
I i
Final

MCR-77-562 Report December 1977

FATIGUE TESTING OF

STAINLESS STEELS

J. A. Shepic

Principal Investigator

F. R, Schwartzberg

Program Manager

MARTIN MARIETTA CORPORATION

Denver Division

Post Office Box 179

Denver, Colorado 80201

15
FOREWORD

This report describes technical activities conducted by the Denver

Division of Martin Marietta from 17 May 1977 through 30 November 1977

under National Bureau of Standards (NBS) contract CST-8426, Item 2.

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 United States

Government,

16
INTRODUCTION

Superconducting machinery will be subjected to cyclic loading at

teii5)eratures down to 4 K. Fatigue data for candidate materials are

relatively sparse below 20 K. At 4 K, most of the fatigue data are

embodied in three papers (references 1, 2 and 3). This work was undertaken to supplement the data already available, and to demonstrate that

a relatively simple testing system for strain-controlled fatigue testing

at 4 K could be developed.

Three alloys (21-6-9, 304 and 316 stainless steel) were evaluated.

No data were available at 4 K for the 316 alloy. Stress-controlled fatigue

data for 21-6-9 were available from a prior program conducted at Martin

Marietta^ Alloy 304L has been previously evaluated using both stressand strain-controlled techniques; however, concern for the effect of

adiabatic heating due to plastic deformation necessitated confirmation

evaluation,

MATERIALS

Stock of each alloy was obtained as plate stock. The 21-6-9 stainless steel was obtained as 3.8 cm (1.5 in.) stock; 304 and 316 stainless

steels were obtained as 2.5 cm (1.0 in.) stock. All material was provided

by NBS.

The 21-6-9 alloy was received in the hot-rolled, annealed and

descaled condition. Data describing annealing and composition are

described in the companion Martin Marietta report on fracture testing

*The designations 21-6-9 and Nitronic 40 are trade names for an alloy with

the nominal composition Fe-21Cr-6Ni-9Mn. Throughout the remainder of this

report it is described by the first designation. The standard stainless

steel alloys AISI 304 and AISI 316 are hereafter referred to as 304 and

316, respectively.

17
of 21-6-9.

The 304 and 316 stainless steels were from the two LMFBR (Liquid

Metal Fast Breeder Reactor) program reference heats received from DOE.

NBS determined the tensile properties of the three materials. Data

are summarized in Table 1,

TESTING TECHNIQUE AND APPARATUS

This chapter describes the specimen geometry, testing apparatus

and technique used to obtain the fatigue data presented in this report.

The fatigue specimen configuration used for this program was similar

to that used by Martin Marietta in previous fatigue studies (references

1 and 4). The design incorporates a smooth gage section with a mild

radius and a bearing shoulder at the end of the threaded section for

alignment and locking, as opposed to aligning on the threads. This

system is advantageous because only two close- tolerance alignment surfaces are required and failure within the prescribed area of the gage

is assured. The specimen, when locked with a nut at each end of the

threaded bar, can be subjected to a fully reversed loading cycle that

is continuous through zero load. Figure 1 gives the specifications

for the specimen.

Fatigue specimens were polished after machining using 240 to 320

grit polishing paper (the coarseness of the initial rough polishing

paper depended on the alloy hardness and machined surface roughness

in the gage section) and then using 320 and 600 grit paper followed

by a felt wheel and Tripoli polishing compound. The polishing paper

was bonded to a rubber sanding disc mounted in a small Dumore grinding tool

18
Table 1 Tensile Properties of Stainless Steel

Alloy and

Heat No.

Temperature,

K

Ultimate

Strength,

MN/m'=^ (ksi)

0.2%

Yield

Strength

MN/m (ksij

Elongation

percent

Reduction of

Area,

percent

21-6-9 293 785 (114) 386 (56) 58 76

[656249-lD] 77 1488 (216) 916 (133) 34 32

A 1'5'57 ('226') 1399 C203') 11 25

304 293 599 (87) 214 (31) 67 67

[9T-2796]

77 1433 (208) 262 (38) 42 65

4 1592 (231) 345 (50) 39 48

316 293 598 (84) 227 (33) 56 73

[8092297]

77 1240 (180) 606 (88) 62 70

19
20
As the specimen is rotated at approximately 350 rpm between lathe

centers, the periphery of the spinning (approximately 20,000 rpm) sanding

disc contacts the specimen gage section with axial strokes. (The same

procedure was used with all grits) , The final polish uses a 2 inch

diameter felt wheel mounted in a Dumore grinding tool and loaded with

Tripoli polishing compound. As the specimen is rotated between lathe

centers at approximately 150 rpm, the spinning (approximately 20,000 rpm)

impregnated felt wheel contacts the specimen gage section with axial

strokes » This polishing technique produces a mirror finish in which

any remaining polishing marks are in an axial direction that will have

minimal effect on axial fatigue test results.

Because of the non-uniform cross- section area of the specimen, it

is necessary to control total strain based on diametral deflection rather

than axial deflection. A diametral strain extensometer was designed

and built specially for this work. This device is a one-piece aluminum

ring with integral strain beams instrumented with nichrome resistance

strain gages. The beam ends are machined with knife edges to grip the

specimen. The knife edge separation distance is controlled to provide

good gripping action on the gage section. The gage grips the specimen

sufficiently tight enough that it remains in the horizontal plane during

filling of the cryostat and subsequent cyclic loading. Because the aluminum is softer than the stainless steel, no nicks or scratches are imparted

to the specimen. Figure 2 gives a closeup view of the extensometer.

The test configuration for 293 K evaluation was quite conventional.

21
Figure 2 Diametral Strain Extensometer

22
As shown in Figure 3, the specimen is installed in adapters that permit

the shoulders to be preloaded; the adapters are then installed in the

testing machine

For 77 K testing, a double- walled, vacuum- insulated cryostat,

described in detail elsewhere (reference 5), was used. As shown in

Figure 4, the specimen is also installed in adapter plates.

Testing at 4 K was performed in a cryostat system prepared specially

for low-cycle fatigue testing. Loading was achieved through two concentric frames, as shown in Figure 5. The smaller frame surrounding the

specimen is constructed from titanium alloy. Note that a specimen

adapter plate is used at the top of the specimen; at the bottom the

specimen shoulder bears against the triangular plate. The remainder

of the frame is constructed from aluminum alloy, A glass vacuum dewar

is used to provide environmental control. Figure 6 shows the load frame

installed in the machine, both without and with the dewar in place.

The liquid helium system is designed so that testing can be achieved

with minimum cryogen consumption. The titanium frame and specimen are

first precooled to 77 K and placed in a liquid nitrogen filled dewar

on the bottom platen. After cooling, the head is raised and a liquid

helium filled dewar replaces the liquid nitrogen dewar. A liquid level

sensor is used to monitor level, and refilling is performed as required.

Fatigue testing was performed in a 222 kN (50 kip) closed- loop

servo-controlled testing machine. Operation was in the strain-control

mode. The relationship between total axial strain amplitude and total

diametral strain amplitude is as follows:

23
Figui'e Z Apparatus for 293 K Testing

24
Figure 4 Apparatus for 77 K Testing

25

27
where = total axial or longitudinal strain amplitude

Xj = total diametral strain amplitude

d

AP = load amplitude at half- life

A = cross- sectional area

E = Young's modulus

u = Poisson's ratio

To calculate the diametral strain amplitude required to achieve a

desired axial strain amplitude, an approximate value for AP was selected,

Following completion of the test, a more accurate value of AP was

selected from the load vs strain hysteresis curves and the ao-tual total

axial strain amplitude value was recalculated. Accuracy of the AP

estimation improved with each succeeding test and, as a result, actual

values were quite close to desired values. The elastic component of the

total strain was calculated using the AP value read from the test record

and the appropriate value of Young's modulus. The plastic component of

strain was calculated by subtracting the elastic strain from the total

strain,

EXPERIMENTAL TEST RESULTS

Fatigue behavior of each of the three alloys was characterized at

293, 77 and 4 K. Because of the extremely high cost of liquid helium

and the large quantities required to obtain long-time fatigue data,

only a limited number of 4 K tests were conducted beyond lO-" cycles.

304 Stainless Steel - This alloy exhibits a significant amount of

discontinuous yielding at 4 K, as shown by prior tensile test investigations. During cyclic loading it is possible that the heat generated in

the plastic range can cause vaporization of the low heat capacity liquid

28
helium and a resultant elevation of the test temperature. This situation

should be rate sensitive and could be eliminated by slower cycling. To

determine whether this was a factor in 4 K testing, a series of tests

were performed at a common strain amplitude and different cyclic frequencies. The data summarized in Figure 7 suggest that fatigue life is

frequency- dependent. For the nominal total strain amplitude of 1.5

percent, it appears that a maximum frequency of 0,1 Hz can be used without

risk of specimen heating.

It was observed that the discontinuous yielding apparent in the

tests performed at or below 0,1 Hz was not noted at 0.3 Hz, This would

suggest that the temperature was greater than 4 K. Without further

extensive testing, it was decided to maintain the test frequency as low

as possible, but consistent with the economics of liquid helium consuraptlono The frequency for all succeeding liquid helium tests was carefully

selected on the basis of strain level.

Experimental data for the 304 alloy exhibited a temperature dependency of life in the low cycle range in which life decreased with

decreasing temperature. A cross-over occured at approximately 1^ percent

strain in which the order reversed. Below this strain level, life increased with decreasing temperature. Figure 8 graphically depicts the

fatigue life as a function of strain. Table 2 gives detailed information

on the diametral and longitudinal strains, stress amplitude, cyclic rate,

and cycles to failure.

The load vs strain hysteresis curves provide a great quantity of

useful information pertinent to the development of an understanding of

cyclic behavior. Figure 9, a typical curve, shows the initial 35 cycles

29

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for a nominal 1«0 percent total longitudinal strain test at 77 K. Note

the first cycle starting from the center of the loop. Loading was initiated in compression to eliminate the possibility of developing a local

necked region during the first tensile loading application of materials

prone to strain softening. The curve shows excellent control of strain

in both tension and compression. Strain hardening in each cycle is

clearly shown.

Figure 10 illustrates strain softening behavior. In this test (0,5

percent strain at 293 K) , the highest load is achieved during the first

cycle. At the conclusion of 10,000 cycles, significant strain softening

is obvious. The figure also illustrates the irregularity (arrow) in the

compression portion of the cycle indicative of cracking in the specimen.

Note the significant decrease in load associated with the cracking.

The liquid helium tests are particularly interesting. Figure 11

presents the loop for the 1,5 percent strain amplitude specimen (304-16)

tested at 0,3 Hz, Note the general similarity to the 77 K data of

Figure 9 with respect to strain hardening and degree of strain control.

Data shown for the specimen evaluated at the same strain level, but lower

frequency (0,02 Hz), are markedly different. Figure 12 illustrates the

first two cycles of this test. The discontinuous yielding is extremely

pronounced. By careful study of this curve, the first two cycles can be

followed. Note the initiation in compression with the initial elastic

portion followed by yielding and then with increased straining the initiation of small discontinuous yield jogs. Further straining results in

increased size of the discontinuities. The loading in tension at the end

of the first half-cycle is characterized by a relatively smooth curve

34
Figure 10 Load-Deflection Curve for Sipeoimen 304-4

35
Figure 12 Load-Deflection Curve for Specimen 304-16

36
Figure 12 Load-Defleotion Cupve for Specimen 304-19 (Cycles 1 and 2)

37
until yielding starts and then an increase in discontinuity activity with

increasing strain. The second cycle is much like the first, with the

discontinuities occurring at similar locations on the curve.

Figure 13 presents the next 18 cycles. Although hard to follow,

the curve does show strain hardening with increased cycling. Despite the

abruptness of the discontinuities, the strain control limits are relatively

well maintained. During the next 20 cycles, the severity of the yielding

discontinuities decreases and the amount of strain hardening with each

succeeding cycle becomes small, as illustrated in Figure 14.

Figure 15 illustrates the behavior at a lower strain level than the

preceeding test. At a level of 1,0 percent strain, the magnitude of the

yielding discontinuities becomes significantly smaller than observed for

the 1,5 percent strain level. It should also be noted that the degree

of strain hardening compared to the comparable test at 77 K (Figure 9)

is significantly reduced and, as a result, an approximate fourfold

increase in life is achieved,

21-6-9 Stainless Steel - Figure 16 summarizes the strain vs life

data and shows a significant temperature dependency of life at strain

levels greater than 1 percent. Below 1 percent strain, a reversal occurs

and an indication of increasing life with decreasing temperature becomes

apparent, A presentation of the data is given in Table 3.

Figure 17 presents a typical hysteresis loop at 77 K. This specimen,

evaluated at 1 percent nominal longitudinal strain, exhibits some strain

softening in the initial cycling and relatively little softening after

1000 cycles; total life was 2443 cycles. The curve is smooth and well

defined, and exhibits good strain control. The corresponding curve

38
Figure 12 Load-BefZeotion Cvacve for S-peo-tmen 304-19 (Cycles S through 20)

41
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43

obtained at 4 K is presented in figure 18, The curve is not quite as

smooth as observed at 77 K, probably as a result of localized heating due

to strain gage excitation, but does not exhibit discontinuous yielding.

By comparison, the 4 K curve obtained at a high strain level (2,5

percent) exhibits significant discontinuities (Figure 19) . The repetitive

trace clearly shows an increase in the magnitude of the discontinuities

with increasing strain level. Although masked by the multiple traces,

the degree of strain hardening observed was relatively small compared to

that exhibited by the 304.

An anomaly noted at 293 K for several tests was a combination of

strain hardening and strain softening occurring at intermediate strain

levels. Although the magnitude of the change was small, it was deemed

unusual. Figure 20 shows this behavior. Strain hardening occurs during

the first 20 cycles, followed by continuous softening until failure occurs,

316 Stainless Steel - The 316 stainless steel exhibits a decrease

in fatigue strength with decrease in temperature from 293 to 77 K for

strain levels greater than 1 percent. At 4 K, the fatigue life is

greater than at 77 K for all but the highest strain level evaluated.

At 293 K, the crossover with the 4 K data occurs at approximately 2

percent strain. Figure 21 presents the data. Table 4 gives detailed

test information.

The hysteresis loops obtained at 293 and 77 K are similar to that

observed for 304 except that much less strain hardening occurred.

Figure 22 presents a typical 77 K curve (1.5 percent strain).

At 4 K, discontinuous yielding was found, but the magnitude was

much less than observed for the 304, Figures 23 and 24 present data

45
Figure 18 Load-Deflection Curve for Specimen 2169-21

46
47
Figure 20 Load-Deflection Curve for Speoimen 2169-4

48

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4-1 o

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cfl <u U 3 o

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50
Figure 22 Load-Deflection Curves for Specimen 316-12

51
Figure 23 Load-Deflection Curves for Specimen 216-21

52

at 1.25 and 2.5 percent strain levels as evidence of the reduction

in yield jogs. The degree of strain hardening was also much lower.

Strain Hardening - A comparison of the strain hardening behavior of

the three alloys at each temperature reveals certain interesting trends.

Figure 25 plots the change in stress amplitude from the first to final

cycle as a function of total axial strain amplitude. The changes in

stress amplitude for the 304 stainless steel at 77 and 4 K is extremely

high. This alloy displays no evidence of strain softening at cryogenic

temperatures. The 316 alloy at all three temperatures and the 304 at

293 K exhibit a modest change in strength; the change is only a fraction

of that observed for the latter alloy at cryogenic temperatures. At low

strain levels, types 304 and 316 exhibit slight strain softening. The

21-6-9 exhibits even less strain hardening than the 316. At 77 K, the

magnitude of strain softening at low strain levels is similar to that of

strain hardening. The 4 K data for 21-6-9 show strain hardening similar

to that observed at 77 K; however, little strain softening is evident.

Transformation - Preliminary attempts to determine the degree of

transformation on the fracture surfaces using a Magne-Gage was unsuccessful. Differences in surface roughness and variations in magnetic attraction associated with local deformation (shear lips) made measurements

difficult. Specimens will be retained for future study using alternative techniques and methods.

With the use of a hand-held magnet, qualitative estimates of

magnetic susceptibility were made. Data are summarized below:

54
0 1.0 2.0 3.0 Total Axial Strain, percent

Figure 25 Effect of Strain Amplitude on Change in Stress Amplitude

55
Temperature (K)

Alloy 293 77 4

304 N to W M M

21-6-9 N N N

316 N to W MM

N = nonmagnetic

W = weakly magnetic

M = moderately magnetic

DISCUSSION OF RESULTS

Data obtained from this work were compared with other available

information to obtain fatigue life curves that provide the best possible

characterization of the behavior of the various alloys.

For the 304 alloy, the strain-controlled data obtained from this

work, the strain-controlled data of Nachtigall (reference 2) , and the

stress-controlled data from a prior Martin Marietta program (reference 1)

were combined to give composite strain versus fatigue life curves. The

stress controlled data from tests conducted in the elastic range were

converted to equivalent total longitudinal strain. The plots for 293,

77 and 4 K show excellent agreement between the three sources and permit

a smooth curve to be drawn at each temperature. Figures 26, 27 and 28

present the fatigue plots. These strain-contolled fatigue data may be

plotted as stress-amplitude-to-failure data by converting the strain

amplitude X to a stress amplitude S using the relation S = E A. where

_L 9. 3 X E is Young's modulus. This form is appropriate for comparison to data

56
as used in the A.S.M.E. Boiler and Pressure Vessel Code (reference 6).

The resulting stresses are, of course, much larger than would occur in

a stress-controlled fatigue tests. No safety or adjustment factors have

been included.

The calculated stress amplitudes are shown on the right-hand ordinate

of the plots in Figs. 26, 27 and 28.

Similarly, data for the 21-6-9 from this work and the prior Martin

Marietta study (reference 1) were combined and give a smooth curve extending into the 10 cycle range. Figures 29, 30 and 31 present these curves.

Again, calculated stress amplitudes are shown on the right-hand ordinate

of the graphs.

57

Bao 'apnaiTdniv ssaaqs paaoxnoxBO o o o o

(1 '

I ^

\

' r

59

61
o o

o O O o

<^ CM i-l

62

j;
1
A comparison of the fatigue strength of the three alloys at each test

temperature was performed. At 293 K, the 21-6-9 alloy exhibits the highest

fatigue strain capability in the low cycle range. 316 exhibits slightly

lower properties. The 304 stainless has significantly poorer low-cycle

4 properties. In the high-cycle range (> 10 ), 316 exhibits the highest

endurance limit. Figure 32 shows the comparative properties of the three

alloys at 293 K. Approximate stress levels calculated as before are indicated on the right-hand ordinate. The stress levels are approximate because

of a difference of about 4% among the Young's moduli of the three materials.

The average of the three moduli was used for the stress calculation.

A similar comparison made for the 77 K data shows the 21-6-9 alloy to

be superior in fatigue behavior for the low-cycle range. 316 is slightly

lower in performance, but crosses the 21-6-9 curve at about 1000 cycles and

exhibits the highest long-tem properties. 304 exhibits the poorest fatigue

properties throughout the test range. Figure 32 presents the composite curve.

Again, approximate stress amplitude levels are indicated on the right-hand

ordinate. At this temperature, the spread of the Young's moduli is about 8%.

At 4 K, 316 is clearly superior to the 21-6-9 and 304 in the low-cycle

range. Insufficient data are available to confirm retention of superiority

in the high-cycle range, but that trend is apparent. The 304 exhibits behavior

generally equivalent to the 21-6-9 (Figure 34). In Fig. 34, as before,

calculated stress amplitude levels are indicated on the right-hand ordinate,

and the spread in Young's moduli among the three materials is about 8%.

A comparison of the effect of temperature for each alloy shows that the

behavior of 304 is relatively independent of temperature. In the low-cycle

range, the 4 and 293 K data are almost identical, and the 77 K results are

quite similar to the other temperatures. Likewise, the 316 exhibits little

64

o o o o

00 \o <r (N o

I

66

j;
67
effect of temperature on behavior. The 21-6-9 exhibits the greatest effect

of temperature on behavior. The low-cycle like is decreased slightly with

reduction in temperature to 77 K, and quite significantly with reduction

to 4 K.

Although differences in fatigue behavior of the three alloys are

apparent, none displays qualities that should affect its candidacy for

service down to 4 K.

REFERENCES

1. F« R. Schwartzberg and T. F. Kiefer, "Study of Fracture Behavior

of Metals for Superconducting Applications", Materials Research

for Superconducting Machinery, Semi- Annual Technical Report IV,

October 1975, National Bureau of Standards, Boulder, Colorado.

2. A. J. Nachtigall, "Strain- Cycling Fatigue Behavior of Ten Structural

Metals Tested in Liquid Helium, in Liquid Nitrogen, and in Ambient

Air", NASA TN D-7532, February 1974.

3. A, J. Nachtigall, S. J. Klima, and J. C. Freche, "Fatigue of Liquid

Rocket Engine Metals at Cryogenic Temperatures to -452F", NASA TN

D-4274, December 1967.

4. E, J. Beck, "Effect of Beta Processing and Fabrication on Axial

Loading Fatigue Behavior of Titanium", AFFDL-TR-69-108, Martin

Marietta Corporation, June, 1969.

5. R. D. Keys and F. R. Schwartzberg, "Axial Fatigue Testing of Sheet

Materials Down to -423F", Materials Research and Standards, May,

1964.

6. "A.S.M.E. Boiler and Pressure Vessel Code, Section VIII, Division 2,"

American Society of Mechanical Engineers, 1977, p. 399-404.

68
Fracture Testing of 21-6-9 Stainless Steel

Martin Marietta

59

Final

MCR-77-561 Report December 1977

FRACTURE TESTING

OF 21-6-9 (Nitronic 40)

STAINLESS STEEL

J. A. Shepic

Principal Investigator

F. Ro Schwartzberg

Program Manager

MARTIN MARIETTA CORPORATION

Post Office Box 179

Denver Division

Denver, Colorado 80201

71
FOREWORD

This report describes technical activities conducted by the Denver

Division of Martin Marietta Corporation during the period 17 May 1977

through 30 November 1977 under National Bureau of Standards (NBS)

contract CST-8426, Item 1.

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 United

States Government.

72
INTRODUCTION

Austenitic stainless steel alloy, 21-6-9 (Fe-21Cr-6Nl-9Mn-0 .3N)

,

(Nitronic 40), is of great interest for cryogenic machinery, particularly

in superconducting magnets, because of its excellent combination of

strength and toughness.

A preliminary evaluation of the fracture toughness of this alloy

was reported by Tobler and Reed^'''^ using laboratory processed material.

The objective of this work is to further characterize the same gage, 3.8

cm (1.5 in.) of commercially produced plate material.

MATERIAL

The 21-6-9 stainless steel evaluated in this program was provided

by National Bureau of Standards (NBS) in the form of 3.8 cm (1.5 in.)

plate. The material had been procured from G.O. Carlson and was identified as heat number 656249. The plate was received in the hot rolled,

annealed, and descaled condition. Final annealing temperature was reported to be 1366-1394K (2000-2050F) for \ hour per 2.5 cm (1 in.) of

thickness, followed by water quenching.

Vendor's analysis showed the following composition for the plate:

Cr = 20.11 w/o

Ni = 6.34 w/o

Mn = 8.97 w/o

C = 0»033 w/o

P - 0.013 w/o

S = 0.005 w/o

Si = 0.56 w/o

N = 0.29 w/o

Fe = balance

73
NBS determined the tensile properties of the material at three

temperatures. Data are as follows:

Temperature, K

293 77 4

Ultimate strength, 785 (114) 1488 (216) 1832 (266)

MN/m2 (ksi)

Yield strength, 0.2% offset 386 (56) 916 (133) 1399 (203)

MN/m2 (ksi)

Elongation, percent 58 34 11

Reduction of area, percent 76 32 25

TESTING TECHNIQUE AND APPARATUS

This chapter describes the specimen geometry, testing apparatus,

and technique used to obtain the fracture data presented in this report,

The J- integral specimen configuration used for this program was a

3o8 cm (1.5 in.) thick compact specimen of the basic geometry described

(2) in ASTM E- 399- 74^ . The specimen was modified, as shown in Figure 1,

to permit loadline compliance measurements to be obtained. Specimens

were machined so that the notch was oriented parallel to the rolling

direction (TL orientation).

Precracking of 293K test specimens was performed at the test

temperature. Cryogenic test specimens were precracked at 77Ko Specimens were precracked in closed- loop servo-hydraulic testing machine at

a frequency of 10 Hz and a load of 22-27 kN (5-6 kips). The maximum

stress intensity (Kf) at the conclusion of precracking to an a/W ratio

74
75
of Oc.65 was a maximum of 42 MN/m-^'^ (38 ksi ^ in.) . Extension of the

crack during cycling was approximately 0.25 cm (0.1 in.). Crack front

curvature was significantly less than the minimum of 90% of average

crack length specified by ASTM E-399-74; typical values were 96 to

98%.

Compliance measurements at 293 and 77K were obtained with a Martin

Marietta built clip gage that permitted loadline displacement of approximately lo2 cm (0,5 in.) to be measured. Figure 2 shows the gage installed at the loadline on a specimen. At 4K, a commercial clip-gage

designed to measure opening up to 0.5 cm (0.2 in.), was used.

Fracture testing was performed in a 448 kN (100 kip) testing machine.

At 77K, a simple foam insulated container attached to the lower loading

rod was used to provide environmental control. For evaluation at 4K, a

more sophisticated cryostat was used„ The system consists of an internal

load frame fabricated from filament wound glass-epoxy rods, a titanium

base plate, and a liquid nitrogen jacketed vacuum dewar.

Figure 3 shows a closeup view of the load train with a specimen

installed. An overall view of the sealed cryostat installed in the

testing machine is given in Figure 4. The cryostat is equipped with

level sensors so that economical use of liquid helium can be assured.

The system is precooled with liquid nitrogen; following cooling, the

liquid nitrogen is removed from the system by pressurizing the cryostat.

Care is taken to ensure that all liquid nitrogen is removed from the

system, because of the deleterious effect on helium resulting from

residual liquid nitrogen. Typical liquid helium consumption for the

system is approximately 40 liters per fill and test of moderate duration.

76


Figure 4 Overall View of Liquid Helium Testing System

79
Testing was performed in accordance with the ASTM recommended

procedure for J-j-^ determination, dated 3/1/77^^-^ . The initial specimen

evaluated at each temperature was loaded to maximum load. Succeeding

specimens were loaded to varying levels of displacement less than that

found at maximum load.

After unloading, each specimen was heat tinted at approximately

800K and then fatigue extended before fracturing. Crack extension was

measured at nine (9) points across the crack front to establish Aa,

the amount of quasistatic crack-growth.

J was calculated from the load displacement record according to

the following:

^ 2A ^ ^ 2P5

•J = "l i5 + ^2

where Cf^ and 0<2 are the Merkle-Corten coefficients (Reference 2) to

account for the tension component of loading

A = area under the load vs displacement curve;

B = specimen thickness;

b = remaining ligament (W - a);

P = final load; and

6 = final displacement.

The J — Aa resistance curve was developed in accordance with the

recommended procedure (Reference 3) with respect to establishment of

the blunting line (J = 2 O^^ ^a.) , number of data points, and size

requirements.

80
EXPERIMENTAL RESULTS

A total of thirteen (13) J integral tests were conducted o Four

each were performed at 293 and 4K; five were conducted at 77K.

The load vs loadline deflection curves are reproduced on a single

sheet for each test temperature and presented in Figures 5 through 7o

A tabular presentation of the data is given in Table 1,

Evaluation of the test result validity, in accordance with the

recommended guidelines, showed that the room temperature specimens

were significantly below the thickness and ligament criteria required

to establish validity. Tests conducted at 77 and 4K satisfy the thickness and ligament criteria.

Figures 8 and 9 present the J vs A a data in the prescribed

manner for determining j , Results were as follows:

Temp erature , K Jj^ , kN/m (ki p /in.)

77 280 (1.60)

4 86 (0.49)

DISCUSSION OF RESULTS

The results clearly show a decrease in fracture toughness of 21-6-9

stainless steel with decreasing temperature from 293 to 4K. At room

temperature, the toughness, in terms of J- integral, would be at least

875 kN/m (5 kip/in.). In terms of K^^ this would represent a value

of approximately 500 MN/m^/^ (450 ksi yirT.) . The J^^ value at 77K is

only 280 kN/m (1<,6 kip/in,), or less than a third of the minimum toughness anticipated. Converting this value to K gives a toughness of

81

83
84
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O O r-1 1-1

CnI CM CM CM

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CM 1-1 O O

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1-1 1-1 CM CO

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85
CRACK EXTENSION. Aa, in

CRACK EXTENSION, Aa, cm

Figure 8 J-Integrdl Data at 77 K

86
Figure 9 J-Integral Data at 4 K

87
approximately 280 MN/m^ (255 ksi yin.) , which is still a relatively high

toughness value.

The 4K data show a further decrease in toughness to less than onethird of that found at 77K. The J^^ value of 86 kN/m (490 lb/ in.)

represents a K_ level of approximately 150 MN/m (140 ksi ./ in. )

Ic V

Although this is a small fraction of the extremely high room temperature toughness, this level nevertheless represents tough behavior.

For example, a semicircular surface flaw defect in a component stressed

to 690 MN/m^ (100 ksi) would have to be approximately 3 cm (lo2 in.)

deep to cause failure under these conditions.

The data obtained in this program are in general agreement with

(1) the results previously reported by Tobler and Reed . A comparison

of results of the two investigations is given below:

Jjp, kN/m

Temperature, ' K MMC NBS

293 >875 ~1400

77 280 310-350

4 86 ~150

Ignoring the room temperature data, for which the results are invalid

and merely estimates of level, we find that the 77K data are quite close,

The 4K results are of similar relative magnitude,

REFERENCES

1, R« L. Tobler and R. P. Reed: "Tensile and Fracture Behavior of a

Nitrogen- Strengthened, Chromium- Nickel-Manganese Stainless Steel

at Cryogenic Temperatures," Materials Research for Superconducting

88
(continued) - Machinery, Semiannual Technical Report IV, October

1975, National Bureau of Standards, Boulder, Colorado.

ASTM Standards, E399-74, Part 10, American Society for Testing and

Materials, Philadelphia, PA, 1974,

Recommended Procedure for Jj^, Determination, ASTM Committee

E24-01-09, March 1, 1977.

89

Toughness, Fatigue Crack Growth, and Tensile Properties

of Three Nitrogen-Strengthened Stainless Steels

at Cryogenic Temperatures

National Bureau of Standards

91
1
TOUGHNESS, FATIGUE CRACK GROWTH, AND TENSILE PROPERTIES OF THREE

NITROGEN-STRENGTHENED STAINLESS STEELS AT CRYOGENIC TEMPERATURES*

D. T. Read and R. P. Reed

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

*Contribution of NBS, not subject to copyright.

Work supported by ERDA/MFE.

93
ABSTRACT

Fracture toughness, fatigue crack growth rate, and tensile property data

for three nitrogen-strengthened stainless steels at 295, 76, and 4 K have been

obtained. These materials were tested to determine their acceptability as

candidate structural materials for cryogenic service. The materials tested

were Fe-18Cr-3Ni-13Mn (Nitronic 33)*, Fe-21Cr-12Ni-5Mn (Nitronic 50), and

Fe-19Cr-9Ni-2Mn (304N). The nitrogen-strengthened alloy Fe-21Cr-6Ni-9Mn

(Nitronic 40) has been tested previously in our laboratory.

Toughness measurements employed J-integral , plane strain critical stress

intensity, and crack-opening-displacement tests. It was found that the

nitrogen-strengthened austenitic stainless steels generally exhibit lower

toughness at 4 K than the widely used 300-series stainless steels AISI 304,

AISI 310, and AISI 316. Of the nitrogen-strengthened steels tested to date,

Fe-18Cr-3Ni-13Mn is the strongest and least tough, and Fe-19Cr-9Ni-2Mn is the

toughnest with lowest strength. Fatigue crack growth rates at 4 K in

the nitrogen-strengthened stainless steels were found to be generally

higher than or comparable to rates for previously tested stainless

steels. Fe-18Cr-3Ni-l 3Mn had rates significantly above previously

tested materials, especially at high stress intensity range levels.

Crack growth rates for Fe-19Cr-9Ni-2Mn for an austenitic stainless steel

at 4 K and the rates for Fe-21Cr-12Ni-5Mn fell approximately in the midrange of previous data.

All studied materials exhibited high strengths at cryogenic temperatures.

After being strained to failure at 4 and 76 K, Fe-18Cr-3Ni-13Mn and Fe19Cr-9Ni-2Mn exhibited some ferromagnetism, indicating that body-centeredcubic strain-induced martensitic transformations had occurred during

straining; Fe-21Cr-12Ni-5Mn exhibited no ferromagnetism within experimental

uncertainty.

*Use of trade names is for clarity only, and does not imply acceptance or endorsement of a particular material by NBS.

94
INTRODUCTION

Nitrogen-strengthened austenitic stainless steels have attracted interest

as possible cryogenic structural materials because of their superior strength;^"^

but data on their toughness at cryogenic temperatures is generally lacking,

r C except for a few studies. ' The availability of high-strength, high-toughness

alloys for cryogenic service would contribute to increased structural safety,

and economy in structural material bulk in devices such as magneticfusion-energy systems.'' Cost savings allowed by reduced nickel content

without loss of low temperature toughness or austenite stability also

would be welcome. Therefore fracture toughness, fatigue crack growth rate,

and tensile property data for three nitrogen-strengthened stainless steels

have been obtained at 4, 76, and 295 K to determine their acceptability as

candidate structural materials for cryogenic service. The specific goal of

the test program reported here was to obtain 4 K mechanical property behavior

of the three tested alloys for comparison to the previously tested alloy

Fe-21Cr-6Ni-9Mn.^

MATERIALS

Two of the alloys tested are sold under the trade names Nitronic 33

and Nitronic 50. The third is usually sold as stainless steel AISI 304N.

The alloys will be denoted in the text by approximate weight percent of

iron, chromium, nickel, and manganese, their major constituents, as follows*:

Nitronic 33 Fe-18Cr-3Ni-13Mn

Nitronic 50 Fe-21Cr-12Ni-5Mn

AISI 304N Fe-19Cr-9Ni-2Mn

Nitronic 40 Fe-21Cr-6Ni-9Mn

*Use of trade names is for clarity only, and does not imply acceptance or endorsement of a particular material by NBS.

95
The chemical compositions of the nitrogen-strengthened steels used

in the present study as supplied by the manufacturers are listed in

Table 1. Also included in Table 1 for reference is the chemical composition of another nitrogen-strengthened alloy, Fe-21Cr-6Ni-9Mn. Fracture

5 and tensile data for that material have been reported separately, but

are repeated here where appropriate for comparison with present results.

These alloys have no familiar designations except trade names, and so

will be denoted in the text by the approximate weight percent of their

major constituents: iron, chromium, nickel, and manganese. Approximate

major constituent content, ASTM designation, relevant ASTM specifications,

and trade names are listed in Table 2. Fe-18Cr-3Ni-l 3Cr was obtained as

a 2.5 cm thick plate, Fe-21Cr-12Ni-5Mn as a 5 cm thick plate, and Fe19Cr-9Ni-2Mn as a short section of extruded tube about 1 m in diameter

by 5 cm wall thickness. Each material was tested as received, in a

commercially-annealed condition. A portion of one of the materials, Fe21Cr-12Ni-5Mn, was given an additional heat treatment, a 1.5 hr anneal

in an 1177°C furnace followed by a water quench. This material is

designated Fe-21Cr-12Ni-5Mn-X below. Final heat treatments, grain sizes

and hardnesses of the materials are listed in Table 3, and representative microstructures are given in Figs. 1-4. It can be seen that the

additional heat treatment applied to the Fe-21Cr-12Ni-5Mn-X alloy brought

its grain size and microstructural appearance close to those of the

other alloys.

SPECIMENS

Fracture toughness and fatigue crack growth data were obtained

using compact specimens of the shape shown in Fig. 5a in the TL orientation. Specimens of Fe-18Cr-3Ni-13Mn were 2.5 cm thick; all others were

3.8 cm thick. Tensile property data were obtained using the miniature

96
round tensile specimen geometry shown in Fig. 5b, oriented transverse to

the rolling direction. The orientation of fracture toughness and tensile

specimens relative to the original plates is shown in Fig. 5c. The

rolling direction of the large tube section was considered to be the

direction of the tube axis, and the transverse direction the circumferential direction.

TECHNIQUES

o Toughness data were obtained using J-integral techniques for all

materials at 295 and 76 K and for Fe-19Cr-9Ni -2Mn at 4 K. A conventional

critical stress intensity measurement technique (the Kj^ test^) was used

for the other materials at 4 K. In the J-integral technique several

specimens are precracked in tension-tension fatigue approximately to a

predetermined crack length. Each specimen is then loaded quasistatically

(load line displacement 50 microns/sec or less) under displacement

control until crack extension is produced. Load is recorded graphically

as a function of load-line displacement during the test. After unloading,

the specimens are heat-tinted to mark the crack extension Aa on the

fracture surface. The specimen is then broken in two and the crack

extension is measured. Reduction of the data consists of analyzing each

test record to determine the area under the load versus load-linedisplacement curve and its complement, calculating the J-integral value,

plotting the several J-Aa data, and determining the critical J value

o Jj^. The approximation often used in previous studies to calculate the

J-1ntegral value from the test data is:

97
where A is the area under the load-displacement curve, B is specimen

thickness, W is the width of the specimen from load line to face side, a

is crack depth after precracking measured from the load line, and b =

W-a. A more elaborate formulation was given by Merkle and Corten^^

0 - lb (2)

where A^ is the complement of the area A, n is a numerical factor of

about 2.2 which depends on the value of a/W, is a similar numerical

factor of about .1, and the other symbols have the same meanings as

before. Values of n and calculated by Merkle and Corten^^ are listed

in Table 4. This procedure has been recommended by the steering committee

of the ASTM E24-01-09 task group on elastic-plastic fracture. It

results in J values slightly higher than those obtained using Eq. (1)

and is considered to be more accurate. The critical value of J is its

value at the intersection of the J-Aa curve with the line given by

J = 2Aaa (3)

where a is the flow strength, taken as the average of the yield and

ultimate strengths. This graphical procedure is illustrated in Fig. 6

for Fe-18Cr-3Ni-13Mn at 76 K. The critical stress intensity factor Kj^

12

is inferred from the critical J-integral value using the equation:

^Ic = ^—2

where E is Young's modulus and v is Poisson's ratio for the material tested,

Values of E, v, and E' = E/(l-v^) used^"^'^^ are listed in Table 5.

98
The load-displacement records were also analyzed to obtain crackopening-displacement (COD) data. The COD value 6 for a given value of

1 ^ load line displacement V is inferred using:

, 45(W - a) rv _ n y > ^

(5a) ^ .45W + .55a e - E'

9 2Ya W

^ = •45(W - a) OV_, V < -p-T^ (5b)

.45W + .55a 4Ya^W

"

where is the yield strength, and y is a numerical factor depending on

a/W, approximately 2.5 for the a/W values obtained here, and E' = E/(l2 15

V ). Values of Y given by Wells are listed in Table 4. COD is extensively

used by the British in elastic-plastic fracture mechanics and in fact a

Draft British Standard exists for design using COD. The fracture resistance

of a material is characterized by its critical COD value. British

practice is to choose the critical value at maximum load or at "pop-in",

a peak in the load-displacement curve followed by a drop and subsequent

rise in load which signifies a sudden but small crack extension. Note

that the critical point on a test record analyzed for critical COD need

not coincide with the critical point determined in a Jj^ test.

Conventional Kj^ tests are performed by precracking compact specimens,

then loading them until fracture. A procedure detailed in ASTM standard

test method E 399 is used to determine a candidate Kj^ value and to determine

whether a successful measurement of Kj^ has been achieved.

Fatigue crack growth rates were also measured using compact specimens.

The specimen is loaded repeatedly (at a frequency of 20 Hz in this test

program) and the compliance, that is, the reciprocal slope of a plot of

load against load-line-displacement, is recorded periodically, at intervals

99
ranging from TOO to 20,000 cycles depending on the rate of crack growth.

1 c The compliance data are converted to crack lengths using the formula

a = W[-.9505 + .3447(EBC)^/^ + .5654(EBC)^/^ - .1 748(EBC)^/^] (6)

where C = displacement/load and E and B have been previously defined. From

these, crack growth rates are easily obtained. The crack length is also used

Q

in calculating the applied stress intensity factor range using the formula

'- (PTMx-Pmin'T?[29-6-185.5{^)+655.7(^)2- 1017(^)3+638.9(^)^] (7)

where P^^^ and P^^^ are the maximum and minimum loads. The end product of

the analysis is a plot of crack growth rate, da/dN, against stress intensity

factor range.

Conventional tensile testing techniques were used to obtain tensile

property data. Yield strength values at 0.2% offset are reported. The

flow strength is taken as the average of the yield and ultimate strengths.

After tensile testing, specimens of each material and test temperature were

checked for ferromagnetism, and thus for bcc martensite content, with a barmagnet, torsion-balance apparatus.

For all tests, temperature control was achieved by immersing the specimen in liquid helium or liquid nitrogen at local atmospheric pressure* or

by testing in laboratory air.

RESULTS

The J-Aa, Jj^, and Kj^ values obtained are listed in Tables 6-9.

Toughness values are listed in Table 10 and plotted against temperature in

Fig. 7. From this figure it can be seen that the toughnesses of all materials

^Typically 620 mm Hg. At this pressure, helium boils at 4 K and nitrogen

at 76 K. 100
decrease with decreasing temperature, as expected. The 4 K toughness of

two of the tested materials, Fe-18Cr-3Ni-13Mn and Fe-21Cr-l 2Ni -5Mn

(commercial heat treatment) were found to be definitely lower than the

value previously obtained for Fe-21Cr-6Ni-9Mn. Fe-19Cr-9Ni -2Mn was

found to have a higher toughness than Fe-21Cr-6Ni-9Mn, and Fe-21Cr-12Ni-5Mn-X

(special heat treatment) was found to have a toughness value approximately

equal (within a rather large experimental uncertainty) to that of Fe-21Cr6Ni-9Mn. The large uncertainty reported for the Fe-21Cr-12Ni-5Mn-X alloy is

due to the large difference between the toughness values obtained for the

two specimens tested, the significant deviation of one of those from the

requirements of ASTM Method E 399, and the fact that the heat treatment was

applied in a small laboratory furnace to machined specimens rather than

plate stock. Measured critical COD values are given in Table 11.

Fatigue crack growth data obtained are plotted in Figs. 8-10. Such

data is commonly represented by a Paris law of the form:

^ = C(AK)" (8)

Best fit (by least squares) Paris law lines are shown on the logarithmic

fatigue-data plots. The parameters of these best-fit lines are given in

Table 12. The fatigue crack growth rate for Fe-18Cr-3Ni-13Mn increased

strongly with decreasing temperature, and at 4 K was significantly higher

than any of the other materials tested and also higher than any other

austenitic stainless steel previously tested in this laboratory. Fe-19Cr9Ni-2Mn had the second highest fatigue crack growth rate at 4 K, with values

about the same as the highest observed for austenitic stainless steels

previously tested in this laboratory. Fe-21Cr-12Ni-5Mn had the lowest 4 K

fatigue crack growth rates of the three measured, in the mid-range of pre101
viously observed values. The fatigue crack growth rate for Fe-21Cr12Ni-5Mn-X was not measured.

Tensile data obtained are plotted in Figs. 11 and 12 and listed in

Table 13. Figure 11 shows that the strong increases of yield and ultimate

strengths expected for these nitrogen-strengthened materials were indeed

obtained. The order of yield strength values correlated inversely with

toughness, for each material as temperature was varied and among materials

at 4 K. The Fe-18Cr-3Ni-13Mn had the highest strength, Fe-21 Cr-1 2Ni -5Mn

(commercial anneal) was next, followed by this same alloy with the

special anneal at a value slightly higher than that observed previously

for Fe-21Cr-6Ni-9Mn. Fe-19Cr-9Ni-2Mn had the lowest 4 K yield strength

of the materials tested.

Results of the magnetic measurements on fractured tensile specimens

are listed in Table 14. Fe-19Cr-9Ni-2Mn fractured in tension at 76 K

was most magnetic (approximately 100% ferromagnetic martensite); the

specimen of this alloy fractured at 4 K was somewhat less ferromagnetic.

The Fe-18Cr-3Ni-13Mn alloy specimens fractured at 76 and 4 K were also

slightly magnetic. However Fe-21Cr-12Ni-5Mn specimens in commercial and

special heat treatments fractured at 4 K and commercially treated material

fractured at 76 K were not ferromagnetic within experimental uncertainty.

Photomicrographs of the fracture surfaces of the toughness specimens

tested at 4 K taken using scanning electron microscopy. Figs. 22a-d,

revealed differences in surface morphology among the specimen materials.

None appeared fully ductile, rather, relatively flat, brittle-appearing

areas were always present. Areas of apparent microvoid coalescence were

present to a significant degree in the Fe-19Cr-9Ni-2Mn specimen, to a

lesser degree in the Fe-21Cr-12Ni-5Mn-X and even less in the Fe-21Cr-12Ni5Mn and Fe-18Cr-3Ni-13Mn. The Fe-21Cr-12Ni-5Mn in its commercial heat

102
treatment contained noticeably more intergranular fracture than did the

same material in the special heat treatment.

Tensile specimens fractured at 4 K were sectioned, and examined using

optical microscopy. All 4 K specimens showed dark bands apparently along

{ill} planes. These are interpreted as due to the stress-induced

martensite e' (hep) and are illustrated in Fig. 23a for the case of

Fe-18Cr-3Ni-13Mn. The Fe-19Cr-9Ni-2Mn revealed, in addition, very small

lath-shaped martensitic phases associated with the a' (bcc) ferromagnetic

phase. The general appearance of these laths was the same as previously

observed for other stainless steels of approximate composition Fe-18Cr-8Ni.

No differences were found between Fe-19Cr-9Ni -2Mn specimens fractured at

76 and at 4 K.

Experimental Uncertainties

Uncertainties in toughness data are of three kinds: first, those

associated with the quantities directly measured, such as load, crack

length, displacement, and so on; second, those associated with the

procedure used to infer toughness values from the measured quantities; and

third, those which arise from specimen to specimen and heat to heat material

variability. Errors of measurement are readily assessed by repeated

measurement, measurement of standard loads or dimensions, calibration, and

so on. Errors associated with inferring a toughness value can be formally

eliminated by adoption of a standard practice. But questions about the

applicability of a value determined according to a standard practice may

persist. Errors associated with variability are the most severe. Evaluation

of material properties for structural design requires careful consideration

of uncertainties due to material variability, through understanding of its

causes and extensive data on its magnitude.

103
Detailed assessment of uncertainties associated with inferring toughness

from measured quantities and with material variability are beyond the scope

of this paper. Briefly, where comparisons have been made, values of toughness

inferred using Eq. 4 have usually been found less than or approximately the

same as those measured by the standard method (ASTM E 399). Values of critical

COD determined using Eqs. 5 can be considered to be correct by definition,

because while the meaning of the actual crack-opening-displacement is not

well-defined, in practice, values obtained using these equations are commonly

taken as correct. Measurements of plane strain critical stress intensity

Kj^ conducted according to ASTM E 399 are considered correct by definition.

Measurements reported here as not in strict compliance with E 399 are

considered to be of value comparable to those with no deviations, because

the deviations from compliance were generally considered minor.

Measurement uncertainties in individual J values are considered to be

of the order of 5% or less, from accumulated uncertainties in load cell and

displacement gauge calibrations and measurements of the area under the loaddisplacement record. Uncertainties in crack extension measurements are

larger, on the order of ten percent at large values and up to .05 mm at

low values due to difficulty in locating exactly the boundaries of the

region of crack extension on the fractured specimens. The cumulative

uncertainty in Jj^ values is estimated to be 10% or less, so the measurement uncertainty in Kj^ values inferred using Eq. 4 is 5% or less. Errors

in fracture toughness associated with material variability are generally

larger than this, and often much larger. Therefore, data reported here are

in general not applicable to materials with different chemical composition

or thermomechanical processing.

Uncertainties associated with inferring property values from measured

quantities are not a problem in fatigue and tensile property measurements,

because standard test procedures are widely accepted. However uncertainties

104
of measurement and material variability remian. The scatter visible in

the fatigue crack growth rate data presented in Figs. 8-10 is indicative

of the measurement uncertainties involved. Standard deviations of 20 to

40% from the average value were typical. Fatigue data is not considered

to be as sensitive to material variability as toughness data, and in

fact each data set in Figs. 8-10 consists of points from two or three

specimens, and all sets appear consistent. However, heat to heat

variations may be larger.

Measurement uncertainties in tensile properties are estimated to be

3 % or less; these are due to uncertainties in specimen dimensions, load

values as read from analog recorder traces, and (for the yield strength)

specimen strains from the calibrated extensometer.

The uncertainties discussed here for fracture toughness, fatigue

crack growth, and tensile properties are considered usual for such

measurements. The fact that many tests were conducted at cryogenic

temperatures did not contribute significantly to the experimental uncertainties.

DISCUSSION

An indication of the effectiveness of nitrogen-strengthening is

1

given in Fig. 13, which shows the yield strengths of AISI 304 and AISI

304N as a function of temperature. A nitrogen content of 0.12 wt % is

seen to increase the 4 K yield strength by approximately a factor of

two, with lesser effects at higher temperatures.

Only a small amount of previous data on the nitrogen-strengthened

austenitic stainless steels used in the present study is available. Montane

reported the tensile properties of Fe-18Cr-3Ni-13Mn between room and

liquid hydrogen temperatures. Figures 14 and 15 allow comparison of his

results to those of the present study. Agreement of strength properties

is good. Agreement of the elongation is poor. A possible reason for the

105
discrepancies is the difference in the specimen geometries used in the

two studies; Montano used a flat specimen with a ratio of width to

3 thickness of about 6 to 1 . Spaeder and Domis have reported tensile

properties of Fe-19Cr-9Ni -2Mn. Figures 16 and 17 allow comparison of

their results to those of the present study. Again, agreement among the

strength property values is good. The elongation and reduction of area

data agree at 77 K but differ somewhat at room temperature. The reason

for the disagreement is unknown.

Figure 18 shows for comparison three sets of toughness results for

Fe-21Cr-12Ni-5Mn: results of the present study for commercially-annealed

and specially annealed materials, and previous results by Dessau. A

significant increase in 4 K toughness due to higher annealing temperature

is apparent in this figure. This increase is believed to be due to

dissolution of precipitates and enhancement of grain size promoted by

higher annealing temperature, because large brittle precipitates at

grain boundaries are associated with intergranular fracture at low

stress intensity values and small grain sizes raise the yield strength,

decreasing crack tip plasticity. The influence of higher-temperature

annealing on mi crostructure can be seen clearly in Figs. 2 and 4 and the

grain-size data of Table 3.

4 K toughness data presented in Fig. 19 allow comparison between

nitrogen-strengthened and conventional stainless steels. As a group,

the nitrogen-strengthened steels are stronger but less tough than the

conventional alloys. Fracture toughness data is used in design to guide

material selection for prevention of brittle failure. For a material of

a given toughness, resistance to brittle behavior depends on section

thickness stress concentrations. Large thicknesses and large stress

concentrations promote brittle behavior. No exact calculation of the

106
maximum safe section thickness is available at present. A simple approximation is that section thicknesses greater than t, where

may be susceptible to brittle fracture. This relation may be used to

determine approximately the fracture toughness necessary to prevent

brittle failure in a material of given thickness and yield strength. In

Fig. 19, lines relating fracture toughness and yield strength according

to Eq. 9 for various values of section thickness t are shown. A material

for which the yield-strength, fracture toughness point lies above the

line given by Eq. 9 with the chosen section thickness should be used.

Figure 19 thus may be used to determine which materials should be considered

for use in sections of various thicknesses.

Figure 20 shows 4 K fatigue crack growth rate data for three nitrogenstrengthened stainless steels and a band which includes all 4 K crack growth

rate data obtained to date in our laboratory for conventional stainless

steels. 4 K fatigue crack growth rates for the nitrogen-strengthened stainless steels tend to be higher than for conventional stainless steels, although

Fe-21Cr-12Ni-5Mn is an exception, its rates falling below the center of the band

which includes the conventional austenitic steels.

Crack-opening-displacements (COD) values are usually related to J1

integral values by an equation of the form

where 6 is the crack opening displacement, is the material yield strength,

and m is a numerical factor with a value between 1 and 2.5 which varies with

(9)

J = (10)

107
material, temperature, and other factors. Equation (10) was tested

using each set of J-integral data obtained by plotting the J values

against COD values obtained from the same test records for each material

and temperature. A typical example of these plots is shown in Fig. 21.

From these plots it appeared that the data fit the generally accepted

linear relationship between J-integral and COD values although some

scatter was usually found. Values of the factor m obtained in the

present study ranged from 1.42 to 2.12; no systematic dependence of m on

any other measured parameters was discovered.

Note that even given the value of m, critical values of COD 6 c cannot generally be obtained from Eq. (9) and the critical value J^^

because the critical point for COD measurement is defined differently

from the critical point for J^^ measurement. The critical COD value

20 occurs at maximum load or pop-in,^ while the critical J-integral value

occurs at crack extension Aa equal to J/2ry. Thus, care must be exercised

in converting J-integral results to COD values.

SUMMARY AND CONCLUSIONS

1. Fracture toughness, fatigue crack growth and tensile property

data at 4, 76, and 295 K for nitrogen-strengthened austenitic stainless

steels Fe-12Cr-3Ni-9Mn, Fe-21Cr-12Ni-5Mn, and Fe- 1 9Cr-9Ni -2Mn have been

obtained and compared with previous data for another nitrogen-strengthened

stainless steel, Fe-21Cr-6Ni-9Mn.

2. Nitrogen-strengthened Fe-19Cr-9Ni-2Mn is tougher at 4 K than

nitrogen strengthened Fe-21 Cr-6Ni -9Mn alloy. Nitrogen-strengthened Fe1 8Cr-3Ni -1 3Mn and Fe-21 Cr-1 2Ni -5Mn , both in the commercially-annealed

condition are less tough at 4 K than Fe-21 Cr-6Ni -9Mn . Fe-21 Cr-1 2Ni -5Mn

with an additional heat treatment has approximately the same toughness

as Fe-21 Cr-6Ni -9Mn , within a rather large experimental uncertainty.

108
3. The 4 K fatigue crack growth resistance ranking of these alloys,

in descending order of resistance, is Fe-21Cr-12Ni-5Mn, Fe-19Cr-9Ni-2Mn,

Fe-18Cr-3Ni-13Mn. Fe-21Cr-6Ni-9Mn ranks between Fe-21Cr-12Ni-5Mn and

Fe-19Cr-9Ni-2Mn.

4. The 4 K strength ranking of these alloys is inverse to their

toughness ranking. The order of decreasing strength is: Fe-18Cr-3Ni13Mn, Fe-21Cr-12Ni-5Mn (commercial anneal), Fe-21Cr-12Ni-5Mn (special

anneal) and Fe-19Cr-9Ni-2Mn. Previously tested Fe-21Cr-6Ni-9Mn has

slightly lower 4 K yield strength than Fe-21 Cr-1 2Ni-5Mn with the special

heat treatment.

Helpful discussions with Dr. H. I. McHenry are acknowledged. The

assistance of Steven Naranjo, David Burkhalter, and Peter Steinmeyer in

testing, and optical and scanning electron nicroscopy is appreciated.

109
REFERENCES

1. D. C. Larbalestier and D. Evans, "High Strength Austenitic Stainless

Steels for Cryogenic Use," Sixth International Cryogenic Engineering

Conference, Grenoble, France, 1976, p. 345-347. IPC Science and

Technology Press Ltd., Surrey.

2. J. W. Montano, "Stress Corrosion Resistance and Cryogenic Temperature

Mechanical Behavior of 18-3Mn (Nitronic 33) Stainless Steel Parent and

Welded Material," NASA TM X-73309, Marshall Space Flight Center, Alabama.

3. C. E. Spaeder, Jr., and W. F. Domis, "Cryogenic and Elevated-Temperature

Mechanical Properties of a High-Nitrogen Type 304 Stainless Steel," Report

No. 44012-062-2, United States Steel Corporation, Pittsburgh.

4. R. B. Gunia and G. R. Woodrow, "Nitrogen Improves Engineering Properties of Chromium-Nickel Stainless Steels," Journal of Materials, JMLSA,

Vol. 5, No. 2, June 1970, p. 413-430.

5. R. L. Tobler and R. P. Reed, "Tensile and Fracture Behavior of a Nitrogen-Strengthened, Chromium-Nickel-Manganese Stainless Steel at

Cryogenic Temperatures," Symposium on Elastic-Plastic Fracture,

American Society for Testing and Materials, Atlanta, November 1977.

6. P. P. Dessau, "LN^ and LH^ Fracture Toughness of Armco Alloy 22-13-5,"

Report No. TID/SNA-2083, Aerojet Nuclear Systems Co., 1971.

7. R. P. Reed, F. R. Fickett, M. B. Kasen, and H. I. McHenry, "Magnetic

Fusion Energy Low Temperature Materials Program: A Survey," 1977, p. 28.

8. J. D. Landes and J. A. Begley, "Test Results from J-Integral Studies:

An Attempt to Establish a Jy Testing Procedure," Fracture Analysis , ASTM STP 560, American Society for Testing and Materials, 1974, p. 170-186.

9. Standard Method of Test for Plane-Strain Fracture Toughness of Metallic

Materials (Designation E 399-74), 1974 Annual Book of ASTM Standards,

Part 10 (1974), p. 432-451.

10. J. G. Merkle and H. T. Corten, "A J-Integral Analysis for the

Compact Specimen, Considering Axial Force as Well as Bending Effects,"

J. Pressure Vessel Tech., Trans ASME No-6, 1974, P. 1-7.

11. Recommended Procedure for J^ Determination, Steering Committee,

ASTM Task Group E24-01-09, Mrch 1 , 1977.

12. J. A. Begley and J. D. Landes, "The J-Integral as a Fracture Criterion",

in Fracture Toughness , ASTM STP 514, American Society for Testing

and Materials, 1972, p. 1-20.

110
13. H. M. Ledbetter, W. F. Weston, and E. R. Naimon, "Low-temperature

Elastic Properties of Four Austenitic Stainless Steels," J. App.

Phys. 46, (1975), p. 3855-3860.

14. H. M. Ledbetter, private communication.

15. A. A. Wells, "The Status of COD in Fracture Mechanics," Proceedings of

the Third Canadian Congress of Applied Mechanics, Calgary, 1971, p. 59-77.

16. H. I. McHenry, private communication.

17. R. P. Reed and R. P. Mikesell, "The Stability of Austenitic Stainless

Steels at Low Temperatures as Determined by Magnetic Measurements,"

Adv. Cry. Eng. 4, (1960), p. 84-100.

18. R. P. Reed and C. J. Guntner, "Stress-Induced Martensitic Transformations

in 18Cr-8Ni Steel", Trans. Met. Soc. AIMI 230, 1713-1720 (1964).

19. Handbook on Materials for Superconducting Machinery, Metals and

Ceramics Information Center, Battel le, Columbus, OH, Table 8.1.2-ME4.

20. Materials Research for Superconducting Machinery I-VI, edited by R. P.

Reed, A. F. Clark, E. C. van Reuth and H. M. Ledbetter. NTIS Order

Numbers AD780596, ADA004586, ADA012365, ADA019230, ADA030170, and ADA036919.

21. British Standards Institution, "Draft for Development on Crack Opening

Displacement (COD) Testing," DD 19 (1972).

Ill
LIST OF TABLES

Table 1. Chemical compositions of the ni tronen-strenqthened stainless

steels used in the present study.

Table 2. Approximate major constituent content, ASTM designation, relevant

ASTM specification numbers, and trade names of materials used in

the present study.

Table 3. Final anneal durations, grain sizes, and hardnesses of the materials

used in the present study.

Table 4. Elastic-plastic fracture mechanics parameters used in calculating

J-integral and COD values.

Table 5. Values of Young's modulus E, Poisson's ratio and plane strairi

modulus E' = E/(l- -2) used in calculations.

Table fi. J-'.a, J, , and K, values for Fe-18Cr-3Ni -1 3Cr.

Ic Ic Table 7. J-;a, J. , and K, values for Fe-?1 Cr- 1 2Ni -5Mn

Ic Ic Table 8. J- '.a, J. , and K. values for Fe- 1 9Cr-9Ni -2Mn

Ic Ic Table 9. K. values for Fe-21Cr-12Ni-SMn-X

ic

Table 10. Final toughness values obtained in the present study.

Table 11. Averaged critical (maximum load or pop-in) COD values obtained

in the present study.

Table 12. Paris-lav; fati gue-crack-grov/th-rate parameters for the studied

materials.

Table 13. Tensile data for the studied materials.

Table 14. Normalized bcc martensite content of fractured tensile specimens

of studied materials.

112
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113
Table 2. Major alloy content, ASTM designations, relevant ASTM plate specifications, trade names for the alloys studied, and a previously

studied alloy.

Fe-18Cr-3Ni-13Mn XM29 A240, A412 Nitronic 33 18-3Mn

Fe-21Cr-12Ni-5Mn XM19 A240, A412 Nitronic 50 22-13-5

Fe-19Cr-9Ni-2Mn AISI 304N A240

Fe-21Cr-6Ni-9Mn XMll A412 Nitronic 40 21-6-9

114
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03 E E

CO cn s: 1 CVJ 1 CTi 1

CVJ oo 1 1 1 1 CJ C_5 o C_J oo

cvj CVJ CVJ 1 1 cu cu cu <u CU

Li_ Ll_ Li_

CO

CO

fO .E

OJ

E:O

CO

CO

to

CU

s~

+-> fOcE

CD

+-> >>o

+-> ro1 ro

CUE CO

•r- 1—

4->

S- to o

4- o +-> E

(U

CVJ E+-> fO

CU

s-

-»-> to

to

CUo JZ

roE T3

S- SZ5 ro 4- T3E E ro -M

U)

CU 1 E EO

1—

115
Table 4. Elastic-olastic fracture mechanics parameters

used in calculating J-integral and COD values.

a/W n ^c Y

r> C . OJL \ .Oc.

'J . LUL.

A

. ^ 9 71 P. C . J \ o u . CUO 9 n/i

A c;

. H 0

0 9P/! C . COH u . cUO

.5 2.262 0.198 2.34

.55 2.238 0.188

.6 2.206 0.174 2.50

.65 2.178 0.156

.7 2.164 0.144 2.50

116
Table 5. Values of Young's modulus E;^ Poisson's ratio v, and plane

strain modulus E' = E/(l - used in calculations.'^*'^

"|- (K)

r-t GPa

V

GPa

re- 1 our-oNi - 1 orin 1 . y/ . ^ou 9 1 "J L,\ O

76 1 .97 .277 2.13

/It 1 QR . 911 LI 1 ? 1 Zl t. . 1

Fe-21Cr-12Ni-5Mn 295 1 .94 .295 2.13

76 2.08 .284 2.26

4 2.09 .283 2.27

Fe-19Cr-9Ni-2Mn 295 1.90 .290 2.07

76 2.05 .278 2.22

4 2.04 .271 2.20

117
Table 6. J-Aa, J. and K. values for Fe-18Cr-3Ni-13Mn.

T J Aa Kj^ Calculated or average measured K

(K) (kJ/m^) (mm) MPa/m (kj/m^) MPav'm

295 658 .66

295 752 .91 665^ 377

295 1140 2.06

295 1520 3.94

76 71.8 .036

76 86.3 .28

76 112 1.14 77 129

76 120 1.85

76 120 1.75

4 70 . 4 70 ' ' IV

4 73 .

^Invalid according to the proposed standard test method because specimens too small.

^Invalid according to ASTM E 399 due to slightly excessive crack front curvature.

This deviation is considered insignificant.

^Using Eq. 4

118
Table 7. J- /a, and KxIc

values for

T J

\ l\U /III /

1 mrn ^

y 1 1 II 1 1 / 1 1 r rx V 1 1

295 313 1 .32

cyO O . JO

fin 295 1211 14.0

-

7 C/D oo o yo . y 76 116 NA3

76 172 NA^

4 NA NA 108

4 NA NA 104^

Calculated or average measured K^^

:J/m^) MPA/m

228 221

87.5^ 141

106

^Fractured before unloading

'^Invalid according to ASTM E 399, because maximum load occured too far

above deviation of load-displacement curve from linear.

Estimate

Invalid according to ASTM E 399, due to excessive crack-front curvature.

119
Table 8. J-Zfi, Jj^ and calculated Kj^ values for Fe-19Cr-9Ni-2Mn

J Calculated K.

^ Ic ^

(kJ/m^) MPa/m

553^ 338

T J

(K) (k J/m^) (mm)

one 295 /I "70 4/2 .33

295 562 .46

r\ r\ t- 295 625 .74

295 794 1.17

295 1015 1 .57

76 448 .38

76 481 .43

76 698 r\ A .94

76 906 2.6

76 1010 3.8

76 1 230 4.9

4 224 .41

4 269 .41

4 354 .46

4 362 1 .78

4 494 2.06

4 723 3.43

486 329

208 214

^Calculated from Eq. 4.

'^Invalid according to the specimen size criterion of Ref. 11

120
Table 9 K^^ values for Fc-21 Cr-1 2N i -5Mn-X

(special heat treatment).

T Average K

(K) (MPav/m) (MPa/m^)

4 . 152

176

4 - 199^

^Invalid according to ASTM E 399 because B < 2.5

121
Table 10. Summary of measured and ca^lculated critical stress

intensity values, in MPa/m.

4 K 76 K 295 K

Fe-18Cr-3Ni-13Mn 71 131 377

Fe-21Cr-12Ni-5Mn 111 141 221

Fe-19Cr-9Ni-2Mn 214 329 338

Fe-21Cr-12Ni-5Mn-X 176+24

Fe-21Cr-6Ni-9Mn^ 182 272 489

^Data taken from Reference 5
Table 11. Average inferred critical Gnaximum load or pop-in)

crack-opening-displacement values, in mm.

4 K 76 K 295 K

Fe-18Cr-3Ni-13Mn .009 .044 .76

Fe-21Cr-12Ni-5Mn .025 .07 .33

Fe-19Cr-9Ni-2Mn .30 .49 1.01

Fe-21Cr-12Ni-9Mn-X .05

123
Table 12. Paris law fatigue crack growth parameters C and n for the studied

materials, for AK in MPav^ and da/dN in mm/cycle.

4 K 76 K 295 K

c r

ll r

U n

re- loCr-JNi - 1 3Nn 1 Otr ., Tr\~ll

. 3d X 10 b. 3^1 3 .15 X 10 3.98 n "7 Tr^^ll 9.7 X 10 3.8t

Fe-21Cr-12Ni-5Mn 3 .81 X 10"^^ 4.66 3 .86 X 10"^^ 3.94 9.8 X 10"^^ 3.8('

Fe-19Cr-9Ni-2Mn 1 .71 X lo'^ 3.46 4 .23 X 10"^ 3.02 3.05 X 10"^ 3.0!

Fe-21Cr-6Ni-9Mn^ 3 .6 X 10"^^ 4.4 1 .9 X 10-^0 3.7 1 .9 X 10"^^ 3.7,

^Reference 5.

= 'i fro

i

124
able 13. Tensile properties of the ni troqen-strenqthened stainless steels used in the present study and a pre- viously tested material

-18Cr-3Ni-13Mn

i

|-21Cr-12Mi- 5Mn

19Cr-9Ni-2Mn

|!-21Cr-12Ni-5rin-X

ipecial anneal

•21Cr-6Ni-9Mn'^

perature

(K)

Yield Strenqth

0.2 offset

MPa Tensile Strength

MPa Flow Strength^

MPa

Elongation

2.5 cm gaqe

length ('

)

Reducti( Area

295 421 787 604 54 53 459 803 632 58_ 51

440 7Qfi D 1 o 56 53

76 1 1 54 1522 1338 19 23 1134 1516 1325 18 25

11 44 1 51 9 1i J1^1.J 1 1 1> £4

4 1 572 1807 1689 5 23 1509 1815 1662 4 29

1 540 181

1 676 lb 295 534 845 690 35 64 522 840 681 35 58 528 842 UC) J jD 6 1

76 1 1 64 1584 1373 28 36

31 1 1 78 1 591 1384 29 1171 1 587 1 379 CO ~<A j4

4 1472 1898 1685 9 34 1422 1889 1655 9 32 1446 1893 1 670 Q 66 295 31 0 654 4f-2 61 70 290 644 467 60 69 300 549 474 DU /U

76 633 1512 1073 48 66 676 1 502 '1089 48 67 655 1 507 i JO 1 r r 66

4 846 1663 1255 37 42 743 1647 1195 32 47 794 1655 1 225 OH 44

4 1 307 1842 1575 16 36

1 353 1853 1604 15 41 1330 1848 1 SRQ 1 D 38 295 350 696 523 61 79 357 7^05 53J 61 78

JO J / (J 1 527 61 78

76 0.913 1462 1 188 42 32 0.886 1485 11 86 43 41 0.899 1 474 11 87 43 37

4 1258 1633 1446 16 40

1224 1 634 1429

1241 1634 1438 T6 40

ow strenqth is taken as the averaqe of the tensile and yield strenqths.

ta from Reference 5.

125
Table 14. Normalized magnetization of the studied materials,

given as (percent bcc martensite)/{percent elongation) from specimens fractured in tension.

T(K) Normal ized magnetization

Fe-18Cr-3Ni-13Mn 295 < .002

76 .05

4 .3

Fe-21Cr-12Ni-5Mn 295 < .003

76 < .005

4 < .012

Fe-21Cr-12Ni-5Mn-X 4 < .005

Fe-19Cr-9Ni-2Mn 295 .006

76 2.1

4 2.6

Fe-21Cr-6Ni-9Mn^ 76 .6

4 .3

Data taken from Reference 5.

126
LIST OF FIGURES

Figure 1. Composite Photomicrograph of Fe-1 8Cr-3Ni -1 3Mn

Figure 2. Composite Photomicrograph of Fe-21 Cr-1 2Ni -5Mn

Figure 3. Composite Photomicrograph of Fe-1 9Cr-9Ni -2Mn

Figure 4. Composite Photomicrograph of Fe-21 Cr-1 2Ni -5Mn-X (special anneal).

Figure 5. a. Fracture toughness specimen geometry used in the present

study (compact).

b. Tensile specimen geometry used in the present study,

c. Relative orientation of fracture toughness and tensile specimens and plate stock.

Figure 6. J-.\a curve for Fe- 1 8Cr-3Ni - 1 3Mn at 76 K.

Figure 7. Fracture toughness results at 4, 76, and 295 K for nitrogenstrengthened stainless steels used in the present study.

Figure 8. Fatigue crack growth rate results for Fe-18Cr-3Ni -1 3Mn at 295, 76,

and 4 K.

Figure 9. Fatigue crack growth rate results for Fe-21Cr-12Ni-5Mn at 295, 76,

and 4 K.

Figure 10. Fatigue crack growth rate results for Fe-19Cr-9Ni-2Mn at 295, 76,

and 4 K.

Figure 11. Yield and ultimate strengths of materials used in the present study,

Figure 12. Elongations and reduction of areas of the materials used in the

present study.

Figure 13. Yield strengths of AISI 304 and AISI 304N (Fe-19Cr-9Ni-2Mn) between

4 and 300 K.

Figure 14. Strength property data for Fe-18Cr-3Ni-13Mn from the present study

and from Montane.

Figure 15. Elongation data for Fe-18Cr-3Ni-13Mn from the present study and

from FIontano.2

Figure 16. Strength property data for Fe-19Cr-9Ni-2Mn from the present study

and from Spaeder and Domis.-^

Figure 1 7 . Elongation and reduction of area data for Fe-19Cr-9Ni -2Mn from the

present study and from Spaeder and Domis.3

127
Figure 18. Three sets of toughness results for Fe-21Cr-12Ni-5Mn; two from

the present study and one by Dessau.^

Figure 19. Fracture toughness plotted against yield strength at 4 K for

conventional and nitrogen-strengthened austenitic stainless

steels.

Figure 20. Fatigue-crack-growth rate data for nitrogen-strengthened

stainless steels obtained in the present study, for comparison with

the range of data previously obtained for conventional austenitic

stainless steels.

Figure 21 .Crack-opening-displacement plotted against J-integral values for

Fe-19Cr-9Ni-2Mn at 76 K.

Figure 22.SEM micrographs of fracture surfaces of toughness specimens

fractured at 4 K. a. Fracture surface of Fe-19Cr-9Ni-2Mn tested at 4 K. lOOOX.

b. Fracture surface of Fe-18Cr-3Ni-13Mn tested at 4 K. lOOOX.

c. Fracture surface of Fe-21Cr-12Ni-5Mn tested at 4 K. lOOOX.

d. Fracture surface of Fe-21Cr-12Ni-5Mn-X tested at 4 K. 500X

Figure 23.Photomi crographs of sectioned tensile specimens fractured at 4 K. a. Fe-18Cr-3Ni-13Mn 450X.

b. Fe-19Cr-9Ni-2Mn 850X.

128
Fe-18Cr-3Ni-13Mn

S

L - Longitudinal or rolling direction

T - Long transverse or width direction

S - Short transverse or plate thickness direction

Figure 1. Composite Photomicrograph of Fe-18Cr-3Ni-13Mn.

129
Fe-21Cr-12Ni-5Mn

(commercial heat treatment)

S

L - Longitudinal or rolling direction

T - Long transverse or width direction

S - Short transverse or plate thickness direction

Figure 2. Composite Photomicrograph of Fe-21Cr-12Ni-5Mn.

130
Fe-19Cr-9Ni-2Mn

.2 mm

L - Longitudinal or rolling direction

T - Long transverse or width direction

S - Short transverse or plate thickness direction

Figure 3. Composite Photomicrograph of Fe-19Cr-9Ni-2Mn.

131
Fe-21CM2Ni-5Mn-X

(special heat treatment)

.2mm

f

S

I - Longitudinal or rolling direction

T - Long transverse or width direction

S - Short transverse or plate thickness direction

Figure 4. Composite Photomicrograph of Fe-21Cr-12Ni-5rin-X (special anneal).

132
Fracture toughness specimen geometry used in the present

study (compact).

133
32 mm

Figure 5. b. Tensile specimen geometry used in the present study.

134
Figure 5. c. Relative orientation of fracture toughness and tensile specimens and plate stock.

135
^ CM O OO CO ^

CM

(jUi/PMlP 3mvA iva93iNi-r

136
0 100 200 300

TEMPERATURE (K)

Figure 7. Fracture toughness results at 4, 76, and 295 K for nitrogenstrengthened stainless steels used in the present study.

137

139
140
Yield' Ultimate

0 100 200 300

TEMPERATURE (K)

Figure 11. Yield and ultimate strengths of materials used in the present study.

141
^ 100

o

80

60

40

— 20

C9

0

El Red^A

Fe-18Cr-3Ni-13Mn o • —

Fe-21Cr-12IMi-5IVIn

Fe-19Cr-9Ni-2Mn A

Fe-21Cr-12Ni-5Mn-X O

Fe-21Cr-6l\li-9Mn v

0 100 200

TEMPERATURE (K)

Figure 12. Elongations and reduction of areas of the materials used in the

oresent study.

. ; 2
1000

CO

800

CO

3. 600

CD

400

200

0

Fe-19Cr-9Ni-2Mn Alloys • AISI 304

(handbook)

AISI 304 N

(Spaeder and Domis^

AISI 304 N

(present)

100 200 300

TEMPERATURE (K)

Figure 13. Yield strengths of AISI 304 and AlSI 304N (Fe-19Cr-9Ni-2Mn) between

4 and 300 K..

143
2000

1500

Qu

LlJ

I—CO

1000

500

Fe-18Cr-3Ni-13Mn

Yield Ult

Present Results

o Montano^

100 200

TEMPERATURE, K

300

Figure 14. Strength property data for Fe-18Cr-3Ni -1 3Mn from the present study

and from Flontano.^

144
100

T

Fe-18Cr-3Ni-13Mn

0 100 200

TEMPERATURE, K

Figure 15. Elongation data for Fe-18Cr-3Ni-13Mn from the present study and

from Montano.2

145
TEMPERATURE, K

Figure 15. Strength property data for Ee-19Cr-9Ni-2Mn from the present study

and from Spaeder and Domis.

146
100

80

20

Fe-19Cr-9Ni-2Mn

Elong,

Present Study

Spaeder and Domis O

Red. of

Area

100 200

TEMPERATURE, K

300

Figure 17. Elongation and reduction of area data for Fe-19Cr-9Ni-2Mn from the

present study and from Spaeder and Domis.

147
CO

CO

C/3

CO

CO

400

300

100

Fe-21Cr-12Ni-5Mn • Present, mill annealed

(112rC)

A Present, specially annealed -

(1177° C)

Dessau6 (1066X)

100 200

TEMPERATURE (K)

300

Figure 18. Three sets of toughness results for Fe-21Cr-12Ni-5Mn; two from

the present study and one by Dessau.^

148
o V)

^4- CO ^ 1

c

-1-ro

+-> 4-> to U) -C O

4-> -p4-> C -p- QJ C

+-> +->

</) (/) O fO r

OJ -o

•I- <u >) c

0)

+-> ^to •<-> C CD

•I- E

«a O) O) s-

(O »->

t/5 o 1 <u c

+-> (U

4-> cn o o

1— s- CL+J

to c

OJ T3 E E

x: »o

CJ) 3 1

O 03

-!-> Eo

OJ -1- S- -M • 3 E CO

+-> <U o > OJ

(0 E <L) S- O +J U- CO

0)

Szscn iZ

149
I II I I—I—{Mill I I I—

o

_e/3 »

C/3

C/3 en

I

ste inle

CO o es

C/3 CO LT) ss O

o U Da us 'E "E

om o 'E

o o "> ^ II o V (/) 1—

c 3

UJ o. CD

UN I I I I Inn liHii i I I I liiil I Mi l CM CO

(a|oAo/uiui) Mp/ep giva HIMOUD >l3Va3

150
CM

<:

a:

I

(uiiu)- 9 lN3W30VldSia-9NIN3dO->l3VyD

o

>

SO)

+->

I

-M

1/1

(13 T3

O)

-t-J O

CD • E ^

O)0 «D

03 1^

Q. 4-> CO fO

•r— a c

1 s:

CD CM C I c ^(V CTi Q. I O SI C_3 ^ C3~> O 1—

(O I S- CU

C_5 Ll_

0)

CD

151
a. Fracture surface of Fe-19Cr-9Ni-2Mn tested at 4 K. lOOOX.

b. Fracture surface of Fe-18Cr-3Ni-13Mn tested at 4 K. lOOOX.

Figure 22.SEM micrographs of fracture surfaces of toughness specimens

fractured at 4 K.
c. Fracture surface of Fe-21Cr-12Ni-5Mn tested at 4 K. lOOOX.

d. Fracture surface of Fe-21Cr-12Ni-5Mn-X tested at 4 K. 500X.

Figure 22 (continued). SEM micrographs of fracture surfaces of toughness

specimens fractured at 4 K.

153
t i. tt 0

' * l» tr $f ,

b. Fe-19Cr-9Ni-2Mn 850X.

Figure 23. Photomicrographs of sectioned tensile specimens fractured at 4 K.

154
STAINLESS STEEL WELDING

155

Evaluation of Stainless Steel Weld Metals

at Cryogenic Temperatures

National Bureau of Standards

157

EVALUATION OF STAINLESS STEEL WELD

METALS AT CRYOGENIC TEMPERATURES

Harry I . McHenry

Cryogenics Division

National Bureau of Standards

Boulder, CO 80302

Austenitic stainless steels such as AISI types 304 and 316 and the

nitrogen strengthened 21Cr-6Nf-9Mn alloy (Nitronic 40) are readily weldable

by all common welding processes, providing the appropriate consummables

and procedures are used. The service experience with welded assemblies

of these alloys has been satisfactory at cryogenic temperatures. A

contributing factor to this success has been low design stresses; generally

the stresses have been limited to about 140 MPa (20 ksi), the ASME

allowable stress for AISI 304 and 316 stainless steels at room temperature.

For large superconducting magnets, the enormous forces that must

be contained and the limited space available for structure may necessitate

operating at significantly higher stresses, stresses based on the strength

of the alloy at the operating temperature. Many of the unwelded austenitic

stainless steels exhibit excellent fracture toughness at temperatures down

to 4 K in the base metal. However, weld metal fracture toughness has not

yet been determined at 4 K. Charpy impact results at 76 and 20 K suggest

that the toughness may be marginal at 4 K, particularly if the stress levels

in the magnets are higher than current practice.

A program is being conducted to evaluate stainless steel welds at cryogenic temperatures. The objectives of the initial phase of investigation are

1. To investigate the metallurgical factors contributing to the toughness, or degradation thereof, of stainless steel weld metals.

2. To develop a high strength, high fracture toughness, filler wire for

welding nitrogen-strengthened stainless steels such as the 21Cr-6Ni-9Mn alloy

159
3. To evaluate the fracture toughness of stainless steel welds at

4 K. The approach has been to draw on the experience and expertise of

filler metal suppliers to develop the high strength, high fracture

toughness, filler wire and to provide welds for the metallurgical and

fracture mechanics evaluations conducted by NBS. The shielded metal arc

process is being used to produce the test weldments because this process

permits control of the weld deposit chemistry by modifying the chemistry

of the electrode coating. Upon completion of the initial phase of the

investigation, the most promising weld metals will be used to evaluate

the influence of other welding processes and process variables on strength

and toughness at 4 K. The test weldments will be supplied to NBS by an

industrial firm with experience in fabricating cryogenic structures.

At the present time, three contracts with industry have been awarded and

the work on each contract has been completed: (1) Armco Steel Corp., for the

development of a high strength filler metal; (2) Arcos Corp., for the preparation of test weldments designed to study the effect of ferrite content, ferrite morphology and sensitization effects; and (3) Teledyne-McKay to prepare

test weldments with AWS 330 filler metal. This report summarizes the work

done on these contracts plus the related work done by NBS.

Metallurgical Effects on Toughness

A joint program is being conducted by NBS and the Arcos Corp. to investigate the effects of ferrite content, ferrite morphology and sensitization on

toughness at cryogenic temperatures. The investigation is being conducted on

316 base plate (25 mm thick) welded with AWS 316L electrodes.

The Arcos program consisted of manufacturing controlled chemistry (within

AWS 31 6L chemistry limits) electrodes, welding a series of test plates and

evaluating the influence of process variables on ferrite content. The

chemistries and ferrite contents of the five electrodes and the test plate

design are given here in the Arcos report. Electrodes 1 through 4 were formulated to systematically vary ferrite content from 0 to 12%. Electrode 5

160
was nominally the same as electrode 2 in chemistry except the carbon content

was 0.059% which is typical of AWS 316 weld metal. Since carbon is an austenite

stabilizer, the ferrite content of electrode 5 was essentially zero, i.e.

comparable to electrode 1.

Arcos welded a series of ten test plates using the parameters summarized

in Table 1. Test plates 1 through 4 were welded used electrodes 1 through 4,

respectively, to systematically vary the ferrite content between 0 and 12%.

Test plates 5, 9 and 10 were welded to evaluate the effect of carbon content

and cooling rate on the degree of sensitization. Plate 5 (C = 0.059%) was

welded using the same procedures used to weld Plate 1 (C = 0.031%), i.e. procedures that result in normal cooling rates. Plate 9 (C = 0.031%) and plate 10

(C = 0.059%) were welded using slow-cooling-rate welding parameters, i.e. high

preheat and interpass temperatures, high current and low travel speed.

Plates 6, 7 and 8 were welded using electrode 3 (9% ferrite) and welding

I parameters designed to significantly alter the cooling rate in order to vary

i the ferrite morphology.

j Arcos also conducted a study of ferrite in 316L weld metal. Ferrite

il content was measured for a variety of welding conditions resulting in difij fering cooling rates. The results are reported in the Arcos report.

The ten test plates have been delivered to NBS and machining of test

|j specimens is in progress. NBS will conduct tensile, Charpy impact and

fracture toughness tests and a metallurgical evaluation on each of the weldments. The metallurgical evaluation of the weldments will consist of

x-ray diffraction measurement of ferrite content, metal lographic examination

of ferrite morphology and sensitization tests on the weld metal and the HAZ.

High Strength Filler Wire Development

Armco Steel Corp. conducted a program to develop a high strength filler

wire for welding nitrogen strengthened stainless steels such as the 21Cr-6Ni-9Mn

161
alloy (Nitronic 40). The results of their program are sunimarized here

in the Armco report. As part of this contract, Armco will supply NBS

with a test plate (25 mm thick, .46 m of weld length) from each of the

three chemistries evaluated. Since nitrogen has an adverse effect on

toughness, the three test plates will all be welded with the minimum

nitrogen level, to permit comparison of the experimental alloys with the

data on AWS 300-series weld metals.

Evaluation of AWS 330

Teledyne-McKay Corp. furnished NBS with two test plates welded with a

modified version of AWS 330 filler metal. The principal modification is in

Mn content, the modified 330 has nominally 4.5% Mn instead of the 2.5% Max

specified in AWS 330. Two joint configurations were evaluated. An open joint

having the AWS specification A5.4 weld pad configuration was used to nive undiluted

weld metal in the test section. A double-Vee joint with a 2.4 mm land typical

of commercial practice was used to give a diluted weld joint. The base metal

was 25 mm thick plate of AISI 304L stainless steel in the annealed condition.

Teledyne-McKay also evaluated welded test pads of the same batch of

electrodes used for the test plates. The test pads were prepared in accordance

with AWS specification A5.4 using 19 mm thick mild steel plates buttered with

the test electrodes. Tests were conducted to determine the weld metal composition, tensile properties at room temperature and Charpy impact toughness at

76 K. The results are summarized in Table 2.

The test plates have been machined into tensile and fracture toughness

specimens for evaluation by NBS. The tensile properties were determined at

room temperature, 76 K and 4 K. The results, summarized in Table 3, indicate

that dilution with the base metal causes a slight decrease in strength. The

strength at 4 K, 860 MPa (125 ksi) yield and 1090 MPa (158 ksi) ultimate, of the

162
diluted weld joint is higher than the strengths of AISI 304 and 316 base

metals at 4 K. Fracture toughness tests are in progress.

Future Work

During FY 78, the evaluation of the test plates furnished by Arcos,

Armco and Teledyne-McKay will be completed. On the basis of the test

results, filler metals will be selected for further evaluation using the

gas metal arc, gas tungsten arc and the submerged arc processes.

163
dz

(U

+-> z1 

>

U-.

Q (L)

in a; o >

t_) i. 1 

1^

jQ

(/) o n 1 (U XJ (/) Q o rrl lU s. CL. o

(U

VD </) to 1 1 CO fOQ.

CO SdJ Q_ 5c S- +J

cu c QJ

f- +J 1—

1 (U

•IP— E 

to

sO D_ <u CI > Q-' <U c

hT3

c/) 4^

It' "q. -M 4J

C/) <U

4-> C

VJ,/ OJ Q4J S- E

C/) s- CO 3o (/)

(U

(/) -oo

S-

+-> /-"«. CO o

<D F—

CO

(—

<:

to <U O -M • fO O

s-

— s: <C D0) 00 -M • QQ fO O

CO

LD CM o CO LD CT> o CTi r— O o CM

>— t— +

CO

ID CM LD CO CO CM o 00 O CO cy> o CM

Ln CO CO cy> CD CM CO

CO CM CM CO CO CO CM CM CM o o o o O o o o o O Ln LD LD LD LO LD LO LO LO CM1 CM CM CM CM1 CM CM CO CD CO o 1 o O O o 1 o o o 1 o O o o O O o o o LO LD LO

CVJ CM CM CM CVJ CM CM LO LD LD

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164
Table 2. Properties of AWS 330 modified test pads measured by Tel edyne -McKay.

Weld Metal Composition:

7oC %Mn %Si %? %S %Cr %Ni %Mo %N

AWS 330 .25 max 2.5 max 0.6 max .04 max .03 max 14 to 17 33 to 37 —

Pad 1 .147 4.47 .51 .019 .007 17.80 34.40 .25

Pad 2 .136 4.70 .45 — .006 17.82 34.36 — .044

Tensile Properties at Room Temperature (.503 inch diameter specimen)

Ultimate Yield % %

MPa/ksi MPa/ksi Elong. m2" RA

585/84.8 382/55.4 45 56

Charpy Impact at 77 K

Energy Lateral Expansion

J Ft-Lbs mm mils

91.0 66.9 1.45 57

90.0 66.2 1.17 46

80.6 59.3 1 .30 51

80.0 58.5 1 .37 54

87.4 64.3 1.35 53

Average: 84.3 62.2 1.33 52.2

165
Table 3. Tensile properties of AWS 330 weld metal.

OCtlll|-' 1 c Temperature

^uts ^^^^^ c "/

10

UNDILUTED

Al DTK 1 66 83 29

B4 RT 58 85 32

A2 1 0

1/i\ 107 142 38

DC 76 K IDA 141

A4 /I"t l\ 128 169 30

R3 4 K 1 3R 1 74 O 1

DILUTED

Bl RT 59.9 84.9 34

A2 RT 55.5 83.1 35

A3 76 K 91.5 131 47

83 76 K 88.9 133 49

A4 4 K 124 157 42

84 4 K 125 159 44

166
Weldability of 21-6-9 Stainless Steel

Armco

167

WELDABILITT OF 21-6-9 STAINLESS STEEL

(ARMCO NITRONIC 40 STAINLESS STEEL)

R. Harry Espy

Armco Steel Corporation

Middletown, OH 4504 3

INTRODUCTION

The 21-6-9 stainless steel is a nitrogen strengthened austenitic

alloy having good mechanical properties over a wide range of temperatures. Its ,2io yield strength at room temperature is about

two times that of Type 304 stainless steel. The alloy is readily

joined by welding using any of the standard processes in use today.

When a matching weld filler is employed (NITRONIC 40W) , sound welds

in all thicknesses having good toughness at temperatures to -100°F

are readily obtained. In welds where thicknesses are less than a

1/4", satisfactory toughness has been obtained when tested at

-320°F. Where thicknesses of the weld are 1/4" or greater, toughness at -320°F as measured by the CVN test average about 10 ft. lbs.

Such a value when compared to Type 304 stainless steel (Type 30^

weld metal - 26 ft. lbs. at -320°F) is not favorable, especially when acceptable limits of 15 to 20 ft. lbs. are used. At

room temperature, the impact strengths of heavy weldments in Type

304 (30S weld metal) and 21-6-9 (NITRONIC 40W - weld metal) are

about the same - 40 to 45 ft. lbs., even though the yield strength

of 21-6-9 weld metal (NITRONIC 40W) is 20,000 psi higher than the

305 weld metal. The loss in impact strength at -320 F for 21-6-9

weld metal is related to the fact that the yield strength at -320°F

is increased more than that of the 30S weld metal as shown in the

table.

Room Temperature -320°F

CVN CVN

.2% Yld. Str. Toughness .2% Yld. Str. Toughness

308 Weld Metal

SMA Process 62,000 43 ft. lbs. 78,000 psi 26 ft. lbs,

21-6-9 Weld Metal

(NITRONIC 40W) 82,000 40 ft. lbs. 160,000 psi 10 ft. lbs

SMA Process

Development work under this contract aimed at producing a weld

having improved toughness and a reasonable strength at -320°F

was divided into 3 areas as shown.

Cr Ni Mn (weight percent)

(1) 13-20-9 + Cb (Ferrite free) N & Si Varied

(2) 20-10-9 (Ferrite containing) N Varied

(3) 1^-16-9 (Ferrite free) N Varied

169
TEST PROCEDURE

Prepare a total of eleven lots of covered welding electrodes

using 3 base core wire chemistries as shown. Each lot designed

to produce the aim weld deposit composition listed.

Type 13-20-9+ Cb (N&Si Varied) Ferrite-Free

C Mn P Q C.-r Ni N Cb

Core Wire

Chemistry Aim .04 9.0 .015 .015 .40 14.0 20.0 low l.S

SMA Weld

Deposit Aim

(A)

(B)

(C)

(D)

(E)

tt

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Type 20-'±\J—y in Q (N Varled) Ferrite Containing

C Mn P S Si Cr Ni N

Core Wire

Chemistry Aim .04 9.0 .015 .015 .40 19.5 9.0 .04

SMA Weld

Deposit Aim

(F)

(G)

(H)

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Type -9 (N Varied) Ferrite-Free

C Mn P S Si Cr Ni N

Core Wire

Chemistry Aim .04 9.0 .015 .015 .40 IS.O 16.0 .04

SMA Weld

Deposit Aim

(I)

(J)

(K)

tt

tt

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tt

tt

tt

tt

It

It

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11

It

II

It

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.14

.20

Examine the SMA weld deposits obtained with each of the eleven

lots of covered electrodes for:

(1) Tensile Strength at 300°K (Room Temperature)

(2) Tensile Strength at 77°K (-320°F)

170
(3) CVN Toughness at 300*'K (Room Temperature)

(4) CW Toughness at 77°K (-320°F)

(5) FN (Ferrite Number) at 300°K (Room Temperature)

(6) Metallurgical Structure

The welds from which the above specimens were obtained were made

according to AWS 5.4-69. The tensile specimens for test at

300°K (room temperature) were .505" diameter and for test at

77°K were .357" diameter. The CVN specimens were 10 mm as per

ASTM E23. The metallographic specimens were sectioned transverse

to weld direction and examined optically up to 500X. Analysis

for composition was by X-ray, emission and Leco combustion

methods calibrated with acceptable standards. Magnetic attraction of test piece fracture faces was measured by magnetized

needle and rated for change as an indicator of austenite

stability during plastic deformation at test temperature.

TEST RESULTS

The test results using the eleven groups (A to K) of coated

electrodes for deposition of welds are shown in Tables I to

IV. Weldability of the coated electrode lots appeared normal,

each performing in a uniform manner when good stainless steel

welding practice was employed.

DISCUSSION OF TEST RESULTS

In comparing properties of the several alloy systems examined in

this project the unwelded 21-6-9 (Armco NITRONIC 40) base metal

and matching 21-6-9 (Armco NITRONIC 40W) all-weld metal are used

as shown below.

21-6-9 BASE METAL PROPERTIES (TESTED UP TO 5" THICK)

Composition

Typical

Tensile

Toughness

Structure

U

Mn Si Cr N

.03 9.0 .025 .015 .60 20.5

Room Temperature

Ni

6.75 .30

-320°F

R Mag

U

103 58 - 70%

205 'ts CVN

Austenite

203 150 - 24%

65 '#s CVN

Austenite

Mag

171
21-6-9 SMA ALL-WELD METAL PROPERTIZS (1" THICX)

Composition

Typical

Tensile

Toughness

Structure

C Jto 5_ 3i Cr

Room TemDerat'jre

T2T

FNT

T

-320°F

S 7^ jez eccs

:to " 109 32 31^ l^of" Change None 20 5 160 29fo 26f5 Change None

40 '#3 C7N No None 10 '#3 CVN

1 to 5/5 Ferrite in a Matriic of Austenita

No None

U - Ultiicate Tensile Strength SSI T - .2^ Tld.- Str. 131

S - Elongation 2" R - Reduction of Area Mag. - Ifegnetisn

The lesser toughness of welds when compared to unwelded base metal

(wrought) is related to the cast structure of the weld plus the

fact that the weld is also higher in strength. This phenomena is

similar to that seen when Type 304 base metal is compared to Type

30^, its matching weld metal,

13-20-9+Cb (With N Varied) Ferrite-Free

The use of Cb in this ferrite-free weld deposit system resulted

in sound defect free welds in all cases with the exception of a

few points of porosity noted in the higher N containing welds.

From the results tabulated below it appears that at -320°F yields

of approximately 100 KSI with toughnesses greater than 15*#s can

be obtained with N contents of approximately .10, The effect of

Si in this Cb containing system was not clear with variations in

the range examined being of no significance.

Effect of N on Properties of

Cb Containing Ferrite-Free Weld Deposits

Room Temp, -320°F

Code N Si

A .05 .30

D .09 .53

B .11 .30

E .16 .44

C .IS .23

op,

T OTT T"^ nW

65 46'#s 91 2g»#s

65 39'#s 94 22'#s

66 3 5*#s 103 15*#s

69 37»#s 9S 19*#s

75 27'#s 117 12'#s

Y - ,2fo Yld, Str, KSI

172
20-10-9 (N Varied) Ferrite Containing

The ferrite levels, obtained with this alloy balance, of 1.4 to

4.2 FN were sufficient in all cases to produce sound defect-free

welds. While strengths were good even at the lowest level of N

used, none of the toughness values at -320°F were satisfactory.

Tabulated below is a summation of results related to N contents.

Effect of N on Properties

of Welds Containing Ferrite

Code N

Room Temp. 77°K (-320°F)

Y__ cw Y m

F .09 74 47'#s 115 9»#s

G .14 73 51*#s 137 10'#s

H .19 go 49 #s 153 9»#s

Y - .2/0 Yld. Str. KSI

1^-16-9 (N Varied) Ferrite-Free

In selecting this chemistry it was known that varying degrees of

grain foundary defects would occur. To determine at what strength

such grain boundary weaknesses would become a problem three levels

of N were examined. Good toughness-strength relationships appear

possible at approximately .15 N with yields at -320°F of approximately 126 KSI and toughnesses of approximately lS*#s.

CONCLUSIONS AND FUTURE WORK

In this project to develop a weld metal for ioining heavy sections

of 21-6-9 stainless steel (Armco NITRONIC 40} three basic alloy

systems were examined and can be commented on as follows. The

ferrite-free lS-16-9 with N at approximately .15 offered the best

combination of toughness (lS«#s CVN) and strength (126 KSI .2 Yield)

at -320°F. The grain boundary problem which had little effect on

CVN toughness is deserving of further evaluation. The ferrite-free

13-20-9+Cb with N at approximately .10 gave sound defect-free welds

with an acceptable toughness (15*#s CVN) at a lower level of

strength (100 KSI .2 Yield) at -320°F. The ferrite containing

20-10-9 welds were all sound and free of defects but did not exhibit any acceptable levels of toughness.

Weld joints in 1" thick 21-6-9 plate will be made using one

chemistry from each of the three weXJ deposit groups and forwarded to the NBS lab at Boulder for J-integral testing.

173
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Ferrite in Type 316L Weld Metal

Arcos

179

FERRITE IN TYPE 316L WELD METAL

R. David Thomas, Jr.

Arcos Corporation

1500 South 50th Street

Philadelphia, PA 19143

BACKGROUND

For the Magnetic Fusion Energy program, the Department of Energy has

authorized an investigation of the materials. Including weldments, to be used

in the construction of large superconducting magnets. Involved in the study

is the notch toughness of the materials when operating at liquid helium temperatures. The structures will involve thick weldments; hence, welding

processes may require large welding energy input. Besides shielded metal arc,

the processes will probably involve submerged-arc and electroslag.

The location of welds may be such that magnetic permeability should be at a

minimum. Most austenite chromium-nickel stainless steel weld metals are designed to

have small amounts of ferrite in order to avoid cracks and fissures, especially in

heavy weldments and with high welding energy input. Control of the amount of ferrite

may become critical in order to minimize the magnetic permeability and minimize the

hazard of weld metal cracks and fissures.

Many development projects have been conducted, and considerable documentation

has appeared, particularly in the past ten years, on the reliability of measurements

of ferrite in weld metals. The effect of compositional changes and the effect of

dilution are reasonably well known. The reliability of measurements under standard

welding conditions has been greatly improved. The effect of welding variables on

the reproducibility of the measurements has not been adequately studied. This area

was studied in a preliminary way and the results reported herein.

ARCOS PROGRAM

Five weld metal compositions were obtained using a single composition of core

wire and adjusting the ingredients in the electrode coverings to give the compositions

shown in Table 1. Electrodes were all of the AC-DC titania type, classified under

the AWS system as "E316L-16" except one with higher carbon would be classified "E316-16".

181
The location of compositions with respect to ferrite number is shown in Figure 1.

Actual measured ferrite under controlled conditions specified in the AWS-WRC procedure

(Ref. 1) appears in Table 1.

The ferrite variation in Type 316L is represented by electrode numbers 1 through

4. Believing that the low carbon grades are more sensitive to fissures than normal

carbon grades when the ferrite is low, one additional electrode. No. 5, was prepared with essentially no ferrite but with a higher carbon level. This test will

help to verify whether the low carbon is essential for fracture toughness.

As directed by the National Bureau of Standards, weldments are being prepared

in accordance with Figure 2, using welding parameters appearing in Table 5.

All electrodes use the welding parameters which are typical of arc

welding with this size (5/32") electrode. Electrode No. 3, having an intermediate

amount of ferrite, is used with parameters giving considerably less and considerably

greater weld energy input. Three of the electrodes, Nos. 1, 3, and 5, are also being

used with high energy input and a high preheat and interpass temperature. These ten

weldments are identified in the contract as Tasks A and C.

Before deciding on the conditions to be used in Tasks A and C, Task B was

carried out to assess the effect of welding variables on ferrite measurements. Weld

beads were deposited under conditions where compositional changes due to dilution would

not be a factor, all using the intermediate ferrite electrode No. 3. An outline of

the types of weld beads being examined is shown in Figure 3.

All testing of the large weldments is to be done by the National Bureau of

Standards, using fracture mechanics type specimens at cryogenic temperatures.

PURPOSE OF TASK B

One object of the task is to determine the reproducibility of the ferrite

readings, using the Magne-Gage testing procedure recommended by AWS-WRC. Whereas

much of the literature indicates wide variations, many of the reports involve multipass welds wherein several welding cycles are involved. All of the tests in

182
Task B were conducted on single-pass welds. Where discrepancies in the data

occur, an attempt was made to determine the cause by considering the factors

which may account for the scatter.

Another object was to consider the effect of wide differences in welding

parameters. In the 5/32" diameter welding electrode, it is possible to vary the

current from 100 amps to 160 amps and still have reasonable stable welding

conditions. It was also possible to vary the welding speed from nominally 3 inches

per minute (actually 3-1/2) up to 12 inches per minute (actually 11-1/2). This

caused the welding energy input to vary from a minimum of 11.7 kJ/in to 71.3 kJ/in.

The simplicity of bead-on-plate deposition allowed us to make a number of tests

with a limited budget. Such welds, however, would not be typical of that which would

be encountered in heavy weldments. Consequently, a few tests were made with beads

deposited in a pre-machined weld groove.

Cooling rate is known to be affected by the ambient temperature of the metal

surrounding the weld bead. By dramatically changing the preheat within limits which

might be encountered in ordinary welding (from room temperature to 600 F) , the effect

of ambient temperature on ferrite measurements was evaluated.

As the data were assembled, it was possible to test theories of the solidification

mechanism of stainless weld metal against results obtained. By this means, it was

hoped that a better understanding of the role of ferrite in welding might be obtained,

possibly to lead to design of welding filler metals and/or welding procedures which

will permit greater control of ferrite level

DATA SCATTER

The ferrite readings obtained in each of the various tests are given in Tables

2, 3, and 4. In the first series of tests five readings were taken along the

length of the weld. Subsequently, ten readings were used in order to obtain a more

183
precise average and standard deviation of the readings. When the averages and

standard deviations were plotted against welding energy (Figure 4), a trend

was indicated; however, a relatively wide scatter in the data was seen. Some of the

scatter appeared to be a consistent change in ferrite readings from the beginning to

the end of the weld bead, as shown in the average of the four to six figures in a

given location. Thus part of the data scatter would seem to be accounted for by progressive changes that occur during the welding operation. The WRC-AWS method calls

for a light mechanical dressing of the top of the weld surface before making the

readings. (Note: The standard procedure calls for the use of a file; we chose to

use a coarse grinding wheel.) It has been reported that this has the effect of

raising the average ferrite reading, and we confirmed that finding. One explanation

has been that the ferrite measuring magnet can more uniformly contact the surface of

the weld when its ripples are removed. If this were the case, one would expect more

uniform data from the ground welds than from the unground welds. The standard deviatioi

calculated on the ground specimens, however, generally scatter as widely as the ungrounc

specimens, raising doubt as to whether greater reliability is actually obtained by

this mechanical dressing of the weld.

In order to get meaningful data to illustrate progressive changes during the

welding operation, all of the readings within a given area were averaged. Thus

the data points appearing in Figure 5 include the significant differences arising

from the mechanical grinding of the second weld bead.

EFFECT OF WELDING PARAMETERS

It is evident (Figure 5) that ferrite readings increase as the current is

raised. In Figure 5 are plotted the data from Table 2, using all of the beadon-plate welds made without preheat. At constant current level, there is higher

ferrite reading with low speed than with high speed. This trend becomes more

noticeable with increasing current.

184
Realizing that welding energy is a major factor in cooling rate, the relationship set forth in pioneering work by D. Rosenthal (Ref . 2) on the theory of heat flow

with a moving heat source (in this case, the arc), the simplest form of relationship is

dT = J. s(T-To)^

dt 60 I V

where dT = cooling rate

dt

K = constant depending primarily on thermal conductivity

s = travel speed, inches per minute

T-To = temperature differential between solidifying metal

and adjacent parent plate

I = welding current

V = welding voltage

60 = conversion of speed from inches per minute to

Inches per second

Notet s 1 ^ ^

- — where J is the Joules per inch welding energy input

In figure 6 it is shown that, as welding progresses, the temperature of the

plate rises. An attempt was made to estimate the temperature (by Tempilstiks)

immediately after the end of the weld. It -was assumed that the temperature

increased uniformly from the original temperature (room temperature) to the final

temperature. The effect of the changing temperature could then be put into the

Rosenthal equation, and the points shown in Figure 6 illustrate the change in

average ferrlte readings from the initial reading to the reading at location 8.

The final readings were felt to be unreliable due to non-uniform conditions at

the end of the weld. Figure 6 shows that cooling rate has a major influence in

the ferrlte reading and supports the view that the increase in ferrlte reading is

caused by the change in ambient temperature as the weld is made.

185
Examination of the bead-on-plate welds made with a 600 F preheat showed

that the f errite-content decreased with increasing preheat temperatures in

contrast to what was shown in Figure 6.

Figure 7 shows the average of all readings under comparable welding conditions for three bead-on-plate welds and one groove weld with and without

preheat. The data in Table 3 and illustrated in Figure 8 indicate that the

increase in f errite-content in welds made with a 600 F preheat remains, as was

seen in the room temperature bead-on-plate tests. The amount of heat build-up

in the high energy tests (80 KJ/in.) is considerably greater than in the low

energy tests (20 KJ/in.). The unexpected results from the preheat test specimens suggest that the f errite-level increase along the length of the weld is

caused by more than one factor.

When the results of the groove welds are examined, little increase in

f errite-content along the weld is seen. The data from Table 4 are plotted in

Figures 9 and 10, wherein the dotted lines show a comparison with bead-onplate tests. These results indicate that the changes in f errite-level along

the weld that appeared in the bead-on-plate welds is caused more by chemical

reactions in the welding arc, than by temperature conditions which are independent of weld-configuration (i.e., groove versus bead-on-plate).

SOLIDIFICATION MECHANISM

The constitution diagrams for the chromium-nickel-iron system have been

well known for over forty years. Much of the welding literature, however,

fails to make use of these constitution diagrams. It is helpful, therefore,

to review the solidification mechanism in the light of the information that

is known about the chromium-nickel- iron system.

186
The first crystals to form from the melt are delta ferrite. The liquid

composition shifts until it reaches a temperature at which a eutectic mixture

of ferrite and austenite form. As the temperature decreases below the solidus,

ferrite transforms to austenite. In compositions where the primary delta ferrite

is sufficiently rich in f errite-stabilizers a network of delta ferrite remains;

as is illustrated schematically in Figure 11.

In Figure 12, an attempt has been made to represent a section of the

chromium-nickel-iron system with a constant iron-content of 62% and chromium and

nickel expressed as their "equivalents," as was done in Figure 1. The five

electrode compositions are located on this diagram to indicate how delta ferrite

may form. In the solidification of a weld from electrode No. 3 (Task B) , the

composition of the primary delta ferrite is illustrated, as well as the transformation of that primary delta-ferrite into austenite at temperatures below the

solidus. The eutectic delta-ferrite crystals are shown as a composition which

readily transforms to austenite. Thus the phase transformations which occur

during cooling of a single weld-pass (from liquid to room temperature) are:

Cooling

L 6 (primary)

L ^ 6 + Y (eutectic)

6 -> Y

Another schematic diagram (Figure 13) is an attempt to present a relationship between temperature and time during solidification and subsequent ferrite-toaustenite transformation. In this diagram, the right-hand dotted line shows a slow

cooling situation representative of that seen in the consitution diagrams. This

may be compared with the left-hand dotted lines which depict rapid cooling situations

representative of that seen during most welding processes. With a high energy input

resulting in slow cooling, a greater amount of primary delta ferrite forms than

when a low energy input resulting in a fast cooling rate occurs. As indicated

in Figure 12, the greater the amount of primary ferrite that forms, especially if

187
it is rich in f errite-stabilizing elements, the less likely the delta-ferrite

is to convert to austenite. This helps to explain the initial results in Figures

4,5, and 6. The high preheat resulting in a lower f errite content is explained

by the longer time that the metal remains in the f errite-to-austenite transformation region. The slight difference in cooling rate in the freezing area, due

to higher preheat, is overshadowed by the much longer times existing at the

temperature below freezing.

Consider the chemical reactions which may occur during welding and influence

the variation in f errite-level along the length of bead-on-plate welds. This

variation was not seen in groove welds. The following reactions were considered:

Effects on Ferrite Reading

0^ (atmosphere) + C ^ CO f Higher

(atmosphere) -> N2 weld metal Lower

CO (coatings) -> increases C Lower

CO (coatings) improves shielding ^ decreases Higher

The extent to which each of these reactions occurs is not known. At the high

welding energy input in the bead-on-plate welds, which have less shielding from

the atmosphere, the start of the weld might have more nitrogen contamination than

the end. Because of self-resistance heating of the electrode-core, a deeper arc

crucible is formed at the end of the electrode and the rate of flow of metal across

the arc is greater at the end of the weld than at the beginning. Both of these

effects lower the amount of nitrogen in the weld metal and increase the ferrite

level. When welding in a groove, convection currents in the weld-atmosphere do

not permit entrainment of as much atmospheric nitrogen. Hence, lesser variations

in f errite-level along the length of the groove-weld, relative to the bead-on-plate

welds, are expected.

188
and

The chemical

are subject to

reaction effects described above are offered as a

verification by accurate analysis for both carbon

hypotheses,

and nitrogen.

STRAIN HARDENING EFFECTS

Austenite will transform on plastic straining to martensite at room temperature. Weld beads cooling under severe restraint may experience enough plastic

deformation to increase the magnetic response of the f errite-measuring instruments.

The magnetic measurements taken on most of the groove welds were noticeably higher

than those on the less restrained bead-on-plate welds made under the same conditions

indicating a higher content of ferromagnetic phase (s) in the former series of welds.

The lower level of nitrogen, a strong austenite-stabilizer , in the groove welds

would explain this observation.

The expectation of a higher cooling rate, because of a larger fusion surface

and a greater volume of metal to conduct away the heat, is not borne out by the

results. The ferrite readings were higher in the groove welds than in the beadon plate welds, which is the opposite of what the cooling-rate considerations

would predict.

The remaining possibility is that there is enough plastic strain during

cooling below the austenite-to-mar tensite transformation temperature to cause a

noticeable increase in formation of strain-induced martensite and in magnetic

response. This appears to be the remaining explanation, as groove welds are

subject to greater restraint than bead-on-plate welds. Also, it should be noted

that magnetic readings measure both delta ferrite and strain- induced martensite.

Also a significant consideration when measuring the magnetic response of weld

surfaces which have been mechanically worked, is the possibility of strain-induced

martensite formation, as was observed in the higher readings on the ground surfaces.

189
SUMMARY

The data available in this study and the conclusions drawn from them are

tentative. The program was relatively small; more work would be necessary to

support quantitative conclusions. A summary of the observations follows:

Accuracy of Readings

a. Standard deviation for 5-11 readings generally lies between

.2 to 1.0 percentage of ferrite.

b. Ground surfaces generally read 0.5 - 1.0 percent higher than

as-welded surfaces.

c. The above generally confirms WRC and IIW round-robin experience.

Effect of Welding Heat Input

a. Higher current and lower travel rate consistently caused higher

ferrite readings.

b. Increasing energy input (hence decreasing cooling rate) gives

higher ferrite readings (presumably by allowing greater amounts

of 6 ferrite to form on cooling)

Effect of Preheat

a. A 600 F preheat consistently caused a drop in ferrite readings.

b. Preheat decreases cooling rate which presumably promotes the

formation of greater amounts of 6-ferrite in solidification,

but a longer time for transformation of 6-ferrite to austenite

in cooling from solidification to room temperature.

c. The increase in ferrite readings from start to finish of a single

weld bead cannot be accounted for by decreased cooling rates,

because the same trend is shown by the 600 F preheat tests whose

overall average shows a distinct drop in ferrite as compared with

room temperature tests. Latent energy released during solidification

may have to be considered in the cooling model. Compositional

changes, i.e., ingress of lesser amounts of nitrogen, an austenitestabilizer may explain the observed increase in ferrite content.

Effect of Weld Shape (Bead-on-Plate vs. Groove)

a. Groove welds had higher f errite-levels than bead-on-plate welds.

b. Cooling rate is expected to be higher than bead-on-plate welds. On

this basis, ferrite content results are contrary to expectation.

A possible explanation is greater shrinkage strains which cause austenite to transform to martensite, increasing the overall magneticresponse of the groove-welds.

190
c. Groove welds show little or no ferrite changes from start to finish

of weld.

d. Observations a. and c. are explainable by the lower level of a potent

austenite-stabilizer , nitrogen, anticipated in groove welds, as compared with bead-on-plate welds.

The Task B test program was designed to examine welds which underwent one heating cycle. Most heavy-section welding involves multi-pass welding, wherein

several welding cycles are observed by any one weld bead. The following phase

changes may be expected to occur in multi-pass welds:

Effect of Reheating Weld Metals (as in Multi-pass Welds)

a. Ferrite will reform close to the fusion line but probably will not

be retained on cooling within any normal weld-temperature sequence.

b. Ferrite will transform to austenite in a narrow band below fusion

line

c. Martensite probably redissolves on each reheating cycle only to

reform during cooling, due to shrinkage-induced strains.

REFERENCES

1. AWS Specification A5.4-7x, WRC recommended weld pad technique, to be published

by the American Welding Society, Miami, FL.

2. Rozenthal, D., Theory of moving sources of heat and its application to metal

treatment. Trans ASME, 68, 618-866 (1946).

191
Table 1: Composition of the Test Electrodes

Electrode WRC Ferrite No. Composition, Weight Percent

No. Diagram Measured C Mn SI S P Cr Ni Mo Cb N

1 2 0.5 .031 1.72 .58 .011 .033 18.47 13.97 2.25 .14 .042

2 4.5 3.5 .030 1.72 .56 .012 .034 18.64 12.97 2.29 .14 .046

3 10 9.0 .034 1.70 .57 .009 .034 19.20 12.01 2.32 .15 .042

4 12.5 12.0 .028 1.76 .59 .008 .029 19.66 11.06 2.18 .15 .067

5 3 0.6 .059 1.75 .56 .009 .033 18.69 12.85 2.25 .15 .044

192
Table 2: Ferrite Readines for Bead-on-Plate Welds - No Preheat

Ferrlte Readings Std.

Weld Ave. Dev.

160 amp - 12 Ipm y^j = 20.3

100 amp - 3 1pm kj = 41.1

1st pass 7.3 8.2 8.4 9.2 9.8 8.58 .86

2nd pass 7.8 7.6 7.6 7.8 7.8 7.72 .10

2nd pass, grd. 7.3 6.5 7.6 8.2 8.2 8.2 7.8 8.6 7.6 7.8 7.78 .56

Ave. 4 7.23 7.90 8.10 8.35 8.25 7.966 .674

160 amp - 3 1pm kJ = 7 1 .

3

1st pass 9.6 9.4 9.2 9.4 9.4 9.40 .13

2nd pass 7.8 9.6 9.6 10.6 8.4 9.2 .99

2nd pass, grd. 8.9 9.6 10.0 10.0 9.8 11.1 11.1 9.6 10.0' 9.6 9.97 .64

Ave. 4 8.98 9.75 9.93 10.18 9.35 9.638 .717

130 amp - 6 ipm kJ = 31.2

1st pass 8.9 8.6 9.2 8.9 9.4 9.00 .28

2nd pass 7.8 7.8 7.8 8.2 6.8 7.68 .47

2nd pass, grd. 8.2 7.8 8. 2 8.9 8.4 8.4 9.2 9. 6 9.2 8.0 8.59 .57

Ave. 4 8 .18 8 .38 8.45 8 .98 8. 35 8.503 1.09

100 amp - 12 ipm kJ = 11.7

1st pass 7.3 6.8 6. 8 7.6 7.3 8.0 8.4 6. 5 8.0 6.5 7.32 .64

2nd pass 8.4 6.5 6. 8 7.3 7.0 8.0 7.6 7. 0 7.6 7.8 7.40 .56

2nd pass, grd. 8.9 6.5 8. 0 8.4 8.4 8.9 9.6 7. 8 8.6 8.2 8.33 .78

Ave. 3 8.2 6.6 7. 2 7.7 7.6 8.3 8.5 7. 1 8.1 7.5

Ave. 6 7 .40 7 .48 7.93 7 .81 7 .78 7.683 .812

1st pass 8.4 8.,4 8.4 8.,4 9.,4 9..4 8.,9 7..8 9. 8 9. 6 8. 85 .63

2nd pass 7.8 8.,4 8.9 7.,8 9.,6 7.,0 9.,4 9.,6 10. 0 11. 6 9. 01 1.26

2nd pass, grd. 8.6 9.,2 8.9 9.,2 9,,4 9.,2 10..2 10.,0 10. 6 11. 8 9. 71 .91

Ave. 3 8.3 8.,7 8.7 8.,5 9,.5 8..5 9..5 9.,1 10. 1 11. 0

Ave. 6 8..46 1.60 9. 0 9,.32 10.57 9 .190 1.038

193
Table 3: Ferrite Readings for Bead-on-Plate Welds - 600 F Preheat

Ferrite Readings

Std

Weld Ave. Dev

100 amp - 12 ipm kj = 41.4

1 pass 5.7 5.5 5.5 6.5 5.5* 5.1* 6.8 5.7 6.8 5.3

1 pass, grd. 6.8 6.3 6.8 6.8 6.8 5.3* 7.3 8.0 7.8 6.5

Ave- 4 6.08 6.40 6.8* 6.95 6.60

130 amp 6 ipm kJ = 3 1 .

1 pass 6.3 7.8 7.3 8.0 7.0 7.8 7.8 7.8 7.0 9.6

1 pass, grd. 7.6 8.0 8.0 8.6 8.6 8.6 8.6 9.2 8.4 9.2

Ave. 4 7.43 7.98 8.00 8.35 8.55

160 amp - 3 ipm kJ = 7 1 .

1 pass 7.3 8.6 8.2 7.8 8.4 8.9 9.4 10.4 9.6 10.0

1 pass, grd. 6.5 8.9 8.6 8.4 9.8 9.6 9.6 10.9 10.9 10.6

Ave. 4 7.83 8.25 9.18 10.08 10.28

Erratic arc-outage — ignored in averages and standard deviation

5.98 .53

7.01 .55

6.53 .70

7.62 .46

8.48 .48

8.05 .89

8.86 .94

9.38 1.77

9.12 1.15

194
Table 4: Ferrite Readings for Groove Welds

Ferrite Readings

Weld Std.

Ave . Dev.

100 amp - 6 ipm - RT

Unground 8.4 7.6 8.9 8.4 8.4 8.2 8.2 8.9 8.4 8.4 3.35 .79

Ave. 2 8.0 8.6 8.3 8.5 8.4

(Compare Bead-on-Plate Welds—Ave. of 3 and 12 ipm 7.84 .61)

100 amp - 6 ipm - 600F

Unground 6.3 7.6 8.0 7.8 7.6 7.3 8.4 7.8 8.2 7.8 7.68 .55

Ave .2 7.0 7.9 7.4 8.1 8.0

130 amp - 6 ipm - RT

Unground 9.2 10.4 9.2 10.0 8.4 9.2 10.4 8.9 8.9 9.2 9.38 .63

Ave .2 9.8 9.6 8.8 9.1 9.1

(Compare Bead-on-Plate Welds - 130 amp - 6 ipm - RT)

1st pass unground 8.9 8.6 9.2 8.9 9.4 9.0 .28

130 amp - 6 ipm - 600 F

Unground 8.2 8.6 8.6 8.2 7.6 8.9 8.2 8.4 8.2 7.8 8.27 .36

Ave. 2 8.4 8.4 8.2 8.3 8.0

(Compare Bead-on-Plate Welds)

1st pass unground 7.0 7.6 7.4 7.8 8.3 7.62 .46

160 amp - 6 ipm - RT

Unground 9.6 9.4 10.0 9.4 9.8 9.8 10.0 9.2 10.0 9.2 9.64 .31

Ave. 2 9.5 9.7 9.8 9.6 9.6

(Compare Bead-on-Plate Welds—Ave. of 3 and 12 ipm 9.02 .33)

160 amp - 6 ipm - 600 F

Unground 8.0 8.0 7.8 8.4 7.6 7.8 8.4 7.0 8.6 6.8 7.84 .56

Ave. 2 8.0 8.1 7.7 7.7 7.7

195
Table 5: Welding Parameters for the NBS Weldments

I II III IV

Current, Amperes 130 100 160 160

Voltage, Volts 24 24 26 26

Speed, ipm 6-8 8-10 4-6 4-6

kJ/in 23-32 14-18 41-62 41-62

Interpass Temp., F 200-250 200-250 200-250 550-650

Table 6

Electrode 1 2 3 4 5

Task

A III I I A II

A III

C IV IV IV

196
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197
198
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Low Speed,

High Speed

Med. Speed

T

High Speed

Low Speed'

Arc

Start

160 amp

r 130 amp

100 amp

Bead-on-pl ate welds

No Preheat

3 4 5 6 7 8

Sequence of Ferrite Readings

Approximate Time

10

Arc

Finish

FIGURE 5. Ferrite content as a function of position in the test plate

201
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203
FIGURE 8. Ferrite content as a function of position in the test plate

for three heat input levels and 600 F preheat

204
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209
I
PHYSICAL PROPERTIES OF STAINLESS STEELS

211

LOW-TEMPERATURE ELASTIC PROPERTIES OF 304-TYPE STAINLESS STEELS

H. M. Ledbetter

Cryogenics Division

National Bureau of Standards

Elastic properties of AISI 304-type (304, 304L, 304N) stainless steels are being studied experimentally to determine:

1. elastic-constant changes between room-temperature and liquid-helium

temperature

.

2. elastic-constant variability due to anisotropy, product form, compositional fluctuations, etc.

Experimentally, longitudinal and transverse sound-wave velocities are measured in several directions in various samples. Four experimental methods

are used:

1. pulse-echo (4-in MHz)

2. pulse-echo superposition for temperature studies (4-10 MHz)

3. continuous-wave resonance with two transducers

4. three-component composite-oscillator resonance (near 60 kHz)

When combined with the mass density, p, these velocities, v, are related to

the engineering elastic constants by:

2 Shear modulus = G = pv^

2 Longitudinal modulus = C,^ = pv^^

4

Bulk modulus ~ ~ 3" ^

Young's modulus = pv^ = 9GB/(3B + G)

Poisson's ratio = (E/2G) - 1

where subscripts £, t, and e denote longitudinal, transverse, and extensional.

The inaccuracy in these methods is typically 1-2%, while the imprecision is

about 10 in the temperature-change measurements.

Elastic constants of a Type 304L stainless steel are given in Table 1 for

selected temperatures. Elastic-constant/temperature curves are shown in

Figs. 1-3. These constants were measured at near-zero stress and represent

the true dynamic elastic constants. For iron at room temperature, dynamic

elastic constants exceed static values by about 2% for B, 1% for v, 0.3%

for E, and 0% for G. At zero temperature all these differences vanish, and

they remain negligible below about 100 K.

Detailed behavior shown in Figs. 1-3 has not been reported previously.

All the elastic constants are regularly behaved between room temperature and

about 75 K. G shows the largest change with temperature, and B the smallest.

213
The reversible anomaly near 45 K is due to a Neel (paramagnetic-antiferromagnetic)

transition. All elastic constants are anomalous except perhaps for B where

the data are scattered. E and G show nearly parallel behavior. At about 16 K,

both and v data indicate that a second reversible transition may occur in

these alloys. This observation is now under study in other specimens.

Besides anomalous temperature behavior, several other observations were

made on AISI 304-type stainless steels: (1) elastic constants show very little

anisotropy in the bar and plate stock studied so far; (2) no evidence was found

for dispersion, sound-velocity variation with frequency, between 2 and 10 MHz;

(3) attenuation has not been measured, but it clearly varies strongly from

sample to sample, due mainly perhaps to grain-size variations.

Variability effects are still being studied in these alloys, and these

results will be reported later.

214
Table 1. Elastic constants of a 1 y Uc o Ld 1 ri 1 p s r> s r.HP i .

T(K) G

do"

B N/m^)E V

(dimensionless)

295 2.526 0.763 1 .509 1 .958 0.284

200 2.587 0.795 1 .527 2 .031 0.278

100 2.629 0.822 1 .533 2 .093 0.273

76 2.632 0.826 1 .530 2 .100 0.271

4 2.611 0.809 1 .532 2 .065 0.275

215
1.05

STAINLESS STEEL 304

T (K)

Figure 1. Temperature dependence of the longitudinal modulus.

216

Stainless Steel 304

TEMPERATURE (K)

Figure 3. Temperature dependence of Young's modulus, the bulk modulus,

and Poisson's ratio.
NONMETALLICS

219

NONMETALLICS

M. B. Kasen

Cryogenics Division

National Bureau of Standards

Nonmetallics research was not funded during FY 77; however, steps were

taken to prepare for work to be undertaken during FY 78. The primary objective

is to provide fabricators of superconducting magnets with a dependable, highpressure industrial laminate system (or systems) for use as electrical or thermal insulation. A major problem in selecting material for such use is the

highly proprietary nature of existing NEMA-grade laminates. These materials

are presently produced to specified room temperature electrical performance

criteria and are not required to meet specific material specifications. Thus,

laminates having the same NEMA designation but produced by different laminators

have been found to have different mechanical and electrical performance at

cryogenic temperatures.

This problem was discussed with the Industrial Laminate Subcommittee of

NEMA, with representation by all U.S. laminating firms. The result was a suggestion that NBS work with a single laminating firm to develop specifications

for laminate systems that meet MFE needs. The selected laminator would act on behalf of the NEMA organization with product formulations and production methods

being made available to all U.S. laminators for competitive production.

The Micarta Division of the Westinghouse Corporation has agreed to work

with NBS in this effort. Following discussions with various knowledgeable people

in the industry, Westinghouse and NBS have agreed that the following series of

laminates would be more appropriate for the initial characterization study:

1. NEMA G-lO/MIL-P-18177 Type GEE

Resin technology to be a published commercial formulation. Fiberglass

reinforcement to be Style 7628 fabric with a silane finish. The m.anu- facturing process will be non-proprietary and fully disclosable.

2. NEMA G-ll/MIL-P-18177 Type GEB

As above but a commercial G-11 formulation using an aromatic amine cure agent.

3. NEMA G-ll/MIL-P-18177 Type GEB with Boron-Free "E" Glass

As (2) above but using boron-free glass reinforcement to evaluate

potential improvement in radiation resistance.

Westinghouse will produce the laminates in their Durham, NC plant.

A sufficient variety of panel thicknesses and sizes will be obtained to

permit a thorough characterization of the mechanical, electrical and thermal

properties of each product at 4 K. Additionally, sufficient material will be

available to facilitate studies of the radiation resistance of the various

221
products at 4 K. The mechanical, electrical and thermal studies will be

undertaken by NBS and by subcontract to industry. It is anticipated that

the radiation studies will be undertaken by a National Laboratory having the

required facilities and will be incorporated as part of the ongoing work of

such a laboratory.

As an independent contribution to the nonmetallics program, the Spaulding

Fiber Company of Tonawanda, NY has volunteered to provide a series of publishedformulation laminates comparable to (1) and (2) above for characterization along

with the Westinghouse material. It is understood that the Spaulding laminates

will be representative of current production material.

As laminate systems (1) and (2) proposed by Westinghouse and/or the com- parable materials from Spaulding Fiber are credible commercial G-10 and G-11

formulations, fabricators of superconducting magnets who have need to purchase

such laminates for immediate use have the option of using these products rather

than proprietary, unspecified formulations. It is hoped that fabricators will

take advantage of this opportunity, as practical experience with the laminates

will contribute to the data base and enhance the credibility of the overall

program.

The Westinghouse laminate system (3) remains experimental and should

not be used for magnet fabrication until the characterization work is completed.

222
HANDBOOK PREPARATION

223

HANDBOOK PREPARATION

F. R. Fickett

Cryogenics Division

National Bureau of Standards

This project started later in the year than the others reported

here and, thus, it is still in the formative stages. The handbook is

being produced jointly by the Mechanical Properties Data Center (MPDC)

of Traverse City, MI and the staff of the NBS Cryogenics Division. The

literature searching and document acquisition is done by NBS. MPDC

extracts the data and prepares worksheet graphs and characterization

tables which are then evaluated by the NBS staff to determine "best

value" curves and appropriate error bands. The handbook will present

"best value" curves for properties in the two categories listed in

Table 1. For the category A properties we will present an error band

which will indicate the spread that one might expect due to material

variability, experimental error, etc. A typical graph is shown in Figure

1. The published graphs will have SI units primary, but common units

will be shown also. Each A category property presentation will also

contain a characterization table giving the appropriate information

(different for physical and mechanical properties) on each of the

original data sources used in constructing the best value curve. The

information to be contained in the table is shown in Table 2. Another

table will be prepared for the "A" properties providing the data from

the best value curve (Table 3). The B category properties will be

presented in smaller graphs as shown in Figure 2. Although not shown on the figure, error and variability estimates will also be shown where

possible. For all properties a complete bibliography will be included.

The materials to be included in the first edition of the handbook are

listed in Table 4. Formats, indexing, covers, etc. have all been

discussed, but no final form has been decided on yet.

225
Table 1. Properties to be presented in the handbook.

A B

Coefficient of Thermal Expansion Bulk Modulus

Crack Growth Rate Compression

Fatigue Strength Creep

Fracture Toughness Electrical Resistivity

Tensile Strength Impact Strength

Tensile Yield Strength Magnetic Permeability

Thermal Conductivity Poisson's Ratio

Young's Modulus Shear Modulus

Specific Heat Capacity

Stress-Strain

Tensile Elongation

Tensile Reduction in Area

Thermal Diffusivity

226
Table 2. Sample characterization table.

I

Author (year)

Grain size

Hardness

Form

Condition

Chemistry

RRR

Test variables

spec, type

loading rate

temperatures of tests

number of spec, per temp.

227
Table 3. Unit conversion table for Figure 1.

' Category structural alloys Class STAINLESS STEELS Material aisi 316

ULTIMATE TEIISILE STRENGTH

Tabular Presentation of Data from Graph

T, K (T,°F) a„, MPa u

(a„, Ksi) ^ u T,"F (T, K) CT„ , Ksi U,, MPa)

4.2 (-452.11431 (207.5) -452 ( 4 .2) 207.5 (1431)

10 (-441 .7) 1409 (204.3) -440 (To. 93) 203.8 (1405)

20 (-423.7) 1391 (201 .7) -420 (22. 04) 201 .2 (1388)

30 (-405.7) 1362 (197.6) -400 (33. 15) 196.2 (1353)

40 (-387.7) 1348 (195.5) -380 (44. 26) 195.0 (1344)

50 (-369.7) 1327 (192.4) -360 (55. 37) 190.0 (1310)

60 (-351 .7) 1298 (188.2) -340 (66. 48) 185.8 (1281)

70 (-333.7) 1266 (183.6) -320 (77. 59) 178.8 (1232)

80 (-315.7) 1227 (177.9) -300 (88. 71

)

174.5 (1203)

90 (-297.7) 1200 (174.0) -280 (99. 82) 169.8 (1170)

100 (-279.7) 1170 (169.7) -260 (110 .9) 165.0 (1138)

no (-261 .7) 1140 (165.4) -240 (122 .1160.1 (1105)

120 (-243.7) 1110 (161 .0) -220 (133 • 2) 155.0 (1069)

130 (-225.7) 1079 (156.5) -200 (144 .3) 150.0 (1034)

140 (-207.7) 1047 (151 .9) -180 (155 .4) 145.0 (999.7)

150 (-189.7) 1016 (147.4) -160 (166 .5) 139.5 (961 .8)

160 (-171 .7) 983.9 (142.7) -140 (177 .6) 134.8 (929.1)

170 (-153.7) 951 .5 (138.0) -120 (188 .7) 128.8 (887.7)

180 (-135.7) 920.5 (133.5) -100 (199 .8) 124.5 (858.4)

190 (-117.7) 884.6 (128.3) -80 (210 .9) 118.8 (818.8)

200 (-99.67) 857.7 (124.4) -60 (222 .1

112.5 (775.7)

210 (-81 .67) 822.5 (119.3) -40 (233 .2) 107.5 (741 .2)

220 (-63.67) 783.9 (113.7) -20 (244 .3) 102.5 (706.7)

230 (-45.67) 750.8 (108.9) 0 (255 .4) 96.25 (663.6)

240 (-27.67) 719.8 (104.4) 20 (266 .5) 90.75 (625.7)

250 (-9.670) 684.6 (99.29) 40 (277 .6) 85.00 (586.1

260 ( 8.330) 647.9 (93.97) 60 (288 • 7) 79.75 (549.9)

270 ( 26.33) 613.2 (88.94) 80 (299 .8) 75.00 (517.1)

280 ( 44.33) 578.2 (83.86)

290 ( 62.33) 546.0 (79.19)

300 ( 80.33)

228
Table 4. Materials to be included in the handbook.

(FY 1977, 1978)

Stainless Steels: 316, 304, 21/6/9

(and L and N modifications)

Aluminum Alloys: 5083, 2219, 6061

229
Category structural alloys Class STAINLESS STEELS Material aisi 316

ULTIMATE TENSILE STRENGTH

0 100 200 300

TEMPERATURE,

Figure 1. Tensile strenath of AISI 316 stainless steel.

230
Figure 2. Presentation of type B properties.

Category structural alloys Class STAINLESS STEELS Material aisi 316 ln

THERMAL EXPANSION

ELECTRICAL RESISTI\'ITY

MAGNETIC PROPERTIES

-0.3

Form:

Diameter, cm (in.):

Condition:

40 80 120 160 200

TEMPERATURE, K

240 280

100

< K 2

o —

liJ CO J

Hi OC ^

Rods (Low Nitrogen)

0.4 (0.157)

10

10 100

TEMPERATURE, K

Annealed (heated at 1348 K (1075 C) for 0.5 hour and

water quenched)

300

Testing Temperature, K (F) RT 77 (-320) 4.2 (-452)

Magnetic Properties

Susceptibility, k 25.77 X

10-"^ 69.14 X 10"^ 113.38 X 10"^

Permeability, /i (H = 3 kOe) 12.60 X 10-^ 12.66 X

10-"^ 12.71 X 10"^

Permeability, m (H = 80 kOe) 12.683 X

10-"^

231

VAIL WORKSHOP

1

233

VAIL WORKSHOP

F. R. Fickett

Cryogenics Division

National Bureau of Standards

A major aspect of the technology transfer part of the program is the

workshop. These meetings are to be held each year with the goal of bringing

together representatives from all parts of the materials community which are involved with low temperature applications. The attendees at this year's

workshop, held at Vail, CO, October 12-14, represented just such a diverse

group including: industry, national labs, MFE projects and universities.

The following pages give the attendance list, the program for the workshop

and a review of the proceedings, which was originally prosented at the 7th

Symposium on Engineering Problems of Fusion Research in Knoxville, TN,

October 25-28.

235
ATTENDANCE AT NBS-ERDA WORKSHOP

Vail , Colorado

J. Benzinger - Spaulding Fiber Co.

P. Bezler - BNL

E. Bobrov - MIT

R. Bourque - General Atomic

R. Braden - Belfour Stulen, Inc.

J. Bronson - LASL

B. Brown - ANL

J. Bywater - ANL

F. Catania - ANL

J. Christian - General Dynamics

T. Collings - Battelle Columbus

E. Dalder - ERDA

R. Espy - Armco

F. Fickett - NBS

R. Foley - LLL

S. Cause - Westinghouse

L. Genens - ANL

R. Gold - Westinghouse

E. Hodge - Battelle Columbus

J. Hust - NBS

H. Johnson - Princeton Plasma Physics Lab.

M. Kasen - NBS

R. Kelsey - Alcoa

V. Kopytoff - LLL

P. Landon - LLL

D. Larbalestier - U. of Wisconsin

H. Ledbetter - NBS

C. Long - ORNL

F. Mazandarany - General Electric

H. McHenry - NBS

H. Mishler - Battelle Columbus

T. Moore - Belfour Stulen, Inc.

J. Morris - Lawrence Berkeley Lab.

R. Nelson - LLL

C. Oberly - AFAPL Wright-Patterson AFB

W. Pelczarski - ANL

D. Read - NBS

R. Reed - NBS

W. Roden - General Dynamics

R, Sampson - Westinghouse

E. Sandersan - Rockwell Int.

P. Sanger - Airco, Inc.

F. Schwartzberg - Martin Marietta

J. Scott - ORNL

J. Shepic - Martin Marietta

T. Siewert - Chemetron

F. Szumachowski - Teledyne-McKay

R. Thomas - Arcos

R. Vandervoort - LLL

W. Vogen - Berkeley

J. Wells - Westinghouse

C. Witherell - U. of California

L. Young - Westinghouse

236
PROGRAM FOR NBS-ERDA WORKSHOP ON MATERIALS AT LOW TEMPERATURES

Vail, Colorado October 12-14, 1977

Wedn esday, 10-12-77

I. Introduction (8:30-9:30)

Welcome

Presentation of materials survey

resul ts

Discussion of data handbook

R. P. Reed

E. N. C. Dalder

R. P. Reed

R. Braden/T. Moore

II. Current Designs and Materials (9:30-12:30)

Large Coil Project C.

TNS

TFTR

MFTF

u . F.L.

R.

H.V.

J. Long

Christian

Mazandarany

Young

Bourque

Johnson

Kopytoff

III. Applications of Nonmetal 1 ics in Magnet Structures (2:00-5:00)

Nonmetal 1 ics at low temperatures

Nonmetal lie cryostats

Industrial laminates

Radiation damage in nonmetal lies

Accelerated aging under stress

M. B. Kasen

W. Vogen

S. Gause

C. J. Long

R, Sampson

NBS

ERDA

NBS

Bel four-Stulen

ORNL

General Dynamics

General Electric

Westinghouse

General Atomic

PPPL

LLL

NBS

U. CA, Berkeley

Westinghouse

ORNL

Westinghouse

Thursday, 10-13-77

IV. High Strength Stainless Steels for Low Temperature Service (8:30-12:00)

Studies of nitrogen strengthened

stainless steels D. Larbalestier

D

R

J

T. Read

P. Landon

Shepic

U. Wisconsin

NBS

LLL

Martin Marietta

GROUP PHOTOGRAPH (12:00)

Thick Section Welding for Low Temperature Applications (2:00-5:00)

Stainless steel welding

Aluminum welding

H.

R.

R.T.

R.

I. McHenry

H. Espy

D. Thomas,

Seiwert

Kel sey

Jr.

NBS

Armco

Arcos

Chemetron

Alcoa

Friday, 10-14-77

Physical Properties of Magnet Materials (8 :30-12:00)

Elastic Properties H. M. Ledbetter NBS

Thermal Properties 0. G. Hust NBS

Radiation Damage B. S . Brown ANL

Magnetic Properties E. W. Collings Battelle

Electrical Properties F. R. Fickett NBS

237
A RCVIEU OF THE "IBS-ERDA WORKSHOP OM

IIATERIALS AT LOU TEMPERATURES*

F.R. Fickett and R.P. Reed Cryogenics Division

National Bureau of Standards

Boulder, CO 80302

Summary

A workshop treating oroblems related to low temo- erature materials applications in superconducting

magnet systems was held at Vail, Colorado October 12-14.

The major topics considered by the participants were:

the recently completed NBS survey of low temperature

materials needs for MFE applications and the production of an appropriate data handbook; materials

choices of recent device designs; nonmetallic materials

applications; high strenath stainless steels for cryogenic service; thici- section welding; and physical

properties of magnet materials. A summary of the presentations and the ensuing discussions is given

along with our assessment of the more pressing of the many low temperature materials research problems considered by the workshop participants.

Introduction

The v;orkshop program on materials at low temperatures is designed to bring together all of the diverse

elements of the materials community for discussions covering problems related to the construction and operation of large superconducting magnet systems for fusion energy. There are a number of nevy nroblems

with fusion systems which have not been encountered

in a serious way in orevious superconducting magnets:

the magnets are very large and must ooerate at stress

levels near the safe limits for many potential structural materials; extensive welding and bolting is essential; significant neutron fluences are found throughout the magnet structure, a situation which

creates serious nroblems for nonmetallic insulators

and low resistivity stabilizing metals; pulsed, high

magnetic fields with lelatively fast risp times are likely to exist. Ono of the goals of the workshop

was to discuss implications of each of these aspects

for materials selection and device desiqn. The only materials specifically excluded from cons i doiM f i on were the superconductors (MbTi and Nb,Sn) themselves,

as we felt they had been adeguately troatcd in a large number of recent conferences.

A major purnose for the workshop was to provide

input and direction to the ongoing NBS-EPDA program

on low temperature materials research for fusion energy applications. This program is now enterinn its second

year and many of the results presented at the workshop

were obtained by the program contractors during the first year of the project.

The fifty-three attendees represented a wide

range of interests: magnet designers and fabricators;

industrial laminate manufacturers and application

engineers; stainless steel and aluminum producers;

welding specialists; material scientists; and data handbook specialists.

^This work was carried out at the National Bureau of Standards under the sponsorship of the Energy Research

and Development Administration.

The six sessions and the topics dicussed (in a total of 28 Papers ) were

I. Survey Results and Handbook Design Materials survey results Discussion of data handbook

II. Current Designs and Materials

Large Coil Program (four papers) TNS

TFTR MFTF

III. Applications of Nonmetallics in Magnet Structures Nonmetallics at low temperatures

Nonmetallic cryostats

Industrial laminates Radiation damage in nonmetallics Accelerated aging under stress

IV. High Strength Stainless Steels for Low Temperature Service Nitroqen stremthened stainless steels

(three papers) Fracture and fatigue of stainless steel

Thick Section Weldinn for Low Temperature

Annl ications Evaluation of weld metal at cryogenic temperatures

Welding nitrogen strengthened stainless

steels Forritr in .uistcnitic stainless steel

wpl ds Impact properties of stainless steel wo 1 d s Aluminui'i alloy nlato for low temperature

appl i ca t i nns

VI. Physical Pronerties of Mannet Materials

Elastic Properties

Thermal Properties Radiation Damage Magnetic Properties

Electrical Properties

In the followinn sections we present a brief summary

of the presentations and some of the conclusions and reccommendations reached during the workshop. Further information regarding specifics can be obtained from

the authors of this paper. Mo oroceedinqs will bp

publ ished

Survey Resu lts a nd Handbook Design

1 The NBS survey considered the detailed materials

choices made in all then-current MFE device designs for materials to be employed at low temperatures. The existing data base was investigated, and areas where more information was needed were determined. The most

critical of these needs were incorporated in a proposal

238
for a five-year research Droqram. Particular attention

was paid to alternative materials and areas where the present materials choices may be questionable due to uncertainties in magnetic field size and oulse time. The results of the survey have been summarized else- where^ in detail, but briefly, for the first year of the research program we chose to- concentrate on the low temperature properties of structural materials for very large load-carrying welded or bolted structures, particularly those using the stainless steels. In the present year of the program, attention is also being

paid to the nonmetallic components.

specifications are hinhlv proprietary formulations at present. At the workshop we were able to reach agree- ment with representitives of the industry to provide

three laminates meeting GIO and Gil sneci f ications with published (i.e. nonoronrietary ) formulations.

The two Gil's will be amine cured and one will contain boron-free E-glass reinforcement. Both of these should have relativel

y

good radiation resistance. The

low temperature properties of these laminates, from industrial product'jn, will be extensively tested in the fIBS program and, hopefully, their radiation resistance can also be determined.

The handbook will contain graphical displays ot best value curves for the properties listed in Table 1. usually as a function of temperature.

Table 1. Properties Covered by Data Handbook

A B

Coefficient of Thermal Expansion

Crack Growth Rate Fatigue Strength

Fracture Toughness Tensile Strength Tensile Yield Strength

Thermal Conductivity

Young's Modulus

Bulk Modulus

Compression

Creep

Electrical Resistivity

Impact Strength

Magnetic Permeability

Poisson's Ratio Shear Modulus Specific Heat Capacity Stress-Strain

Tensile Elongation

Tensile Reduction in Area

Thermal Diffusivitv

Category A properties are considered of primary importance. Those data will also be given in tabular form

and the display will be accompanied by a characterization table describing the original data. Hnly the graphical display will he used for Category B. In both categories an attemnt will be made to show the extent

to which the general curve displayed represents the actual data. Materials chosen for inclusion in the first two years are 316, 304 and 21/6/'^ striinless steels, and 5083, 2219 and 6061 aluminum allocs.

Current Desi g ns and Materials

These devices have been described in some detail at this and previous conferences. The LCP presentations indicated that a wide range of structural materials are being used and that some mechanical properties

data is still needed on the chosen materials, i.-hich tend

to be modifications of the conventional austenitic

stainless steels and one aluminum alloy. The TFTR presentation was particularly interesting in that, while not a cryogenic system, it is one that is actually

being built and the problems and solutions occur much more rapidly.

Applications of Nonmetallics in Magnet Structures

The presentations indicated conclusively that large nonmetallic dewars can be produced. However, strict attention must be paid to problems of helium permeability through certain materials and the possibility of severe creep near room temperature. The use

of industrial laminates as load-bearing insulation and thermal standoffs is an integral part of all magnet

designs. Figure 1, from M. B. Kasen's presentation, com- pares several composites with welded stainless steel in terms of the thermal conductivity-to-strength ratio.

All of the NEMA grade laminates meeting the GIO or Gil

10 100 TEMPERATURE. K

Fiq. 1. 'Thermal conductivity to strength ratio for several uniaxial composites compared to welded stainless steel. From M. B. Kasen

NBS.

It seems clear that the radiation resistance of organic materials of all types must be determined

soon. The very sketchy available data indicate that these materials may be significantly degraded by the flux levels presently nronosed as operating conditions.

Most available work was not done below room temperature and has looked at the mechanical properties nearly

exclusively. Dearadation of the electrical properties

mav well occur at lower fluences. The fact that these materials are at a temperature of 4 K is a much more

serious change of conditions for oraanics than

for metals. The only existing research program is very

small when one considers the myriad of questions to be resolved. We do hope, however, that this program will

be able to determine the seriousness of the situation

in the near future. Perhaos then it will be possible

to Pursue a larger scale effort to determine the limits

of existing materials and to start the development of organic materials more resistant to radiation for future appl ications

.

High Strength Steels for Low Temperature Service

''li troqen-strengthened AISI 304N fracture toughness

239
and tensile tests at 4 K indicate that the yield

strength is doubled from its room temperature valve

and that the toughness remains high. Use of the nitrogen-strengthened austenitic Cr-Ni steels is very promising, as most current projects do not contemplate

using extremely strong alloys, such as the Nitronic

series, and the development of intermediate strength,

tough austenitic weld filler rods is closer to fruition.

Figure 2, from the presentation by D. T. Read and

R. P. Reed shows the interesting relationship between strength and toughness at 4 K for a number of stainless

steels. The lines represent a simple calculation of the maximum section size to permit plastic deformation

prior to unstable crack growth. The higher the fracture toughness the larger the thickness can be and

still achieve elastic-plastic fracture characteristics.

YIELD STRENGTH, a;, (MPa)

Fig. 2. 4 K fracture toughness versus yield strength

for some nitrogen-strengthened and conventional austenitic stainless steels. From

D. T. Read and R. P. Reed - MBS.

The next 4 K strength and toughness measurements

on this material class will be on AISI 316, in partic- ular AISI 316N and AISI 316LN will hp PxaminpH. AISI 310S with high nitrogen may be tested also if it becomes

available. Those in attendance felt that this program

direction, providing preliminary strength data on nitrogen-strengthened Fe-Cr-Ni austenitic steels, could

lead to a new higher-strength, tough alloy class for

use in superconducting magnet applications.

Thick Section Welding for Low Temperature Applications

Austenitic stainless steels such as AISI types

304 and 316 and the nitrogen strengthened 2lCr-6Ni-9f1n alloy are readily weldable by all common welding pro- cesses, providing the appropriate consumables and procedures are used. The service experience with welded assemblies of these alloys has been satisfactory

at cryogenic temperatures. A contributing factor to Ihis success has been low design stresses; generally

the stresses have been limited to about 140 MPa (20 ksi),

the ASME allowable stress for 304 and 316. For large

superconducting magnets, the enormous forces that must

be contained and the limited space available for structure may necessitate operating at significantly

higher stresses, stresses based on the strength of the

alloy at the operating temperature. Many of the un- welded austenitic stainless steels exhibit excellent

fracture toughness at temperatures down to 4 K in the base metal, as shown above. However, weld metal frac- ture toughness has not yet been determined at 4 K. Limited Charpy impact results at 76 and 20 K suggest

that the toughness may be marqinal for these materials

at 4 K, particularly if the stress levels for the magnets become higher than current practice allows.

The proDosed explanations for this behavior have included weld rod composition, ferrite content and microstructural variations. Welds with varying ferrite content have been prepared from 316 coated electrodes and are now being investigated metallurgically. Some preliminary toughness values of weld materials investigated for this program, and reported

bv T. A. Seiwert and R. H. Espy, are given in Table 2.

Table 2. CVN Toughness of Welds

Material -Toughness in joules (ft. lbs.) at (nominal composition) Room Temperature 76 K

308 58 (43) 35 (26) 310 42 (31) 330 90 (66) 13Cr 20Ni QMn + Ch 6? (46) 38 (28) 13Cr 20 Ni I'-'n * Cb 37 (27) 16 (12) 20Cr lONi fi'ln 64 (47) 12 ( 9) 20Cr lOMi grin 67 (49) 12 ( 9) 18Cr IfiMi gMn 72 (53) 37 (27) 18Cr 16Mi q'ln 60 (44) 18 (13)

No fracture toughness data at " K are yet available, but measurements are now being made.

Physical Properties of Magnet Materials

It was shown that measurenient of the dynamic

moduli of structural alloys provides an easy and inexpensive means of determining the behavior of the elastic oroperties over a wide ranqe of temperature.

Some results are shown in Figure 3, from the presentation of H. M. Ledbetter. Determination by static

m

TEMPERATURE (K)

Fig. 3. Young's modulus, the bulk modulus, and Poisson's ratio for 304 stainless steel. From H. M. Ledbetter - NBS.

240
measurement at these many points would be prohibitively

expensive. Agreement with static data is ciood at all temperatures where comparisons have been made.

The answer to the question "what is the thermal conductivity of copper at low temperatures?" is shown

in Figure 4, from J. G. Hust's presentation. Obviously

it depends on the specific material at hand. The parameter necessary for determining the conductivities

is the residual resistance ratio (RRR). The purity

(in "nines" or similar designations) cannot be used!

4 6 10 20 40 60 100 200 300 TEMPERATURE K

Fig. 4. The thermal conductivity of copper as a function of Residual Resistance Ratio [R(295K)/

R(4K)]. From J. G. Must - NBS.

can cause significant temperature excursions. The problem in superi nsul ations has been solved by etching

fine strips of the aluminum coating away so as to pro- duce a grid pattern.

References

1. R. P. Reed, F. R. Fickett, M. B. Kasen, "Magnetic

Fusion Energy Low Temperature Materials Program:

A Survey," Report to ERDA/DMFE from NBS Cryogenic

Division (March 1977).

2. F. R. Fickett, M. B. Kasen, H. I. McHenry and

R. P. Reed, "A Low Temperature Materials Research

Program For ^lagnetic Fusion Energy" paper BA 1, International cryogenic Materials Conference,

Boulder, Co, August 1977. To be published in: Advances in Cryogenic Engineering, Vol. 24, Plenum Press, New York (1978).

The effects of neutron damage on the electrical resistivity of the pure stabilizer metals, copper and aluminum, are reasonably well understood. Some questions remain regarding the effect on the thermal con- ductivity, the magnetoresistance and the implications

of the metals being under stress during irradiation at low temperatures. We feel that the usual prediction

schemes, such as the Wiedemann-Franz law relating electrical and thermal resistivity and modifications of Kohler's rule which predicts electrical resistance in

a magnetic field, will probably be valid for the case

of radiation induced defects. However, these assumotions are of such great importance for the overall magnet design that some attempt should be made to verify

them by experiment.

The discussion of the magnetic properties of the stainless steels indicated a complex behavior for these

alloys: transitions from room temperature paramaqnetism to a low temperature antiferromagnetism occurs

in some alloys; other alloys show characteristics of superparamagneti sm at low temperatures. Ferromagnetism

is associated with the stress induced martensitic trans- formation and with the presence of the ferrite phase

in welds. For the most part, none of these effects appear large enough to have serious consequence for the structural elements. Some questions were raised reqarding the source of the often stated requirement that the relative permeability of the structual steels not exceed

1.02, but no answers were nrovided.

The eddy current losses in both structural elements

and pure metals found in the conductor composite and

as coating for superinsulation were discussed. The

losses can be significant and, coupled with the very low specific heat values at cryogenic temperatures, they

241
I
SURVEY OF LOW TEMPERATURE MATERIALS

FOR MAGNETIC FUSION ENERGY

243

Magnetic Fusion Energy

Low Temperature Materials Program

A SURVEY

March 1, 1977

BY

R. P. Reed, F. R. Fickett, M. B. Kasen, H. I. McHenry

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder / Colorado 80302

TO

Division of Magnetic Fusion Energy

Energy Research & Development Administration

Washington, D.C. 20545

245
TABLE OF CONTENTS

Page

I. INTRODUCTION 1

II. MFE MAGNET SYSTEMS 3

A. Operating Conditions 6

1 . Thermal 6

2. Mechanical 7

3. Electrical 9

4. Magnetic 11 5. Radiation 11 B. Design Considerations 14

1. Codes and Specifications 14

2. Fabrication 15

3. Inspection 16

C. Materials and Components 18

1 . Material Overview 18

2. Component Research Summary 23

III. MATERIALS TECHNOLOGY 27

A. Structural Alloys 28

1. Austenitic Stainless Steel 28

2. Aluminum Alloys 36

3. Alternative Alloys 40

B. Structural Composites and Bulk Insulators 41

1 . Overview 41

2. Experience 41

3. Properties 43

4. Specifications, Availability and Cost 43

5. Fabrication 44

6. Alternative Materials 44

C. Insulating Films, Coatings and Inorganics 46

1 . Overview 46

2. Experience 46

3. Properties 46

D. Conductors 48

1 . Overview 48

2. Superconductor 48

3. Stabilizer 49

4. Strengthener 51 5. Other Materials 51

6. Conductor Composite 51

246
Page

IV. RECOMMENDATIONS 53

A. Program Objective ...... 54

B. MFE Research Requirements ....... 54

C. Summary of Materials Considerations ......... 56

1 . Structural Al loys . . 56

2. Thermal Insulators . . . 56

3. Insulating Films, Coatings and Inorganics .... 56

4. Structural Composites and Bulk Insulators .... 56

5. Conductors 5g

6. Adhesives ' 57

7. Demountable Coils * 57

D. Specific Materials 58

E. Program ........ 60

V. REFERENCES 75

Appendix A: Summary of Project Materials and Components . . A-1

Appendix B: Project Information Sheets B-1

Appendix C: Research Capabilities ! !

' ' C-1

247
I. INTRODUCTION

The present knowledge concerning magnetic-fusion energy (MFE)

devices indicates that economical reactors will require superconducting magnets operating near the temperature of liquid helium. In this

temperature region, many materials exhibit properties significantly

different from those at room temperature. Sometimes these altered

properties work to the advantage of the magnet system designer and

sometimes not. In any event, the very high initial cost of the

magnets and the difficulty of repair makes it essential that the

behavior of magnet construction materials be well understood at the

operating temperature and under operating conditions. Our objective

is to define the materials problems and to organize a plan of effective materials research and technology transfer that will provide the

materials information necessary to insure economical construction and

reliable, safe operation of the low temperature components of MFE

systems. Both short-term and long-term needs are addressed. The

materials considered are structural, thermal and electrical insulation

and normal-metal conductors; superconductors while of obvious importance, are not assessed in this survey.

The survey first considers ^fE design constraints on material

requirements. Materials proposed by the current MFE reference design

projects are reviewed by project and material function. The second

section presents an overview of the present state of relevant materials

technology, discussing candidate materials, experience, property data

base, fabrication, availability, costs, inspection and specifications.

In these two sections we focus on the low temperature material design

requirements and applicable modern materials technology, permitting a rapid appraisal of these two essential concerns and the development of

a research program to meet the needs of MFE superconducting magnet

technol ogy

The final major section (blue) presents recommendations for the

research program. The entire scope of the work is summarized and the

most cost-effective program at a 300-500 K$/year level is itemized.

Descriptions of each program element are included.

A complete reference list of the documents used for this survey

is provided. Appendices are included that itemize system components

and materials of the various major MFE design projects. Also, in the

appendices, a number of ongoing research projects that are relevant to

this program are described.

We thank the many people engaged in the fusion energy program who

contributed to this survey. We solicit their continued interest and

request that suggestions be communicated to the authors.

248
II. MFE MAGNET SYSTEMS

In this section we first describe the extent to which specific

operating conditions and design considerations place constraints on materials used in cryogenic components of existing MFE reference designs.

We then present an overview of the proposed uses of four classes of

materials: alloys, composites, insulators and conductors. Finally, we present a tabular summary of materials suggested for various components

of the conductors, the cryostats and the support structures of the

various designs.

The designs included in this survey are:

ANL: EPR-I, -II

GA: D-III, EPR-I, -II; TNS (with ANL), DEMO

MIT: HDTR

ORNL: EPR-I, -II; TNS, EBTR, DEMO

PPPL: TFTR, PLANT

U. WISC: UWMAK I, II, III: TETR

LLL: BB-II, MX, FERF

Table II-l gives an outline of the schedule proposed by ERDA for

the development of a number of the fusion devices treated here. Table

II-2 gives some general parameters of the toroidal system designs.

249
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251
II. A. Operating Conditions

The following five subsections describe the effects of the MFE designs

on the properties required of the various materials. In general the

magnets are very large, cold structures operating In a relatively severe radiation environment in the presence of high forces and large magneticfields both of which may change rapidly with time. The aspects of the

environment considered here are: thermal, mechanical, electrical,

magnetic and radiation.

II . A. 1

Thermal, because large ''1FE magnets must operate at very low

temperatures, thermal effects are a primary consideration. There are four distinct areas of concern and we will discuss them separately here.

a. Thermal Contraction

Very large forces can result from differential contraction of the

elements of the magnet structure on cooling to the operating temperature. For example, the copper in the UWMAK I TF coil, under no stress

at room temperature, acquires a tensile load of 21 MN/m2 (3000 psi) on cooldown. In an unrestrained part, the total dimensional change can

be quite significant. A round stainless steel coil form 20 m in

diameter will show a decrease in its circumference of about 20 cm on cooling to helium temperature. The case of metallic composites, such

as the conductor, with components which contract at different rates

often requires some careful thought. Fortunately, the calculations

involved in predicting the forces are relatively easy to make and, in

most cases, adequate data exist.

b. Heat Transfer

The removal of thermal energy from the magnet to the coolant is across an interface whose transfer properties are quite complex. The

various designs assume a heat generation rate of ^ 0.2-n.4 W/cm2 of

copper in contact with a pool boiling system. The permissible value in

force cooled systems, such as proposed by ORNL, is still a topic for

quite a bit of research at present. In the pool boiling case, the

obtainable heat transfer rate depends on a number of variables including

the type and condition of the surface, its physical orientation and temperature distribution.

c. Thermal Transport in Solids

Probably the most severe demand on materials in a thermal application comes when they must provide the mechanical means by which forces

on the cold magnet are transmitted to the room temperature support

structure. Non-metals, usually fiberglass epoxy, are preferred for this

application by all the designs, but, even then, rather clever techniques

must be developed to prevent excessive heat input to the cold region.

One example is that used by UWMAK of insulating rods which contact the

cold surface only in the event of a magnet quench, thereby supporting

the large additional loading caused by the quench. The low temperature

thermal conductivity of homogeneous solids is an area in which a large

amount of data exists. Some specific materials and complex composites

including superconductors and intact coil structures need to be investigated.

252
The thermal diffusivity , the ratio of the thermal conductivity to

the product of the specific heat and density, is an important concept

in the magnets. It determines, in effect, how rapidly thermal energy

can be moved through the structure. It is a field dependent quantity

in metals and, again, its behavior in complex structures needs to be

investigated

.

d. Dewar Insulation

Nearly all magnet systems in the MFE designs propose evacuated

super-insulation, usually aluminized mylar, for this application. In some instances a copper shield cooled to an intermediate temperature is

included. There is some indication that radiation damage may make the

mylar insulation of questionable value. Even a slight change in reflectivity could represent a serious problem here. Whether or not this is

the case must be determined and, if it is, an investigation should be

made of the rather wide range of alternate approaches to the insulation

problem. ANL suggests that eddy current heating in the aluminum layer

could result in unacceptable thermal losses for pulsed coil dewars. The

aluminum layer tends to be 0.05-0.1 ym thick. The eddy current calculation requires many assumptions but it seems that this, indeed, may be a general concern. Again, a simple experiment should provide the answers.

II .A. 2. Mechanical. The most common mechanical stress, the maximum

tensile hoop stress (omax) of a cylindrical coil can be approximately

calculated using the expression:

2 2 r + r

2 ^2

-

1 where K is a constant proportional to length, B is the maximum field; I is the associated coil current, r2 is the outer'^coil radius and V] is

the inner coil radius. Obviously, the higher field magnets, coupled

with severe space limitations, results in maximization of the coil hoop

stress

.

In Tokamak designs the pure-tension TF coils push against the internal

support cylinder, with a force (F) equal to:

B I F = ^

(£)

where i is the total length of magnet in contact with the central cylinder.

These forces are also quite large, but there is greater bearing area to

reduce the radial stress on the cylinder.

Conductor tensile forces from a given current and from a parallel con- ductor are significant in conductor cable and coil design but are not discussed here.

253
Table II A-1.

Approximate structural weights and forces

Project, Magnet, Total Weight Selected Approximate

Maximum Height . in^iKc dn^vn^ Forces

10" lbs (MN)

MX, Yin-Yang, 5m 430 (195) 31 (140)^

EPR (AND, TF, 13m 416 (189) 23 (102)^

EPR (ANL), Poloidal,

2-20m - 22 ( 98)^

UWMAK III, Poloidal,

9m

517 (235) 21 ( 95)^

UWMAK, II, TF, 30m 1562 (710) 46

29

(207)^

(129)^

UWMAK, I, TF, 30m 1980 (900) 71 (315)^

EPR (ORNL), TF, 12m 20

79

( 91)^

(350)^

LCP (ORNL), TF, 4m 22 (100)^°

Force to open up Yin-Yang coil.

From inadvertent radial coil translation of 2 cm. Vertical force on poloidal coil.

Maximum coil tension.

'Radial from TF coil to central support cylinder.

'Radial i n TF coi 1

TF coil maximum hoop force.

'Average tensile force in D-shaped coil.

'tF coil centering force in D-shaped coil.

*^Maximum radial force on central support cylinder.

254
Forces on shields from rapidly-pul sed coils must be considered.

Shields will probably be required if the pulse-time of some poloidal

coils remains short (probable for TNS-type reactors). Forces from the

toroidal field on the induced eddy currents of the shield will be of the

order of 4.4 MN/m, which will require extensive structural alloy reinforcement or composite construction.

From the design parameters discussed in the recent reports on TNS,

EPR or power reactor studies, approximate structural weights and forces

have been summarized in Table IIAl. These data are extracted from the

project reports as indicated and are, undoubtedly, very approximate.

However, several conclusions are obvious from this table:

(1) The magnetic field-induced forces are very large. This indicates that care must be taken to carefully choose the proper structural

members.

(2) The gravitational forces from magnet weight are quite small,

compared to the fi el d- induced forces.

(3) While radial forces on the central support cylinders from each

D magnet in the tokamak design are the largest, a layer bearing surface

is also assured between the cylinder and the magnet.

(4) The use of a higher allowable design stress will significantly

reduce the amount of structural material and permit more operating

space. If one assumes pure tensile or compressive forces and an allowable design stress of 30 ksi, the minimum cross-sectional areas range

from 700 to 2600 in2 (4500-15,300 cm2) to accomodate these forces. It

is obviously advantageous to use materials with larger allowable stresses;

for instance if a system is permitted to use 120 ksi as a 4K design

stress, the structural members are tremendously reduced.

(5) Yin-Yang coil restraining members are very critical, for there

is a large force which tends to open up the coil.

In the tokamak design inter-coil torque between TF coils is significant and arises from the interactions between the poloidal field and

the TF field and from possible malfunctions of a TF coil. Torque shells

or other structures are mandatory to alleviate these loads.

II .A.

3

. Electrical. The electrical properties of materials which are

of concern in fusion devices can be separated into three groups. Here

we give a brief description of each group and how it relates to the

reactor.

a. Conductors of Magnet Current

Our major interest is with superconductors, although copper and,

occasionally, aluminum are sometimes proposed for cryogenical ly cooled

magnets. The superconducting material itself has been adequately described

in many reports. Most designs propose to use the ductile NbTi . The Oak

Ridge EPR design does use Nb3Sn in the TF coils.

The stabilizing metal is usually copper, although aluminum is

proposed by the recent University of Wisconsin designs. In general, the

higher the purity of the stabilizer, the better, although there are

255
practical limits due to problems of fabrication, cost and strength. The

designs generally assume minimum resistivity values of ^ 3 nfi cm for Cu

and '^1 nfi cm for Al at 4 K. Magnetoresistance effects are calculated

to be large in copper, small in aluminum. Somewhat the inverse is true

of radiation damage effects. In some instances, such as in the pulsed

coils, it would be desirable to combine the stabilizing and strengthening function in one metal. Cold worked copper has been proposed for

this role by ANL and UWMAK. Various dispersion hardened Cu and Al

materials are being discussed also, but have not found their way into

the designs yet.

A completely demountable superconducting coil joint may be required

for access purposes. The problems involved would be immense. In any

event, joining materials, principally solders are involved in all real

coil systems and surprisingly little is known about them.

b. Eddy Current Losses and Control

The uncertainty in the fast field-rise time in tokamaks makes an assessment of eddy current effects somewhat difficult. It appears that

a rise time of 1 sec to fields approaching 7 T is reasonable for near- term devices. The major problem here lies in the possibility of excessive heating of metals operating near liquid helium temperature, such as the conductors, dewar walls and coil forms. Eddy current shields within

the superconductor composite used for pulsed coils appears mandatory.

Copper-nickel is always the choice. The use of nonconducting composite

materials for dewar walls is proposed by GA and may well become desirable or even necessary for the pulsed magnets. The option of shielding

the (dc) toroidal field coils with a cryogenical ly cooled pure metal

shield is considered by about half of the designs. One major advantage

is that a given refrigerator power can remove a great deal more thermal

energy from a 20 K region than from one at 4 K. Eddy current loss

calculations are strongly dependent on geometry and quite complex. In

general the power dissipation is proportional to the square of the field

rate of change and varies inversely with the electrical resistivity. In large reactors, those that produce net power, the suggested field rise

times are longer, of the order of 10 sec. This slower time dramatically

reduces the concern for eddy-current heating; suggestions of need for

non-metallic dewars and supports are replaced by suggestions for using

aluminum alloys, which are cheaper and lighter, but have high conductivities.

c. Electrical Insulation

If it were not for the radiation environment, electrical insulation

research for these large magnets would be reduced to measurements of the

effects of stress and temperature on conventional organics. This is a topic already being studied in connection with superconducting powerline

research and is a significant problem in itself. The maximum magnet

terminal voltage at quench has been calculated to be as high as 10 kV,

while most reasonable (organic) insulation schemes at present can withstand about 1 kV. Unfortunately, most existing information indicates

that both the mechanical and electrical properties of the organics will

be degraded by the radiation environment in the coils. Nevertheless all

extant designs use organic insulators in one form or another and all

256
indicate concern. There is no question, though, that a radiation resistant,

high voltage insulator is needed. It may be that some inorganic, or organic-inorganic, system will be necessary to satisfy these requirements.

II .A. 4. Magnetic. In nearly all applications, it has been assumed

desirable to avoid highly magnetic materials. As well as the possibility

of field distortion and significantly increased forces, there is a hysteretic heating effect in hard ferromagnetic metals. However, some ferritic steels may offer advantages of cost, availability and fabricability and these questions need to be investigated in more detail

before ferromagnetic materials are dismissed. In particular, it has

been suggested that ferromagnetic materials in the magnet region would

not noticeably (< 0.5%) perturb the fields at the plasma. On the subject

of hysteretic heating, a cursory look at the literature indicates that

very little information exists on the low temperature magnetic properties of materials, such as 400 series stainless steel and conventional

steels, whose applications are not directly magnetic. However, what

information does exist seems to indicate very narrow hysteresis loops.

Since the energy loss is proportional to the area of the loop, one should not expect a problem in general. A calculation for pure iron, a relatively very high loss material in this group, shows a heating of

80 X 10"^ J/cm3. In a 12 m magnet with a 0.03 m2 cross section, this

represents a loss of less than 1 W over a 100 s burn cycle. Note that

the loss per cycle is not dependent on the rate of change of the field

and is generated pretty much uniformly over the burn cycle of the

reactor. Eddy current losses, on the other hand, depend directly on the

the cycle frequency. It, thus, seems that the hysteretic losses are probably not a major concern where long burn times are involved.

A topic which may require special consideration is the matter of

ferrite in austenitic stainless steel welds. Typically the welds contain 4-10% ferrite, a ferromagnetic phase, to avoid hot cracking. In

nearly all designs such welds may be very close to the conductor. It

needs to be determined if magnetic effects from the ferromagnetic phase

can cause serious problems in these situations.

II .A.

5

. Radiation. The prospect of damage to the magnet system by

radiation is a concern for all but the largest device designs. Unfortunately, it is difficult to describe a general radiation environment for

the cold system with any accuracy. The type and amount of radiation and

its energy spectrum are all determined by the exact details of the

tritium breeding and blanket and shield systems. It appears, however,

that the neutron spectrum seen by the innermost magnet coils will be

similar to the fission spectrum. The typical neutron wall loading of

1 MW/m2, corresponds to a 14 Mev neutron flux at the first wall of 4.5

X 10^3 n/cm2/s. This flux will be attenuated by 10-5 -10-7 in its trip

to the magnets depending on the particular design. In general, the

smaller the device, the smaller the attenuation.

257
The neutrons produce damage by a variety of processes: neutron

capture with subsequent y oi^ B emission (n, y) and recoil of the

emitting nucleus, hydrogen production (n, p), helium production (n, a)

and damage by secondary charged particle cascades resulting from an elastic interaction between the neutron and atoms of the host material.

The cross sections for these interactions are highly dependent on neutron

energy and the interaction depends, also, on the presence of particular

elements in the system. It is important to realize that the energy

deposited in the magnets by some of these interactions may exceed that

of the incident neutron flux, perhaps by an order of magnitude in the case of the (n, y) reaction.

Magnets operating at liquid helium temperatures are very susceptible to thermal input as the heat capacity of copper at 4 K is more than

three orders of magnitude below its room temperature value; 1 W of heat

input to the helium bath requires about 250 W of room temperature

refrigeration capacity. Thus, even if the actual damage due to the

radiation is considered acceptable, for example a 5% increase in

stabilizer resistivity, the thermal input to the refrigerating system

may not be. Furthermore, the possibility of (n, p)-produced hydrogen

clogging the liquifier systems must be considered in some instances.

In terms of potential for radiation-induced problems in cold magnets

and their surroundings, the materials appear to fall in the following

order: superinsul ation , organic electrical insulation, stabilizer

superconductor structural material. There is little agreement on the

exact magnitude of the problem regarding the insulations. It seems most

prudent to assume that it is severe, that it is strongly dependent on the type of radiation and that the effects at low temperatures and

during annealing cycles will be quite different from those observed at room temperature. Some guidelines exist in the literature regarding

polymer susceptibility to energy deposition but direct extrapolation of

this information without experimental verification is not wise. Frequent

lip service is given to the use of inorganic insulations with all their

attendent structural problems, but little hard information exists. It seems imperative to determine quite soon whether or not organics may be

used.

The low temperature resistivity of the pure metals used to stabilize

the superconductor increases fairly rapidly under neutron irradiation

due to the production of vacancy-interstitial pairs. The result is

usually expressed in displacements/atom (dpa). The relationship of this

parameter to incident neutron energy depends on the metal, but typically

lO's of eV are necessary for one displacement. The net effect is quite

different for the two common stabilizers, aluminum and copper and there are various complex tradeoffs to be made involving magnetoresistance,

strength and radiation effects. It appears that the existing data are barely adequate under a number of assumptions about the neutron spectrum. Furthermore, it seems that room temperature annealing of the

magnets about every year is likely to be necessary at the neutron flux

levels presently being considered.

258
Radiation damage of the superconductor itself and the structural

materials does not appear to be a serious concern in the designs. The

materials do provide sources of secondary radiation, and a heat load and

gas load on the refrigerating system. From this point of view, their

impurity element chemistry may be important.

259
II. B. Design Considerations

The MFE project reports to date have dealt with conceptual designs

of superconducting magnet systems. Little consideration is given to the

exact sizing of structural details, fabrication methods and inspection.

In this section, design, fabrication and inspection considerations that

influence material selection are discussed.

1 1 . B. 1 . Codes and Specifications. A code is needed for the design,

construction and inspection of magnets for MFE devices. Adherence to

this code should provide assurance that the magnet system provides the

intended field and operates in a safe, reliable and efficient manner. Establishment of such a code for large magnets must await further developments in the technology. However, parts of the code should be set forth

as preliminary guidelines to facilitate the orderly development of

magnet systems. In this section, the guidelines and specifications

needed for proper selection, procurement, processing and inspection of

materials are identified.

Design guidelines needed for selection of materials are definitions

of operating loads and conditions and criteria for establishing allowable working stresses for candidate materials. For magnet systems,

description of the operating loads should include the anticipated load

history for normal operation, loads due to credible fault conditions and

procedures for treating combined loads. The operating conditions of

interest are the thermal environment, radiation dosage, magnetic field

profile and design life. Allowable working stresses should account for

the materials strength variability, the fatigue strength for a specified

scatter factor on fatigue cycles, fracture mechanics analysis and a factor of safety based on uncertainties in the loads and stress analysis.

Material procurement and processing should be controlled by a set

of specifications which are sufficient to preclude use of materials in

the magnet systems which have properties inferior to those assumed in

design. Material procurement specifications should be prepared for each

material and product form. These documents should specify minimum

acceptable values for tensile properties and, in certain cases, for

fracture toughness and for physical properties such as residual resistance ratio or magnetic permeability. Requirements will be established

for melting and primary processing, heat treatment, chemical composition,

ultrasonic quality and dimensional tolerances. Quality assurance provisions, testing procedures and reporting requirements should be described.

Microstructural requirements should be established as applicable, e.g.

ferrite content, grain size, inclusions, etc. Care should be taken in

the preparation of these specifications to avoid stipulating requirements that add cost but do not contribute to magnet performance.

Processing specifications are needed for processes such as heat

treatment, casting, forging, welding, adhesive bonding, soldering and

composites manufacturing. Requirements are to be established for equipment control, processing procedures, acceptance standards, quality assurance

provisions and workmanship. Where applicable, mechanical properties,

I

260
chemical compositions, surface requirements and dimensional tolerances

should be defined. Procedures for items such as surface preparation,

rework and repair and temperature control should be detailed. Applicable

testing procedures, subcontractor provisions and reporting requirements

should be covered.

Inspection specifications and standards are needed to provide a high

degree of confidence that no defects are present in the raw materials, or introduced during fabrication, which could cause unsatisfactory performance of the magnet system. Receiving inspection should verify that the

materials and components meet the requirements of the procurement specifications. Detail part inspection should be conducted in accordance with

specific instructions, dimensional tolerances and NDE standards applicable

to the part. Final assembly inspection should verify conformance to the

material, dimensional and installation requirements specified on the

engineering drawing.

II . B.2 . Fabrication. The size and complexity of proposed designs cause many problems unique to the construction of MFE devices. Materials

considerations related to fabrication of the conductors, coil forms,

eddy-current shields, dewars and magnet-support structure are identified

in this section. In each case, the problems associated with basic

production are not discussed; e.g. it is assumed that the conductor will

be supplied, but problems associated with winding the conductor are considered.

a. Conductors

The conductor is a complex assembly consisting of some of the

following elements: the superconductor, the stabilizer, a strengthener

such as stainless steel and channels or interstices to permit the flow

of liquid helium. This assembly must be co-wound into coils with an insulating layer and in some cases a strengthening member. Sufficient

winding tension must be used to assure that the relatively stiff assembly

conforms to the desired coil shape, maintains that shape during winding

and does not move excessively when the coil is energized. The winding

tension must be limited to avoid exceeding the strain tolerance, joint

strength or crushing strength of the assembly. Fabrication problems may

also arise due to the enormous size of the magnets, e.g. perimeters up

to 200 m for PF coils.

b. Coil Forms, Central Column and Magnet Support Structure

The structural elements of the magnet systems are large, subjected

to large forces and must fit within limited space. The sizes will

require build-up by joining many pieces. The forces necessitate use of

thick sections; the manufacturing, joining and inspection of thicksection parts require further development, particularly for stainless

steels and fiber reinforced plastics. The space limitation may necessitate the use of high strength materials, which are generally more difficult to fabricate.

261
c. Eddy-Current Shields

Special fabrication procedures may be required to preserve the high

RRR of the pure Al or Cu used in eddy-current shields.

d. Dewars

Space limitations and accessabi 1 i ty problems may hinder fabrication

of dewars. Leak-tight closure of the dewars may require welds with

limited inspectabi 1 i ty and high restraint. Force transmission from the

magnet to the support structure may require close-tolerance assembly to

avoid overloading the dewars. Accessabi 1 i ty requirements during assembly

and service may require fabrication of a removable dewar.

e. Magnet System Assemblies

Cantilever- type loading cranes (we think) have a present capacity

of about 300 tons. Giant shipyard gantry loading cranes, fixed on two

sides, have capacities up to 900 tons. The magnet systems are extremely

large and weigh in excess of 200 tons. The handling of these large

assemblies during erection may be prohibitive if the entire magnet system

is completely fabricated as a subassembly. It may be necessary to split

the coil into segments and assemble the segments into the final configuration on site. This will require development of new coil fabrication

concepts and a breakthrough in joining methods. Similarly, dewars may

have to be assembled around the coil after installation.

II .B. 3

Inspection. A quality assurance system is required for major

MFE devises to ensure that the materials, components and final assembly

of the devise conforms to the engineering drawings and quality specifications. Procurement controls, inspection standards and process controls

will be defined as the detail design and manufacturing plans are developed.

The technology to permit specification of reasonable quality standards

should either be available or developed for the principal elements of

the MFE magnet system. In general, the controls needed to construct

large, complex devices are well understood on the basis of military,

NASA and nuclear-power-plant experience. However, each new system

presents unique needs for quality standards, testing requirements and

non-destructive evaluation. The design features of large magnets for

MFE devises that may require the development of improved inspection

methods and standards are identified in this section.

a. Conductors

The conductors are complex assemblies that will be mass produced

and supplied to the magnet fabricator. Procurement quality standards,

sampling criteria and receiving inspection tests are needed to assure

that consistently satisfactory conductors are used to build the magnets.

The standards and tests must ensure that the conductor exhibits the

electrical, thermal and mechanical properties needed to function properly

in the magnet. During fabrication, controls must be developed to ensure proper fit of the conductor in the winding, complete coverage of the

insulation, adequacy of the conductor joints and, where applicable,

satisfactory incorporation of the strengthening elements. Upon completion of the magnet assembly, tests may be needed to ensure satisfactory

performance of the magnet prior to installation of the magnet into the

262
MFE devise - a procedure that is nearly irreversible in many of the

designs

.

The complex nature of the conductors and the diverse requirements

for satisfactory performance may eliminate the possibility of routine

NDE. Instead, performance tests on representative samples that require

removal of the sample from the lot of material may be essential for

receiving inspection of conductors. Inspections during magnet winding,

e.g. of joints, may be inadequate for complete assurance and thus, there

may be a need for stringent fabrication controls that virtually assure satisfactory performance. The current practice of strain gaging the

coil form to monitor winding tension is helpful, but further controls

may be needed.

b. Coil Forms and Magnet Support Structure

From an inspection standpoint, the choice of structural materials,

the thick section requirements, the overall size of the magnets, the

need for structural efficiency and the limited accessibility during

assembly create inspection diff icul i ties in many MFE conceptual designs.

The likely material choices for principal load carrying members are austenitic stainless steel and, where eddy current losses are a consideration, fiber reinforced plastic. These materials are not intrinsically

difficult to inspect, however the thick section requirements and the

need for joints in stainless steels arising from the large overall size

may cause inspection problems. Ultrasonic inspection is not satisfactory

for thick section joints and thus reliance is on radiography - a method

that does not reliably detect crack-like defects. The need for structural efficiency leads to higher operating stress levels, and consequently,

to greater defect sensitivity, more stringent defect acceptance standards

and more rigorous inspection. The limited accessibility and other space

constraints lead to designs with limited inspectabil ity, e.g. welded corners and blind installations.

c. Dewars

The dewars are likely to be a particularly complex inspection

problem because of the variety of materials used to transmit the forces

from the magnet to the support structure, to provide the thermal gradient from 4 K to room temperature and to contain the liquid helium

without leakage. Proper load transmission and thermal performance

requires care in assembly and bonding of many dissimilar materials. The

no leakage requirement may require blind closure joints in areas where

poor heat-input control may degrade the insulation.

253
II .C. Materials and Components

In this section we list the projects reviewed for this study and

present a brief summary of the uses proposed by these projects for four

classes of materials: alloys, composites, insulations and conductors.

Tables summarizing all of the materials choices by project are included

as Appendix A. Detailed discussions of the more important materials

choices for particular applications are given in Appendix B. These

listings include the requirements which must be met by the material.

They also include some assessment as to what research will be required

in order to produce design data for the application.

II .C.

. Material Overview. We present here an overview of the proposed

cryogenic uses for alloys and composites in MFE systems, plus an over- view of the materials proposed for insulators and conductors.

' ' a. Structural Alloys

Stainless steel and aluminum alloys are used exclusively for

metallic structural applications in all of the MFE conceptual designs

reviewed. Stainless steels are used for the coil forms (ANL, ORNL, GA)

the Yin-Yang coil case, the central support cylinder of tokamak designs

(UWMAK I and II, ORNL, ANL), the C-clamp structure for FERF, the support

structure (LLL, ORNL/EBT, GA) and dewar walls (ORNL/EBT, GA, UWMAK I and

II). Aluminum alloys are used by ANL for dewar outer-walls, external

support structure and the inner cylinder of the central support, and

by Wisconsin for the coil form, dewar walls, central support and coiljoint bolts for UWMAK III.

Generally the specific alloy grades are not designated, however,

most design groups had some preference. For stainless steels, type 316

is preferred by ANL and Wisconsin (for UWMAK I and II); Nitronic 40 is

preferred by LLL; ORNL is considering 304/304L, 310S and Nitronic 40.

For aluminum alloys, ANL prefers 7075-T6 for the external support structure which is at room temperature; and Wisconsin prefers 2219.

In an ORNL report. Long (1976) considers superalloys to be one of

the most promising classes of alloys for MFE magnet applications.

Inconels 706, 718, 750 and A286 are mentioned as candidates. No specific

applications for these alloys were included in the conceptual designs

of EPR, Demo and EBT.

b. Structural Composites

Non-metallic composite materials fill two primary structural functions in current MFE conceptual designs:

1. Reducing the refrigeration costs by providing supports transitioning from 4.2 K to 300 K. 2. Reducing eddy current losses (and thus refrigeration costs) due

to interactions between the pulsed magnetic fields of the poloidal magnets and the surrounding metallic structure.

Thermal Standoffs: All Tokamak designs visualize the entire toroidal

magnet system plus the central support column to be maintained at 4.2 K. The ORNL EPR reference design cites the need to support the total weight

264
on insulators having load carrying ability but without specifying the

insulator type. ANL proposes to support the central column on a 6FRP

cylinder, while the TF coils receive additional support from an external

G-10 laminate structure loaded in compression and transmitting the load

through the vacuum vessel walls. The warm-reinforcement system of BNL

supports their TF magnets with an 0.5 m thick thermal barrier of interleaved layers of stainless steel and GFRP (1:2 ratio). The barrier

modulus is controlled by the stainless, while the overall thermal conductivity is controlled by the GFRP. Thermal shorts are provided at 77 K

and 20 K. GA proposes the general principle of supporting the TF coils

on long, slender rods of low-thermal conductivity material, providing

both gravitational and lateral stability. Stacked ceramic discs and a "bicycle-chain" tension support were considered, however, the GA DEPR

design settles for a hollow, non-metallic composite cylinder supporting a stainless steel ball joint which provides some lateral adjustment. No

specific material is discussed.

A common problem is transfer of gravitational and fault loads from

the TF magnets through the dewars to external supports. PPPL uses G-10

pads inside the dewar which contact the opposite wall only in the presence

of fault loads. UWMAK inserts a large number of GFRP struts between the

TF magnet dewar walls, 10% of which continuously bear on the opposite

wall through conical spring caps to provide a centering force, while the

rest contact only under fault loading. The UWMAK diverter coil dewars

are similar in design, using Micarta spacers.

ORNL has an EF coil design of concentric stainless steel pipes, the

inner of which is the conductor jacket and the outer annul us of which

contains LN2- The coil must be isolated from the surrounding warm structure with load-bearing insulating struts capable of sustaining cyclic stresses.

ANL considers the alternatives of constructing the central support

column of their Tokamak from either concrete or a GFRP composite such as G-10. They are concerned, however, that the low modulus of the latter

(2 X 106 psi) will result in excessive conductor strain in the TF magnet.

Reduction of Eddy Current Losses: The ORNL, ANL and GA Tokamak

designs anticipate unacceptable AC losses due to eddy currents induced

in metallic structures near the pulsed poloidal coils. Substitution of

non-metallic composites for metal may be mandatory if refrigeration costs

are to remain reasonable.

Fiber-reinforced plastic is proposed for use as poloidal coil mandrels (ORNL, GA), for at least the inner wall of the poloidal magnet

dewars (ORNL, ANL, GA) and for structural casings surrounding the superconducting bundles in poloidal coils (ONRL, GA). The latter also function

iS the helium cyrostat, requiring minimum helium permeability.

The ORNL EPR design proposes the GFRP mandrels be divided into 12

coupled groups, each group made up of seven segments, layer wound with

enough tension to hold the poloidal coils in place without the need of

an outer case. GA proposes that the central OH coil solenoid be wound

on a GFRP spool and prestressed with stainless steel bands and inserted

into the central bore formed by the mating vertical sections of the TF coil.

265
ANL proposes non-metallic poloidal coil dewars with a 5-niil metal

helium permeability barrier, citing a McDonnel 1 -Douglas automatic glassfilament braiding method as a means of obtaining the desired result. GA

proposes a GFRP helium container (inner dewar wall) for their OH-coil

cryostat, preventing AC losses at the ends of the coil solenoid where a variable magnetic flux crosses the cryostat wall. Outgassing of the

composite is considered a negligible problem due to the low temperature.

Good electrical insulation between the OH coil and the outer dewar wall

is an added advantage. For the superconducting magnetic energy storage

coils in their Scyllac Fusion Reactor, LASL proposes a series of 128

cylindrical GFRP liquid helium dewars, each 1.07 m i.d. by 6.96 m deep,

operating at a pressure of 5 atmospheres. S-904 glass filament-winding

is proposed using a flexibilized epoxy resin.

ORNL would like to reduce AC losses n their poloidal field coils

by surrounding transposed conductor bundles with leak-tight GFRP tubes

serving as both a structural member to relieve the stress on the conductor and to contain the supercritical helium used as a coolant. ORNL cites

lack of knowledge of the elastic response to load, of failure mechanisms

related to load regimes and of response to dynamic fatigue at 4.2 K as major problems facing designers wishing to use GFRP materials for critical

applications in superconducting magnets. ANL arrives at a similar solution

to eddy current losses in their OH coil, differing from ORNL only in

proposing to use a rectangular G-10 tube with a thin-wall stainless liner.

Radiation Effects: T"^-? comments under Insulators (see the next

section) apply equally well to structural composites.

Alternate Materials: Suggestions for use of industrial laminates

reflect familiarity with such materials in normal magnet construction.

However, such materials have not been well characterized at cryogenic

temperatures, particularly in a radiation environment. Furthermore,

industrial laminate tradenames refer to a wide variety of products

produced by a given company. The National Electrical Manufacturers

Association (NEMA) specifications for industrial laminates are more meaningful; however, the specifications require only functional performance at or slightly above room temperature and can therefore be met by

manufacturers products differing significantly in type and weave of the

reinforcement or in the matrix resin. Existing NEMA specifications are inadequate for demanding cryogenic laminate performance. It is also

highly probable that the less expensive borosilicate glass used in

industrial laminates will lead to more rapid degradation of the laminates

under neutron irradiation than would be the case with the more expensive

(but higher modulus) silica-alumina-magnesia S-type glass used in more specialized composites.

Whenever section stiffness becomes a sufficient problem, higher

modulus fibers such as Kevlar, graphite or boron are readily available

for substitution for glass at some increase in cost, although hybrid

construction can often be used to advantage. Obviously, boron fiber is

restricted to areas of low neutron fluence.

These subjects are discussed in more detail in Section III B of

this report.

266
c. Insulators

GA mentions the possibility of using polyethlene, mylar or glassfiber reinforced plastic (GFRP) for electrical insulation in superconducting coils, while UWMAK suggests Kapton for this purpose. However, the

materials most often specified are the NEMA G-10 (glass-epoxy) or the

micarta (cellulose-phenolic) industrial laminates.

The possibility of combining electrical insulations with structural

support of the superconducting coils is recognized in all designs. However,

ORNL believes that the magnet stresses will be too high for adequate strain

limitation by low-modulus GFRP restraints, leading to their decision to

separate the load-carrying and the insulation function in the coils of

their EPR devices.

ANL proposes to spirally wrap the TP coil cables and supporting stainless steel plates with G-10 glass-epoxy, providing nonvapor-locking helium

cooling channels. The insulation is given as 1 mm thick, 1 cm wide, with

50% coverage. The 200 V dielectric breakdown required for the application

is believed to be well within the capabilities of the insulation. The

overall coil form is stainless steel insulated with micarta or G-10, 1 cm

thick. ANL notes that the multilayer OH and EF coils require large gaps

between layers and proposes to fill them with G-10. Concern is expressed

that pinholes and cracks which exist in the material or which form during

cooling, or in service, will degrade the dielectric properties of the

insulator when the magnet is energized. Such damage is irreversible and,

since insulation is almost impossible to replace in a completed magnet,

insulations will require much more study. ANL also proposes to "jelly

roll" band the 150 basic cables in their poloidal coils with GFRP 2 mm

thick, 2 mm wide and with a 3-inch pitch. No stainless steel reinforcement

is used, the stresses being transferred through the composite to external

support rings. Insertion of 2-cm thick G-10 or Micarta laminates every

10-15 turns is proposed to provide additional resistance to the com- pressive (repulsive) forces and to relieve stress on the cables. Concerning energy storage and transfer superconducting magnet systems, ANL

proposes to separate pancake-type coils by GFRP boards 1.25 cm thick with

50% coverage. The basic conductor support is a U-shaped GRFP channel,

3 cm high, 12 cm wide with an 0.5 x 10 cm slot. A 200-strand, transposed

superconducting cable wound around a GFRP board as a coil form is inserted

into the slot. Again, ANL expresses concern about dielectric strength

and possible arcing problems in use, perhaps accentuated by a radiation

environment. In the absence of hard data, ANL tentatively specifies an epoxy set with an aromatic curing agent, reinforced with glass and filled

with granular alumina.

The UWMAK design contemplates using GFRP to support and insulate the

superconducting windings in the preformed stainless or aluminum discs which

form the main structural element in their TF magnets, using Micarta spacers

between the discs. Divertor and transformer coils would utilize Kapton

between conductors and Micarta spacers between coil pancakes. UWMAK also

proposes to spiral wrap the poloidal coil surfaces with B-staged GFRP to

50% coverage.

267
LLL proposes a polyimide or a polyimi de-glass laminate in the form

of slotted sheets for interlayer insulation in their FERF (mirror) superconducting magnets. Interturn insulation would be spaced, laminated dots

held on a film substrate.

Finally, the DC homopolar energy storage system proposed by LASL

would use GFRP struts to insulate and support the superconducting excitation coils in the stator iron and would use a GFRP strut as a double-acting

thrust bearing to maintain the axial ity of the rotor with respect to

the superconducting coils.

Radiation Effects: All designers of MFE systems express concern over the possible adverse effects of neutron and secondary radiation on the integrity of organic materials. ANL believes that the epoxy insulator

in the magnet will see 3.5 x 108 rads at a lifetime integrated wall load

of 2.5 MW-yr m-2. This is believed to be well below the lO^-lO^O rads

and 10 MW-yr m-2 level at which the properties of epoxy are expected to

be significantly damaged. Extrapolation of data from higher temperatures

has lead ANL to a choice of the aromatic base epoxies as the organic

material most resistant to radiation damage -- perhaps 10-20 times more resistant than mylar. ANL evidence indicates that mylar could only be

used up to 1 MW-yr m-2 in EPR designs. There are no 4.2 K radiation

data. However, ANL believes that the damage to GFRP structures will be

primarily manifested in strength reduction rather than in dielectric

property changes.

These views are not necessarily shared by LLL who considers that

radiation damage might preclude all organic insulators in MFE superconducting magnet systems. LLL states that polyimide is the most radiation

resistant polymer presently available and should be used in place of

epoxy. Designers of the UWMAK system are somewhat more optimistic about

the chances of mylar as an insulator, calculating that the insulation

would be exposed to 60 x 10^ rads in 30 years in their design, while

citing a previous analysis as showing the mylar damage threshold to be

at 30 X 106 rads with 1.2 x 10^ rads being required for a 25% decrease

in tensile strength. UWMAK also assumes a threshold for GFRP damage

at 8 X 109 rads, with 5 x lOlO rads required for a 25% strength decrease,

suggesting that the problem should be minimal.

A major unknown is the extent to which release of energy stored

in polymers during 4.2 K radiation will result in damage when warmup

releases the energy (Wigner energy).

Alternate Materials: If radiation damage to unreinforced polymers

becomes a significant problem, reinforcing the polymer with S-glass

(silica-alumina-magnesia) is expected to double the radiation resistance.

The use of the less expensive E-glass may cause a problem, as its

lime-alumina-borosil icate composition is sensitive to heating and

degradation in a neutron radiation environment.

Should a higher modulus in the plane of the reinforcement be required,

fibers such as Kevlar or graphite might be substituted for glass. Graphiteepoxy composites could be substituted for stainless steel support plates

268
around which the conductor is wound, providing equal or higher stiffness, higher strength and better insulating qualities thus reducing

the probability of turn-to-turn shorts. The very low thermal con- traction of graphite composites would have to be carefully considered

in this and similar applications.

Care should be taken in referring to industrial laminates by

company tradename or by NEMA specifications, ^'either are sufficiently

precise for specifying demanding cryogenic performance. This subject

is addressed further in Section III B.

d. Conductors

In nearly all proposed applications in MFE designs, NbTi cryogenically stabilized with copper is the superconductor of choice. In applications which are basically dc and involve the largest forces, such as TF coils in tokamaks, the conductor is strengthened by stainless steel,

usually of significant cross section. Allowable strain values tend to

be ^ 0.2%, although it seems this constraint is somewhat artificial. It

has been suggested that a copper alloy might serve both the stabilizing

and strengthening function in some instances. In other designs, such

as Yin-Yang magnets and some TF coil constructions, where the load can be

transferred to the coil case, the strength of the pure copper stabilizer

or a cold worked copper is assumed adequate. When pulsed coils are con- sidered, the coils generally are more complex and contain less stabilizer

along with a high resistivity alloy, usually CuNi , which breaks up potential eddy current paths. NbTi is still the superconductor. The conductor

structure often is a braid containing a stainless steel wire or strip for

strength. Stress levels in general are not as high for these coils as for the dc ones.

Current carrying capacity of the conductors nearly always falls in

the 5-20 kA range and most are cooled by pool boiling helium, although

forced flow cooled conductors are also under serious consideration as they greatly simplify the overall magnet structure.

Radiation effects expected in the conductor composite depend strongly

on the particular design, but it appears that, in all cases, degradation

of the low temperature conductivity of the stabilizing metal will require

periodic annealing of the magnets at temperatures near room temperature.

The effect on the superconductor itself is not expected to be severe.

II .C. 2. Component Research Summary. Table II C-1 presents a general

summary of the material choices listed by component. Alternate materials are suggested and the required data is summarized in abbreviated

form. We have also indicated the project class for which we feel the

information will first be necessary and class of data. Detailed

listings of the materials selection for the components by each MFE

reference design are given in Appendices A and B.

269
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272
III. MATERIALS TECHNOLOGY

The previous Section of this report addressed the cryogenic materials problems as reflected in existing MFE design. Only general materials classes were usually suggested for a given component, e.g., stainless steel, Micarta, aluminum, etc. We now consider which specific

alloys, composites, etc., are the best current selections based upon

contemporary and past experience, on the existing data base, availability

and cost. We consider fabrication problems with specific materials and

how control over the quality of a material and of components can be

established. Finally, specific alternative materials are examined. The

Section is divided into four general material classes: alloys; structural composites and bulk insulators; insulating films, coatings and

inorganics; and conductors.

273
III. A. Structural Alloys

Alloys will be used for the main structural members of the superconducting magnet systems needed for MFE devices. Designs of tokamak

and mirror machines that require superconducting magnets are currently

in the conceptual stage (except for MX), but many design features that

strongly influence materials selection are reasonably well established.

The principal features governing alloy selection for structural applications are the enormous size and stored energy of the magnet systems, the

extremely high forces exerted by the magnets, the massive structural

elements needed to restrain these forces, the limited space available

for the structure, and the need for accessabi 1 i ty to install and periodically remove the blanket and shield systems -- systems that are com- pletely surrounded by magnets and the support structure. These features

result in the need for alloys with sufficient strength and stiffness to

perform the required structural functions within the space available,

sufficient fatigue and fracture resistance to operate safely, and sufficient fabricabil ity to permit manufacture and assembly of the components.

The choice of materials is further restricted by the adverse operating

environment of the magnet system. The alloys are subjected to temperatures

down to 4 K and must not be susceptible to low temperature embrittlement.

The presence of high magnetic fields cautions against the use of ferromagnetic materials in certain application because they would be subjected

to higher forces due to the magnetic field. The rapid magnetic field

changes associated with cyclic operation of the poloidal field coils cause heat inputs to the structure from eddy current losses. Consequently,

alloys with high electrical resistivity are preferred for use at 4 K. The radiation environment eliminates use of alloying elements, such as cobalt, which transmute to long half life products. With this exception,

radiation does not appear to be a significant factor in the selection of

structural alloys because dosage levels sufficient to degrade structural

performance probably cause unacceptable deterioration of other system

materials. Finally, the microstructure of the material must be stable

when exposed to the operating strains, temperature, and radiation; i.e.

the alloy should not transform to either a brittle or a ferromagnetic

state.

The constraints on materials selection caused by size requirements,

space limitations and operating environment significantly reduce the

number of candidate alloys. Further limitations arise due to considerations related to experience, cost and availability. Currently, the

leading candidates are the austenitic stainless steels and aluminum

alloys. The candidate alloys are discussed in detail in the following

sections

Ill .A.lAustenitic Stainless Steels. Austenitic chromium-nickel stainless steels are the most widely proposed alloys for structural applications in FIFE magnet systems. These steels have the best combination of

strength, stiffness and toughness at K. The physical properties also

offer advantages over competing materials. Specifically, the modulus is

high, the thermal expansion is close to that of the copper-stabilized

conductor, magnetic permeability is low and the electrical and thermal

conductivi tes are low. The principal disadvantages are the high cost of

construction and problems arising from microstructural stability and

control

274
a. Preferred Grades

For superconducting magnet systems, the preferred stainless steels

are AISI grades 304 and 316, and alloy 21/6/9 (Nitronic 40 and Pyromet

538). AISI 304 contains 18-20% Cr and 8-12% Ni; it has a typical yield

strength of 413 MPa (60 ksi at 4 K. AISI 316 contains 16-18% Cr, 10-14%

Ni and 2-3% Mo; it has a typical yield strength of 551 MPa (80 ksi) at 4

K. Alloy 21/6/9 contains 19-21.5% Cr, 5.5-7.5% Ni , 8-10% Mn and 0.15-

0.4% N; it has a typical yield strength greater than 1033 MPa (150 ksi)

at 4 K. These three alloys were selected on the basis of strength,

toughness, and austenite stability at 4 K, and upon consideration of

cost, availability, experience and f abricabi 1 i ty . Each of these factors

is discussed in detail in the following paragraphs.

Several other 300-series stainless steels have been widely used in

cryogenic applications, but are not preferred for magnet system applications for a variety of reasons. Lower nickel grades such as type 301

and 302 have insufficient austenite stability. Type 310 has superior

austenite stability but is is expensive (70% more expensive than 304)

and not readily weldable in heavy sections. The extra-low-carbon grades

(types 304L and 316L; .03 max C) may be used interchangeably with the

standard grades (304 and 316, respectively) providing the lower strength

(about 20% less) is acceptable. The stabilized grades (321, 347, and

348) are not generally used below 76 K (INCO, 1974); the reason for this

is not clear.

For applications requiring higher strengths, several alloys other

than Nitronic 40 were considered. Alloy A-286 was eliminated from

consideration because heat treatment is required and weldability is

poor. Several Ni-Cr-Mn-N stainless steels - alloy 22/13/5 (Nitronic

50), alloy 18/3/12 (Nitronic 33), Cryogenic Tenelon, Kromarc 58, 216

(ASTM 240, Grade XM17) and others -offer great potential for MFE applications, but there is currently insufficient information available on these

alloys to permit a reasonable assessment of the potential. Similarly the

nitrogen strengthened stainless steels, 304N and 316N, have greater strength

and austenite stability than 304 and 316 respectively, but little is known

of their properties at 4 K.

b. General Information

Experience: The most widely used stainless steel for cryogenic

service are AISI types 304 and 304L. These alloys have been used for

structural members in most of the superconducting magnets built for high

energy physics devices. NASA used 304 nearly exclusively in ground

installations for piping systems and containment vessels for liquid

hydrogen at 20 K. The experience has been satisfactory.

Type 316 is a commonly used stainless steel, but experience with

this alloy in liquid helium applications is nil. The only known application of type 316 at 4 K is the ANL U-25 superconducting dipole magnet

for an MHD device. The heavy section 316 had an abnormally large grain

size and exhibited lower strength than expected at 4 K (Reed and Read,

1977).

275
Alloy I'XI^I^ is a relatively new alloy but it has been used by LLL

for the Baseball II superconducting magnets. In the Baseball II application, the 21/6/9 is only stressed to 207 MPa (30 ksi) and consequently

the successful operating experience is not indicative of this material's

performance as a high strength alloy, i.e. in highly stressed applications (Henning, et al., 1969).

Specifications. The ASTM and ASME specifications for various

product forms of AISI types 304, 315 and several other 300-series

stainless steels are summarized in Table III A-1 . The Federal specifica

tions for 300-series stainless steels are QQ-S-766 and QQ-S-763. Alloy

21/6/9 listed as grade XM-11, is covered by ASTM specifications A412 for

plate, sheet and strip and A276 for base and shapes.

Many of the ERDA laboratories have specifications for procuring

stainless steels for specialized applications. A modified specification

for type 304 which requires the Ni content to be at the high end of the

allowable range has been used for superconducting magnet applications.

Special specifications are also used to purchase high purity stainless

steels manufactured by either the vacuum-i nduction-mel t/vacuum-arcremelt (VIM/VAR) process or by the VIM/electro-slag-remelt (VIM/ESR)

process. For procurement of 21/6/9 certain specifications stipulate

tighter controls on allowable nitrogen content, e.g. LLL limits the N range from 0.20 to 0.25 instead of 0.15 to 0.40 as specified in the ASTM

specifications

Cost: Cost information for 304, 316, 21/6/9 and several other

stainless steels are summarized in Table III A-2. Costs for round bar,

flat bar and plates are included to provide information regarding the

influence of product form on price. The cost data are applicable to

orders over 4500 kg (10,000 lb) delivered in February 1977. The information was obtained from Armco Steel Co. and G. 0. Carlson Co.

The cost data of Table III A-2 are applicable to stainless steels

purchased to ASTM requirements. These steels are produced in electric

furnaces and are not vacuum processed or remelted to improve purity.

Many of the ERDA laboratories have more stringent specifications for

procuring stainless steels for specialized applications. The processing

chemistry and metallurgical requirements of these specifications can result in significantly higher prices and in some cases limit the number

of suppliers and availability.

Availability: The availability of type 304, 316 and 21/6/9 stainless steels is essentially unlimited in standard product forms at the

present time. Standard product forms include plate, sheet, strip, bar

and forging billets. Each of the alloys can be cast, extruded or welded

into pipe or tubing. Since these alloys contain 30 to 40% of Cr, Ni, Mn

and Mo, the long term availability may limit applicability in a fusionpower economy. However, there should be no problem for the next twenty

years because of the limited number of devices planned.

275
Table III A-1

ASTM and ASME specifications for various forms of

austenitic stainless steels (including AISI Types

304, 304L, 316 and 316L)

.

Product Form

ASTM

specifications

A c Mr~ ASME

specifications

Plates A240 SA-240

Forgings A473 SA-182

Bars and sections A479 SA-479

Seamless tubes A213 SA-213

Welded tubes A249 SA-249

Seamless and welded tubes A269

Bolts, screws, studs, stud

bolts

A320 SA-320

Nuts for bolts A194 SA-194

277
Table III A-2.

Cost Data for Stainless Steels

(Dollars per pound)

Grade

pi Flat Bar

1", Round Bar Plate^

304 1 .29 1.15 .93

OUM-L QQ

O 1 D 1 fin 1 . HO

316L 1 .66 1 . 52 1 .27

Ni tronic 33 1 .26 1.12 1 .00

Nitronic 40 1 .69 1 .55 1 .32

Ni tronic 50 1 .94 1 .80 1.85

^

5 X 10 cm flat, hot rolled, annealed and pickled

^ 3.8 to 12 cm round, centerless ground

^ 1.9 to 7.6 cm thick, 122 to 244 cm wide, 305 to 1070 cm

long plate; hot rolled, annealed and pickled

278
c. Properties

Tensile: Typical yield strengths of AISI 304, 316 and Nitronic 40

are 413 MPa (60 ksi), 551 MPa (80 ksi) and 1033 MPa (150 ksi), respectively. In these alloys, the tensile strength and to a lesser degree

the yield strength increase with decreasing temperature. This is particularly true for 21/6/9 and other nitrogen-strengthened stainless

steels. Although tensile tests have been conducted on many austenitic

stainless steels at 4 K, there are insufficient data on any of the

alloys to establish an allowable stress based on the minimum credible

yield strength.

Compression: The compressive yield strength of AISI 304 has been

measured at temperatures to 4 K. The yield strength in compression

appears to be significantly higher, 930 MPa (135 ksi) in one investigation (NBS, 1977) and 827 MPa (120 ksi) in another (INCO, 1974), than

the tensile yield strength. This is in marked contrast to the behavior

of most structural alloys which exhibit approximately equal yield

strengths in tension and compression. Tests should be conducted to

verify the compressive yield strength in AISI 304 and to determine if

similar behavior occurs in AISI 316 and Nitronic 40. An increased

allowable stress in compression would significantly reduce some section

sizes if either 304 or 316 is selected because of the compressive nature

of the centering forces.

Shear and Bearing: The shear ultimate strength and the bearing

yield and ultimate strengths have not been determined for any alloys at

4 K.

Fatigue: Smooth and notched (Kt =3.1) fatigue tests at 4 K have

been conducted by Schwartzberg and Kiefer (1976) on 304L, 310, 21/6/9

and A286 stainless steels under fully reversed (R =-1) stress cycling.

Strain cycling fatigue tests at 4 K have been conducted by Nachtigall

(1974) on 304L and 310 stainless steels. In both investigations, the

fatigue life was better at 4 K than at room temperature for endurances

greater than 1000 cycles. At low cycle lives (less than 500 cycles),

Nachtigall observed a crossover, i.e. less fatigue life at 4 K than at room temperature. Nachtigall compared his experimental data with that

predicted by the Manson (1965) method of universal slopes and showed

that the fatigue behavior of these materials at 4 K can be predicted

within a scatter factor of three from material tensile properties

obtained at 4 K.

Fatigue Crack Growth: The fatigue crack growth behavior of several

austenitic stainless steels has been determined at 4 K by Tobler and

Reed (1976). Alloys tested include 304, 304L, 316, 310, A286 and 21/6/9.

Their results indicate that fatigue crack growth rates are generally

slower at 4 K than at room temperature. The crack growth rates generally decrease with increasing austenite stability, i.e. AISI 310 and

A286 are the best, followed by AISI 304 and 316; alloy 21/6/9 which

exhibits faster growth rates at 4 K than at room temperature, is the

worst.

279
Fracture Toughness: Fracture toughness tests at 4 K have been

conducted by Tobler (1976) on 304, 316, 310, A286 and 21/6/9. The

results scale with yield strength. The 300-series alloys have low yield

strength and excellent toughness at room and cryogenic temperatures.

A286 has a relatively high yield strength and low toughness at room temperature and 4 K. Alloy 21/6/9 exhibits a substantial increase in

yield strength as temperature is reduced to 4 K; the fracture toughness

exhibits a corresponding decrease with decreasing temperature.

Physical Properties: The physical properties of austenitic stainless steels are good relative to other alloy systems for MFE-magnet

structural applications. The elastic moduli are relatively high, limiting elastic displacements and improving buckling stability. The thermal

and electrical conductivities are low resulting in reduced thermal

losses. The thermal expansion is close to that of copper thereby minimizing thermal stresses due to cooldown. The magnetic permeability is

low relative to other iron or nickel base alloys resulting in minimal

interactions with the applied magnetic field.

d. Fabrication

The fabricabil i ty of austenitic stainless steels is certainly

sufficient for MFE magnet systems as attested by the widespread use of

stainless steels in a variety of industries. Fabrication characteristics

that should receive special consideration for MFE-magnet applications

are briefly reviewed in this section.

Welding: Stainless steel alloys 304, 316 and 21/6/9 are readily

weldable by all common welding processes providing the appropriate

consummables and procedures are used. Filler wires are generally formulated to provide 4 to 10% delta ferrite which tends to reduce the

incidence of hot cracking. For magnet applications, the ferrite phase

is objectionable because it is ferromagnetic and because it may result

in some brittleness at low temperatures. If a wholly - austenitic weld

is desired, the chemistry of the weld deposit must be adjusted - usually

by increasing the Ni or Ni-equivalent as shown by the Schaeffler constitution diagram for stainless steel weldments. Microfi ssuring can be

minimized by a small increase in carbon, a substantial increase in

manganese and increasing the nitrogen content. Thus, the Mn-N steels,

such as 21/6/9 may be better suited than 300 series steels for obtaining

sound welds with low permeability. Nickel -based filler wires are some- times used to obtain sound weld deposits with excellent low temperature

toughness under conditions of maximum constraint - such as in the welding of heavy sections. The main disadvantages of the Ni-base filler

wires are high cost and low yield strength; the latter can be very

important at 4 K where 21/6/9 and to a lesser extent AISI 304 and 316

have substantially higher yield strengths.

An important problem in stainless steel weldments is sensitization - the grain boundary precipitation of Cr carbides at temperatures between

about 500-800°C which locally depletes the Cr content and therby reduces

corrosion resistance. In MFE magnet applications there is not expected

to be a corrosive environment. Limited impact tests indicate that

sensitization does not significantly reduce fracture resistance at

280
cryogenic temperatures. However, our recent experience on the U-25

dipole magnet bore tube AISI 316 alloy urges caution: this material

exhibited all the characteristics of highly-segregated grain boundaries,

having large grain size, intergranul ar tensile fracture at 4 K, and

reduced mechanical properties. Chemistry changes to prevent sensitization, such as specifying stabilized grades (321, 347) and extra-lowcarbon grades (304L and 31 6L) should be avoided for more reasons than

simply the added cost. Certain stabilized grades contain Cb and/or Ta

which tend to promote microf i ssuri ng in whol ly-austenitic welds (Linnert,

1967). The extra-low-carbon grades have lower strength, particularly at

4 K, and less austenite stability.

Heat Treatment: The best possible heat treatment is no heat treatment. One of the principal advantages of wrought-anneal ed-austeni ticstainless-steel s is that heat treatment is generally not required. The

possible exceptions are annealing at 1000 to 1100°C following severe

forming operations and stress relief at 420 to 480°C to reduce peak

residual stresses and improve dimensional stability.

Machining: Annealed-austeni tic stainless steels are more difficult

to machine than carbon steels or aluminum alloys because of their relative-ly high tensile strength, work hardening rates, and low thermal

conductivity. These properties cause poor finishes, heat buildup and

excessive tool wear. As a result, it is necessary to use heavier feeds,

slower speeds, better quality tool bits, more rigid machines and more

power.

Forming: The candidate stainless steels have excellent formability.

Alloys with more austenitic stability than types 304 or 316 may be

needed if severe forming operations are required. Alloy 21/6/9 is

reported to have excellent austenitic stability, even after 60% reduction at room temperature, the permeability is less than 1.02.

Forging: The austenitic stainless steels can be forged into shapes

suitable for many MFE-magnet applications, e.g. forged pancake billets

for ring rolling coil forms. Stainless steels are more difficult to

forge than most common structural alloys because of its high strength at

elevated temperatures. Within the family of austenitic stainless steels,

the forgeability of types 304 and 316 is rated good (Jensen, 1966).

Casting: The austenitic stainless steels are widely used for

manufacturing corrosion resistant casting. The casting alloys equivalent to AISI types 304 and 316 are ACI (Alloy Casting Institute) types

CF8 and CF8M respectively. The ferrite content of these alloys may vary

from 0 to 40% due to chemistry variations within the allowable ranges.

The ferrite content of 21/6/9 casting could reach 20%. Ferrite levels

in this range are likely to be unacceptable because of low fracture

toughness and high permeability. Ferrite-free castings require use of

higher nickel content alloys such as ACI grades CK 20 (corresponding to

AISI grade 310) and CN-7M. The CN-7M grade, which has a nominal composition of 28Ni, 18Cr, 3.5Co, 2.5Mo and 0.07C, is widely used for certain

corrosion-resistance applications and should be a suitable ferrite free

casting alloy for cryogenic applications. The problem with using ferritefree castings is that the cost benefits of using castings may be lost

because of the high alloy content.

281
e. Inspection

During fabrication of MFE magnet systems, the stainless steel

material and components will be inspected using conventional methods.

The only problem peculiar to the proposed designs is the wide spread use

of heavy-section stainless steel welds. These welds cannot be effectively inspected for the presence of cracks or crack-like defects.

Radiography is of limited use because of its inability to reliably

detect cracks. Ultrasonic methods don't work effectively because of trie

high noise levels and large and variable attenuation arising from the

large-grain-size welds. Standards are needed to assure that the defect

sizes that can be reliably detected do not adversely affect structural

integri ty

In-service inspection of magnet-system components will be particularly difficult because of the low temperatures and radiation.

II I .A. 2 . Aluminum Alloys. Aluminum alloys have been proposed for many

structural applications in MFE magnet systems. Their principal advantages

are: low as-fabricated cost, lightweight, non-magnetic behavior, stable

microstructure, and good retention of strength and toughness at cryogenic

temperatures. The main disadvantages of aluminum alloys ar- low strength

in weldments, high electrical and thermal conductivity and an unfavorable

elastic modulus (too low) and thermal expansion (too high) for many

appl ications

a. Preferred Alloys

For cryogenic service, the preferred aluminum alloys are 5083, 2219

and 6061. Alloy 5083 is generally used in the annealed condition (5083-

0), has a nominal composition of Al-4.5 Mg-0.6 Mn; and has a typical

yield strength of 172 MPa (25 ksi) at 4 K. Alloy 2219 is generally used

in a heat treated condition (T 851, T 87 etc.), has a nominal composition

of Al-6.3 Cu-0.3 Mn, and in the T 851 temper has a typical yield strength

of 482 MPa (70 ksi) at 4 K. Alloy 6061 is usually used in the heat

treated condition (T 6, T 651, etc), has a nominal composition of AllMg-0.6Si-0.3Cu-0.2Cr, and in the T-6 temper has a typical yield strength

of 380 MPa (55 ksi) at 4 K. Although each of these alloys is readily

weldable, only 5083 retains full strength in the welded condition and is

the best choice for most large welded components. Alloy 2219 has the

highest strength of the three alloys and is recommended for bolttd

construction where space or weight need to be conserved.

Several other aluminum alloys have excellent properties at 4 K and

may be used for special applications. Alloy 3003 has been widely used

for cryogenic applications and is the best choice when brazing is used

for joining. The high strength alloys such as 2024 and 7075 (in an overaged temper such as T 73) may be used for nuts and bolts.

b. General Information

Experience: The most widely used aluminum alloy for cryogenic

applications is 5083. Thousands of tons have been used for LNG tanks in

plates up to 18 cm thick. Alloy 2219 is used in aerospace applications

282
including cryogenic tankage for the space shuttle. Alloy 6061 is used

primarily for piping systems. Although 2014-T 651 has been successfully

used by NASA for propel 1 ant tanks (LH2 and LOX), it is not recommended

for CTR applications because of its marginal weldability and toughness.

Specifications: The Federal, ASTM and ANSI specifications for

aluminum mill products of alloys 5083, 2219 and 6061 are summarized in

Table III A-3. For a more complete listing of specifications refer to

the Aluminum Standards and Data Book (Aluminum Assn., 1976). Procurement of heavy sections may require additional specification requirements

on through-thickness ductility.

Cost: The cost of 2.5 cm-thick aluminum alloy plates are as follows:

$1.83/kg for 5083-0, $2.10/kg for 2219-T 851 and $1.90/kg for 6061 -T 6.

Notice that cost do not significantly vary for the various grades as they do for stainless steels. The costs are applicable to orders over

13,600 kg delivered in February 1977. The information was obtained from

Alcoa.

Availability: The availability of 5083, 2219 and 6061 is essentially

unlimited in standard product forms at the present time. Standard

product forms include plate, sheet, strip, bar and forging billets.

Pipe and structural shapes are readily available in alloy 6061. The

long term availability of these alloys is excellent because of the

abundance of aluminum and the principal alloying elements, Mg and Cu.

c. Properties

Tensile: Typical yield strengths of 5083-0, 2219-T851 and 6061 -T6

at 4 K are 172 MPa (25 ksi), 482 MPa (70 ksi) and 380 MPa (55 ksi),

respectively. In these alloys, the tensile strength and to a lesser

extent the yield strength increases with decreasing temperature. Many

tensile tests have been conducted on 5083 to 77 K, but there are insufficient data on any of the alloys to establish an allowable stress at 4

K based on the minimum credible yield strength, particularly for heavy

sections

,

Compression, Shear and Bearing: Values for these properties have

not been determined for any aluminum alloy at 4 K.

Fatigue: Strain cycling fatigue tests have been conducted by

Nachtigall (1974) on 2219 and 2014 alloys. The results exhibit the same trends as discussed for stainless steels, i.e. improved life at endurances

greater than 1000 cycles and reasonable correlation between Manson's

theory and the test results.

Fatigue Crack Growth: The fatigue crack growth behavior of 5083

(Tobler and Read, 1976) and a Al-6Mg alloy (McHenry, et al., 1976) have

been determined at 4 K. The results indicate that growth rates are slower at 4 K than at room temperature.

Fracture Toughness: Fracture toughness tests at 4 K have been

conducted on 5083 and 2219 (Tobler and Reed, 1976, Read and Reed, 1976).

Alloys 5083 and 2219 have greater toughness as measured by the J-integral

method at 4 K than at room temperature.

283
Table III A-3.

Specifications for Aluminum Mill Products

P1 rnrl IkJLiUV-'l-' 1 jr t Form 1 Willi Fprlprrl 1 c vj c 1 1 a 1 ASTM ANSI

Qhppf ?ir\(\ nlr^fp v^v^ r\ L. J\j 1 \j H38. 2

Wire, rod, bar, shapes, tube QQ-A-200/4 B221 H38. 5

Forgings and forging stork QQ-A-367 B247 n JO • po

Oiictru uiiu place R?nQ noo

c

InlllC, lUU, UCir, lUllfcrU Ul LUlU R?1 xjC 1

1 1 n JO

nt

) 1 1 1 1 o 1 1 cu

MInlliC, "i K'P v^nrl lUU, Uoi •, pcaLiUUcU Y "h K" 1 irl oH R221 n JO • J Fnminnc: anH fnrninn <;'t'nv"l<' R247 n JO • Qo

6061 Sheet and plate QO-A-250/n B209 H38. 2

Wire, rod, bar; rolled or cold QQ-A-225/8 B211 H38. 4

f i ni shed

Wire, rod, bar, shapes, tube QQ-A-200/8 B221 H38. 5 Structural shapes QQ-A-200/16 B308 H38. 10

Pipe B241 H38. 7 Forgings and forging stork QQ-A-367 B247 H38. 8

Structural pipe and tube; extruded B249 H38. 16

284
Physical Properties: With the exception of magnetic permeability,

the physical properties of aluminum alloys are not as good as those of

stainless steels for structural applications at 4 K. The elastic modulus

is relatively low and thus when used for coil case applications, more of

the load must be carried by the conductor assembly. The thermal and

electrical conductivities are relatively high, resulting in increased

thermal loads particularly from eddy current losses.

d. Fabrication

The relative ease of fabricating large aluminum structures is the

principal advantage of aluminum over stainless steels. Fabrication

characteristics that should receive special consideration for MFE-magnet

systems are briefly reviewed in this section.

Welding: Aluminum alloys 5083, 6061 and to a lesser extent 2219

are readily weldable by the gas metal-arc and gas tungsten-arc processes.

Alloy 5083 retains full strength in the as-welded condition; 5083 welds

have higher strength than either 2219 or 6061 welds unless a post-weld

heat treatment is used.

Heat Treatment: Aluminum alloys are generally heat treated by the

producer unless fabrication requires heavy forming operations or a post

weld heat treatment. Alloy 2219 has exceptionally good response to heat

treatment, sections up to 18 inches thick have been produced that develop

strengths only about 5% below those of 2 to 3-inch-thick sections (Cryogenic Engr. News, 1966). Post-weld heat-treatment, where feasable, can restore strength levels in 2219 welds to nearly 100% of the base metal

values and in 6061 welds to strength levels comparable to 5083.

Machining: Aluminum alloys, particularly the heat treated grades,

have excellent machinabil ity . Machined and bolted assemblies of high

strength aluminum alloys can frequently replace more costly welded

stainless steel assemblies.

Forming: Aluminum alloys have excellent formability in the annealed

or the solution treated conditions.

Forging: Aluminum alloys have the best forgeability of all structural alloys because of their very low strength at forging temperatures

of 370 to 510°C. The size limit on aluminum forgings is significantly

greater than that on stainless steels because of the reduced pressures

required.

Casting: Many aluminum casting alloys should be suitable for

cryogenic applications. However, limited use of castings is anticipated

because wrought products have better properties, can easily be machined

into most shapes and limited quantities of a given component are required.

e. Inspection

Aluminum alloys and their weldments are readily inspectable by

conventional methods.

285
Ill .A.

. Alternative Alloys. Alternative alloy systems do not provide

candidate materials for structural applications at 4 K. Consider the

following systems:

Nickel Base Superalloys - Ferromagnetic at 4 K, expensive, difficult

to fabricate

Nickel Alloys - Ferromagnetic at 4 K, expensive

Titanium Alloys - Brittle at 4 K, expensive, difficult to fabricate

Cobalt Base Superalloys - Transmutes to long half-life reaction

products, expensive, difficult to fabricate, generally ferromagnetic at

4 K Carbon and Alloy Steels - Ferromagnetic, brittle at 4 K Ferritic and Martensitic Stainless Steels - Ferromagnetic, brittle

at 4 K

Copper Alloys - Acceptable, but generally not competitive with

austenitic stainless steels

Magnesium Alloys - Low strength, brittle at 4 K

Further consideration of alloys from any of these systems is not

warranted at the present time.

286
III.B. Structural Composites and Bulk Insulators

III .8. 1

.

Overview. Cost, availability and performance of structural

composite materials useful in MFE cryogenic systems scale in the general

order: industrial laminates (NEMA types or resin-impregnated wood),

specialty fiberglass composites and advanced, high-modulus, composites.

Industrial laminates are commonly referred to by tradenames (Micarta,

Textolite, Phenol ite, Permali, etc) which are a general term for a company product line. Specifications established for filamentaryreinforced polymer industrial laminates (NEMA publication Li-1 (1971),

ASTM D709-72 Part 40, MIL-P and LP) are room-temperature electrical

performance specifications rather than material specifications and are therefore inadequate for cryogenic MFE applications. The industrial

laminates on Table IIIBl are the most likely candidates for cryogenic

MFE use.

Improved radiation resistance is expected from specialty fiberglass

composites using boron-free S-type glass reinforcement in radiationresistant polymers such as aromatic-base epoxies or polyimides. The use of impregnated and partially-cured (B-staged) composite tape would likely

fall into this category.

Polymer-matrix composites reinforced with the higher-modulus fibers

(graphite, boron, Kevlar) can be effectively used at 4 K, offering the

possibility of meeting or exceeding strength, fatigue resistance and

modulus of stainless steel while providing electrical and thermal insulation plus reduced weight.

III.B.

2

. Experience. The use of composites as structural materials in

superconducting magnets has not been extensive. Existing magnets do not

suffer eddy-current losses, and heat leaks through metallic supports has

been accepted, although excessive helium consumption is a common problem. Industrial laminates of the NEMA LE or G-10 types are commonly

used as bulk electrical insulators in the magnet coils with apparent

success, although internal shorts remain a problem. There is no experience

operating a coil in a radiation environment.

Glass-reinforced specialty composite straps and cones are presently

being used as primary load-carrying thermal standoffs (4.2-300 K) in two

designs of a 3000 hp DC superconducting motor, providing substantial

savings in refrigeration cost. The only attempt to restrain superconducting coil magnetic forces by composites was in a lightweight MHD magnet for

airborne use. This magnet failed because the low-modulus glass-epoxy

support permitted excessive superconductor motion. High-modulus fibers

should have been used.

Non-conducting composite helium dewars have been hand-made for

laboratory use. Results indicate that the concept is feasible, major

problems being in construction methods rather than in the fiberglassepoxy structural material. This technology remains in its infancy.

287
Table III B-1

Types and Grades of Industrial Laminated Thermosetting Materials

of Primary Interest in Magnetic Fusion Energy Systems

111

ASTM/NEMA

Designation

NEMA/ASTM LE

NEMA/ASTM G-10

NEMA/ASTM G-11

NEMA/ASTM C

NEMA/ASTM G-5

Wl, W2 EH

MIL-P and LP

Designation

MIL-P-15035, Type FBE

MIL-P-18177, Type GEE

MIL-P-18177, Type GEB

MIL-P-15035, Type FBM

LP 5094

Descri ption

Cellulose fabric fine-weave, phenolic resin

Glass cloth, epoxy resin ''

Glass cloth, high-temperature epoxy resin

Cellulose fabric, medium weave, phenolic resin

Glass cloth, melamine resin

Beechwood veneer, phenolic impregnated

288
Ill .B. 3. Properties. A fairly extensive mechanical and physical data

base exists for industrial laminates, specialty fiberglass composites

and advanced composites at room temperature (manufacturers data sheets,

NEMA publications, MIL-HBK 17 (a), USAF Advanced Composites Design

Guide). However, relatively few efforts have been made to extend these

data to cryogenic temperatures. The current state of knowledge of the

effect of cryogenic temperatures on the mechanical and physical properties of specialty glass and high-modulus composites has recently been

reviewed by Kasen (1975 a-c). Results are encouraging, indicating that

cryogenic temperatures are seldom deleterious to strength properties,

and usually result in substantial improvement. Unfortunately, most of

the available data cannot be generalized; only recently has composite

theory permitted predictions of off-axis composite performance to be

made from relatively simple uniaxial laminate test data. A conspicuous

deficiency is the lack of cryogenic property data on the industrial

laminates, hundreds of metric tons of which are proposed for use in

current MFE reference designs. The effect of radiation at 4 K has not

been studied on any composite materials system.

Ill .8.4 . Specifications, Availability and Cost.

a. Specifications and Quality Assurance

Functional performance specifications for room temperature electrical insulating use of industrial laminates have been established by ASTM,

NEMA and by MIL specifications. These are of limited use for cryogenic

applications. Material specifications must be established for each

constituent of the industrial laminates listed on Table III B-1 unless

it can be shown that a general class of constituents is acceptable for

MFE cryogenic service. Pre-inpregnated tape commonly used in making

specialty fiberglass products require added specifications such as fiber/resin ratio, filament alignment and tape tolerances. The final

composite product requires specifications on reinforcement orientation,

overall fiber/resin ratio, void content and significant property minimums. Presently, each organization planning to use composite materials

prepares its own specifications for internal use. Typical examples are given in Vol. Ill, USAF Advanced Composite Design Guide (1973).

To insure successful application of composite structures it is

necessary to establish a quality assurance program encompassing all

aspects of the design and fabrication of parts manufactured from intermediate products such as prepreg tapes. Unlike metals, quality control

on the composite material itself is inadequate. Quality assurance

procedures are discussed in detail in Vol. Ill, USAF Advanced Composites

Design Guide (1973).

b. Availability and Cost

Neither short-term nor long-term availability of current types of

industrial laminates appears to be a problem and the amounts required

for MFE applications is well within the capability of the industry. However, if it becomes necessary to impose a tight specification on the con- stituent components of a given laminate, availability problems may arise

should a specific component be taken off the market. Current costs

range from about $1.70/lb for NEMA C (cotton-phenolic) to $3. 80/1 b for

NEMA G-11 (glass-epoxy ) . A tight component specification will increase

these costs.

289
Material shortages are not expected to limit the availability of

speciality glass-reinforced composites. However, the tonnage production

of these materials is far less than for industrial laminates. Industrial

S-glass roving costs about $2.25/lb with the preinpregnated tape form

costing $5-6.00/lb. The 1 ime-alumina-borosil icate/type F qlass is some- what less expensive but less radiation resistant.

Ill .B.5. Fabrication. Industrial laminates are produced in monolithic

plate, sheet, block, rod and tube forms. Special channel , tee and zee sections are also available. The material is machined to final shape

following the procedures of metals technology.

A large degree of flexibility in fabrication method is available

with specialty fiberglass composites depending on the final result

desired. Preinpregnated tapes can be used to form complex shapes using

relatively simple molds if they can be obtained in widths appropriate

for jelly-roll winding of conductors or for tension-wrapping of coil

segments. Either cloth or uniaxial fiber orientation is available. Some

large parts, such as dewars, might be more efficiently made by hand layup

of the woven cloth reinforcement and resin, sacrificing some strength

reliability for fewer leaks through joint defects. Large symmetrical

cylinders can be automatically braid-wound. The final cure can be at room temperature for uncritical applications or by vacuum-bagging and

autoclaving for maximum strength and minimum property variability. Virtually any length of fiberglass angles, beams and channels can be con- tinuously produced by pultrusion in which glass roving is pulled through

a resin bath and heated die.

Preimpregnated tapes are used to good advantage in fabricating web

or honeycomb type structures in contrast to monolithic structures of

metals or industrial laminates. This approach can often reduce weight

and produce economies compared to monolithic structures, despite higher

base material costs. Fabrication with composites is reviewed in V III,

USAF Advanced Composite Design Guide (1973); design with composites is

reviewed in V I of that series.

Ill .B.6. Alternative Matrials. High-modulus graphite-reinforced composites could be effectively substituted for many stainless steel structures in present MFE superconducting magnet designs, the primary obstacle

being cost and lack of design experience. The future of these materials

is bright because, unlike metals, graphite fiber costs are dropping

rapidly and future availability will not be limited due to scarcity of

natural resources. While large-scale structural use of such materials

remains in the future, certain components can be identified for which

graphi te-epoxy materials should be presently considered. There are applications where structures fabricated from state-of-the-art, uniaxial,

B-staged graphi te-epoxy tape can be substituted for stainless steel to

obtain an immediate improvement in performance or reduced overall com- ponent cost. Existing systems can provide uniaxial ultimate tensile

strengths in the :range of 69 MPa (100,000 psi ) with moduli from 275-324

GPa (40-47 x 10^ psi) or, alternatively, strengths of about 138 MPa (200

X 103 psi) with moduli of about 137 GPa (20 x 10^ psi) combined with

290
excellent fatigue properties and negligible thermal contraction on cooling. Candidate components would be coil forms, non-metallic loadcarrying helium containers and stiffening filaments interwoven with

the conductors in superconducting cables. Current prices for graphite

fiber produced from rayon or polyacrilonitrile precursors are $30-

40/1 b. A reduction to $8-in/lb is considered realistic as production

increases and as the less expensive pitch precursor systems become

perfected

.

The future for boron-reinforced composites is less favorable due

to the continuing high cost of the fiber $200/lb) and increased

difficulty in fabrication. The high neutron cross section is also a disadvantage in MFE systems. Nevertheless, the high state of development of this system, its excellent cryogenic performance, particularly

its extremely high compressive strength, its insulating qualities and

a thermal expansion closely matching that of stainless steel may

suggest some special applications.

The aramid fiber Kevlar 4Q is a likely replacement for glass in

components where the 90% higher fiber tensile modulus and 57% decrease

in fiber density is advantageous. This relatively low-cost fiber is

already being considered for laboratory helium dewar construction.

High-modulus fiber-reinforced metal composites may have applications where high electrical conductivity must be combined with high

strength and stiffness as, for example, in eddy-current shields for

poloidal magnets. Boron, graphite, or FP (fire polished) aluminareinforced aluminum are candidates. Boron-aluminum is commercially

available and graphite-aluminum composites are in a late state of

development. The FP-aluminum composites are in the initial stages of

development but hold promise for a relatively low-cost product having

a lower density than aluminum but with equal or higher strength and a modulus higher than that of steel. This product has future potential

as a replacement for some of the stainless steel support structures.

291
III.C. Insulating Films, Coatings and Inorganics

II I . C.

. Overview. We consider here the unrelnforced polymers, oxides,

silicates, etc., used primarily as electrical insulators within magnets

but having some uses as thermal standoffs. Polymers may be either

applied in the fluid uncured state followed by curing, as films, or as enamel coatings. Oxides can be indigenous to the material to be insulated (anodized aluminum) or bulk oxides, fired or unfired. Coatings

can also be sprayed onto the conductor or fired on at elevated temperatures. The particular material and method of application will be con- ditioned by the design and by the radiation resistance required. Commonly

used Insulating films, coatings and organics are listed on Table III C-1

III. C. 2. Experience. The organics are favored as they combine strength

and flexibility required for ease in fabrication. To accomodate larger

radiation fluences the more radiation-resistant polymers (amine-cured

epoxy) are impregnated with inorganics (such as AI2O3). When these

become unsatisfactory (> 109 rads), inorganics without organic binders

must be used. Here, the design has to account for the brittleness of

the Inorganics by providing a rigid supporting structure, either by

mechanical clamping or by spraying or firing the inorganic directly onto

the conductor. Performance of the inorganic insulators in radiation

environments has been generally favorable provided sufficient care was given to subsequent assembly so as not to damage the fragile inorganic

material

III.C . 3. Properties. The wide range of properties obtainable by stretching organic films such as PET to obtain differing degrees of crystal

Unity complicates a discussion of mechanical and physical performance.

A recent compilation of published data on PET, PPMI, PPX and PC (Schramm,

et al . , 1973) contains a large amount of room-temperature data, but

comparatively little of cryogenic Interest. It is probable that a similar situation exists for the other materials listed on Table III C-1,

although the authors have not surveyed the relevant literature. It is

evident that studies intended to evaluate the effects of irradiation at

4 K on the properties of polymeric films and coatings must start with

well characterized material, carefully selected to meet all obvious

design parameters. Subsequent attention must be given to the variability

in cryogenic performance created by processing variables.

Experience has thus far Indicated that the radiation flux likely to

exist at the MFE superconducting magnets will not substantially degrade

the properties of the metallic oxides, suggesting that room-temperature

data can be used for design purposes. This assumption deserves verification in critical components.

292
Table III C-1

Commonly Used Insulating Films, Coatings and Inorganics

Generic Name Common Name

Polyethylene Terephthal ate (PET) Mylar

Polypyromel 1 itimide (PPMI) Kapton

Polyvinyl Formal Formvar

Polyamide Nylon

Polyurethane

Polyparaxylylene (PPX) Parylene

Polycarbonate (PC) Lexan

Aluminum Oxide (Al^O^) Alumina

Magnesium Oxide (MgO) Magnesia

Porcelain

Fused Quartz (SiO^)

Mineral Silicate Mica

293
III.D. Conductors

II I . D. 1

Overview. As in the proposed designs described earlier, in

existing superconducting magnet technology the material used most often

is one of the niobium-titanium alloys. Most existing magnets tend to be

quite small and are not cryostatical ly stabilized. This leads to a quench behavior which would be totally unacceptable in a fusion magnet.

The few very large dc magnets presently in operation use heavilystabilized NbTi conductor in very simple configurations and are cooled

by pool boiling. Stainless steel is used for reinforcement. Maximum

magnet current values range from 1.5-2.0 kA. Pulsed superconducting

magnets of a size anywhere near those proposed for MFE do not exist at

present. The few large (by current standards) pulsed magnets have shown

training effects and most have not reached design specifications.

We have purposely ignored the NbsSn superconductor up to this

point. There are a number of programs to procure and evaluate large

conductors 10 kA) containing this material. A few MFE designs con- sider it as an option for the future and, in a couple of cases, it may

be essential to the design. However, with the exception of the ORNL/EPR,

it is not proposed for any of the major projects evaluated in this

report. The major difficulties at present appear to be strain sensitivity, which limits the design strain to < 0.1%, and the problem of

producing very long wires with uniform characterisitics. The material

is also quite expensive. If tape conductor is acceptable, the objections are not as serious. However, the tape configuration may lead to

unacceptable losses in pulsed fields. Some basic research is being done

on strain-insensitive A15 composites but it is very limited for a topic

of such potential importance.

In the following subsections we treat each of the components of the

conductor composite in more detail and end with specific comments on the

composite itself.

III. D.2. Superconductor. Niobium titanium alloys are readily available

with 45, 46.5, 48 and 55% Ti by weight. Currently the 46.5% Ti alloy is

the most popular. Other alloy compositions are available if the order

is for 3000 lbs of material. The group at Oak Ridge has worked on detailed specifications for NbTi. It is available from only two sources

in the U.S., Wah Chang-Teledyne and Kawecki -Beryl co. At present the

cost is $45/1 b in rod form. The mechanical properties of a number of

compositions have been measured recently by Read (1976). The superconducting properties are available for a wide range of alloy compositions

(Roberts 1972, 1974).

The properties of Nb3Sn and the other A15 compounds are highly

dependent on the precise details of their preparation. Exact stoichiometry is not common and rumor has it that the addition of small quantities

of other elements greatly improves the properties. Data are available

on both the mechanical and electrical properties of the compounds, but

they are not of interest to the wire user.

294
III.D.3. Stabilizer. The choice of stabilizing metal involves a number

of concerns. The effect on the electrical and thermal resistivity of

the high magnetic fields, strain and radiation all must be considered.

The only two serious candidate materials are copper and aluminum.

Copper, processed by "oxygen free" techniques, is the most often used

material at present. Residual resistance ratio values* of 100-500 are available at a cost of about $0.75/lb. Very pure copper (RRR ^ 2000) is

available, but costs about $50/lb. Pure aluminum (RRR 3000) is available at ^ $10/lb. These very high costs could easily drop by an order

of magnitude if large quantities were involved.

In the final analysis, of course, the resistivity of the metal in

the operating magnet is the parameter of greatest concern. The trans- verse magnetoresisti vi ty of copper has been measured by several investigators. It appears to obey Kohler's rule which allows prediction of

p(H) from the RRR value (Fickett, 1972). For oxygen free copper at 4 K

this can be written very roughly, as

A (H) - p^(l + 5-).

CU^ ' 0^ ^' The situation is not as clear for aluminum, but it certainly has a much

smaller magnetoresistance at high fields which, for modest purity materials, saturates at a value 2-3 times that at zero field (Fickett 1971).

Surprisingly little information exists on the effect of strain on the

electrical resistance of either metal. Aluminum is the most sensitive

to strain, but if strain levels 0.1% are maintained it will not degrade

significantly.

More work is needed on a variety of prospective materials. Radiation damage causes the resistivity of either Cu or Al to increase quite

rapidly due to production of vacancy-interstitial pairs. A graphical

representation from the UWMAK III report (1976) is shown in Fig. III.D.l.

Again, more work is needed, but it appears that at any reasonable neutron

flux design level, magnets will either need to have large amounts of excess stabilizer or must be annealed more often than desirable. Recovery

of radiation-induced resistivity on annealing at room temperature appears

to be complete for aluminum and about 80% for copper (Thompson, 1969).

The times required are quite short, on the order of minutes. The effect

of radiation on the magnetoresisti ve properties must also be considered.

In the case of copper it seems that the correlation with RRR still

holds, although some recent work by Coltman (1976) at Oak Ridge indicates

that the constant may be somewhat higher than given above. Aluminum is

not as easy to make predictions about because it does not obey Kohler's

rule in general

.

Finally there are the thermal properties. Recent work indicates

that, at least for commercial copper, the magnetic field effects the

thermal resistivity qualitatively the same as it does the electrical

*A common measure of purity for low temperature conductors, the RRR value

is the ratio of the room temperature resistivity to that measured at

T = 4 K. The higher the value, the more pure the metal.

295
DAMAGE LEVEL (dpa)

Figure III.D.1. Radiation induced resistivity of copper and aluminum.

From UWriAK III.

296
(Sparks and Fickett 1975). No data exist for aluminum. The bond between

the stabilizer and both the superconductor and the strengthener is also

a potential problem area for heat transfer, particularly for aluminum.

Ill . D.4 . Strengthener. This component is not always present in the

conductor, particularly in the case where the conductor is embedded in a disc or some such structural member. Where it is used, it is nearly

always stainless steel strip or wire. The cost probably will be ^ $2/lb

when the fabrication is considered. Most of the necessary material properties are available. The possibility of neutron radiation inducing

local magnetic regions (Stanley and Garr, 1975), while remote at the

anticipated radiation levels, should be considered because of the unknown

effect of magnetic field and the proximity to the superconductor. The

particular function of the steel is to keep the conductor strain to an acceptable level, usually < 0.1%. To this end, various clever cabling

techniques have been described.

The possibility of replacing the steel -stabil izer system with a single component, usually cold worked copper has been proposed (Wilson

and Walters, 1976). Such an arrangement might be able to decrease eddy

current losses in the magnet considerably.

III.D .5. Other Materials. As long as NbTi is the conductor choice, the

only other material involved in the conductor will be the eddy-current

shielding barriers which are always proposed to be copper-nickel alloy.

In tube form the alloy costs about the same as stainless steel. Its

electrical and thermal properties at zero magnetic field have been

measured. Magnetoresisti ve effects have not been studied in any detail,

nor has the magnetothermal conductivity. The alloy has already been

used successfully in conductors for a number of pulsed magnets (Coupland,

1976).

In Nb3Sn wire conductors the alloy CuSn is an essential component

as it provides tin for the reaction of the niobium filaments to the

intermetal 1 ic. In existing conductors, alloys containing from 8-13% Sn are used. After reaction, ^ 0.3% Sn remains in the bronze. The bronze

is not a good enough conductor to perform a stabilizing function. In

order to separate the bronze from the pure copper stabilizer a barrier

of tantalum is used. This prevents contamination of the stabilizer with

tin from the bronze.

III. D.6 . Conductor Composite. The materials described above are com- bined in the composite conductor. Existing commercial superconductors

tend to have critical currents of several hundred to a thousand amperes

and to be several millimeters in dimension. It is difficult to state

prices with any certainty, but NbTi conductor tends to be $0.10/ft and

Nb2Sn wire $l-2/ft. Superconducting wire and tape is produced by only

a few companies in this country. Most processes are highly proprietary

and, thus, any properties except those specified by the buyer usually

must be determined by measurement.

297
There have been many successful applications of NbTi superconductor

in relatively large magnets. The same is not true of Nb3Sn, particularly

in wire form. Quench initiation and the training phenomenon in medium

and small size magnets is still not understood. Much of the problem

may lie within the conductor itself.

Detailed specifications for some superconducting wire have been

made, but buyers generally specify only critical current at field. NbTi

wire is quite readily available, but the industry is small and very

large orders could cause delays. NbsSn, while available, is still

produced in relatively limited quantities.

Properties data on the composite conductors is being acquired slowly.

Work has been done on the effect of stress on critical current (Ekin and

Clark, 1976; Easton and Koch, 1977), the effect of fatigue is the subject

of the current NBS program. Neutron irradiation experiments (Guess, et al

1975) have indicated that NbTi critical current is degraded by less than

10% under fluences up to ^ 10l8 n/cm2. Annealing at near room temperature

repairs the damage. NbsSn shows an improvement in properties to a fluence

of 10^8 n/cm2. More measurements are needed on all of these topics as the available data are often conflicting. The general area of ac losses

in superconducting composites is a complex subject but one which has

been investigated in some detail. Brechna (1973) presents a description

of the various loss mechanisms and some sample calculations.

Complete fabrication of these materials is nearly always performed by the manufacturer. The joining of conductors by the user is

often a problem, particularly where very high currents are involved.

The inspection or NDE of wire is an area where little is now done and one in which work is needed.

298
IV. RECOMMENDATIONS

In this Section our recommendations are presented for an effective program of materials research and technology transfer to insure

economical construction and reliable, safe operation of low-temperature

components of magnetic fusion energy systems. The program objective

is stated and the entire scope of required materials research from

Section II and the candidate materials from Section III are summarized.

Finally the recommended program is discussed. To assemble this, we have performed exhaustive analyses of present designs, visited every

principal ERDA-MFE group concerned with superconducting magnets,

summarized available materials technology and drawn extensively on our

own experience in low temperature materials research.

Our present recommendations are a direct reflection of the materials problems and design constraints defined in the MFE documents

referenced in this report. We recognize that progress in understanding

the physics of fusion power may result in new MFE designs which will

require a shift in emphasis in the materials research program. The

present program is designed to be responsive to such changes.

There are some fundamental physics decisions which affect, con- siderably, the thrust of a low temperature materials research program.

For instance, if the smaller prototype reactors continue to use relatively fast rise times for the ramping cycle of the ohmic heating

coils, then low conductivity materials will be almost certainly be

required to reduce eddy-current heating losses. Cheaper cryostat

materials, such as 5083 aluminum, will not be used. Secondly, if the

ohmic heating coils are placed immediately adjacent to, or within, the

central support cylinder, it may be necessary to use a non-conductor

material for the support cylinder, again if the rise times are fast.

It is likely that a major influence on shield design and system warmup will be from two material considerations: (1) the degradation and

Wigner storage energy characteristics of the electrical insulators and

(2) the radiation-enhanced resistivity and subsequent recovery at room- temperature of the purity-cold work dependent resistivity at 4 K of OFHC copper. Additionally, ultimate reactor design may be directly

influenced by the maximum allowable design stress of the structural

magnet coil form and casing: higher allowable design stresses will

permit less material, thus conserving working space.

We have attempted to consider the impact of ongoing research programs in the formulation of the recommended program. Our compilation

of these programs is presented in Appendix C. The potential overlap

of these projects is pointed out in the discussion of the individual

projects.

299
IV. A. Program Objective

To assist the design, construction, and safe operation of lowtemperature magnetic fusion energy systems, especially superconducting

magnets, through effective materials research and materials technology

transfer

IV. B. MFE Research Requirements

A survey of material requirements for all of the proposed MFE

magnet systems has been performed. Design analyses, discussions and

survey trips regarding all concerned laboratories (ANL, BNL, LASL,

LLL, ORNL, GA, U. Wise, MIT, PPPL) have been summarized. Working

data sheets for each major component of each major project are all

included in Appendix A and are summarized in Section II. C. 2. Table II C-1 provides the information needs to which the research program must

be responsive. Since it is convenient to address research programs in

terms of material categories, as opposed to system components, it is

more descriptive to rearrange these tables into research needs for the

major material and material development groups. This is presented in

Table IV B-1

The major material technology categories which have been identified

are:

(1 ) structural al 1 oys

(2) thermal insulators

(3) electrical insulators, films and coatings

(^) structural composites and bulk insulators

(5) conductors (non-superconducting)

(fS) adhesi ves

Some general considerations regarding applications of these

material groups are given in the next Section.

Specific materials within each category are listed in subsection,

IV D.

Perusal of Table IV B-1 will provide an accurate assessment of the

research required for proper design and material selection of superconducting magnets for MFE.

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301
IV. C. Summary of Materials Considerations

Presented here are some general conclusions which we have reached

following our review of the existing MFE designs and the environmental

conditions as outlined in the last several sections. Specific materials and recommendations are considered in IV. D and IV. E.

IV.C.l

Structural Alloys. Radiation effects are not serious.

Alloy requirements vary depending on the rise time of the pulsed

(tokamak) fields. Smaller devices (TNS) will require steels while

larger power plants will be able to make use of aluminum alloys. The

allowable operating stress is important in determining space available;

it is thus desirable to design with 4 K data. Regardless of the material, thick section welding and inspection techniques are an area of

prime importance for study. The relatively inexpensive ferritic steels

are conceivable as structural materials. The additional magnetic

forces must then be considered in the design. In some instances con- crete may be an acceptable, economic replacement for steels.

IV. C. 2. Thermal Insulators. These materials, primarily aluminized

polymers, may be very sensitive to mechanical and thermal degradation

by radiation.

IV .0.3. Insulating Films, Coatings and Inorganics. An effort to screen these materials as to radiation resistance, particularly to

gamma rays, at low temperatures and on repeated cycling to room temperature is urgently needed. Once this is done, the low temperature

physical and mechanical properties of the most resistant materials must

be measured. There is little low temperature data on these materials.

IV. C. 4. Structural Composites and Bulk Insulators. Again, radiation

damage effects are likely to be of prime importance in organic matrix

composites and screening measurements to access the severity of the

problem are urgently needed as input to this program. For most of

these materials there is no low temperature data base. Specifications

for industrial laminates (NEMA/ASTM) are generally electrical in nature

and mechanical specifications may be necessary additions for low temperature applications. The advanced (high-modulus) composites and

speciality fiberglass composites may be needed for low temperature structural applications near pulsed coils. Very few 4 K data exist.

IV. C. 5. Conductors. Degradation of the low temperature conductivity of

the stabilizing metal by neutron radiation will be a serious problem

for economical power systems because of the loss of operating time

involved in warming to room temperature to remove the damage. In nearer

term systems, the effect of purity, processing and radiation on the

stabilizing function of copper and aluminum needs investigation.

302
IV. C. 6. Adhesives. The data base on the electrical, thermal and mechanical properties of solders is surprisingly small. A similar situation

exists for epoxies used as adhesives.

IV. C. 7. Demountable Coils. Problems of tokamak construction and maintenance would be greatly simplified if demountable joints could be made

in large superconducting magnets. We recognize that the task seems nearly impossible, but the possible advantages indicate that the concept

needs some investigation.

303
IV. D. Specific Materials

This section briefly summarizes the specific materials which are recommended for consideration in this program.

Structural Alloys

AISI 304 and 316

21/6/9

5083, 2219, and 6061

stainless steels which are extremely tough

at 4 K, relatively easily welded, high

modulus, AISI 304 is relatively low cost

nitrogen-strengthened stainless steel having

much higher yield strength but less toughness, high modulus

aluminum alloys have greater likelihood

of use in slow rise-time reactors, lower

cost (5083 is cheapest of all alloys in

this list), lower modulus and yield strength,

5083 is easily welded

Possible candidates for high-strength alloys:

18/3/8

AISI 304N

22/13/5

Thermal Insulators

Superinsul ation

Reinforced polyurethene foam

nitrogen-strengthened steel that is cheaper

than 21/6/9, most properties at 4 K unknown

nitrogen-strengthened AISI 304, has considerably less yield strength, but greater

toughness than 21/6/9

nitrogen-strengthened steel that has comparable

strength and more stability compared to 21/6/9,

properties at 4 K unknown

(1) aluminized mylar sheet (2) hollow glass

beads, aluminized coating

(20-300 K)

Insulating Films, Coatings and Inorganics

Polyethylene Terephthalate

Polypyromel 1 i timide

Polyvinyl formal

Polyamide

Polyurethene

Polyparaxylylene

Polycarbonate

Aluminum Oxide (Al20q)

Magnesium Oxide (MgOj

Porcelain

Fused Quartz (Si02)

Mineral Silicate

(Myl ar

(Kapton)

(Formvar)

(Nylon)

(Paryl ene)

(Lexan)

Al umina

Magnesia

Mica

304
structural Composites and Bulk Insulators

Industrial Laminates

NEMA/ASTM Grade LE

NEMA/ASTM Grade G-5

LP Spec. Grade Wl , W2

NEMA/ASTM Grade G-10

NEMA/ASTM Grade G-11

NEMA/ASTM Grade C

cellulose fabric, fine weave, phenolic resin

glass cloth, epoxy resin

glass cloth, high-temperature epoxy resin

cellulose fabric, medium weave, phenolic resin

glass cloth, melamine resin

beechwood veneer, phenolic impregnated

Specialty Fiberglass Composites

S glass/aromatic base epoxy

S glass/polyimide

Advanced (High-Modulus Composites)

Graphite, Boron, Kevlar 49/aromatic-base epoxy, polyimide

Graphite, Boron, FP-alumina, Stainless Steel, tungsten/aluminum, copper

Conductors

OFHC Copper stabilizing material

Cold worked OFHC Copper stabilizing material

ETP Copper room temperature magnet-coil material

CuSn Alloys partial stabilizing material (Nb^Sn)

CuNi Alloys eddy-current shield material

Al alternate stabilizing and shield material

Adhesi ves

SnPb solders

Epoxies

common conductor adhesive agent common potting agent or adhesive agent

in magnet winding

305
IV. E. Program

The recommended program is outlined below. Ongoing research sum- marized in Appendix C and available research facilities (also in Appendix

C) have been considered. The initial three years of the program are devoted to supplying information for TNS-type structural design and

construction. The following two years concentrate on TNS follow-up

research and on the anticipated needs of future years. Structural

alloys, stainless steels are emphasized. Research on structural

alloys will insure structural safety and present the opportunity for

significant cost and space savings. Aluminum alloy research is delayed

for application after TNS; TNS devices presently could better use a low resistivity alloy. This emphasis will be pursued unless there are major changes in the MFE magnet requirements or unless present insulators become unuseable at low levels of radiation, requiring insulator

development.

This program has been developed on the premise that radiation

research, very important for organic materials, be conducted by the

National Laboratories under separate funding. Other projects (12-14) are listed for which funding is recommended, but which fall outside

the monetary limitations of the present budget.

PROJECTS SCHEDULE

1/yr 2/yr 3/yr 4/yr 5/yr

1. Screening of Nitrogen-Strengthened X -- — -- --I Stainless Steels

2. Base Data of Selected Materials

Structural alloys X X X X X Thermal insulators

Electrical insulators, films, and -- XX X X coatings

Structural composites and bulk -- X X X X insulators

3. Thick-section Welding XXX —

4.

5. Technology

Radiation Effects

Transfer

***** X X X X X 6. Project Management X X XX X 7. Material Procurement Specifications -- -- X -- X 8. Screening, Characterization of Concrete -- -- -- X X 9. Applications of Advanced Composites -- -- -- X X 10. Inspection Standards -- -- -- X X 11. Design Data for Codes -- -- -- -- X

Other Possible Projects

12. Heat

steel

Product-Form

s Variability (stainless XXX

)

13. Copper

Stabilizing

Processing

and Strengthening

and Purity Effects

Material

for XXX

(with INCRA)

14. Copper Specifications for Copper Magnets XX--

(with CDA)

Performed by National Laboratories, additional funding will be required

for this program.

306
Proposed Complementary Research

Ferritic alloy development (U. Cal -Berkel ey ) . This program, now initiated

on Fe-Mn alloy systems by Prof. Morris, is of concern to future MFE

systems in that it may result in the development of a less expensive ferromagnetic, ferritic alloy.

High strength/high conductivity composite development (U. Wise . -Madi son )

.

Under the guidance of Prof. Boom, the development of a high strength, high

conductivity material could produce long-range material advances. Processing equipment, aluminum experience, and an explosive bonding of aluminum

to stainless program indicate that the ingredients are available for significant progress.

Quality requirements for 5083 weldments and thick-section weld development (ALCOA)

The following programs are not funded, but are suggested as possible areas of fundamental research that should, ultimately, impact on materials technology for MFE superconducting systems.

Radiation effects in polymers (fundamental studies - ERDA)

Materials research related to joining (fundamental studies - ERDA)

Effects of magnetic fields on low temperature deformation (fundamental

studies - ERDA)

Austenite phase stability as influenced by low temperature, deformation

and magnetic fields (fundamental studies - ERDA)

There are projects at the National Laboratories and at NBS which

complement this program; these are listed in Appendix C.

307
Project 1: Screening of Nitrogen-Strengthened Stainless Steel

Evaluation of available high-strength structural alloys for potential use in large superconducting-magnet structures has led to the

selection of the stainless steel 21/6/9 (21% Cr, 6% Ni , 9% Mn, Fe

balance) to fill this role. However, there are other nitrogen-strengthened,

austenitic steels which have some advantages over 21/6/9. Alloy 18/3/12

(Nitronic 33) is cheaper, while alloys 22/13/5 (Nitronic 50) and AISI

304N are considered to have much better fracture toughness at 4 K and

the alloy 22/13/5 is expected to be completely stable with respect to

the martensite transformation (therefore non-magnetic). All materials

have higher yield strengths than AISI 304 or 316.

Therefore, the purpose of this screening program is to compare the

4 K mechanical property behavior of these three alloys with alloy 21/6/9.

From this comparison, and accounting for material cost, availability,

and product forms, a program decision will be made to discontinue testing these alloys or to include one of the alloys in the structural alloy

group.

I. Conduct tensile and J-integral fracture toughness tests at 4 K on the three alloys 18/3/12, 22/13/5, and AISI 304N.

II. Evaluate these alloys with respect to the 21/6/9 alloy.

308
Project 2: Base Data of Low Temperature Materials

One primary program objective must be to insure that reliable,

relevant data are available for the major construction materials used

in superconducting magnets. This program is designed to supply this

information for a variety of materials. These data will permit efficient design and careful material selection. They will be available

as a standard reference source for RFP preparation and evaluation.

I. Compile from the recent "Handbook of Materials for Superconducting Applications," from the ARPA-sponsored materials research at 4

K, and from other recent investigations the relevant properties of the

selected materials for use in MFE superconducting magnets. Some

suggested properties and material categories are presented in the

accompanying Table IV E-1 . Prepare, for inclusion in the new MFE-Low

Temperature Handbook, graphs of typical property values versus temperature.

II. Measure, when not available from I, the properties required

as identified in Table IV E-1. The major, significant exclusion from

this program are radiation effects; these are included in Project 4.

To obtain a perspective on structural alloy material variability, the

properties should include several product forms in the annealed condition and weldments (mechanical and magnetic properties only). For

structural composites data must be available as a function of orientation. Specifically for composites we recommend determining the mechanical and thermal properties of cellulose-phenolic and glass-epoxy

types of industrial laminates and of S glass-aromatic epoxy and S

glass-polyimide specialty fiberglass composites. No data are presently

available. Data from these measurements should be included in the new Handbook. Specifically for structural alloys we recommend that

fatigue tests (21/6/9, AISI 304 and 316), multiple product and heat

tensile tests (AISI 304, 316), tensile tests, fracture toughness tests

(21/6/9) be conducted. In alloy 21/6/9 the fracture toughness is

marginal compared to AISI 304 and 316; therefore more tests are required to reliably describe the fracture behavior.

309
Table IV. E-1 . Test flaterial Matrix for Base Data Program

PROPERTIES MATERIALS

Structural Structural composites Thermal Insulating films

Alloys & bulk insulators insulators coatings & inorganii

Mechani ca

Elastic

Tensi 1 e Coiiipressi op

Shea r Fat i que

Fracture tougnness

Fatigue crack growth rate

X?

X1

T

"X11

y

XX

X

X

Thermal

Thermal conductivity

Thermal expansion

'Nnprifir hpflt

X

X

X A

X

Electrical

Resistivity

Magneto-resistivity

Dielectric

X s X

Magnetic

B-H curves ®

Permeabi 1 i ty to He X

(^' - includes weldments

X , - function of fiber orientation

310
Project 3: Thick-Section Welding

The construction of large superconducting magnets involves the

accomodation of large, field-induced loads; thick-sections of structural

alloys are required (Section II. A. 2). Significant unknowns are the

effect of delta ferrite (which probably must be present to prevent hot

cracking in stainless steel thick-section welds) on weld fracture

behavior at 4 K, the variability of design properties at 4 K which thicksection processing introduces and the degree to which specifications

must be controlled to prevent erratic property behavior at 4 K. This

information should be obtained on AISI 304, 316 and the stainless steel

21/6/9; the alloys presently thought useful for TNS type system construction.

I. Conduct fracture toughness tests at 4 K on restrained, thicksection butt welds (1-6 inches thick) of AISI 304, 316 and 21/6/9. The

amount of ferrite must be varied, to obtain the dependence of fracture

toughness on ferrite concentration. Several weld processes should be

evaluated, J-integral fracture toughness procedures will probably be

required.

II. Assess tensile property (4 K) variations as a function of

thick-section cross-section, in longitudinal and transverse directions,

for AISI 304, 316 and 21/6/9 stainless steel.

III. Conduct tensile and/or fracture toughness tests at 4 K as a function of time/temperature of holding in the temperature range 900-

1500°F to assess severity of solute grain boundry segregation (sensitization) on low temperature mechanical properties.

IV. Develop a higher strength weld consumable for use with

nitrogen-strengthened 21/6/9 stainless steel or similar alloys.

LLL (Wi therspoon , Landon) is considering thick-section welding of

21/6/9 for MX construction. HEDL has performed ferrite-related research

on AISI 304 and 316. Rennselaer Polytechnic Institute (Nippes, Savage)

have a program on ferrite elimination in stainless steel welds. Rockwell Intl. (Rocky Flats Facility, Liby) is planning to evaluate the

influence of composition variations on the 6-ferrite content of 21/6/9

welds.

311
Project 4: Radiation Effects

It is necessary to examine the effects of neutron, beta and gamma

radiations at 4 K on material properties at 4 K and on material properties after subsequent warming to room temperature. Little information on material other than the noble metals now is available. Especially

it is important to assess very carefully the radiation effects on organic

materials which are thought necessary for use in the magnet system, as there is concern regarding their usefulness following radiation at 4 K and warming to room temperature.

This program is best carried out at the National Laboratories (ANL,

BNL, ORNL, LLL), where facilities now exist and are available, to

produce useful neutron spectra and fluences and to test at cryogenic

temperatures prior to recovery. As discussed earlier (Section II. A. 5)

in this report it is thought that fission spectra are useable, especially

for material screening testing. Furthermore, comparison of data with

the LLL rotating target 14 Mev neutron source will probably allow the use of a fission spectra (for superconducting-magnet system materials)

in the characterization phase of this program.

Phase I - Screening

A. Screen industrial laminates and films, thermal insulators,

epoxies, and glass-reinforced plastics by irradiation at 4 K (seeSection IV. C. for complete listing of suggested materials);

followed by (1) warming to room temperature and observation

of physical condition and (2) flexture (or tensile) strength at

4 K before and after irradiation and after thermal cycling to

300 K. At present there is one program initiated in this area

at ORNL (Long), however the candidate lists are long and mechan-

' ical property testing at 4 K is desirable, leaving ample research

and persisting urgency for another program.

Phase II - Base Data of Selected Materials

A. Measure effect of radiation dosage and thermal cycling to

300 K on thermal conductivity of selected thermal insulator(s)

B. Measure effect of radiation dosage and thermal cycling to

300 K on tensile, compressive and dielectric strength of

selected industrial laminates specialty fiberglass composites

and films (electrical insulators).

C. Measure effect of radiation dosage and thermal cycling to

300 K on magnetoresistance of selected conductor materials.

D. Confirmation of structural alloy performance by tensile tests

at 4 K after irradiation at 4 K of selected alloys in this

program.

312
Project 5: Technology Transfer

I. A new handbook, MFE Low Temperature Handbook, should be

started. It is recommended that the Handbook include both material

properties and relevant helium fluid and heat transfer properties. The

materials properties will be prepared from six principal sources during the first three years:

(1) the Handbook on Materials for Superconducting Machinery

(2) the recent ARPA-sponsored program on 4 K materials research

(3) the Base Data of Low Temperature Materials Program recom- mended in this report

(4) the Radiation Effects Program recommended in this report

(5) the Structural Alloy Thick-Section Welding Program recom- mended in this report and

(6) other ongoing research programs at 4 K throughout the world.

There should be the following subdivisions of the materials section

of the new Handbook:

(1) Base Data - figures of significant properties versus temperature

(2) Specifications - any material procurement specifications developed

for MFE will be included in a separate section

(3) Other Specifications and Allowable Stresses - current ASTM, NEMA,

etc. specifications and current code stress allowables will be

included in this section

(4) Design Data - all available data will be presented in this section,

to establish, eventually, design allowable values.

II. Once a year a Workshop will be held to accelerate communication

between the participants of the program recommended in this report and

the National Laboratory MFE project personnel. Proper programming will

insure this to be worthwhile to the ML personnel and encouragement of

feedback will permit program reeval uation

.

III. Twice a year project reports should be assembled, printed and

distributed. These reports could include contributions from Project 4 as well as from the complementary programs.

IV. Consultation, either from individuals or from program teams,

must be available to advise or assist RFP writing and evaluation, project

material selection, design troubleshooting or failure analysis, and in

specification writing and code development.

Project 6: Project Management

The program, although small, involves an inordinate amount of

interfacing, limited subcontracting with attendant project monitoring,

and semiannual report editing and assembly. A small amount of project

management funds is required for these functions.

313
Project 7: Material Procurement Specifications

Projects 1-4 will result in a considerable amount of information

regarding product performance from a material property vantage point.

The project must translate this information to specifications and,

additionally, conduct more property tests, if needed, to adequately

define the material product for purchasing. The objective of this

project is to produce material procurement specifications when applicable and significant to insure adequate low temperature performance.

All material classes will be considered and these are some obvious examples:

(1) Weld rod specifications for proper ferrite and hot-cracking

control may have to be rigidly controlled. Project 3 results, manu- facturer consultation, and, perhaps, additional rod sampling may have

to be performed.

(2) NEMA Grade G-10 and G-11 specifications were developed for room temperature insulating performance, and cover a wide variety of

glass fiber/epoxy lay-ups. Projects 2 and 4 may indicate that additional control is required to insure adequate 4 K performance. If so,

subsequent sampling and testing at 4 K will be required to identify

closely the parameters to be controlled (specified).

(3) Thick-section processing and chemistry control of alloy 21/6/9

will probably be identified from Project 3 data. Subsequent control,

through material specifications, will then be required. Consultation

with suppliers and knowledge of their capabilities, consideration of

time, costs and product forms will result in new specifications for

purchase.

(4) It is quite possible, if f iberglass/epoxy dewars are con- structed, that Project 4 results will imply better performance of

specific fiber or epoxy mixes. Additional specimens may be required

to compare these component combinations and lay-ups with others. These

tests would be followed by material specifications for non-metallic

dewars

LLL (Landon) and Rockwell Intl. (Rocky Flats, Liby) have previously

considered 21/6/9 procurement for low temperature use.

314
Project 8: Screening, Characterization of Concrete

For magnet systems following TNS, it is conceivable that one of

the better support materials will be reinforced concrete, especially for

the tokamak central support cylinder, for Yin-Yang central support

(already recommended by Bechtel for this application) and for tokamak

support columns for the gravitational weight of the TF coils. Major

advantages are low cost and maintenance with good compressive strength,

and low thermal conductivity and good radiation resistance. Disadvantages are lack of mobility and the probable requirements of thick

cross-sectional areas, using valuable space. Preliminary research on concrete to 76 K has been performed by the University of Wisconsin.

I. Screen various mixtures (aggregate type, water content) for

tensile and compressive properties at 4 K. Assess increase of properties at 4 K with post-tensioned steel.

II. Measure elastic, tensile, compressive and fatigue properties

and thermal conductivity of selected concrete composites at 4 K.

315
Project 9: Application of Advanced Composites

Seven areas may be defined where advanced, high-modulus composites offer the potential of increased performance, comparable cost

and reduced consumption of scarce resources:

(1) Poloidal Coil Dewars and Structural Helium Containers. Reduction of ac losses suggest the use of non-metallic composites for at

least the inner wall of the OH poloidal coil dewars and for load-bearing

helium containers closely surrounding bundles of OH coil poloidal field

conductors. Here, stiffness may dictate the use of high-modulus composites.

(2) Coil Forms. Reduction of ac losses promotes the use of non- metallic composites for the OH poloidal magnet coil form. Again, stiffness may require the use of advanced composites.

(.3) Nonmetallic Strengthener within Superconducting Cables.

Replacement of stainless steel internal conductor reinforcement with an advanced high-modulus composite can provide higher strength, higher

modulus, lighter weight and a reduced probability of electrical shorts.

This application applies to both solid reinforcement and to filamentreinforcement of braided conductors.

(4) External Reinforcement of Superconducting Coils. Replacing

stainless steel tension bands with a high-modulus composite band can provide higher strength, lighter weight, a reduced probability of shorts

and facilitate field erection of very large coils, and may act as an electrical insul a tor- strengthener

(5) High-Modulus Superconductor Stabilizing Material. Reinforcing

high-purity copper or aluminum with high-modulus fibers may substantially

stiffen the stabilizing material without significant degradation of its

electrical conductivity, thereby reducing strain limitation imposed by

the superconductor.

{f>) Eddy-Current Shields. Replacing bonded stainless-copper with

aluminum or copper reinforced with a high-modulus fiber can substantially

increase the strength, stiffness and fatigue resistance of eddy-current

shields in poloidal -coil systems where the shield is subjected to pulsating electromagnetic stress.

The data base required for the use of industrial laminates and

specialty fiberglass composites in thermal standoffs, poloidal coil

dewars and coil forms is encompassed in the recommended base data program. This should be followed by prototype dewar construction to determine the least expensive method of fabricating very large compositemetal dewars.

The use of high-modulus composites as load-carrying helium con- tainers or for internal or external superconductor reinforcement requires

a screening program to select the optimum high-modulus composite on the

basis of uniaxial fatigue performance under radiation.

Initial studies on fiber-reinforced stabilizing materials and eddycurrent shields need only exploit existing state-of-the-art metal -matrix

technology to demonstrate feasibility of reinforcing high-purity metals.

Favorable results should be followed by prototype testing.

316
Project 10: Inspection Standards

Subsequent to research primarily directed toward TNS design and

material selection and procurement, inspection becomes significant.

Weld quality standards, based on ferrite and defect control, will be

needed; measurement technique and quantitative standards pertaining to

ferrite and defects will be essential to insure reliable weld joint

performance. Additionally, multi-pass stainless steel welds may result

in some sigma phase formation; requirements may be needed for this

phase concentration. Secondly, if non-metallic dewars are constructed,

standards and measurement techniques must be stipulated to insure

adequate composite properties at 4 K. Composite-to-metal joints will

have to be inspected and/or tested. Thirdly, in prototypes extensive

strain-gaging is required to compare material performance with design

expectations. The magnetoresi stance contribution to the strain-induced

resistance change of the strain gages must be known. Currently, there

is some research on this problem at ILL.

Therefore, the major program thrusts should be:

I. Weld inspection measurement techniques and standards for

ferrite defects, and, possibly, sigma phase for stainless steels.

II. Inspection measurements and standards for porosity or void

content, fiber/resin ratios, fiber orientation, and interlayer debonding

in non-metallic composites - if used for dewar construction.

III. Weld inspection measurement techniques and standards for

defects in aluminum alloys - if used.

317
Project 11: Design Data for Codes

At some stage in the development of MFE, codes are inevitable.

Perhaps, even at the TNS stage very careful and accurate assessment

of 4 K (or 300 K) properties used in design that are critical to

structure safety (such as the tensile yield stress) will be required.

Furthermore, as is pointed out in the variability discussion, permission for use of 4 K property data in design will surely be contingent on more-than adequate characterization. The structural alloy

portion of the program concentrates on the subject of material and property variability, but a complete answer to fully characterize this

subject is quite expensive. Heat and product form variability studies,

such as performed on the LMFBR program for AISI 316 and 304 at room

and elevated temperatures, demands the acquisition of multiple heats,

the processing into several product forms, close control and characteri

zation of chemistry, and extensive testing.

Therefore, the purpose of this program is to continue variability

studies on structural alloys and/or other materials that are (1)

anticipated to fall under regulatory codes or (2) that careful documentation of 4 K properties will permit the use of those properties in

design - thus saving material cost, permitting the use of less material

and conserving open space.

Actual program plans must be negotiated with concerned regulatory

bodies or with project personnel concerned about decisions pertaining

to operating stress levels.

318
Project 12: Heat, Product Form Variability

Thorough documentation of low temperature properties relevant to

design, especially the tensile yield strength, should provide sufficient

incentive to permit the use of low temperature design allowables. The

use of low temperature property data will be of tremendous assistance in

meeting space requirements. For instance, the permission to have the

design allowable based on a 4 K yield strength of 140,000 psi , as opposed

to the room temperature strength of 55,000 psi, will result in substantial material reduction. Both cost and space will be conserved. Therefore, there is ample incentive to characterize, reliably, the extent of

the variability of low temperature properties, as influenced by heat,

product form and chemistry.

However, as extensive tests involving a number of material heats

and/or product forms are needed, the program is very expensive. Some of

the needed variability data will accrue from Projects 1-3; however, it is

estimated that for complete examination of these factors, additional

funds will be required.

I. Tensile tests at ^ K on variations of 12 heat/product forms

per alloy (three alloys, transverse, short transverse and longitudinal

for plate, 4 specimens each).

II. Compression and shear tests at ^ K on variation of 3 heat/

product forms per alloy (three alloys, longitudinal for sheet and

plate, 2 specimens each).

III. The product performance variability of industrial non-metallic

laminates must be determined whenever such variability impacts on the

laminate application.

ORNL (Brinkman) and HEDL on the LMFBR program have already greatly

assisted this project by their measurement and compilation of room temperature tensile properties for AISI 304 and AISI 316 stainless

steel s.

319
Project 13: Low Temperature Physical Properties of Copper

and Copper Alloys for Conductor Applications

This program is anticipated as a cooperative effort with the International Copper Research Association (INCRA).

Large amounts of copper and copper alloys are proposed for use in

current MFE magnet designs as a stabilizer for the superconductor and

as a strengthening element. It is essential that the role of purity of

OFHC copper (RRR % 200-500) be identified in affecting recovery at room temperature and the final RRR after conductor processing and after

irradiation at 4 K. Knowledge of these variables should result in the

determination of magnet warmup schedules subsequent to irradiation at

4 K and may result in specification of a minimum purity of OFHC copper.

Secondly, for pulsed magnets the effects of mechanical fatigue on magnetoresi sti vi ty , as a function of purity and radiation, must be

understood. NBS (Clark, Ekin) is currently conducting a limited fatigue

program, but the results are likely to apply only to small conductors

and to Cu/NbTi composites.

Project 14: Copper Specifications for Copper Magnets

This program is anticipated as a cooperative program with the

Copper Development Association (CDA).

Recent experience of several manufacturers of large copper magnets

indicates that existing product specifications accepted by the copper

industry are not adequate when the product is to be used in high current

magnet applications such as those required by the controlled thermonuclear research devices now under construction. As normal -metal copper

magnets will continue to contribute to the MFE program, it is expected

that this demand for copper will increase in the near future. Significant economy could be achieved if available large copper plate products were compatible with magnet design requirements.

This program would rely on a strong cooperative effort between the

industry, the magnet designers and builders and appropriate research

laboratories to arrive at a mutually-acceptable solution to the problem.

Its objective would be to set specifications for thick-section copper

plate for MFE magnets that were compatible with both MFE requirements

and producer capabilities. Inspection procedures would also be formulated.

320
V. REFERENCES

The references listed here are, for the most part, those used

in this materials review in one way or another. The first group consists of the laboratory design reports and closely related documents.

These are listed under the appropriate laboratory. Next we have included

a list of general references on fusion. Specific materials references

are listed next with a separate listing of those specifically treating

radiation effects. In Section III of the report specific references are given in the format (Jones, 1976). These refer to the two materials

sections in this listing.

Argonne National Laboratory (ANL)

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(1976) W. M. Stacey, Jr., et al

.

"A Tokamak Experimental Power Reactor," ANL/CTR/TM-47 (1975) W. M. Stacey, Jr. , et al

.

"Tokamak Experimental Power Reactor Conceptual Design," ANL/CTR-76-3,

Vol. 1 and 2 (1976) W. M. Stacey, Jr., et al.

"Tokamak Experimental Power Reactor Studies," ANL/CTR-75-2 (1975)

W. M. Stacey, Jr . , et al

.

Brookhaven National Laboratory (BNL)

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.

"Program Plan for CTR Magnet Safety Studies," BNL 19620 (1975)

G. Danby, et al

.

"Warm Reinforcement and Cold Reinforcement Magnet Systems for Tokamak

Fusion Power Reactors: A Comparison," BNL 17434 (1972) Engineering

Division/Department of Applied Science.

"Summary of Existing Superconducting Magnet Experience and Its Rele- vance to the Safety of Fusion Magnet," IEEE 75 CH 1097-5-NPS (1975)

S. Y. Hsieh, et al.

"BNL Cryogenic Data Handbook," (1966) J. Jensen, R. Stewart, W. Tuttle.

"Workshop on Structural Analysis Needs for Magnetic Fusion Energy

(MFE) Superconducting Magnets," BNL (1976) proceedings to be published.

Energy Research and Development Administration (ERDA)

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Confinement Experiments," ERDA 76-34 (1976) Confinement Systems Program.

321
"Neutral Beam Energy and Power Requirements for the New Generation

of Tokamaks," ERDA 76-77 (1976) Division of Controlled Thermonuclear

Research.

"The 1974 Review of the Research Program," ERDA 39-75 (1975) Division

of Controlled Thermonuclear Research.

"Programs of the Materials and Radiation Effects Branch," ERDA 76-83

(1976) Division of Magnetic Fusion Energy.

"Materials Sciences Programs," ERDA-76/123 (1976) Division of

Physical Research.

"Physical Research Program: Research Contracts and Statistical

Summary," ERDA 63-75/UC-2 (1976) Division of Physical Research.

"An Assessment of the Role of Magnetic Mirror Devices in Fusion

Power Development," ERDA 76-74 (1976) ERDA.

"Report of the Ad Hoc Panel on RF Heating in Tokamaks," ERDA 76/115

(1976) ERDA.

"Fusion Power by Magnetic Confinement," ERDA 110/0-4, Vol. 0-4

(1976) ERDA.

"Fusion Power Research and Development Program C5-Year Program

Budget and Milstone Summaries," ERDA, Vol. 1-3 (1976) ERDA.

"New Directions in Fusion Magnet Development," Proc. Appl . Superconducting Conf. (1976) C. D. Henning, to be published.

"Applications of Superconductivity in the Controlled Thermonuclear

Research Program," Proceedings of 1974 Applied Superconductivity

Conf. (1975) G. Hess, et al

Los Alamos National Laboratory (LASL)

A 100-MJ Superconducting Magnetic Energy Storage System," LA 5882-MS

(1975) W. V. Hassenzahl.

"Superconducting Magnetic Energy Storage," LA 5935-PR, Vol. 1 and 2 (1975) W. V. Hassenzahl.

"An Engineering Design Study of a Reference Theta-Pinch Reactor (RTPR)

Environmental Impact Study," LA 5336, Vol. 1 and 2 (1975) LASL.

"Proposed METS-FTR Coupled Superconducting Prototype System," LA 5918

(1975) J. D. Rogers, et al

"Conceptual Engineering Design of a 1-GJ Fast Discharging Homopolar

Machine for the Reference Theta-Pinch Fusion Reactor," EPRl ER-246

(1976) K. Thomassen.

"Conceptual Design Study of a Scyllac Fusion Test Reactor," LA 6024

(1976) K. Thomassen.

322
Lawrence Llvermore Laboratories (ILL)

"Conceptual Design of a Mirror Reactor for a Fusion Engineering

Research Facility (FERF)," UCRL-51617 (1974) T. Batzer, et al.'

"The MX Major Project Proposal," LLL-Prop-142 (1976) F. W. Coensgen.

"The MX Magnet System," UCRL-78335 (1976) R. H. Bulmer, et al

.

"The Field Reversed Mirror and Power Reactor," Presented at Richland ANS Meeting (1976) W. Condit, et al

.

"High Field Test Facility Major Project Proposal," LLL-Prop 130

(1975) D. N. Cornish, et al

.

"2X11B Mirror Machine," UCRL-52000-75-1 1 (1975) LLL.

"An Evaluation of the Technical and Economic Feasibility of Mirror

Fusion Devices," UCRL 13695 (1976) LLL.

"Magnetic Fusion Energy Program," UCRL-52000-76-6 (1976) LLL.

"Program for Development of High-Field Superconducting Magnets for

Fusion Research," MISC-2007 (1975) LLL.

"Conceptual Design of a Mirror Hybrid Fusion-Fission Reactor," UCRL76515 (1975) R. Moir, et al

.

Massachusett Institute of Technology (MIT)

"A High Density, High Field Tokamak Demonstration Power Reactor,"

Proceedings of the Meeting on the Technology of Controlled Nuclear

Fusion, Richland, WA (1976) D. R. Cohn, et al

.

"Annual Report," (1976) Francis Bitter National Magnet Laboratory.

"The Elmo Bumpy Torus Reactor," Proceedings of the Meeting on the

Technology of Controlled Nuclear Fusion, Richland, WA (1976)

D. G. McAlees, et al

.

"Promising New Results from Alcator," Physics Today 18^ (Jan 1976).

"A Maanet Field System for a High Density Tokamak Reactor," Paper

(1976) J. Williams.

Oak Ridge National Laboratory (ORNL)

"High Beta Flux-Conserving Tokamaks," ORNL/TM-5429 (1976) J. F. Clarke.

"Conceptual Design of the Blanket and Shield Region and Related

Systems for a Full Scale Toroidal Fusion Reactor," 0RNL/TM-3096 (1973)

A. Fraas.

323
"Mechanical Engineering Design Considerations for a Tokamak Scientific

Feasibility Experiment and Plasma Test Reactor," ORNL/TM-4371 (1973)

A. Fraas , et al

"Program for Development of Toroidal Superconducting Magnets for

Fusion Research," Paper (1975) H. Long, et al

"Oak Ridge Tokamak Experimental Power Reactor Study - 1976 Part 3

Magnet Systems," ORNL/TM-5574 (1976) J. W. Lue, et al

"The Elmo Bumpy Torus Reactor (EBTR) Reference Design," ORNL/TM5669 (1976) D. McAlees.

"ORNL-EPR Study - Results and Implications," 76-WA/NE-lO (1976)

D. McAlees.

"Annual Progress Report - Solid State Division," ORNL-5135 (1975)

ORNL.

"Annual Progress Report - Thermonuclear Division," ORNL-5154 (1975)

ORNL.

"OR Tokamak Experimental Power Reactor Study Reference Design,"

ORNL/TM-5042 (1975) ORNL.

"Superconducting Magnet Development Program Annual Report," ORNL/TM5019 (1975) ORNL.

"Research, Development and Demonstration Needs," ORNL/TM-5577 (1977)

M. Roberts , et al

"The Effect of High Magnetic Fields on Metal Foil Strain Gauges at

4.2 K," Cryogenics (May 1975) P. Walstrom.

"A Summary of Tritium Handling Problems in Fusion Reactors," ORNL/TM4022 (1972) J. Watson.

"ORNL Fusion Power Demonstration Study: Arguments for a Vacuum Building in which to Enclose a Fusion Reactor," ORNL/TM-5664 (1976) R. Werner.

Princeton Plasma Physics Laboratory (PPPL)

"Analytic Solutions for Constant-Tension Coil Shapes," Journal of

Applied Physics, Vol. 47, 2710 (1976) S. L. Gralnick, et al

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(1976) D. L. Jassby.

"A Fusion Power Plant," MATT-1050 (1974) R. Mills.

324
"Poloidal Diverter Experiment," Physics Today (April 1976) Physics

Today - FCB.

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Fusion," Special Report 71/3 (1971) Princeton U. Plasma Physics Lab.

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WA (1976) F. H. Tenney.

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General Atomic

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Engineering Staff.

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(1975) Fusion Engineering Staff.

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.

University of Wisconsin

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R. Boom, et al

.

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.

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.

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325
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C. Maynard, et al

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University of Wisconsin.

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University of Wisconsin.

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326
General

F. R. Fickett, "Controlled Thermonuclear Reactors," Report to

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W. Gough, "Prospect of Fusion Power," Sci . American 224, 50

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R. Hirsch, "Status and Future Directions of the World Program in

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Paper (1976).

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327
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328
Material

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Cryogenic Temperatures," ASTM STP580 (1975) p 586.

V. I. Kononenko and V. V. Pustovalov, "Effect of a Constant Magnetic

Field on the Yield Stress of Nickel at 4.2° K," Soviet Journal of

Low Temperature Physics, 2^, 132 (1976).

330
C. A. M. vander Klein, "The Organic Insulation in Fusion Reactor

Magnet Systems," Reactor Centrum Nederland, Report No. 240 (1975).

V. I. Kozlovskaya, "Weldability and Mechanical Properties of Joints

in Martensitic Steel 03Khl2N10MT at Low Temperatures," Science and

Technology (1975).

G. Kulcinski, "Materials Problems and Possible Solutions for Near

Term Tokamak Fusion Reactors," EDM-186 (1976). (Univ. of Wisconsin)

G. L. Kulcinski and N. M. Burleigh, Eds., "Proceedings of the

Second Topical Meeting on the Technology of Controlled Nuclear

Fusion," ERDA/CONF-760935-Pl , Four Volumes (1976).

D. Larbalestier, et al., "High Strength Austenitic Stainless Steels

for Cryogenic Use," RL-76-048 (1976).

G. E. Linnert, "Welding Characteristics of Stainless Steels,"

Joining of Stainless Steels, American Society for Metals, Metals

Park, OH (1967).

C. J. Long, "Structural Materials for Large Superconducting Magnets

for Tokamaks," ORNL/TM-5632 (1976).

S. S. Manson, "Fatigue: A Complex Subject - Some Simple Approximations," Experimental Mechanics, 5^, 193 (1965).

H. I. McHenry, S. E. Naranjo, D. T. Read and R. P. Reed, "Low

Temperature Fracture Properties of a USSR Aluminum-6% Magnesium

Alloy," Proceedings of the Applied Problems of Low-Temperature

Materials and Welding Cryogenic Structures, Kiev, USSR (1976).

A. J. Nachtigall, "Strain Cycling Fatigue Behavior of Ten Structural Metals Tested in Liquid Helium (4 K) , in Liquid Nitrogen

(78 K), and in Ambient Air (300 K)," NASA TN D-7532, (1974).

L. E. Morgan, et al., "Material Fabrication Study for Fusion Power

Reactors," CASD-ERR-76-019 (1976). (General Dynamics, Convair

Division)

NASA, "The Stress Corrosion Resistance and the Cryogenic Temperature Mechanical Behavior of 18-3 MN (Nitronic 33) Stainless Steel

Parent and Welded Material," NASA TM X-73309 (1976).

NBS/ARPA, "Handbook on Materials for Superconducting Machinery,"

(1976).

NBS/ARPA, "Materials Research for Superconducting Machinery," Six

Volumes (1974-1976).

NBS Cryogenic Data Center, "Bibliography of References - Mechanical

Properties of Chromium-Nickel-Manganese Steels at Cryogenic Temperatures," (1975).

331
ORNL, "Controlled Thermonuclear Materials Technology Program Annual

Progress Report," ORNL-5082 (1975).

ORNL, "Magnetic Fusion Energy Materials Technology Program Annual

Progress Report," ORNL-5189 (1976).

D. T. Read, "Summary of Metallurgical Effects in Bulk NbTi," ASM

Conf. on the Manufacture of Superconducting Materials (1976).

R. P. Reed and D. T. Read, "Summary of NBS-Boulder Test Results for

U-25 Bore Tube Material," (1977) (Test Report from NBS to ANL).

B. W. Roberts, "Properties of Selected Superconductive Materials,"

NBS Tech. Note 724 (1972).

B. W. Roberts, "Properties of Selected Superconductive Materials - 1974 Supplement," NBS Tech. Note 825 (1974). \

L. H. Rovner and G. R. Hopkins, "Ceramic Materials for Fusinq,"

Nuclear Tech. 29, 274 (1976).

SAE/ASTM, "Unified Numbering System for Metals and Alloys,"

J1086/ASTM DS-56 (1974).

R. E. Schramm, et al . , "A Compilation and Evaluation of Mechanical,

Thermal and Electrical Properties of Selected Polymers," NBS Monograph 132 (1973).

F. R. Schartzberg and T. Kiefer, "Study of Fracture Behavior of Metals

for Superconducting Applications," NBS/ARPA Materials Research for

Superconducting Machinery, 4^ (1975).

L. L. Sparks and F. R. Fickett, "Magnetothermal Conductivity," NBSIR

74-393 (1974).

R. L. Tobler, "Fracture of Structural Alloys at Temperatures Approaching Absolute Zero," NBS/ARPA Materials Research for Superconducting

Machinery, 6, 279 (1976).

R. L. Tobler and R. P. Reed, "Fatigue Crack Growth Resistance of

Structural Alloys at Cryogenic Temperatures," NBS/ARPA Materials

Research for Superconducting Machinery, 6^, 256 (1976).

R. L. Tobler and R. P. Reed, "Fracture Mechanics Parameters for 5083-0

Aluminum Alloy at Low Temperatures," NBS/ARPA Materials Research for

Superconducting Machinery, 6^, 1 35 (1976).

L. A. A. Warnes and H. W. King, "The Low Temperature Magnetic

Properties of Austenitic Fe-Cr-Ni Alloys," Cryogenics 1^, 659 (1976).

J. Wells, "Evaluation of Nitronic 33 SS for TFTR Device Applications,"

(1975).

J. M. Wells and M. W. Hagadorn, "Evaluation of Nitronic 33 Stainless

Steel for TFTR Device Applications," 75-8C3-WETW0-R1 (1975) (Westinghouse Research Labs).

332
M. Wilson and C. R. Walters, "Development of Superconductors for

Fusion Technology," RL-76-038 (1976). (Rutherford Laboratory).

W. Witzke, et al . , "Effect of T^inor Reactive Metal Additions on Reactive Toughness of Iron-12% Nickel Alloy at -196°C and 25°C,"

NASA TN D-8232 (1976).

H. Yoshimura, et al., "Microstructures , Low Temperature Toughness

and Thermal Expansion Coefficients of High Manganese-Chromium Austenitic Steels," Transaction ISU, Vol. 16 No. 2 (1976).

333
Radiation Effects

H. Brechna and W. Maurer, "Irradiation Effects in Superconducting

Synchrotron Magnets," KFK-1468 (1971). (Kernforschungstentrum

Karlsruhe)

B. S. Brown, et al . , "Critical Current Changes in Nb^Sn Irradiated

with Fast Neutrons at 6 K," Journal of Applied Physics, 46 (1975).

B. S. Brown, et al . , "Radiation Effects on Superconductivity," ASM

Seminar (1975).

R. Bui lough, "A Systematic Approach to the Radiation Damage Problem

in Fission and Fusion Reactor Structural Alloys," FDM-184 (1976).

(University of Wisconsin).

M. Couach, et al . , "Effects of Fast Neutron Irradiation at Low Temperature on the Properties of Nb-Ti Superconducting Wires," IEEE

Trans, on Magnetics 1J_, 170 (1975).

J. F. Guess, et al . , "A Survey of Radiation Damage Effects in

Superconducting Magnet Components and Systems," ORNL-TM-5187

(1975).

J. F. Kircher and R. E. Bowman, Eds. Effects of Radiation on Materials and Components , (Reinhold Publishers Co., New York, 1964).

Lockheed Nuclear, "Effect of Nuclear Radiation on Materials at

Cryogenic Temperatures," NASA CR-54881 (1965), NASA CR-72056 (1966),

and NASA CR-72332 (1967).

R. McDaniel, et al . , "Combined Effects of Reactor Radiation and

Cryogenic Temperatures on Nerva Structural Materials," FZK394-1

(1973). (General Dynamics, Convair Aerospace Division).

P. L. Merz, "Interim Report on Evaluation of Dielectric Materials

Under Irradiation," (1976) (General Atomic Co.)

ORNL, "Summary of U.S. LMFBR Programs on High Temperature Structural Design and Associated Materials Testing," ERDA-76/146 (1976).

D. M. Parkin, et al . , "Neutron Irradiation of Nb^Sn and NbTi Multifilamentary Composites," IEEE Trans, on Magnetics VI_, 166 (1975).

N. L. Peterson and S. D. Harkness, Eds., Radiation Damage in Metals , (ASM, Metals Park, OH, 1976).

W. Weleff, et al . , "GTR-16 Radiation Effects Test on Structural

Materials at -423°F," NO. RN-S-0290 (1966). (NERVA Program, Aerojet General Corp. )

C. Restat, et al . , "The Effect of Radiation on Superconducting

Coils," CERN ISR-MA/75-20 (1975).

334
L. W. Ricketts, Fundamentals of Nuclear Hardening of Electronic

Equipment , (Wiley-Interscience, New York, 1972).

H. Schonbacker, et al . , "Radiation and Fire Resistance of CableInsulating Materials Used in Accelerator Engineering," CERN 75-3

(1975).

E. Seibt, "Irradiation and Annealing Effects of Neutron Irradiated

NbTi and V^Ga Mul til filamentary Composite Wires at Low Temperatures,"

IEEE Trans, on Magnetics V\_, 174 (1975).

M. Soell, et al., "The Influence of Low Temperature Neutron Irradiation on Superconducting Magnet Systems for Fusion Reactors," IEEE

Trans, on Magnetics ri_, 178 (1975).

J. T. Stanley and K. R. Garr, "Ferrite Formation in Neutron Irradiated

Type 316 Stainless Steel," Metallurgical Transactions 6A^, 531

(1975).

A. R. Sweedler, et al . , "Neutron Induced Disorder in Superconducting A-15 Compounds," IEEE Trans, on Magnetics 1J_, 163 (1975).

M. W. Thompson, "Defects and Radiation Damage in Metals," (Cambridge

Press, 1969).

F. L. Vook, "Effects of Radiation on Materials," Physics Today, 28,

34 (1975).

F. L. Vook, et al . , "Radiation Effects on Materials," Reviews of

Modern Physics, 47, 51 (1975).

M. Van de Voorde, et al . , "Selection Guide to Organic Materials for

Nuclear Engineering," CERN 72-7 (1972).

J. M. Williams, et al., "A Method of Transferring Irradiated Samples

in Liquid Helium," ORNL 76-22 (Jan. 1976).

335

Appendix A. Summary of Project Materials and Components

The tables presented here summarize the materials choices made by

the various MFE projects evaluated for this report. We have attempted

to achieve a uniform format for the tables and, as a result, it is

probably not perfect for any project. The major component material

choices are described in more detail in the next section.

A-1
\\\ ucvxLu UUrlrUlNclN 1 J /AooULlMItU Wlln

SUPERCONDUCTING MAGNFTS DR rRYDGFNTr MAGI\IFT<; 1

.ANI (FPR)

I. Magnet Coils TF cm (EE)

A. Conductor Components

1. Normal magnets —

1 — 1

NbTi NbTi

OFHC Cu OFHC Cu — —

Diffusion barrier

Eddy current barrier _ _

Strengthening material 316 SS oo Wirt: Ur UW LU

wire

PbSn solder PbSn solder

Metallic potting (cables) — Soft solder

B. Electrical Insulation

Formvar Formvar

FEP FEP

. . - - ---

C. Bonding Agents

1. Insulation to coil

2. Insulation to confining structure

A-2
D. Current leads -- --

E. Spacers, Shims and Coolant Channels Mi carta G 10

F. Coil Form and Confinement

2. Joining materials (bolts, etc.)...

316 SS

--

SS

--

II. Eddy Current Shields Al/SS panels

III. Magnet Dewars

A. Inner Wall — —

B. Insulation Superinsulation Superinsulation

C. Outer Wall Al alloy Al alloy

D. Thermal Shields

IV. Support Structure

A. Magnet to Dewar Inner Wall FEP FRP

B. Dewar Inner Wall to Outer Wall FEP FEP

C. TF Coil Central Support

SS(Shield to OH coi'

Al ( Inside cyl inder)

FRPf Cylinder to flo(

)

>r)

D. External Support Structure (v/arm) FRP 7075 Al

A-3
LLL Summary (MX, FERF)

Component

Conductor Components

Conductor

Stabil izer

Potting

Electrical Insulation

-layers

-turns

-corners

Cryostat

Coil Form

Insul ation

Thermal Sink

Outside Siructural Supports

Minor Axis Reinforcement

Interconnecting Core (C-clamps,

FERF)

Magnets to Floor

Magnet- to-Magnet and Floor

MX Materials and Form (FER

in parens when distinct

NbTi filaments

OFHC Cu wire, sheet

Soft solder

NEMA Grade G-11 sheet (polyimide or polyimide-gl ass laminate)

NEMA Grade G-11 buttons

Mylar sheet

21/5/9 stainless steel plate

and rings

Vacuum and multi -layer

(LN, cooled Cu sheet (FERF))

21/6/9 stainless steel plate

21/6/9 stainless steel plate

21/6/9 stainless steel columns

Post-tensioned concrete
CTR DEVICE COMPONENTS ASSOCIATED WITH

SUPERCONDUCTING MAGNETS OR CRYOGENIC MAGNETS

ORNL (EPR)

I. Maqnet Coils TF OH (EF)

A. Conductor Components

1. Normal magnets — --

Conductor -- --

Eddy current barrier — —

2. Superconducting magnets Hi-field Lo-field

Nb3Sn NbTi NbTi

Cu Cu Cu

Cu-Sn --

Diffusion barrier Ta

Eddy current barrier -- CuNi

Strengthening material Stainless steel

ilnininn matprial Qrk 1 H o oU 1 Uc

1

1 Hoy*

Al-Alloy conduit SS tube conduit

^ FRP nntpntiflllv^

B. Electrical Insulation

"Insulation" "Insulating matrix"

C. Bonding Agents

1 . Insulation to coil

2. Insulation to confining structure

A-5
D. Current leads Vapor cooled Vapor cooled

E. Spacers, Shims and Coolant Channels

F. Coil Form and Confinement

2. Joining materials (bolts, etc.)...

Stainless steel FRP

II. Eddy Current Shields

III. Magnet Dewars

A. Inner Wall

B. Insulation

1

C. Outer Wall

D. Thermal Shields

IV. Support Structure

A. Magnet to Dewar Inner Wall

B. Dewar Inner Wall to Outer Wall

C. TF Coil Central Support Stainless steel FRP

D. External Support Structure (v/arm)

A-6
CTR DEVICE COMPONENTS ASSOCIATED WITH

SUPERCONDUCTING MAGNETS OR CRYOGENIC MAGNETS

QEMI. - ICP fdata from RFP)

I. Magnet Coils

A. Conductor Components

1. Normal magnets.

Conductor.

Eddy current barrier.

2. Superconducting magnets,

Superconductor.

Pure metal stabilizer

Alloy for diffusion process

Diffusion barrier

Eddy current barrier

TF

NbTi or NbsSn, a 7 T

a 7 T field region.

Calculate effect of

(E > 0.1 Mev)

strain > .05% in

0.5 X

10^'' n/cm^

Strengthening material

Joining material

Metallic potting (cables)

Pulsed to ^ 0.15 T

conductor in 1 sec,

OH (EF)

field normal to

B. Electrical Insulation

1 . Turn to turn

2. Coil to confining structure

Max dose 1.0x10 rad

100 kr at .8 X max voltage

100 kr at 4.2 x max ^/oltage

C. Bonding Agents

1 . Insulation to coil

2. Insulation to confining structure

A-7
D. Current leads

E. Spacers, Shims and Coolant Channels

F. Coil Form and Confinement

2. Joining materials (bolts, etc,)...

Material s wi th know 1 properties

II. Eddy Current Shields

III. Magnet Dewars

A. Inner Wall

B. Insulation

C. Outer Wall

D. Thermal Shields

ORNL provides vacuu

at p < 10-5 torr wi

wal 1

n enclosure

th 80 K cold

IV. Support Structure

A. Magnet to Dewar Inner Wall

B. Dewar Inner Wall to Outer Wall

C. TF Coil Central Support Provided by (ORNL)

D. External Support Structure (v/arm) Torque rings (provi ded by ORNL)

A-8
CTR DEVICE COMPONENTS ASSOCIATED WITH

SUPERCONDUCTING MAGNETS OR CRYOGENIC MAGNETS

General Atomics EPR - I . II

I. Magnet Coils TF OH (EF)

A. Conductor Components

1. Normal magnets ETP Cu (EF)

Conductor

Monolithic NbTi Braided NbTi

OFHC Cu OFHC Cu

Diffusion barrier

Eddy current barrier

Strengthening material Stainless Stainless

Metallic potting (cables)

B. Electrical Insulation

Mi carta

GFRP

Epoxy

C. Bonding Agents

2. Insulation to confining structure

A-9
D. Current leads

E. Spacers, Shims and Coolant Channels Micarta, GFRP

F. Coil Form and Confinement

1. Structural material

2. Joining materials (bolts, etc.)'

GFRP

II. Eddy Current Shields

III. Magnet Dewars

A. Inner Wall Stainl ess GFRP

B. Insulation

C. Outer Wall Stainless Stai nl ess

D. Thermal Shields

1. Within insulation space

2. Nitrogen jacket

IV. Support Structure

A. Magnet to Dewar Inner Wall

Stainless, GFRP

ceramics, epoxy

B. Dewar Inner Wall to Outer Wall

C. TF Coil Central Support

D. External Support Structure (v/arm) Stainl ess

A-10
CTR DEVICE COMPONENTS ASSOCIATED WITH

SUPERCONDUCTING MAGNETS OR CRYOGENIC MAGNETS

, General Atomics DEMO (DFPR)

I. Magnet Coils TF OH (EF)

A. Conductor Components

1. Normal magnets

Conductor Cu (EF)

Eddy current barrier

NbTi NbTi

Cu Cu-CuNi

Alloy for diffusion process

Diffusion barrier

Eddy current barrier

Strengthening material Stainl ess Stainless

Joining material

Metallic potting (cables)

B. Electrical Insulation

C. Bonding Agents

2. Insulation to confining structure

A-11
1

D. Current leads

E. Spacers, Shims and Coolant Channels

F. Coil Form and Confinement

2. Joining materials (bolts, etc.)...

Stainl ess

II. Eddy Current Shields Al umi num

III. Magnet Dewars

A. Inner Wall

B. Insulation

C. Outer Wall

D. Thermal Shields

j

IV. Support Structure

A. Magnet to Dewar Inner Wall

;

B. Dewar Inner Wall to Outer Wall

C. TF Coil Central Support GFRP

D. External Support Structure (v/arm)

A-12
CTR DEVICE COMPONENTS ASSOCIATED WITH

SUPERCONDUCTING MAGNETS OR CRYOGENIC MAGNETS

II XI

I

A. Conductor Components

1. Normal magnets.

Conductor.

Eddy current barrier.

2. Superconducting magnets

Superconductor.

Pure metal stabilizer

Alloy for diffusion process

Diffusion barrier

Eddy current barrier

Strengthening material

Joining material

Metallic potting (cables)

TP

TINb

Cu (I, II)

Cu + pure Al (III)

OH (EF)

TiNb

Cu (I, II)

Cu, work hardened

(III)

Sol der

3. Electrical Insulation

1. Turn to turn I No turns in contact | Kapton

CuNi sheath on sc

f i 1 aments

316 SS in some cases

Sol der

2. Coil to confining structure

C. BcndiiKj Agents

1 . Iiisul otion to coi 1

PEP FEP

?. Insulation to confining structure

A-13

Epoxy

Epoxy

Not given

Not given
D. Current leads Vapor cooled Vapor cooled

1

E. Spacers, Shims and Coolant Channels Mi carta Mi carta

F. Coil Form and Confinement

2. Joining materials (bolts, etc.)...

316 SS (I, II)

2219 Al (III)

Al alloy bolts

None

Al alloy bolts

II. Eddy Current Shields

III. Magnet Dewars

A. Inner Wall SS SS

B. Insulation Aluminized Mylar Aluminized Mylar

C. Outer Wall SS (I, II)

Al alloy (III)

SS (I, II)

Al alloy (III)

D. Thermal Shields

IV. Support Structure

A. Magnet to Dewar Inner Wall Micarta Micarta

B. Dewar Inner Wall to Outer Wall GFRP struts Probably GFRP str

C. TF Coil Central Support SS (I, II)

Al alloy (III)

D. External Support Structure (warm)

fvl indpr-oiston

elements. Secondary vacuum wal

"Some vertical

structure"

(II, III)

A-14
CTR DEVICE COMPONENTS ASSOCIATED WITH

SUPERCONDUCTING MAGNETS OR CRYOGENIC MAGNETS

^iJyiQ^-Bum py .Torus

T. Magne t Co i 1

s

A. Conductor Components

1. Normal magnets.

Conductor.

Eddy current barrier.

2. Superconducting magnets

Superconductor

Pure metal stabilizer

Alloy for diffusion process

Diffusion barrier

Eddy current barrier

Strengthening material

Joining material

Metallic potting (cables

B. Electrical Insulation

1 . Turn to turn

TF

NbTi

Copper

Solder

Kapton

2. Coil TO confining structure j " Insul ation"

i C. Bonding Agents

1 . Insul ation to coi 1

2. Insulation to confining strncture

A-15

OH (EF)

NbTi

Copper

Sol der

Kapton

"Insulation"
D. Current leads —

E. Spacers, Shims and Coolant Channels GFRP GFRP

F. Coil Form and Confinement

2. Joining materials (bolts, etc.).-.

SS

: .

SS

II. Eddy Current Shields

III. Magnet Dewars

A. Inner Wall Austenitic SS Austenitic SS

B. Insulation Aluminized Mylar Aluminized Mylar

C. Outer Wall Austenitic SS Austenitic SS

D. Thermal Shields

Copper W/LN piping Copper W/LN pipini

IV. Support Structure

A. Magnet to Dewar Inner Wall SS tendons

304L

SS tendons

304L

B. Dewar Inner Wall to Outer Wall

C. TF Coil Central Support

_ —

D. External Support Structure (v/arm) Cone re '..e Concrete

A-16
Appendix B. Project Information Sheets

The sheets in this appendix are summaries of the work sheets which

we used in our project evaluations. They are included here by laboratory.

The units used are those listed by the original report. The sheets cover each major component of the cold magnet system and give a description of the material chosen as well as information on the component environment. In many instances a listing of the data which we feel is required

for the material in this application is given. In a lesser number of cases possible alternate materials are described.

ARGONNE NATIONAL LABORATORY

TF Conductor Composite

ANL - EPR (1976)

Carry current

NbTi filaments

Mm @ 10 - 3 T field (a 3 K

Component : TF Conductor Composite - Conductor

Project : Function : Descri ption : Requirements : Critical current = 0.70-2.5 x 10

Operational current = 60 kA/turn

^ Total conductor length/coil = 151 x 10 A-m

Smallest radius of curvature = 2 m 0 10 T field

Ave. hoop force/turn = 133 x 103 lbs

Ave. turn cross-section = 40.6 cm^

NbTi cross-section = 1.2 - 0.3 cm2/turn

NbTi volume/coil = 0.267/m3

Radiation: large doses of 10 n/cm reduce I by 10-15%

Forces: gravitational (bottom) - 208 tons weight

D straight section - 8 ksi compression

Component : TF Conductor Composite - Stabilizer

Project : ANL - EPR (1976)

Function : Carry full current if conductor goes normal

Description : Copper wire and copper sheet

Requirements : Cryostatic stabilization

Average stress = 14.5 ksi in wire

Radiation-induced resistivity saturates at 3 x 10" cm Radiation-induced resistivity of 6 x 10"'^ cm with 1 MW-yr integral

wall loading • _o

Using safety factor of 2.5, designed to 1.5 x 10 cm

If 0.5 MW/m^ wall loading and 50% capacity factor, 4 yrs to reach

1.5 X 10~8 cm, radiation removed if warmed to 300 K Copper thickness/turn = 0.46 0.23 cm/turn

Copper volume/coil = 6.82 m3

Sheet - 30 to 50 cm wide and 0.43 cm thick

Other forces: gravitational (bottom) - 208 tons weight

D straight section - 8 ksi

Data Required : Magnetoresistivity as a function of radiation, purity,

stress and cold work.

B-1
TF Conductor Composite - Strengthener

ANL - EPR (1976)

Carry conductor load

Stainless steel strip

Average stress = 26 ksi

smallest radius of curvature 2 m, thickness varies from 0.31 cm/turn

at 10 T field to 0.54 cm/turn at 3 T field, width varies between 30 to 50

Forces: gravitational (bottom) - 208 tons weight

, • D straight section - 8 ksi compression

Component : Project : Function : Description

Requirements stainless steel/copper ratio = 1.5

Component : Project : Function : Description:

Requirements

copper sheet.

Forces

TF Conductor Composite - Potting

ANL - EPR (1976)

To grip conductor to copper sheet stabilizer

SnPb soft solder

Cu/NbTi conductor is pretinned, then soft-soldered to

gravitational (bottom) - 208 tons weight

D straight section - 8 ksi compression

Component : TF Conductor Composite Electrical Insulators

Project: ANL - EPR (1976)

Function : To electrically insulate conductors, turn to turn

Wetted f i berglass-epoxy strip

Thickness of 1 mm, width of 1 cm, covers 50% of conductor

Description : Requirements

area, electrical insulation, turn to turn, for 1 mm thick is ^ 200 V. Cure of 150-200° F needed for prepreg, aromatic-type curing agent.

Radiation: thought to degrade at 10^-10^0 rads designed to receive

3.5 X 108 rads at 2.5 MW-yr/m^.

Data Required : Influence of radiation; compressive strength (4 K), thermal

conductivity, specific heat, thermal expansion, electrical properties

(300-4 K).

Component : Project : Function : Description:

Requirements

TF Conductor Composite - Electrical Insulators

ANL - EPR (1976)

Reduce ac losses, insure magnet charging

Formvar coating on conductor

Not discussed

Data Required : Influence of irradiation

Component : TF Cryostat - Coil Form

Project : ANL - EPR (1976)

Function : Support conductor

Descri ption : Stainless steel containment case

Requi rements : AISI 316 plate

Thickness of 1 . 25 cm Width of 68-104 cm, height of 68 cm, length in D shape about 36 m Strengthening rib to divide case allows use of 1.25 cm thick stainless, instead of 2.5 cm thick plate.

Data Required : Weld specifications

B-2
Component : TF Cryostat - Coil Form Electrical Insulation

Project : ANL - EPR (1976)

Function : Electrically insulate coil form from radiation assembly

Description : NEMA Grade G-10 or micarta sheet

Requi rements : Sheet thickness = 1 cm Grooved sheets for LN2 channels for venting

Data Required : Tensile, compression (4 K), electrical and thermal (4 K),

irradiation effects, purchase specifications, joining to stainless steel

Component : TF Cryostat - Eddy-current Shield

Project : ANL - EPR (1976)

Function : Shield conductor components from pulsed magnet fields

Description : Aluminum-stainless composite panels

Requirements : RRR aluminum ^ 5000 (2 x 10~ ^-cm at 4 K in 4 T field, 5 x 10" ^

at 18 K in 4 T field.

Operating temperatures vertical D-coil section = 3, 4.2 K, 3-10 T

and peak fields: outside D-coil section = 12.5-23.5 K; 1.5-5.0 T with

0.5 T ac field

Thickness of aluminum at vertical D-coil section = 4 cm, stainless

= 0.25 cm Thickness of aluminum at outside D-coil section = 5 cm, stainless

= 1.0, 0.25 cm Aluminum explosively joined to stainless (estimated stainless stress

= 60 ksi, aluminum strain = 0.17)

Data Required : Joining techniques

Irradiation effect on electrical resistivity

Composite, aluminum-base exploritation

ac losses = 18 K portion =8.5 KW/coil

3 K portion =0.3 KW/coil

Component : TF Cryostat (Outer section) - Insulation: Shield to

outside wall

Project : ANL - EPR (1976)

Function : Thermal insulation of aluminum shield

Description : Superinsulation in vacuum Requirements : Radiation loss 300 K wall to 18 K shield = 64.5 W/coil

Multiple layer superinsulation

Thickness = 3.5 cm

Data Required: Effect of radiation on superinsulation

Component : TF Cryostat (Outside section) - Outside wall

Project : ANL - EPR (1976)

Function : Enclose coil and provide support

Description : Aluminum alloy plate casing

Requirements : Aluminum alloy plate, 3.5 cm thick

B-3
Component : to OH coil

Project : Function : Descri ption : Requirements

thickness of

Load on Circumferential

TF Coil Cryostat (Vertical section) - Support: Shield

ANL - EPR (1976)

To absorb compressive stresses from D coils

Stainless steel cylinder

Approximate 4 m diameter stainless steel /cyl i nder with

22 cm.

cyl inder = 8.6 x 10" lbs.

stress 60 ksi

Compressive stress - 7.7 ksi

Data Required : Thick-section processing

mechanical properties of welds (4 K)

and welding of stainless steel

TF Coil Cryostat (Vertical section) - Insulation: Support

ANL - EPR (1976)

Electrically insulate OH coil from support cylinder

Cyl inder

Probable NEMA G-10 sheet, thickness = 5 cm. Data Required : Compressive strength at 4 K, effects of irradiation, electrical and thermal properties under load.

Component : to OH coil

Project : Function : Descri ption

Requirements

Component : TF Coil Cryostat (Vertical section) - Insulation: OH

coil to LN2 shield

Project : ANL - EPR (1976)

Function : Thermally insulate OH coil from LN^

Descri ption : Vacuum and superinsul ation layer

Requirements : Superinsulation thickness - 0.5 cm, vacuum spacing = 2.5 cm Data Required : Effects of irradiation

Component : Project : Function : Descri ption

Requirements :

TF Coil Cryostat (Vertical section) - Inside cylinder

ANL - EPR (1976)

Structural support on inside for double-walled dewar

Aluminum alloy inner dewar wall cylinder

Aluminum alloy vacuum wall thickness = 3.5 cm at 300 K Data Required : Thick-section welding procedures

Component : TF Outside Support Structure - Coil to Coil

Project : ANL - EPR (1976)

Function : Provide support for coil-to-coil torque loads

Description : Composite support panels

Requirements : Operate within TF cryostat at 4 K Positioned along outside perimeter of D-coils

Data Required : Compressive and fatigue properties at 4 K of fiberglassepoxy composites, thermal expansion and specific heat (4-300 K).

B-4
Component : TF Outside Support Structure - Support Cylinder to Floor

Project : ANL - EPR (1976)

Function : Carry gravitational forces through temperature gradient

(4-300 K)

Descri ption : Composite cylinder

Requirements : Temperature gradient (4-300 K)

Load of 8.5 x 10^ lbs (assuming 1/2 of total TF coil load)

Glass-reinforced epoxy material

Data Required : Compressive properties (4-300 K)

Thermal conductivity (4-300 K) under load of FRP

Creep, fatigue properties (4-300 K) of FRP

Thermal expansion, specific heat (4-300 K) of FRP

Component : OH/EF Conductor Components - Conductor

Project : ANL - EPR (1976)

Function : Carry current to produce magnetic field

Descri ption : NbTi filaments

Requi rements : Wire diameter - 0.3 mm

Basic strand Cu/NbTi ratio = 2 to 1 Overall Cu/NbTi ratio = 11 to 1 , ^ NbTi cross-sectional area = 2.3 x 10 cm Operate at 4.2 K, pool boiling

Iq ~ 117 KA/cm2, basic strand carry 27 A

Basic strand 1171 filaments, 5 y filament diameter

Component : OH/EF Conductor Components - Stabilizer

Project : ANL - EPR (1976)

Function : To carry current if superconductor locally goes normal

Descri ption : Cu wire and CuNbTi strands

Requirements : Cu wires of 0.3 mm diameter interwoven with Cu/NbTi (2 to

1 strands to produce Cu/NbTi ratio of 11 to 1 . Operate at 4.2 K, cryostatic stabilization.

Data Required : Magnetoresi stance and magnetothermal conductivity of oxygen

free Cu under stress. Effect of purity on final RRR of oxygen free Cu

fatigue strength.

Component : OH/EF Conductor Components - Strengthener

Project : ANL - EPR (1976)

Function : Provide strength to conductor composite

Descri ption : Stainless steel or cold-drawn Cu wire

Requirements : For large hoop stress: 1.2 mm diameter stainless steel wire

For small hoop stress: 1.2 mm diameter cold drawn Cu wire

Cu and Cu/NbTi strands woven on strengthener; this cable will carry

270 A at 5 T. Data Required : Fatigue strength of stainless and cold drawn Cu at 4 K.

B-5
Component : Project : Function : Description : Requirements : strands with

Data Required

OH/EF Conductor Components - Potting

ANL - EPR (1976)

Provide structural stability to Cu and Cu/NbTi strands

Soft solder

Soft solder, high resistivity, will be used to join

structural -reinforcing wire.

: Shear strength at 4 K.

Component : Project : Function : Description : Requirements

Data Required: None

OH/EF Conductor Components - Electrical Insulator

ANL - EPR (1976)

Provide strand to strand electrical insulator

Coating on Cu/NbTi strands

Not discussed

Component : Project : FunctionOH/EF Conductor Components - Electrical Insulation

ANL - EPR (1976)

Provide turn-to-turn electrical insulation

Description : NEMA Grade G-10 wetted strip

Requirements : Strip - 2 mm thick, 2 cm wide, covering 2/3 area of cables

Wetted, must be heated to 150-200° F for setting

Data Required : Electrical and mechanical properties at 4 K

Component : OH/EF Cryostat - Supports

Project : ANL - EPR (1976)

Function : Provide internal support about coil form

Description : Inner and outer internal FRP supports, separated by

permeation barrier.

Requirements : Must have high electrical resistivity to reduce ac losses

and good strength. Stainless steel permeation barrier to reduce helium

permeabil i ty.

Data Required : Tensile and fatigue strength at 4 K, joint behavior at 4 K,

need for permeability barrier.

Component : OH/EF Cryostat - Insulation

Project : ANL - EPR (1976)

Function : Thermal insulation between 300 and 76 K

Description : Multi-layer insulation

Requirements : Reflective coatings must be segmented, designed to minimize

eddy-current heating. Used in vacuum.

Data Required : Proper design of multi-layer insulation to minimize eddycurrent losses.

B-6
7075 aluminum alloy

Component : External Support Structure

Project : ANL - EPR (1976)

Function : To support magnets and forces between beam-panel structure

magnets

Description : Requirements : Use 7075-T6 aluminum alloy, bolted together.

McDonnel -Douglas study selected this alloy - comparing many, but

excluding normal plate steel.

Design with allowable stress level either 0.6 tensile yield, 0.4

tensile strength (shear), and 0.6 tensile yield for buckling.

Fatigue allowable for 7075 at 10^ cycles is about 0.6 tensile yield.

Data Required : None

B-7
LAWRENCE LIVERMORE LABORATORY

Component : Conductor Components - Conductor

Project : LLL - MX (1976)

Function : Carry current to produce magnetic field.

Descri ption : NbTi filaments

Requi rements : 2 Current density - 3400 Al cm Maximum field = 7 - 5 T Conductor design current = 6.8 kA

^ Conductor mass, both coils = 60 x 10 kg

Cu/NbTi conductor =^ 6 x 6 mm square cross-section

Operate at 4.2 K, cryostatic stability

Component : Project : Function : Descri ption : Requi rements : Superconductor element has Cu/NbTi ratio of 1-1.5 to 1 Other 3 quadrants of conductor are OFHC Cu, therefore,

overall ratio of Cu/NbTi - 8 to 1

Conductor I^R loss =1.9 W/cm

Cooling surface =5.3 cm^/cm

^ Required heat transfer rate = 0.36 W/cm

Total conductor mass is 30,300 Kg

Data Required : Cu magneto-electric and thermal properties.

Conductor Component - Stabilizer

LLL - MX (1976)

Carry current in event superconductor goes normal

Cu wire and sheet

the

Component : Project

Function : Description

Requi rements

Data Required

Conductor Component - Potting

LLL - MX (1976)

To hold 4-segments of conductor together

Soft solder

Shear strength at 4 K

Component : Conductor Component - Electrical Insulation

Project : LLL - MX (1976)

Function : Electrically insulate between turns.

Descri ption : Discs of NEMA Grade G-11

Requirements : Discs - 1 mm thick, about 1 cm diameter to be joined together

to prevent migration. NEMA Grade G-11 will probably be used. Maximum

compressive stress = 6 ksi. Voltage about 1 KV.

Data Required : Compressive and dielectric properties at 4 K.

Component : Project : Function

Conductor Component - Electrical Insulation

LLL - MX (1976)

To electrically insulate between layers

Description : NEMA Grade G-11 sheet

Requirements : NEMA Grade G-11 sheet - 2 mm thick. Sheets separated to

provide cooling channels. Maximum compressive stress = 6 ksi. Voltage

about 1 KV.

Data Required : Compressive strength at 4 K. Dielectric breakdown under

stress.

B-8
Component : Yin-Yang Cryostat

Project : ILL - MX (1976)

Function : Structural support

Descrl pti'on : Casing for magnet windings

Requirements : Casing thickness about 3 inches thick of 21/6/9 stainless

steel. Reinforced in some areas. Forces on casing are between -1.25 to

27.7 X 106 lbs. Maximum design stress, combined, = 50 ksi. 300,000 lbs

needed for magnets. (Baseball operates at combined design stress of about

15 ksi.)

Data Required : Properties of thick-section welds at 4 K. Thick-section

processing and welding specifications.

B-9
OAK RIDGE NATIONAL LABORATORY

Component : Conductor and Coil Form for Poloidal Field Coils

Project : ORNL - ERR

Function : Provide the primary flux swing that establishes the plasma

current; the vertical fields that stabilize the plasma position in the

TF coils and the magnetic fields that determine the shape of the plasma

cross-section.

Description : The PF coil systems consists of the shield-vertical field

(S-VF) and decoupling coils (resistive-coils), the trim vertical field

(T-VF) coils and the ohmic heat (OH) coils. The conductor for the T-VF

and OH coils is a round bundle of 37 strands of NbTi superconductor in a Cu stabilizer with CuNi eddy current barrier. The strands are transposed,

enclosed in a stainless steel tube, and surrounded by an insulating matrix.

The conductor has a rectangular cross-section measuring 4 x 4.44 cm and is

wound on a fiberglass coil form.

Requirements : The OH coils in the central solenoid sustain the highest

fields (7 T) , the largest rate of field change (7 T/s) and the largest field

swing (-7 T to + 7 T). The principal requirements for these coils are to

minimize eddy-current losses and to remove the heat produced by these

losses. The conductor bundle must be sufficiently flexible to permit wrapping the central solenoid and strong enough to sustain an average hoop

stress of 12,500 psi. The rectangular cross-section insulating matrix

must be able to tolerate the radiation environment.

Data Required : The mechanical, thermal and electrical properties of the

insulating matrix and GFRP coil forms should be characterized along with

the effects of radiation and fatigue on these properties.

Component : Conductor and Coil Form for TF coils

Project : ORNL - EPR

Function : Produces plasma confining field

Description : Four conductors are simultaneously wound in parallel,

insulated and enclosed in L-shaped stainless steel segments that are welded to each other and to the coil case. Each conductor consists of

93 triplex strands (2 superconductor composite wires and 1 Cu or Al wire

all soldered together) encased in an Al alloy conduit. NbsSn and NbTi are used for high field and low field regions of the magnets, respectively.

Requirements : The 93 strands within the conduit must permit cooling by

the forced-flow of supercritical helium. The conductor/conduit assembly

must be sufficiently flexible to permit winding into a coil that is tight

enough to resist movement when the coil is energized. The insulation sur- rounding the four parallel conductors must survive the welding operation

used to enclose the conductors in stainless steel. The stainless steel

is stressed to 45 ksi. The maximum field at the winding is 11 T at a cur- rent density of 6200 A/cm2.

Data Required : Properties of the welds (laser welding suggested by ORNL)

used to join the L-shaped stainless reinforcement without damaging the

insulation should be determined.

The thermal and mechanical properties of the solder, the Al alloy for

the conduit and the stainless steel structural elements and weldments need

to be characterized.

Mechanical properties of the thick-section stainless steel weldments

used for the coil form must be determined.

B-10
GENERAL ATOMIC COMPANY

Component : Normal EF Coils

Project : GA - EPR

Function : Plasma field shaping

Description : Electrical tough-pitch copper, water-cooled. Each conductor

5x5 cm, insulated either with f iberglass-epoxy or with oxides or silicates (mica) if radiation is a problem. Located inside TF coils. Total

of 22 coils of 155 turns each, segmented into 6 sections each, bolted

together. Welded steel casing on each coil. Aluminum also considered

but rejected for EPR because of larger volume required. May be used on larger reactors. Specifications require 2 KV turn-to-turn and ICQ KV

layer- to-layer insulating capability.

Superconducting coils are likely to be used in commercial plants.

Dynamically-stabilized Nb^Sn would be required along with improved shielding by the blanket.

Alternate Materials : None

Data Required : Properties of thick-sections of Al and Cu as affected by

irradiation.

Component : Superconducting TF Coil Composite

Project : GA - EPR

Function : Plasma driving

Description : NbTi in copper matrix. Stainless steel strip interleaved

with conductor.

Requirements : Maximum magnetic field in TF coils is 8 tesla, 4 tesla

at plasma major radius. May use either Nb-45% Ti or Nb-55% Ti , (over

atomic fraction Ti for higher fields. Either copper or aluminum could

serve as stabilizing materials; the lesser magnetoresistivity of aluminum

being an advantage. Nb3Sn is rejected due to inferior state of development and necessity to restrain larger forces at higher fields. Conductor is monolithic with about 1060 200 ym diameter NbTi filaments.

Conductor is jacketed by copper stabilizer and bonded to S.S. strengthener. Of 17 radial conductor layers, 3 nearest plasma are of 8.0-T

grade, next three are 7.1-T grade and remaining 11 are 5.7-T grade.

Alternate Materials : None

Data Required : Radiation effects on NbTi and Nb-^Sn conductors.

Component : Superconducting OH Coil Composite

Project : GA - EPR

Function : Ohmic heating of plasma

Description : NbTi interwoven with stainless steel to form a braid.

The central solenoid would be wound on a glass-epoxy spool and prestressed

by 2.5 mm thick stainless steel bands on the outside of the cable. The

coil would be prestressed vertically and radially. AC losses sustained

at solenoid ends are reduced by subdividing the stainless reinforcing

bands into ribbons. Internal pressure from the radial magnetic load

would be about 13.2 MPa (1900 psi) along most of the OH SC solenoid,

increasing 30% at the ends. Magnetic loading reduces the compressive coil

stress to 3.6 MPa (530 psi) and increases the tensile load on the stainless bands by 109 MPa (15.8 ksi) to a maximum overall stress or 226

MPa (32.8 ksi).

B-n
They originally considered normal water-cooled coils of semi-hard

ETP copper each conductor 5 x 5 cm square. Each turn typically 20 x 40 cm reinforced with adjacent prestressing belt of high-strength steel. Insulation to withstand 500 V conductor to conductor, 2 KV turn to turn and

100 KV layer to layer. Insulation not stated - probably same as EF coil,

glass-epoxy or mica. Most of the OH turns are inside the central bore

with several at outer periphery of the torus. Aluminum could replace

copper in large reactors.

Alternate Materials : None

Data Required : Properties of thick sections of Cu and Al for normal

coi 1 s

Component : Superconductor Stabilizer

Project : GA - EPR

Function : Superconductor stabilization

Descri ption : OFHC copper is specified for EPR due to availability;

however, aluminum is an acceptable alternative, superior to copper at

cryogenic temperatures due to lower magnetoresi sti vi ty . Copper stabilizer

must withstand total fluence of 10" n/cm2-sec at inside of TF coil (year

operation) and integrated fluence of 2 x 10^8 n/cm2 without excessive

resistivity increase. Radiation-induced resistivity changes in stabilizer

will most likely dictate acceptable radiation levels. Limited data indicate this fluence will reduce the critical current density by 10%,

recoverable 70% by 300 K annealing. May be a problem in a commercial

reactor, but not in EPR due to frequent warm-ups. Stresses on the copper

stabilizer are 76.4 MPa (11 ksi) compared to the 85 MPa (12 ksi) at which

increased resistivity is noted.

Alternate Materials : Filamentary or dispersion strengthened copper or aluminum to increase the strength without increasing the resistivity.

Data Required : Effect of 4.2 K neutron irradiation and of strain on the

resistivity of copper and aluminum of purities suitable for stabilization.

Exploratory work on reinforced high-conductivity Cu and Al

Component : Strengthening Material - Superconducting Coils

Project: GA - EPR

Function : Resisting static and cyclic forces on superconductor

Description : A stainless steel support structure encloses the superconductors in the TF coils to resist the Lorentz conductor force resulting

from the TF coil current and torrodial magnetic field. Calculated 134.9

MPa (21.6 ksi) stress carried by the stainless steel to reduce the stress

in the copper stabilizer to values below that causing resistivity increases

(85 MPa, 12.3 ksi). Also, an austenitic stainless steel strip is interleaved with the NbTi conductor in the TF coils. The operating stress in

the reinforcement is calculated to be 198 MPa (28.7 ksi). Associated

stress in the stabilizer is given at 112 MPa (16.2 ksi) - at variance

with the maximum value given above but perhaps reflecting preliminary

thinking

B-12
Conceptual design of a superconducting OH coil postulates NbTi

interwoven witli stainless steel to form a braid. The central OH solenoid

is prestressed to 117 MPa (17 ksi) by 2.5 mm thick stainless bands on the outside of the cable. The bands are subdivided into ribbons at the

solenoid ends where higher AC losses are sustained. Magnetic loading

increases the stainless stress to 226 MPa (32.8 ksi).

A1 ternate Materia! : Graphite or boron-reinforced composites might be

used for banding, for interleaving with the conductor or for a structural

form on which the conductor is wound.

Data Required : Fatigue of stainless steel at 4.2 K. Fatigue of graphite

and boron-reinforced composites under irradiation at 4.2 K. Effect of

Wigner energy release on the mechanical and thermal properties of composites.

Component : Electrical Insulation

Project : GA - EPR

Function : High-voltage breakdown barrier

Description : Polyethylene, mylar, reinforced epoxies and inorganic

insulators mentioned as possible alternatives for the TF coil insulation.

More specifically, the NbTi conductor is helix wound with 0.05 cm micarta

spacers between adjacent turns and 0.10 cm thick glass-epoxy strips between

radial layers covering half the surface area to form helium cooling passages. Space between the end TF coil windings and cryostat wall is filled

with epoxy. Epoxy also used to insulate and to smooth the irregular exposed

surface of the final layer which bears against the inner cryostat wall.

No details on OH coil insulation in the superconducting variant.

Alternate Materials : Polyimide-matrix composites for probable superior

radiation resistance.

Data Required : Mechanical and physical properties of organic insulators

under neutron irradiation.

Effect of Wigner energy release on insulating properties of polymers

and composites.

Component : Coolant Channel

Project : GA - EPR

Function : Superconductor cooling

Description : TF coils - coolant channel formed by 0.05 cm micarta spacers

between adjacent turns and by 0.10 cm thick epoxy-f ibergl ass strips between

radial layers covering half the surface of the helically-wound conductors.

OH coils - coolant flows through small radial grooves carved in the

turn-to-turn spacers (material not given).

Alternate Materials : None

Data Required : Electrical properties of micarta, fiberglass epoxy, etc.,

at 4 K under fatigue and radiation.

Effect of release of Wigner energy on the probability of coolant

channel blockage by composite breakup.

B-13
Component : Coil Form - OH Magnet

Project : GA - EPR

Function : Support of ohmic heating coil

Description : 15.3 m high, 2.75 m o.d., 1.95 m i.d. epoxy-f iberglass

spool inserted into central bore formed by mating of straight sections

of TF coil cryostats. Subjected to prestress both vertically and radially

by stainless straps pressing on NbTi conductors. Maximum radial com- pressive stress 22 MPa (3,200 psi) reducing to 3.6 MPa (530 psi) when

coil is operated. Spool sees axial torque producing shear stress of

about 1 MPa (140 psi) with 1° angular deflection. Will see fatigue, but

analysis for 500,000 cycles at 4 K indicates no problem. Prestressing

expected to compensate for differential thermal expansion coefficients

during cooldown. Vertical prestressing by steel bolts through a top

fiberglass cap and penetrating the fiberglass core. Axial compressive

stress 376 MPa (40,000 psi), coil form (and coil) shrinks 2 mm. Energizing

coil reduces axial stress to 2.9 MPa (420 psi). Epoxy-f ibergl ass coil

form is selected to minimize ac losses.

Aternate Materials : Kevlar or graphite-reinforced polymer composites.

Data Required : Mechanical, thermal, electrical properties of thick section

fiberglass at 4 K including fatigue, thermal cycling effects and radiation.

Feasibility of replacing solid composite section with web of a higher

modulus composite. All above data required.

Component : Inner Wall - Magnet Dewars

Project: GA - EPR

Function : Vacuum containment, helium containment

Description : OH coil helium pressure vessel made of glass-epoxy to

prevent ac losses from occurring at 4.2 K at the ends of the OH solenoid

where variable magnetic flux crosses the dewar (cryostat) wall. Negligible outgassing expected due to low temperatures. Inner dewar wall of TF

coil is austenitic stainless steel.

Alternate Materials : Glass-polyimide, Kevlar- or graphi te-epoxy or polyimide.

Data Required : Permeability of composites to helium at 4.2 K. Effect of

irradiation on permeability at 4.2 K. Composite-metal joints at 4.2 K. Possible permeability barrier development.

Component : Magnet Dewar Insulation

Project: GA - EPR

Function : Reduce thermal losses.

Description : OH dewar uses superinsulation (multilayers). Laminar

insulation in TF coil dewar.

Alternate Materials : 3-D reinforced form if thermal short is used in

the dewar.

Data Required : Radiation effects on thermal and mechanical properties.

Component : Magnet Dewar Thermal Shields

Project : GA - EPR

Function : Minimize radiation losses.

Description : OH coil: liquid nitrogen cooled shield located in vacuum

region between outer wall and helium container, no material stated.

TF coil: 0.5 cm copper thermal radiation shield cooled by gaseous

helium at 77 K.

B-14
Alternate Materials : None

Data Required : Effect of radiation at 77 K on thermal conductivity of

shield materials.

Component : Outer Wall - Magnet Dewar

Project : GA - EPR

Function : Vacuum containment, warm structure

Description : Stainless steel in all dewars. The TF dewars see an average bearing stress of 28.6 MPa (3,900 psi) from the centering

forces on the straight sections of the coils.

Alternate Materials : Aluminum, FRP

Data Required : Outgassing of FRP candidates under irradiation.

Component : Support Structure - Cold-to-Cold

Project : GA - EPR

Function : Contain electromagnetic forces, dewar supports.

Description : Magnet to dewar inner wall: fi berglass-epoxy or ceramic

standoffs covering less than 1% of dewar surface area support the inner

dewar walls from the TF coil cryostat. The space between the TF coil

and the cryostat wall is to be filled with epoxy to assure a uniform

bearing surface for transferring large conductor forces to the cryostat.

Magnet-to-magnet: out of plane loads on TF coils due to pulsed

poloidal fields are resisted by tie rods between the coils and contained

in the TF coil dewar. Four or five rods between adjacent magnets.

Referred to as hollow cylinders 50 cm o.d. and 40 cm i.d., no material

specified. Maximum cyclic stress is less than 140 MPa (20 ksi) so stainless, aluminum or composites are candidates.

Alternate Materials : Advanced composites if design requires excellent

fatigue life under high compressive stress at 4.2 K. Data Required : Cyclic fatigue at 4.2 K under irradiation - all materials

composite-metal joints and fatigue performance at 4.2 K.

Component : Support Structure - Cold-to-Warm

Project : GA - EPR

Function : Thermal standoff structural support transitioning from

300 K through the magnet dewars to the magnet coils.

Description : Primarily gravity support for TF coil. Alternatives are ceramic wafers in compression, a long hollow cylinder loaded in tension

or a long "bicycle chain" loaded in tension. Example given of two

stainless steel pipes, 18 cm i.d., 20 cm o.d., 100 cm long.

Alternate Materials : Industrial laminates, GPRP or advanced composites.

Data Required : Radiation effects on creep of metal supports. Radiation

effects on long term strength (aging effects) of composites.

B-15
Component : TF Coil Central Support

Project : GA - EPR

Function : Support of OH solenoid

Description : Central bore formed by butting together tapered sections

of the TF coil cryostats. OH coil and dewar placed inside bore,

apparently supported by a warm structure in contact with the OH dewar

walls. No distinct, load-bearing cold central support column is used.

Alternate Materials : None

Data Required : None

Component : EF Coils

Project : GA - DFPR

Function : Plasma field shaping

Description : Normal EH coils were selected for the initial conceptual

design but consideration was also given to superconducting coils.

Normal coils: similar to the EPR design, copper, water or oil cooled.

Anticipate radiation damage will require annealing at about 300°C at regular

intervals. Expect 10% increase in resistivity for neutron dose of 6 x 1020 n/cm2 at 80°F. 20 MPa (2900 psi) maximum stress on conductor. 32

coils, 150 turns of 5 x 5 cm conductors. Total copper weight 3576 metric tons^'

Superconducting coil: Also 32 coils, NbTi superconductor with 2-

component Cu/CuNi stabilizer. Superconducting composite 0.5 mm diameter,

braided. Current density maintained less than 90% of NbTi critical current

at 5.2 K based on radiation damage and combined toroidal and poloidal field

strengths at the coil. Stress on superconducting composite kept below

83 MPa (12 ksi) so added reinforcement is not needed. Low ac loss construction. Conductor is force-cooled inside thin 0.05 cm stainless steel casing.

Alternate Materials : Nb^Sn

Data Required : None

Component : TF Coil Conductor Composite

Project : GA - DFPR

Function : Plasma driving magnet coil

Description : Monolithic NbTi/Cu composite, 100 kA conductor current,

cryostatic stabilization. Conductor spirally wrapped around thin 120 cm wide, 0.275 cm thick OFHC copper strip which provides added stabilization,

14,000 NbTi filaments, 0.0175 cm diameter required for 100 kA and 30

kA/cm^.

Alternate Materials : Nb^Sn conductor, aluminum stabilizer.

Data Required : None

Component : OH Coil Conductor

Project : GA - DFPR

Function : Plasma ohmic heating

Description : NbTi with Cu/CuNi stabilizer braided into cable with stainless steel wire reinforcement. Average stress on the OH solenoid is

92.7 MPa (13.4 ksi). Design concept same as in EPR.

Alternate Materials : Nb3Sn conductor, graphite-epoxy filaments replacing

stainless steel for higher modulus, higher strength and reduction of

electrical shorting problems.

Data Required : Feasibility study of braiding conductor with graphite

epoxy reinforcement.

B-16
Component : Superconductor stabilizer

Project : GA - DFPR

Function : Cryogenic stabilization or conductor

Description : Two component Cu/CuNi stabilizing matrices are used in the

OH coils and in the EF coil if the latter is superconducting. Both use braided cables. The monolithic conductor for the TF coil apparently uses only copper for stabilization. The monolithic conductor is inset into a copper stabilizing element and the resulting conductor is spiral wrapped

around an OFHC copper strip 120 cm wide, 0.275 cm thick which adds additional stabilization. Apparently, this layer stabilization is required

for the 100,000 amp conductor rating.

Alternate Materials : Aluminum; fiber or dispersion-strengthened copper

or aluminum.

Data Required : Feasibility study on strengthening copper and aluminum

while maintaining conductivity.

Component : Strengthener for Superconducting Composite

Project : GA - DFPR

Function : Relieve stress on the superconductor

Description : OH Coil: Stainless steel wires intermixed with superconductor wires in the braided cables. Some indication of additional support in form of tube surrounding conductor bundle.

TF Coil: Stainless steel interleaved between layers of the 100,000 A

conductor elements. Entire coil inside stainless steel helium vessel.

Alternate Materials : Graphi te-epoxy strands replacing stainless wire for

improved strength, modulus and to reduce shorting problems. Interleaved

stainless in TF coil could be replaced by a high-modulus graphite-polymer

composite to achieve similar benefits.

Data Required : Combined effect of radiation and cyclic loading at 4.2 K on the mechanical and electrical properties of graphite-polymer composites.

Component : Electrical Insulation

Project : GA - DFPR

Function : Electrical standoffs turn-to-turn, layer-to-layer, coil-support

Description : Conventional EF Coil: mica, data indicate little degradation of mechanical properties at doses to lO^"^ n/cm^.

Superconducting EF Coil: combination of mica and dry fiberglass

cloth for turn-to-turn insulation.

Superconducting OH Coil: no information.

Superconducting TF Coil: no information

Alternate Materials : Glass or linen-reinforced industrial laminates.

Data Required : Effect of release of Wigner energy during warmup on the

electrical properties of insulators.

Component : Coil Forms

Project : GA - DFPR

Function : Support of windings, perhaps electrical insulation function

Description : Superconducting TF Coil: stainless steel helium vessel

also serves as coil form.

Superconducting EF, OH Coils: no information, probably same as EPR.

Alternate Materials : Composite materials for coil form of OH coil.

Data Required : Effect of irradiation, cyclic loading and release of

Wigner energy on structural and electrical integrity of composites.

B-17
Component : Eddy Current Shields

Project : GA - DFPR

Function : Reduction of ac losses due to pulsed coils.

Description : Aluminum at 15 K proposed as magnetic flux shield sur- rounding TP coils. Time constant of shield must be long compared to

poloidal field flux swing and long compared to shield hysteretic losses.

Alternate Materials : Copper, fiber-reinforced aluminum or copper to

resist electromagnetic stresses.

Data Required : Effect of 4.2 K radiation on electrical properties of

reinforced and unreinforced copper and aluminum.

Component : Magnet Dewars and Cyostats

Project : GA - DFPR

Function : Thermal insulation of superconducting coils

Description : TF coil dewar must be remotely opened for reactor maintenance. No material discussion, selection would be similar to the EPR

(stainless steel plus a non-metallic inner dewar wall for superconducting

OH coils).

Alternate Materials : Aluminum

Data Required : Helium permeability of non-metallic composites as a function of temperature from 4.2 K and as affected by irradiation including

release of Wigner energy on warmup.

Component : Support Structures

Project : GA - DFPR

Function : Resistance to electromagnetic and gravitational loads.

Description : TF coil gravity support: Fiberglass cylinder with length

greater than its cross-sectional area. A stainless steel ball joint on cold side provides limited radial coil movement.

Central support cylinder: In contrast to EPR, DEMO uses a 25 cm thick support cylinder which also provides partial gravity support for

the TF coils. Material not stated, but cross-sectional drawings imply

glass-reinforced epoxy. Centering force on cylinder given as 36,800

MT per coi 1

Out-of-plane loads: Resisted by circular tie rods between adjacent

TF coils enclosed in common TF coil dewars. Structural fatigue is a major

concern. No material stated.

Alternate Materials : Stainless steel, high-modulus composites

Data Required : Properties of thick-wall glass-epoxy including irradiation

effects and aging effects under sustained loading (creep). Fatigue

behavior of stainless steel and composites under irradiation at 4 K including Wigner energy release damage to composites.

B-18
UNIVERSITY OF WISCONSIN

Component : Superconductor composite for TF coils

Project : UWMAK I, II and III and TETR

Function : Produces toroidal field

Description : NbTi with copper stabilizer (I and II).

NbTi in copper surrounded by pure aluminum to stabilize (III and TETR).

Requirements : Mostly described in (I). Conductor is wound into a groove

in a disc. Maximum strain level allowed for the conductor is 0.002. It occurs where the conductor crosses the disc and where it leaves the winding, both are low field regions. This system puts the copper into yield

but the resulting resistivity is acceptable. Prestress of the conductor

is not necessary; it happens when the magnet is activated. Maximum stress

levels in the copper of the conductor and the stainless steel of the coil

form were calculated for (I). They are:

Magnet Condition Copper S. S.

N/m2xl0-b psi N/m2xl0-6 psi

After assembly 0 0 0 0

After cool down 20.7 3,000 -20.7 -3,000

After magnetic loading 103.4 15,000 413.7 60,000

Cold after removing -68.9 -10,000 68.9 10,000

magnetic loading @ 3.83 T central field

After warm up to -75.8 -11,000 75.8 11,000

room temperature

The amount of superconductor is chosen so as to carry full magnet current

at 5.2 K. This maximum temperature is that which the magnet will reach

with all current in the stabilizer, in liquid helium with a heat transfer

coefficient of 0.4 W/cm^ and edge cooling of windings. A large number

(34) of 10 kA superconductor joints are required in each coil. The

plan calls for lapping and clamping of conductor for 20 cm with a resulting maximum current path length in the copper of 0.5 cm. Joints are located in low field regions and have a predicted resistance of 10"^ Q. for a total heat loss of 376 W. There is no description of the solder

to be used.

Neutron damage is expected to be negligible in the TF coils for all

designs except I and TETR. Damage levels in II and III are predicted

10-S dpa/yr. They indicate that 10"^ dpa is necessary to degrade Ic

by 10% in NbTi and 5 x 10~5 dpa is necessary to cause a 10% increase in

the resistivity of OFHC copper at 4 K. Radiation levels of this latter

order may occur in TETR. In I additional copper was provided as was a regular schedule of annealing to handle the predicted 3 x 10-4 dpa/yr.

Data Required : Strain vs Ic characteristics of the conductors.

The combined effect of stress and neutron damage on the resistivity

and magnetoresisti vi ty of copper and aluminum.

An estimate of the pulsed loads due to the poloidal fields is

needed followed by fatigue studies on the superconducting properties

of the composite.

Thermal diffusivity of the conductor composite.

Magnetic diffusivity of the conductor composite.

Heat transfer coefficients.

Stress and resistivity characteristics of joints.

B-19
Component : Superconductor composite for OH and EF coils

Project : UWMAK I, II and III and TETR

Function : Produces the plasma driving field, the stability and

director fields.

Description : NbTi/Cu/CuNi composite strengthened with either cold

worked copper or stainless steel in I, II and III. Cryogenical ly cooled

(20 K w helium) pure Al on a stainless steel backing for some TETR magnets

Requirements : Stress effects are not as important as in TF coil design.

Effects due to the rapidly changing fields predominate. Details of the

separate designs are:

(I) EF coils are outside TF coils. Depending on the particular

location of the coil, it uses Al or Cu or Cu-SS to provide strength and

stability. Yield strengths assumed are 4000 psi for aluminum and 12000

psi for copper. The coils are layer wound, edge cooled, pancakes. Structure of superconductor to eliminate ec effects pretty much ignored except

to say CuNi will be used along with fine, twisted and transposed filaments

Rise time is 100 sec, dB/dt = .076 T/sec.

(II) EF coils are inside TF coils. Eddy current shielding treated

in detail requires CuNi clad filaments to be transposed inside a copper

matrix! Very little stainless steel used in the coils. Rise time is

10 sec, dB/dt = .57 T/sec. Design stress levels depend on type of winding

but are well within reasonable limits. No specific neutron damage calculations for these coils but the TF coil treatment indicates that they

do not expect a problem.

(III) EF coils are inside TF coils. Three conductor types con- sidered. A design much like ILL four components conductor. The conductor

component has the CuNi shielded superconductor in work hardened OFHC

copper for all designs. The other components are either work hardened

copper or 3/6 stainless steel depending on which conductor type is used.

Design stress at 4.2 K:

10% cold reduced OFHC = 1.24 x 10^ N/m^ (18 ksi)

316 ss = 2.11 X 10^ N/m^ (30 ksi)

8 2 One conductor type allows 2.07 x 10 N/m (30 ksi) in the copper and

3.5 X 10^ N/m2 (51 ksi) in the steel. The parts of the conductor are soldered together on 50% of their adjoining surfaces. The use of the

higher resistivity, worked copper greatly reduces eddy current losses

(p = 1.86 X 10-10 m (3 300 K, RRR - 110 at H = 0 and 30 at H = 9 T. They point out that if one designs to 60% of yield stress for a Cu/SS

composite, the copper can be in yield. Rise time is 15 sec, dB/dt = 0.57 T/sec.

Data Required : All data listed for the TF coil conductor.

Eddy current and hysterisis effects in the conductor.

Electrical resistivity and thermal conductivity as a function of

stress for the various coppers and aluminum which have been proposed.

B-20
Component : Coil bond to TF coil form

Project : UWMAK I, II and III

Function : Hold coil in groove in stainless steel coil form.

Description : Epoxy, filled around B-stage FEP-wrapped conductor.

Requirements : Bonding failure will permit conductor motion during field

changes leading to frictional heating. Quality control is essential here.

Micarta spacers must provide a clamping action to stop the spread of a crack.

Compressive stress < 8000 psi but not stated.

Threshold for observable radiation damage is 8000 M rads and 25%

degradation occurs at 50000 M rads thus they forsee no problem with

radiation effects on the mechanical properties. Because the coil is

wound into the form, each turn sees the maximum voltage to ground and

this can be as high as 10 kV ( 1 1 1-IV-A-9) . The effect of radiation

is less certain here.

Data Required : Tensile and shear strength.

Thermal contraction of metal to epoxy bond.

Effect of radiation on electrical properties.

Crack initiation and propagation.

Component : Spacers

Project : UWMAK I, II and III, all coils.

Function : To transfer loads between windings and to provide

cool ing channel s

.

Description : Micarta spacers (linen phenolic)

Requirements : TF: Need to provide a clamping action to keep conductor epoxy

bond to the stainless steel form in good shape. Prevent cracks from

spreading through the epoxy.

OH & EF: Micarta block spacers between pancakes. Whole structure

bolted together with Al bolts.

Data Required : Micarta is used in almost all low temperature magnet structures and

very little information has been gathered on its properties. All mechanical

and thermal tests are needed.

Component : Coil form

Project : UWMAK I, II and III and TETR

Function : Winding form for TF coil

Description : A series of pancakes grooved to accept the conductor.

Requirements : The TF conductor description sheet gives a complete

stress cycle. The solid pancake design is used because: electrical

connections can be easily made on the outside; internal heating due to

relative motion of conductor and coil form is minimized; stresses on the insulation are low; coolant channels are stable; bolting is easy.

Forging, machining, heat treating and winding techniques are unknowns.

(I and II) To limit design strain to 0.2% in the copper stabilizer

requires the use of AISI 316 with a design stress of 4.14 x 108 N/m2 (60000

psi). (Ill and TETR) allow 0.38% strain in the Al stabilizer and use Al

2219-T8T for the form. The design stress in the aluminum form is 3.11

X 108 N/m2 (45000 psi).

B-21
Data Required : Fabrication data appear to be most critical here

Eddy current heating of the aluminum alloy could be significant

Low temperature mechanical properties

Component : Insulation for dewars

Project : UWMAK I, II and III. Most detailed treatment is in II.

Function : Thermal insulation, not load-bearing

Description : Aluminized mylar (superinsulation)

Requirements : Will be subject to both neutron and gamma radiation.

(II) predicts a X energy of 0.137 ev/DT neutron and 0.884 ev from the

neutron. The maximum exposure is 0.9 Rad/sec or 690 M Rad in 30 yrs.

They feel that actual levels may only be 10% of this ( II-IV-H-57) . The

threshold for damage is given as 30 M rads and a 25% decrease in tensile

strength occurs at 120 M rad.

Data Required : This is one of the most serious materials problems. Immediate work is needed to determine the changes in thermal properties under

irradiation. There are indications that it may be much less radiation

tolerant than this design assumes.

Component : Magnet to outer dewar wall support

Project : UWMAK I, II and III, TF coils.

Function : Transfers lateral loads, i.e. those in the toroidal direction, from the magnet coil to the outer shell of the dewar when necessary.

Description : FEP struts

Requirements : If one magnet fails (I) those on either side would experience

a force > 8 x lO'^ kg tending to push them away from the failed magnet.

Smaller forces can occur in a normally operating system if exact symmetry

is not achieved.

Cold supports are not practical - too many are needed to prevent shear

at micarta spacer - as stainless steel coil form interface.

FEP struts are connected to outer dewar wall. 9 of 10 clear magnet

by 1 cm, the 10th makes contact. When magnet failure is detected (?),

external pistons drive all struts to full contact (see external support

structure IVD writeup). Wall at long strut is designed to deflect 1 cm elastically. This involves use of a nonlinear conical spring cap which

must be lubricated to slide against outer dewar wall in a vacuum!

(II-III-A-21 ).

Data Required : Compressive stress of composites

Spring material

Thermal conductivity

Component : External support structure

Project : UWMAK, I, II and III, TF coils

Function : Keeps TF coils from moving laterally.

Description : A cylinder piston element, driven by gas or fluid, which

pushes the outer case of the dewar against the linear (reversibly) in the case of loss of one coil. (II and III) use a continuous wall between the

TF magnets which serves as a secondary vacuum barrier as well as a support

for the coil system.

Requirements : A complex mechanical system

Data Required: No cryogenic parts

B-22
Component : Central support for TF coils

Project : UWMAK I, II and III, TF coils

Function : Supports weight of coils and radially inward force due

to magnetic field interaction.

Description : A massive stainless steel or aluminum alloy cylinder to

which TF coils attach.

Requirements : Supports TF coils by their straight sections.

Is cold, '^^ 4 K, and thus, must be isolated by a thermal barrier - not discussed in UWMAK design. The column also supports some of the

OH and VF coils.

TF coils hang on with a shear beam-shear pin attachment. Shear

stress is kept to < 1000 kg/cm2 in pins. Also stress on welds attaching

shear beam to magnet discs will be kept < 1000 kg/cm2. Bottom shear beam

does not have pins - used only for alignment.

When magnets are energized, the compressive force of magnets against

the central support exceeds the outward radial tensile force on the upper

shear beam. When unenergized, though, the total force will be ^ 615 M tons.

More massive structure in (II) because D's are not constant tension

and forces are larger and are often out-of-plane. Also there are radial

outward forces now. Limits for II: 414 x 10^ N/m2 (60 ksi) in tension

207 X 106 N/m2 (30 ksi) in shear. Shear pins ("high strength alloy")

345 X 106 N/m2 (50 ksi), shear stress, will se 3.1 x I08 in shear and 3.9

X 108 in tension.

Forces on the central cylinder: effective internal pressure 1.24 x 108 producing a tangential stress of 2,9 x I08 and a radial stress of

1.24 X 108. There is also an axial force of 20.8 x I06 kg/magnet to be

transferred to straight leg of magnet by the support beam resulting in a shear stress of 1.9 x I08 N/m2.

(Ill) central cylinder is aluminum alloy and coils, which are smaller and lighter than the other two designs, are attached by a key

system.

Data Required : Mechanical and fatigue properties of welds and shear pin materials

at low temperatures.

Eddy current heating in the cylinder.

The general problem of transfering forces from the TF magnets to

the central column is one that needs a good deal of investigation.

B-23

Appendix C. Research Capabilities

In the course of this survey a number of on-going, relevant lowtemperature materials research programs have come to our attention.

In addition a number of research laboratories which are interested in

this program have been identified. This section lists these and briefly

explains the type of research. We realize that we have probably missed

some projects at the national laboratories and have not done justice to

the many fine low temperature facilities at universities across the

country with whom we have not come in contact - regarding this program.

We urge any interested reader to call these deficiencies to our attention.

National Laboratories

Argonne National Laboratory (ANL)

Superconductor joining techniques

(J. Hafstrom)

ultrasonic welding - MHD magnets,

NDE procedures for joints,

acceptance criteria

Tensile, fatigue at 4 K: (E. Fisher)

Radiation of superconductors

(B. Brown)

Brookhaven National Laboratory (BNL)

Dielectric tape development:

(A. Muller)

Strain-insensitive sc conductors

(M. Suenaga, B. Sampson)

Thick epoxy castings:

(A. McNerny)

Radiation of materials:

L. Snead, A. Sweedler)

Los Alamos National Laboratory (LASL)

Electrical insulation:

(P. Chowdhuri)

effects of high cycle fatigue

on critical current of Cu,

Nb-Ti, and Nb^Sn

studies (fission, CP-5 reactor)

on Nb3Sn, intiated international

conference on subject in 6/77

hi-modulus polyethylene tapes

for sc cable; electrical, thermal,

mechanical property measurements

on insulators for screening

development of composite (NbsSn)

permitting more strain

large (10 ft x 1 ft-dia) castings,

insensitive to thermal shock, to

be used as insulators, matching

thermal expansion of Al

proton irradiation (30 Gev)

of superconductors, glassepoxies, and insulators

dc power transmission, consideration of cellulose paper and

plastics -dielectric breakdown

C-1
Composite dewars:

(C. King)

Conductor testing:

(R. Bartlett, D. Carlson)

Lawrence Livermore Laboratories (ILL)

Superconductor development and testing

(D. Cornish)

Welding, processing, specifications

(P. Landon)

Radiation

(D. Vandervoort, R. Scanlon)

Welding:

(C. Witherall)

Oak Ridge National Laboratory ORNL)

Solders:

(A. Morehead)

very large composite dewar performance and specifications for

energy storage program

Nb3Sn critical current testing,

effects of prestrain, acousticemission monitoring

Nb3Sn development, critical current, stress measurements - 3000 A capacity

research on 5083 Al and 21/6/9

stainless steel for specifications

for use at low temperatures

rotating target (14 Mev neutron)

irradiation of Nb-Ti, Cu,

and AISI 316 (300 K only)

Nb^Sn

weld rods and processes for

thick-section 21/6/9 stainless steel

shear tests of variety of solders

for superconductor bonding, also

NDE-related research

Stress-effects in superconductors

(D. Easton)

Radiation effects in insulators:

(J. Long)

Radiation effects in conductors

(R. Coltman)

critical current measurements as

a function of stress, field, and

conductor

assess gamma and neutron damage,

using fission spectrum, BSR reactor,

of electrical insulators and

potting compounds

irradiated Cu at 300 K and low

temperatures, using BSR reactor,

measured low temperature magnetoresistance

Conductor characterization:

(W. Fietz)

studies of conductor variability,

including critical current,

dimensions

C-2
stainless steel variability:

(A. Morehead, C. Brinkman)

variations of mechanical properties of multiple heats of AISI

304 and 316 for LMFBR (no low

temperature measurements)

Low Temperature Research at Other Laboratories

Closely-Associated with ERDA-MFE

Rocky Flats (Golden, CO)

Welding

:

(A. Liby)

University of Wisconsin

Fatigue, electrical

:

(H. Segal)

Thermal conductivity:

{I. Stone)

Dispersion strengthening:

(T. Hartwig)

Al stabilization:

(R. Boom)

University of California - Berkeley

Alloy development:

(B. Morris)

NBS Program

Structural materials for superconducting

machinery

(R. Reed, A. Clark, M. Kasen,

R. Tobler, et al .

)

Welding of structural alloys

(H. McHenry, R. Tobler)

chemistry, weld process, and

base-metal effects on welding

of 21/6/9 stainless steel

effect of plastic deformation

on electrical resistivity of Al

thermal conductivity of fiberglass/epoxy and selected non-metals

hi-strength, lo-resisti vity Al

development using precipitation

conductor design and processing

varying NbTi-Cu/hi -purity Al/Al

alloy ratios and geometries - for energy storage program

development of ferritic, lowtemperature alloy, mechanical

properties

large program (involving subcontracts) designed to screen structural alloys, provide base data

(mechanical, electrical, thermal)

of structural alloys and composites,

developed test procedures at 4 K,

published handbook, held workshops

five separate programs on weld

properties at cryogenic temperatures, including Ti alloys, Al

alloys, steels and Fe-Ni alloys

C-3
Properties of superconducting magnet:

materials, conductors and coils

(A. Clark, J. Ekin)

program for NSRDC superconducting machinery project; measured

critical current as function of

stress, field and temperature; now measuring fatigue properties of

conductor composites and critical

current, temperature and field

Superconductor standards:

(A. Clark, D. Read)

Cu-Nb3Sn processing:

(F. Fickett)

Composite development:

(M. Kasen)

establishing definitions and stan- ]. dard measurement procedures }5

research to assess thermal -mechanic,

processing effects on NbsSn precipi

tation in Cu, CuSn matrices

mechanical properties, development of graphite/epoxy structural dewars

Other Low Temperature (4 K) Test Facilities

ALCOA Center

(R. Kelsey, G. Barthold, E. Spuhler)

Battel le Memorial Institute;

(T. Collings, F. Jelinek)

Boeing Aerospace Corporation : (J. Boysen)

General Dynamics : (J . Christian, M. Campbell)

Martin-Marietta Corporation : (F. Schwarzberg)

Westinghouse Electric Corporation

(J. Wells)

mechanical properties, thicksection welding, composite development, processing, alloy development

thermal expansion, specific heat,

thermal conductivity, magnetic

properties, stress-fatigue, superconductor processing

mechanical properties, compositefracture experience

thermal expansion, thermal conductivity, specific heat, mechanical

properties, insulation-compositejoining-processing experience

mechanical properties; compositefracture experience

mechanical, magnetic properties;

processing- joining-fracture

experience

C-4
1'IBS-l 14A (REV. 7-73)

U.S. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO.

NBSIR 78-884

2. Gov't Accession

No.

3. Recipient's Accession No.

4. TITLE AND SUBTITLE

Technical Reports on Materials Studies for Magnetic Fusion

Energy Applications at Low Temperatures - I

5. Publication Date April 1978

6. Performing Organization Code

7. AUTHOR(S)

F. R. Fickett and R. P. Reed

8. Performing Organ. Report No.

9. PERFORMING ORGANIZATION NAME AND ADDRESS

NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. Project/Task/Work Unit No. 7360531

11. Contract/Grant No.

12. Sponsoring Organization Name and Complete Address (Street, City, State, ZIP)

Department of Energy

Division of Magnetic Fusion Energy

Washington, D.C. 20545

13. Type of Report Sc Period Covered

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

16. ABSTRACT (A 200-word or less factual summary of most si^ificant information. If document includes a significant

bibliography or literature survey, mention it here.)

The reports presented here summarize the work of the low temperature materials

research project for the first year of the program. The various projects are outlined and the research results are presented. The major thrust of the measurements has been the evaluation of the low temperature properties of stainless steel

base metal and welds, with particular emphasis on the nitrogen strengthened stainless steels. Some initial work has also been done on the production and properties

of nonmetallics, primarily industrial laminates, for low temperature applications.

A handbook of material properties is also planned. A survey of low temperature materials needs and problems related to magnetic fusion energy, performed by

NBS as groundwork for the program, is included as is a brief description of the

first workshop held in October 1977.

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

)

Data compilation; low temperatures; mechanical properties; physical properties;

stainless steel; welding.

18, AVAILABILITY [x] Unlimited

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I I Order From Sup. of Doc, U.S. Government Printing Office Washington, D.C. 20402, SD Cat. No. C13

I X ! Order From National Technical Information Service (NTIS)

Springfield, Virginia 22151

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(THIS REPORT)

UNCL ASSIFIED

20. SECURITY CLASS

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UNCLASSIFIED

21. NO. OF PAGES

364

22. Price

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