NBS TECHNICAL NOTE 609

National Bureau

of Standards

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Study of Cryogenic Storage

Tank Fatigue Life

Low Temperature Mechanical Testing

of AISI 304 and 310 Stainless Steels
NATIONAL BUREAU OF STANDARDS

The National Bureau of Standards1 was established by an act of Congress March 3, 1901. The Bureau's overall goal is to strengthen and advance the Nation's science and

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10MAL BUREAU Of STANBttK

10

UNITED STATES DEPARTMENT OF COMMERCE

Maurice H. Stans, Secretary U. 5 .NATIONAL BUREAU OF STANDARDS.* Lewis M. Branscomb, Director

NBS TECHNICAL NOTE 609

ISSUED AUGUST 1971

Nat. Bur. Stand. 'U.S.l, Tech. Note 609, 86 pages Aug. 1971 CODEN: NBTNA

Study of Cryogenic Storage Tank Fatigue Life

Low Temperature Mechanical Testing

of AISI 304 and 310 Stainless Steels

R. P. Reed, R. L. Durcholz, R. E. Schramm, and T. J. Patrician

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

Sponsored by

Naval Air Engineering Center

Philadelphia, Pennsylvania

under Project Order 9-8029

-v*

^ik

*"»eAu of

NBS Technical Notes are designed to supplement the Bureau's regular publications program. They provide

a means for making available scientific data that are

of transient or limited interest. Technical Notes may

be listed or referred to in the open literature.

For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, D.C. 20402

(Order by SD Catalog No. C13.46:609), Price 75 cents

Contents

Page

1 . Summary 1

2. Introduction 3

3. Test Materials, Specimen Preparation and Equipment ... 6

4. Measurements and Calculations 11

5. Experimental Results 15

5. 1 Mechanical-- Impact 15

5.2 Mechanical- - Tensile 41

5. 3 Transformation Products in AISI 304 43

5.4 Microscopic Observations 46

6. Discussion 57

6. 1 Impact- Fatigue Equipment 5 7 6.2 Fatigue Strength 57

6.3 Welding Effects 59

6.4 Thermal Cycling Effects 60

6.5 Cold Working Effects 60

6. 6 Fatigue Fracture Stress 61

6. 7 Deformation Energy Density 61

6. 8 Martensite Formation in AISI 304 62

6.9 Tensile Data 62

6. 10 Microscopy 63

6. 1 1 Possible Additional Programs 64

7. References 65

Appendix: "Apparatus for Impact- Fatigue Testing" 68

in
List of Tables

Page

1. Nominal composition of AISI 304 and 310

Stainless Steels 6

2. Impact-fatigue results, AISI 310 16

3. Impact-fatigue results, AISI 304 18

4. Tensile results 42

5. Summary of deformation structure and phases

present for deformed alloys 56

IV
List of Figures

Page

1. Sample geometries for tensile and impact tests 7

a. Uniaxial, thickness = 0. 1 in.

b. Biaxial, thickness = 0. 1 in.

c. Triaxial.

2. Electron microscopy specimen, thickness = 0. 1 in. ... 10

3. Determination of E, a , and <J from a stress

u y

strain curve 12

4. S-N curve for AISI 310, uniaxial 2 3

5. S-N curve for AISI 310, biaxial 2 3

6. S-N curve for AISI 310, triaxial 24

7. S-N curve for AISI 304, uniaxial 24

8. S-N curve for AISI 304, biaxial 25

9. S-N curve for AISI 304, triaxial 25

10. S-N curve for AISI 304, uniaxial, welded 2 6

11. S-N curve for AISI 304, biaxial, welded 26

12. Fatigue life at 10 2 cycles 27

13. Fatigue life at 103 cycles 2 7

14. Fatigue life at 104 cycles 28

15. Comparison of impact-fatigue data for annealed

AISI 304 with that obtained in tests using tensioncompression at low strain rates (Refs. 15-19). The

17

triaxial samples of Krempl had a stress concentration factor, K , of 3. 3. ... 29

16. Fatigue-life data (Ref. 20-22) for cold-worked (to 210,000

psi tensile strength) AISI 304. Tests were conducted by tension-compression at low strain rates. The biaxial samples

had a stress concentration factor, K^., of 2. 75 to 3. 9 . . 29

17. Impact -fatigue data for cold-worked uniaxial specimens.

The 304 contained 50% martensite after straining .... 30 v
Page

18. Fracture stress for AISI 310, uniaxial 31

19. Fracture stress for AISI 310, biaxial 31

20. Fracture stress for AISI 310, triaxial 32

21. Fracture stress for AISI 304, uniaxial 32

22. Fracture stress for AISI 304, biaxial 33

23. Fracture stress for AISI 304, triaxial 33

24. Fracture stress for AISI 304, uniaxial, welded 34

25. Fracture stress for AISI 304, biaxial, welded 34

26. Deformation energy density for fracture of

AISI 310, uniaxial 3 6

27. Deformation energy density for fracture of

AISI 310, biaxial 36

28. Deformation energy density for fracture of

AISI 310, triaxial ... 37

29. Deformation energy density for fracture of

AISI 304, uniaxial 37

30. Deformation energy density for fracture of

AISI 304, biaxial 38

31. Deformation energy density for fracture of

AISI 304, triaxial . 38

32. Deformation energy density for fracture of

AISI 304, biaxial, welded 39

33. Summary of deformation energy density for

fracture at 103 cycles 39

34. Elongation in one inch for fracture at 103 cycles 40

35. Uniform reduction of area for fracture at 103 cycles. . . 40

36. Parameters during an impact -fatigue test of

AISI 304at297K. 41
Page

37. True stress-true strain curves for tensile tested

AISI 310 44

38. True stress-true strain curves for tensile tested

AISI 304 44

39. Representative tensile test data for annealed AISI 310 . . 45

40. Representative tensile test data for annealed AISI 304 . . 45

41. Formation of martensite during tensile and impact tests . 46

42. Temperature dependence of martensite formed after

impact-fatigue fracture in AISI 304 47

43. 304 specimen strained ~ 3%; "A" grain with slip on primary

slip system only, "13" grain showing slip on two systems.

200X 50

44. 310 stainless steel strained~ 3%. Arrows indicate slip

systems; (a) are dislocation interactions at dislocation

pile -up where two slip planes intersect and (b) is a stacking fault separating two partial dislocations.

40, 000X 50

45. Cell dislocation structure of 310 stainless steel after

strain of ~ 9%. 40, 000X 52

46. Deformation twin in 304 stainless steel afters 9% strain.

100, 000X 52

47. 304 stainless steel strained ~ 9%. (a) Electron micrograph at 100,000X; (b) selected area diffraction,

e (hep), a' (bec), and y (fee) 53

48. 304 stainless steel strained- 3% at 76K; &' and Y present but no s . 16.000X 54

vn
STUDY OF CRYOGENIC STORAGE TANK FATIGUE LIFE

Low Temperature Mechanical Testing of AISI 304 and 310 Stainless Steels

R. P. Reed, R. L. Durcholz, R. E. Schramm, and T. J. Patrician

Two 300 series stainless steels were tested for impact-fatigue

life and tensile properties as a function of temperature, percent transformed phase (martensite), stress level, and specimen geometry.

These alloys were studied to predict the fracture characteristics of

parent material and weld joints in large cryogenic dewars which are

subjected to periodic stresses. Normally, AISI 304 is employed in

the construction of such dewars. Under fatigue loading conditions at cryogenic temperatures, the behavior of AISI 304 is complicated by

the gradual transformation to a martensitic structure. The influence

of this transformation on the fracture characteristics was studied.

For this purpose, a new impact-fatigue test and necessary equipment

were developed.

The fatigue strength of AISI 304 and 310 exhibit similar trends,

but at a given fatigue life and temperature, the fatigue strength of

AISI 304 is slightly superior. AISI 310 displays an endurance limit

of about 20, 000 psi;* the 304 limit is 25-30, 000 psi. However, the

lowest endurance limit (weakest) specimens appear to be the triaxiallyloaded AISI 304 weld specimens which have an endurance limit near

10, 000 psi. Little temperature-dependence of the impact-fatigue properties was observed.

Key words: Fatigue; fracture; impact; low temperature; mechanical

property equipment; stainless steel.

* For ease in interpretation all stress values in the text are expressed

in units of psi. Conversion to other stress units are

10 3 psi = 0. 703 kg/mm 3

= 6. 90 X 10 6 N/m3

.
1. Summary

The principal results of this report are briefly summarized

below.

1. The impact-fatigue strength of AISI 304 is superior

or equal to the impact-fatigue strength of AISI 310

under all conditions in the temperature range 7 6 to 294K.

2. The endurance limit generally appears at about 10

cycles. At all temperatures (294, 195, 76K), the

limit can be roughly approximated by using the room

temperature tensile yield strength. Possible critical

exceptions are under triaxial stresses at 195 and 76K,

where the endurance limit is less than the room temperature yield stress. In these cases, a considerable extra

safety factor in designing is desired.

3. Welds exhibit less strength under impact-loading

conditions than the parent material: under uniaxial

loading conditions, approximately a 10% reduction was found; under biaxial conditions about a 15%

reduction was found. Although triaxial loading

conditions could not be tested, on the basis of our

results, it is expected that this configuration

(weld, triaxially applied stress) at low temperature

would prove the weakest link in any structure.

4. Cold working (11%) at 76K produces a 35-50% increase

in AISI 310 fatigue strength and a 40-55% increase

in AISI 3 04 fatigue strength.

5. The transformation to martensite during repeated impact

cycles has little deleterious effect on fatigue life. Under

uniaxial and biaxial loading conditions, the transformation
is beneficial, increasing the fatigue strength.

6. In the case of AISI 3 04, there is a good possibility of

monitoring the fatigue life by measurement of the

amount of ferromagnetic phase (martensite) by simple

magnetic techniques.

7. Two studies remain in order to insure the mechanical

stability at cryogenic temperatures of AISI 3 00 series

stainless steels, undergoing partial martensitic transformations. These are: the impact-fatigue strength

in the temperature range 4 to 76K and the influence on martensitic transformations and mechanical stability

of differences in weld techniques.
2. Introduction

The objective of this study is to predict the life of AISI 300

series stainless steel parent material and weld joints when subjected

to periodic large, sudden stresses in the temperature range 300 to 76K.

At present, knowledge concerning the endurance life and effects of martensitic transformation on fatigue strength is not available. To enable

safe transport of liquid oxygen in large quantities (400 gallons and up),

such knowledge must be obtained» since most dewar systems are now

fabricated with AISI 304 stainless steel. Actual conditions were simulated to provide the most useful information to dewar design.

Brittle failure, from analysis of accident investigations, is

listed as a factor of primary importance in causing system failures in

cryogenic equipment. The continual exposure of liquid oxygen tanks to high stresses (for example during the carrier on-deck delivery) poses

a serious question of possible fatigue -transformation failure. Failure

may occur for two reasons:

(1) When normal 304 stainless steel is stressed above its

yield strength at temperatures below about 200K, a partial

crystallographic transformation from austenite (face-cen>-

tered cubic structure) to martensite (body -centered cubic

structure) occurs. There is an associated local volume

expansion of 4% caused by the transformation. This local

volume expansion promotes very high stress concentrations in material that is not ductile, such as weld areas.
(2) Normal fatigue failure, under periodic stress above the

yield stress, also may occur. The endurance or fatigue

limit of the combination of 304 steel, weld joint, and large

stresses has never been tested; and to our knowledge, no

similar circumstances have been duplicated in previous

usage. Therefore, prediction of tank and safety is very speculative.

Various types of fatigue tests have been devised and are identified in three general categories.

1. tension-compression

2. bending

3. torsion

The test specimen is subjected to constraints which fall into one of two

c la s s e s

,

1. constant load, bending moment, or torque.

2. constant strain, bending deflection, or twist.

To keep test times to a reasonable duration, the repetition frequency is

2 4: / usually on the order of 10 - 10 cycles/min. High frequencies can

result in an artifically short life due to sample heating. There are,

however, indications that a low repetition rate, without any heating

problem, can also lead to a short life. 2-13 Many investigations have shown that the strain rate to which a metal is subjected does affect the microstructure and this, in

turn, is reflected in the mechanical properties. A review of many of

14 these influences has recently been published. Of course, there are

actual situations which are approximated poorly or not at all by the

Superscripts indicate literature references listed at the end.
standard tests. Among these is the sudden impact loading or jolt delivered at a frequency of about one per hour or day, such as might

happen during handling, transportation, or in the transfer of fluid

through pipes.

To approximate better the actual environmental loading conditions, a new type of fatigue test was devised. This new test was performed on a commercial impact machine, modified to apply repeatedly

the same impact load to the specimen (see Appendix). This report describes the tests carried out in the temperature range 76 to Z97K,

documents all test results, and discusses the significance of the data

in terms of use in designing cryogenic equipment.
3. Test Materials, Specimen Preparation and Equipment

We tested two austenitic stainless steels, AISI 304 and 310;

Table 1 gives the nominal composition of these alloys. AISI 310 is a crystallographically stable alloy with a face-centered cubic (fee) structure, while annealed AISI 304 is known to undergo a martensitic phase

transformation from fee to hexagonal close-packed (hep) and bodycentered cubic (bec) structures. This transformation always occurs

on deformation of the 304 steel below room temperature, and, depending on the exact composition, may occur spontaneously on cooling to

5 the vicinity of 7 6 K.

Table 1. Nominal Composition of AISI 304 and 310 Stainless Steels

AISI Composition (%)

Type

304

310

Cr Ni Mn

0.08 max 18.0-20.0 8.00-11.00 2. 00 max

0.25 max 24.0-26.0 19.0 -22.0 2.00 max

Specimens were tested in three configurations:

1. flat, smooth, (uniaxial).

2. flat, notched (biaxial).

3. round, notched (triaxial).

The notches cause the load to be broken down vectorially into longitudinal and transverse components. In this way the flat specimens are

subjected to biaxial stresses while the round specimens see three-way

or triaxial stresses. Dimensions are shown in Figure 1. The flat

samples were cut from sheet parallel to the rolling direction, and the

round specimens were produced from bar stock. By computing the

stress concentration factor, Kt

, from the formula
f

-10 NC Thread

Figure 1. Sample geometries for tensile and impact tests: a. Uniaxial,

thickness = 0.1 in; b. Biaxial, thickness = 0.1 in; c. Triaxial.
/

~ distance between notches K; = J-2.

t notch radius

it has the value 4. 2 and 4. 5 for the biaxial and triaxial geometries,

respectively. After machining, the flat samples were mechanicallyground and polished. All samples were then vacuum annealed at 1000°C

for one-half hour and air cooled. The majority of tests involved these

materials as annealed, but some samples were subjected to further

treatment to check the effect of chemical and physical variables on the

impact fatigue life.

AISI 304 uniaxial and biaxial specimens, prepared from a melt

which transformed spontaneously on cooling, were thermally cycled

ten times between 297 K and 76 K. This produced about 2-7% martensitic products prior to the impact tests. Also, several 304 uniaxial

specimens were held at 81 3K for three hours in air to test possible

effects of sensitization, or the formation of intergranular carbides.

To get a quantitative idea of the influence of cold-work on impact-fatigue life, uniaxial specimens of both alloys were mounted on a tensile-testing machine and stressed at 76 K to 11% strain; in the 304

this produced about 50% martensite. Those samples to be tested at 76K were maintained at that temperature until broken; the samples

for the 297 K tests were warmed just prior to use. Welding introduces another very important variable. Welded

plates were prepared from which uniaxial and biaxial samples were

machined. Notches were ground both in the weld area and at two positions in the heat- affected zone. The technique was as follows:

material: AISI 304 stainless steel, 1/8 inch thick plate;

AISI 308 stainless steel weld rod, 0.035 inch

diameter wire.

preparation: plates were machined, deburred, electro-
cleaned, rinsed, then electro polished. The

caustic electrocleaning solution consisted of 8 oz.

of Oakite #160 per gallon. This was run at 160°F (344K)

at 6 volts. The plates were electro polished in 70%

phosphoric acid, 30% sulfuric acid at 140°F (333K)

using 9 volts for 4 minutes. Finally, the plates were

again rinsed in water.

welding procedure: weld was butt- type in a flat position,

using machine welding techniques. One pass, using an average welding speed of 18 inches per minute, was made over a length of 2 inches. A 0. 09 inch

gap was maintained by tacking 3. 5 inch intervals.

The plates were held down pneumatically and a Pandjiris weld positioner was employed.

welding atmospheres: a mixture of 3 0% He and 7 0% Ar

at an average flow of 10 cubic feet per hour was used.

inspection and finishing: weld beads were inspected using

standard radiographic techniques for porosity. Welds

of poor quality were rejected. Weld beads were then

machined off, the sample was polished, and magnetic

permeability measurements made using a Magne- Gage.

No further heat treatment was used prior to testing.

The samples used in the microscopic analysis were of the univ Trademark. Precise specification of the material employed has been

necessary to make the description of this apparatus sufficiently meaningful. Identification of this material or its manufacturer by the National Bureau of Standards in no way implies a recommendation or en- dorsement by the Bureau. Furthermore, use of other trade names in

this paper is for the sake of clarity and does not in any way imply a recommendation or endorsement by the Bureau.
axial geometry, but thinner and with a wider reduced section (Fig, 2).

They were annealed as noted for the impact specimens.

The impact-fatigue tests were conducted using a Riehle impact

machine which had undergone major modifications; the details on the

automation of this device and the cryogenic system used with it are presented in a paper to be published in Section C, Journal of Research, NBS

(See Appendix). A standard Instron testing machine served for the

tensile tests. The amount of ferromagnetic bcc martensite produced

daring these tests was recorded by measurements with a Magne-Gage,

an instrument consisting of standard bar magnets affixed to torsion

balance equipment.

The microscopic studies were done with a standard optical metallograph and electron microscope.

375 Dia.

Figure 2. Electron microscopy specimen, thickness - 0. 1 in.

10
4. Measurements and Calculations

In the documentation and description of our data, a number of

symbols and definitions are employed. The following discussion identifies these.

Definitions of Symbols Used

A Sample area outside necking region after fracture.

AF Sample area at fracture.

A Initial sample area. E Elastic modulus.

EF Absorbed energy density required to deform sample.

e Elongation in 1 inch.

Li Sample load.

t Gauge length just before necking.

Hjf Gauge length after fracture.

t Initial gauge length.

N Number of impact cycles to fracture.

RA Reduction in area before necking.

RAN Reduction in area at neck.

T Temperature.

V Sample volume after fracture.

V Initial sample volume.

e Strain rate.

e

E Engineering strain.

e-|- True strain.

Qa Applied or engineering stress. aF Fracture stress.

CTT True stress.

c^ Ultimate or tensile strength.

Oy Yield strength (. 002 offset).

11
Direct measurements are made on: A, AF , A , L, l F , l , N, T, and e During the tensile tests, a strip chart recorder was used to con

tinuously note the engineering strain, £

E , as a function of the sample

load, L. The engineering stress, cr A , is given by

L

CTa = A~

The stress-strain curve was then easily constructed from the strip

chart record and used to determine the ultimate and yield stresses,

<7U and cr y) and the elastic modulus, E. Figure 3 illustrates this

graphically.

°"a

Fracture

E = slope of elastic portion of curve

-*| ^—0.002

*i >

Figure 3. Determination of E, o , and o fr om a stress-strain curve.

12
Reductions of area and elongation are determined directly from

sample measurements.

RA = ^A = ! . A

Ao Aq

RAN = ^ = 1 " - Ao Ao

In the notched samples, both biaxial and triaxial, little, if any, necking occurs so A = AF and RA = RAN.

Fracture stress is L

CTf= a;

For the tensile tests this is most easily determined from

Of 1-RAN

While it is true that dislocations and vacancies are introduced

into the crystalline structure during deformation, it is nearly true that

the volume of the Sample remains constant, i. e.

So

V = V • IA = 4 A '

a _ Ao

in '' A

Since: £E = TT = ^

- 1

it follows that 1 + e E

and e E =

1-RA

RA

1-RA

13
From this the true strain can be determined.

The true stress is L

GT = A •

Again assuming constant volume:

or = ^ ^ =CTa 7" = a A(l+e E ) . Aq A *

There are basically two stages in the fatigue life of a sample:

1. Uniform strain.

2. Necking followed by fracture.

Most of the energy is absorbed during the uniform strain portion.

Once necking has begun, almost all further energy is concentrated in that

small volume and fracture quickly follows. To make calculation possible, it is assumed that the only significant absorbed energy is that

required to just bring the sample to the point of necking. The density

of absorbed energy required to deform the sample is then:

L(4-4 ) = L(X-J& B ) = _L l-l

F V V© Ae ip

ct a e E

14
5. Experimental Results

5.1 Mechanical- -Impact

The impact -fatigue data for AISI 304 and 310 stainless steels are presented in Tables 2 and 3 and are plotted as S-N curves (applied

stress vs. number of cycles) in Figures 4-11. The fatigue life curves

in Figures 12-14 give a comparison of these two alloys for durations of

10 s

, 10 , and 10 4 cycles. Some results are immediately apparent from

these curves:

(1) A decrease in temperature increases the impact -fatigue life.

(2) Fatigue life of uniaxially stressed specimens is generally

longest, followed by biaxially stressed specimens. Specimens

stressed under triaxial conditions are usually weakest.

(3) In the uniaxial configuration, the fatigue strength of 304 is superior to that of 310. At room temperature, the difference is about 5-15% and at 76K it is 25-30%. In Figures 7 and 8, the

data for the 3 04 samples that were sensitized or precycled to

7 6K fall in line with the annealed points, indicating very little or no effect on the impact-fatigue behavior by these thermal

treatments. This agrees with previous work on sensitized 3 04

23 subjected to high strain rates.

(4) The shape of the curves in Figures 10 and 11 for the 304

weldments exhibit the same general behavior as the annealed

material. In the uniaxial geometry, the strength is slightly re- duced by welding; however, this decrease depends on temperature - about 20% at 297K and 8% at 76K for a life of 10 ? cycles.

The results are similar for a 10 4 cycle life (Figure 14). The

strength of biaxial samples is nearly unchanged at room temperature but reduced at 76K. The most remarkable thing here is the insensitivity of fatigue life to the relative position of weld

and notch. In Figure 1 \ a single curve at each temperature

could be drawn through the data points for all three notch positions.

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22
260 x

10'

220

Ql

co" CO

UJ

or

I-COQ

UJ_l

0_

0-<

180

140

100

60

AISI 310

Uniaxial

a - 76 K A - 195

K

O - 297 K

Fi:

I 10 10 10 10 IMPACT CYCLES to FAILURE

rure 4. S-N curve for AISI 310, uniaxial.

'if 140

120

180

160

100

80

CO

CO

UJ

or

i-00Q

UJ_l

Q_

Q_< — 60

!0~

240 xlO3

200

160 CO

CO

UJ

or

h-coQ 120

0_

o_<

80

40

...

Al 31 3IC

axial

) a

A

- 76

- 195

- 297

K

...

B o

KK ... r 95 K

....

Nj

76 1 <.

...

C]S N

....

' ~^^z

c 1 -

2£ 7 <—

J ...

I 10 I0

C

I0 J 10 IMPACT CYCLES to FAILURE

Figure 5. S-N curve for AISI 310, biaxial.

140

20

00

CO

CO

UJ

orCOQ

UJ - 60

CL

Q.<

-.:: 40

io23
200 x 10'

160

$ 120

L±J (T

I-coS

80

_l

0.<

40

300x10'

260

220

CO

co

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or

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a.<

180

140

100

60

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...

Al S! 3I<

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D —

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- 76

- 195

- 297

K o K -

/ —76 K

...

~lx J— 195 K

...

/""~»«^

...

2 97 K—J r h4sITM^V.

...

...

140

120

:::: 100 — 80

40

CO

CO

UJ

or

i-coQ

UJ

_l

o_

o_<

10 10 10 10 IMPACT CYCLES to FAILURE

Figure 6. S-N curve for AISI 310, tr iaxia 1.

10

i AISI

i i i 30

ii 4 ...

c] U liaxio. 1 D - A - 76 K

...

O - 297 K

1

1 1

3 H V " i b K, sensmzea ® - 297 K, sensitized

7(~ 1/ _ - 1-Ji

....

1 • - 297 K, pre ;ycledto76K

...

z sJH ....

s v.A

* A

sS y

"l"

U

,-19!)K

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-- 200

180 - 160

CME

E

140

120

CO

CO

UJ

or

h-coa

ioo y

o_

0_<

80

- 60

10 10 10 10 10 IMPACT CYCLES to FAILURE

Figure 7. S-N curve for AISI 3 04, uniaxial.

24
240

10 10 10 10" IMPACT CYCLES to FAILURE

Figure 8. S-N curve for AISI 3 04, biaxial.

200 x I0~

160

CO

CO

Ld

cc

\-coQ

UJ

_l

Q.

CL<

120

80

40

._.

Al SI 30

riaxial

4 -

T - 76

A - 195 O- 297

K

K ...

76 K

....

-IS)5 K -

291\( '

....

--

10

140

120

100

E

E

CO 80 CO

Ul

q: 60^

Q

UJ 40 cl

0-<

20

I0

2

I0

3

I04

I05 IMPACT CYCLES to FAILURE

Figure 9. S-N curve for AISI 3 04, triaxial.

25
260 x

10'

220

CO

co

UJ

crH

COQ

UJ

_l

QL

a.<

180

140

100

60

Al

.- 1 lr

SI 30-

liaxial elded

4

.... urw

1 D

... A

76

- 195

- 297

K

...

I \ [ o K

1 ....

l< 55 K -76 K

..

A ...

( i

> v A

....

:

2£>7 K—y h ...

I

160

i40

0>

CO

UJ

or -#oo toa

UJ 80 a.

o_<

60

10 10 10 10 10 IMPACT CYCLES to FAILURE

Figure 10. S-N curve for AISi 3 04, uniaxial, welded.

240x10"

200

ina.

CO

co

UJ

or

\-coQ

UJ

_1

0_

D_<

160

120

80

40

AISI 304

Biaxial,

Welded

- 76 K O - 297 K A = l/8

" off weld

" B = '/|6 off weld

10" 10 10 IMPACT CYCLES to FAILURE

--- 160

::::

| 40

CVJE

E

120

100

-- 80

:: 60

CO

CO

UJ

(Z

h-coQ

UJ

_l

0_

0_<

40

ioFigure 11. S-N curve for AISI 304, biaxial, welded.

26
200x10

100 200 TEMPERATURE,

K

300

Figure 12. Fatigue life at 1 0~ cycles.

190x10"

[70

150

130

COQ

y no

90

70

50

AISI 304 .Annealed

1 3

3 Cycles AISI 3I0 ,Ann ealed

uniaxial

-

s

biaxial

^-N.

""N., ^^ ^vJ

1 riaxial

-E"TM

~~ —-- "— ^T ^^^^

I30

20

I00

90

CO

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flO cr

1-coa

70 UJ

1 na 60 <t

50

40

100 200 TEMPERATURE,

K

300

Figure 13. Fatigue life 10 cycles.

27
140x10"

120

10 Cycles

I

I AISI 304, Welded

AISI 304 .Annealed

AISI 310 .Annealed

- 90

- 80

70

- 60

--•«-.,. = 50 -j

40

30

20

100 200 TEMPERATURE,

K

300

Figure 14. Fatigue life at 10 cycles.

Figure 15 allows comparison with some previous fatigue life 15-19 , , measurements that were conducted in tension-compression.

While the impact and conventional fatigue curves do have the same

general structure, there are some differences:

(1) The impact S-N curves fall at a higher stress level.

(2) The fatigue limit on impact, uniaxial stress appears to occur at or above the yield strength (0. 002 offset) as determined from tensile tests; in conventional fatigue the limit is well below this level.

These factors are probably related to the increase in yield strength

with increasing strain rate. 20-22

Data from the literature show that cold-work has a significant effect on the mechanical properties of 18 Cr-8 Ni steels, and

Figures 15 and 16 show this effect on fatigue life. To make a controlled

28
260 x I0 3

220

180

to

co

UJ

°£ 140

COQ

UJ 100

60

20

III

jj AISl 3U4

Annealed -

1 |:;| r— Band includes all present / impact fatigue measurements

-

N

— M i -L: m

I

I A

I i

V

w~~

\ \ jj.r-Fushimi, \r\/ uniaxial,stal. 104 K

L--AIS 1 304

7*s McDowell I i i " t'

Berling.etal.

uniaxial, 2951 J

**

uniaxiah

4 Brie

[flJAISI 304 L

k.etal. uniaxial / ^233K +

tr ax iaj, 295 V-< r—

V 295 K

80

160

140

120

100

80 S

bO

40

20

10 10 10 10 10 10 10 IMPACT CYCLES to FAILURE

Figure 15. Comparison of impact-fatigue data for annealed AISI 3 04

with that obtained in tests using tension-compression at low strain rates

(Refs. 15-19). The triaxial samples of Krempl17 had a stress

concentration factor, K , of 3. 3.

260 x I0 3

10 10 10' 10" IMPACT CYCLES to FAILURE Figure 16. Fatigue-life data (Ref. 20-22) for cold-worked (to 210, 000 psi

tensile strength) AISI 3 04. Tests were conducted by tension- compression

at low strain rates. The biaxial samples had a stress concentration factor,

Kt

, of 2.75 to 3.9.

29
check of the effect of cold-work on 3 04 and 310, uniaxial specimens

were strained 11%, as noted above, and then impacted in the usual

manner. Figure 17 shows the results and comparison with Figures 4

and 7 indicates a considerable strengthening at both 297 and 7 6K. At

a life of 10s cycles, the improvement in the applied stress is about

40-50%. Also, the superiority of AISI 3 04 over 310 is again evident.

Figures 18-25 show that the fracture stress for all materials

and all configurations follows a pattern very similar to the applied

stress with two exceptions:

(1) The 310 biaxial data (Figure 19) lie on a single curve for

all temperatures.

(2) The 3 04 biaxial data (Figure 22) show a small separation

between the precycled and annealed room temperature curves.

350xl0~

300

Uniaxial

(strained 11% at 76K)

250

CO

CO

l±J <r 200 Q

UJ

150

100

~

HI D 76K, AISI 304

76K.AISI 310 O 297K, AISI 304 • 297K.AISI 310 I 1 |.

•\ I

-76K, AISI 310

III! -29 *IO 1 D4h

-29"'K , ) fr Alt>l ill

240

220

200

iE

180 E

160 _ co

CO 140 l±J rrCO 120

100 £

a.<

80

60

io icr io io IMPACT CYCLES to FAILURE

Figure 17. Impact-fatigue data for cold-worked uniaxial specimens.

The 3 04 contained 50% martensite after straining.

30
560 x I03

480

400

320

U- 240

160

Iffl AISI 31

Uniaxia

1

1 A

O

- 76 K

n

195 K

- 297 K ks

-76 K

...

195 K

1

1

1 I—

i

...

i

297K--4

c '^ -

...

380

340

10 10 10 10 IMPACT CYCLES to FAILURE

300

en 260 en

-Q* 220

80 H

o

<

140

10

Figure 18. Fracture stress for AISI 3i0, uniaxial.

400x10-

320

$2<

Ld

CCCO

UJ

cc 160

<

Ll I

Tn 280

k

A, SI 310

axial

- 240 \

B D- 76 K A- 195 K O - 297 K

::: 200 \\ I Nl - 160 IN - 120

C

... C 4 80

... 40

-u

I 10 I0C10 10 I0J IMPACT CYCLES to FAILURE

Figure 19. Fracture stress for AISI 310, biaxial.

31
400 x 10'

320

CO 240

Ld

CE

bJ 160

a:

h-o

<

ff 80

T

J- AISI 310

Triaxial

- 76 A - 195 O - 297

K

K I

T

^76 K

•

/

}; /J >97 h

J- -195 K

'

280

10 10* 10 10* IMPACT CYCLES to FAILURE

240

200

160 CO

COUJ

(T 120 CO

LiJ rr

=> 80 H

O<

cr 40^

I03

Figure 20. Fracture stress for AISI 310, triaxial.

560 x io480

inCO

LJ

CEH

co

Ld

cr

h-o

<

cr

Li.

400

320

240

160

a

Mill 76 K

-

c O - 297 K \ ® - 297 K, sensitized _

• -

i i 297 K, prec ycled 1o76K

cv A 76 K

: m

A

C : AISI 304

Uniaxial y

-2 97 1

fil - ^•v \, l

T

-

380

340 — 300 -*

260

220

180

140

10 10' 10° !0H IMPACT CYCLES to FAILURE

Figure 21. Fracture stress for AISI 304, uniaxial.

32
400 x 10-

320

<f> 240

CO

LU

or

GO

Ld

(TID

l~O

<

160

80

AISI 304 1 -

Biaxial A- 195 K O - 297 K

- 76 K, pre

1 1 II f

cycled to 76 K _

--H95K -

I 2S 7Y T

•a. /-76K

|

..Trf - 297 K, precycled at 76 K—

280

240

200

160

120 GO

LU

or

ID 80 <

cr 40

I 10 10^ 10° 10* I0D IMPACT CYCLES to FAILURE

Figure 22. Fracture stress for AISI 304, biaxial.

400 x 10

10 io" io" io IMPACT CYCLES to FAILURE

Figure 23. Fracture stress for AISI 304, triaxial.

33
480x10-

400

320

CO

Ld

O. 240 H

O

<

rr 160

80

Al 51 304

liaxial, elded

...

[ — 76 K

T" Uf

195 K

...

'il

'a*

P- 76 K A - 195 K O- 297 K 297 K- -'

Illlll

...

(Did not break at weld) Y

1

1

1 1 1 1 M II

340

- 300

220

180 CO

UJ

or 140 ^

O<

rr H ioo u.

60

10 10^ 10° 10 IMPACT CYCLES to FAILURE

Figure 24. Fracture stress for AISI 3 04, uniaxial, welded.

400 x 10-

320

240

160

80

AISI 304

Biaxial,

Welded

1.

I

B

- 76

- 297

* = '/e' i='/|6

: = on

K \ C

I off wel

off wel weld

d [.

-76.K C

A .1. c^

n*~ c

J" A c A D 29 7 K

1 '-—a BA C J'

C

'

280

240

200

160

120

80

40

CO

CO

UJ

rrH

CO

Ld

rr

=>HO

<

rr

Lu

10 10 10° 10 IMPACT CYCLES to FAILURE

Figure 25. Fracture stress for AISI 304, biaxial, welded.

34
Another technique to characterize the mechanical behavior is to consider the applied energy needed to fracture specimens. This maybe considered, to a first approximation, as the product of the applied

stress and specimen strain (A more detailed approximation is described in Section 4 and used in all calculations). This computed deformation energy density, Ef, is plotted versus the log of the number of cycles to failure, N, in Figures 26-32. In the range tested, these curves

can be approximated by an exponential relation,

-mE F N = No exp kT J , between the number of cycles to failure, N, and E F , where N3 and m are

material constants for a given stress configuration and k is Boltzmann's

constant. In all but one case, the amount of energy that must be absorbed by the sample before fracturing decreases with the temperature. The

one exception to this decrease in absorbed energy with temperature is

the AISI 310 uniaxial (Figure 26). Figures 29 and 30 indicate that the

presence of martensite in AISI 3 04 prior to testing decreases the energy

absorbed before fracture. The relative position of weld and notch has

very little effect on absorbed energy at 297K, but does exert a considerable influence at 76K (Figure 32). Figure 33 displays summary

curves at 10 cycles and the deviation of the AISI 310 uniaxial is obvious.

The ductile behavior of the two steels can be seen in Figures

34 and 35. Again, the room temperature differences are small but in- crease considerably as the temperature decreases. The low ductility

of AISI 3 04 at 7 6K is probably due to the large amount of martensite

present.

35
12x10

10 10^ 10° 0° 10' I0n

I0J IMPACT -CYCLES to FAILURE Figure 26. Deformation energy density for fracture of AISI 310, uniaxial.

I2xl03

in

-z. o

>

oH

<

oro

Q

10 10 10 10 10 10 10' IMPACT CYCLES to FAILURE

Figure 27. Deformation energy density for fracture of AISI 310, biaxial.

36
io" icr 10" io' 10' IMPACT CYCLES to FAILURE

Figure 28. Deformation energy density for fracture of AISI 310, triaxial.

I2xl03

10

Q 6 >

orUJ

-z. Ld -4

CCO

L.

LUQ

AISI 304

Uniaxial o - 76 K ^ - 195 K D - 297 K

I V - 76 K, sensitized at 813 K ® - 297 K, sensitized at 813 K • - 76 K, precycled to 76 K -297 K, precycled to 76 K

1.0

.6 <St

UJQ

IMPACT CYCLES to FAILURE

Figure 29. Deformation energy density for fracture of AISI 3 04, uniaxial.

37
I2xl03

>-

>-O

rr

UJ

<

oro

UJo

10 icr io 10° J

io"* io" IMPACT CYCLES to FAILURE

Figure 3 0. Deformation energy density for fracture of AISI 3 04, biaxial.

rnm 12x10 L0

^ 10

IUJQ

>-O

or

<

orou_

UJo

10" 10" 10' 10 IMPACT CYCLES to FAILURE

Figure 31. Deformation energy density for fracture of AISI 3 04, triaxial.

38
12 x I03

£ 10

O - 76 K D - 297 K A = Notch '/8 from weld B = Notch V| 6 from weld

C = Notch on weld

297 K, All notch locations

1.0

CO

LUQ

CD

o

h-

<t

rro

Q

I 10 \0* I0

3 10* I03 I0b

I0

r IMPACT CYCLES to FAILURE

Figure 32. Deformation energy density for fracture of AISI 304,

biaxial, welded.

7x10'

LU 4 Q

>

or

LU 3

o

5

2 s

a:O

Ll. LU I a

AIS 304 .Annealed AIS 310 ,Annealed

l( )3 Cycles

un axial -\ ~^>^

^**** r

-'""" - _ -

**/

** .^ "~"--

bia xial -/

tri(ixiai —

(^ _^_.^-^*"

0.5

0.4

0.3 V

LU

- 0.2

100 200 300 TEMPERATURE,

K

Figure 33. Summary of deformation energy density for

fracture at 1 3 cycles.

39
80

70

60

50

40

30

AISI 304 , Annealed

I0 3 Cycles

AISI 310 .Annealed

"""

100 200 300 TEMPERATURE,

K

Figure 34. Elongation in one inch for fracture at 10" cycles,

100 200

TEMPERATURE,

K

Figure 35. Uniform reduction of area for fracture at 10 cycles

300

40
The progress of one impact test on a uniaxial 304 specimen

at Z97K was interrupted frequently to make measurements on the

elongation, reduction of area, and amount of martensite formed,,

From the results in Figure 36, it appears that all three quantities

reach a plateau and then increase abruptly before fracture.

5.2 Mechanical — Tensile

The results of the tensile tests on both steels are presented

in Table 4. The sample elongation shows that the ductility does de- crease as the strain rate increases. The considerable increase in

yield strength noted for the moderate strain rate of 20 min may possibly be due to a response time in the strip chart recorder too low to properly follow the load extension curve at this higher crosshead speed.

The inverse relationship between strain rate and ductility has

9 11 23

been observed previously. An increasing strain rate has also

80 <

LlI

Z

o

I-o

z>Q

UJ

a.

60

40 o

3

Q

Z<

zo

r-<o

zo

_lUJ

20

!

i AISI 304

Annealed

Uniaxial 297 K

1

1

1 • 1

Elongation -\

Uniform reductio

in area

n —\/

1 '

1

Martensite -x

1 1

1 1

20

15

UJ

u~> 10

rr<

I 10 I0

£ 10° 10* NUMBER OF IMPACT CYCLES

Figure 36. Parameters during an impact-fatigue test of AISI 304 at 297 K.

41
P4

to

<u I—I ^ CM

M o r

_l O in

r

_( ^_, in in * CO ^ r- ^ o O

—i --O

w ^ ^D * rM ro CM (J rM

M tM tM M< "* rM r- ^ co vO in ro

*-"

^_ IT) CM CO r> cm CO ro to CO 00 -O

p

_l r> r- so

m cm r~ co <y ^ 00 LO ro CO rM cm rM NI CM tM in ^o rM h m tn CO vO ^ "** rt z -. <M * * -o o ^f ro r- * o in vO rt O CO < t£ r> r- sD m in in in sO vO sO in in in in -O K —

S£

sD * in r- CO CM o a- rM rM -o CO * vD -.O in o o * ro (M cm ~ CM " " (M tM ro CO CM in ro CM ^ ^ ro ^ O CO * _ CO CM rs] rM CO ^ O o rt

<U ^ r- in ^h Tf ro ro ro * •* ^ ^O r- CO

en

m in ro o> o tM CO co 00 CM O r- in ro a- CO r> ~o CO ^° ^ iTJo

cm CM rM rM rM rM ro ro ro ro

~-"

'S rt CO -< ro o o vD r- rM o O in o ro

O ro r- r- nir-'*

a r— in CO rM r- o (M ro o^ -o vD o tM

O in ro o O co ^h

m °~« iri in -o ro o cm Tf CO * rM O CM CO O ro co co r> co o * o ro co rM * -* ^h •* (M rM rM CO ^ CO rj -O rM rM ro —i (M CM

H- -HU) o in r- CM in NO o ro r- Tf CM O —

00 (M -O * ro in o co = o. o cm"

-^'

O ro ro 00 ro r> o CO in CO in CO o in ^ ro°

H rf"

D ^ o o CT>

O o o o O CO o r- CO -D CO

O o co rO tM [- o cm tM rM —i —i CM

-.* . ^_ ^ 4- *-

CM -— O 'to o CO ro -1 in rm o o O ^r CO r> in rt

°.

a CO CO •>D r>

m rM a* CO CO ro xf" CO ro o r> co r- ^ in o ro CO r- 00 00 ^ r- ~^ro

>. o O —

M, CMo

rMOrMO

CMO rM rM rM rM rM rM tM tM • <> c o cm O cm O CM O tM o O o o o O O O 1 o o O o o o o o" o o o' o° O o* o" o o o o o o

rM cm rq CM CM

H S?

r- r- r- ~o vO vD >o r> r- r> -X) sO --D r- -D rm so rm -o o r> O- r- c- r- r- r> o o r- r- r- a- r~ o o r- r> r> rrM rM eg tM rM rM rM rM ** rM —i

T3 -a V tu

a T) T3 o

1) 0)

cd ? s ^d

00 cd rd cd rd rd id nj rrj ctl rtj rd rd cd rt) rt) ,ii

r

_) , t f—t 2 2 3

nj rd rtj

CU

XXXX

XX

X

X

X

XX

XX 'xX

cd ni ni ni ni td id rd ni rtj nj a! cd nl tvj 'x 'x 'x 'x 'x 'x

•a

'3 '3 '3 '3 '3 '3 '3 '3 '3 '3 '3 '3 '3 '3 .2 .2 .2 2 2 2 d d 3 3 d d d d d d d d d d d J3 's> !q ^a 'A Xi * ^h 5f •* ^< * * o o o o o o * * ^ ^ ^ o o o

•-< >N o o O o o o o o o o o o

<;

h

co co co ro 00 ro CO ro ro ro ro ro ro ro ro ro ro ro co ro ro

rd S

tu —i 3 ^

42
been shown to decrease the amount of material transformed in meta3 9 24 25

stable alloys. ' ' ' In turn, the amount of martensite has been

related directly to the tensile and yield strengths of the metal; generally they increase together. As would be expected from these

two observations, the tensile strength will tend to decrease as the

strain rate is increased.

True stress-true strain curves for both alloys have been re- corded in Figures 37 and 38. These were constructed from the loadextension curve obtained during the tests, with the assumption that the

elongation is equal to the engineering strain; this assumption should be

valid until necking occurs shortly before fracture. The spread of some

5 27-29

representative tensile test data ' is shown in Figures 39 and 40,

and the present data show general agreement with these results with

the exception of the yield strength at 20 min" as noted above.

5. 3 Transformation Products in AISI 304

The amount of martensite transformation product formed in

304 is a function of strain, strain rate, temperature, and chemical composition. The amount of bcc phase present is most easily determined by measuring relative magnetic permeability. This can be

performed quite rapidly with a Magne- Gage, a device which measures

the force required to pull a standard magnet from the surface of the

sample. Using this technique, careful observations have been made

previously on transformed martensite during the progress of tensile

5 tests. In a similar manner, as noted above, we recorded measurements for an AISI 304 annealed, uniaxial sample during its impactfatigue life at 297K. Figure 41 indicates that, as in tensile tests, the

percent martensite increases with the strain. In this same figure,

point data are included from measurements on samples after being

impacted to fracture. In general, the results at room temperature

43
425 x I03

400

350

250

CO

co

UJ

or 200

I-co

UJ

Z> 150

cc

\-

100

50

76 K, e =0.2 min"'-\ JX.

-

Al SI 310 76 K, e = 0.002 min"P"

^^ 195 K, k

= 0.005 rr in

1 ^s^

^-297 K, k

= 0.2 min

^—297 K,e = 0.002 min 1

- 300

EE

v. 200 £T

50

COCO

UJcr

hoo

100 =

u .05 .10 .15 .20 .25 .30 .35 .40 45 .50 .55 .60 TRUE STRAIN, e T Figure 37. True stress-true strain curves for tensile tested AISI 310.

_

280 AISI 304 / 76 K, e = 0.2 min" 1 "

76 K, k

-- 0.00 2 min" 1

240

Q.

~~

,1 200

b -

297 K,e = 0.002 min" 1

co" uj 160

l-

~ =*""^

co

UJ 120

Z>

CE

297 K, e = 0.2 min"' -

80

40

^00** - 40

.05 .10 .15 .20 25 .30 .35 .40 .45 .50 .55 .60 TRUE STRAIN, e T Figure 38. True stress-true strain curves for tensile tested AISI 304.

44
wuj/6>rsS3UlS

1 1 1 / /

i

/ 1 1

1

C

I / I

/ o

' o

d 7

I

/ / / : /

/

o X / 1 /

!Sd 'SS3aiS N0UV9N013 % V3uv jo Nouonasa %

7ujui/6)('sS3aiS

1 1 1

1

1

1

I

/ i ' /

/

/ ^

J / 3 /'

5 /

o / /

/

2

QJ

i—

l

o

ro

1—

1

fl CO

<U 1—1

+-> en

(1} >

i—

i

•rH •n ^ i—

I

u ti <u <D CO r^ <D fj M <ti ft

<]) ^ K

. o>

CO

<uu3M

isd 'SS3B1S N0I1V9N013 % vsav 30 Nononasa %

45
-d—r

Alloy A • Impact 297 K Impact 76 K

( Impact ) 1 . } 297 K

Alloy B < A Tensile

J , (a Tensile 76 K

100 STRAIN , %

Figure 41. Formation of martensite during tensile and impact tests.

seem to be much more dependent on alloy composition than they are at

76K.

In all of the 304 samples, the martensite fraction was recorded

after fracture. The results in Figure 42 indicate that this value falls

within rather narrow limits for each temperature and geometry.

5.4 Microscopic Observations

Tensile specimens of 304 and 310 stainless steels, both impact

and slowly loaded (0.002 min" ), were examined using techniques of

optical and transmission electron microscopy. Specimens were

deformed in tension at either 76 or 297K to produce strains of approximately 3% or 9%. After a specimen was strained.it was examined

using optical light microscopy to determine the nature of the slip traces

resulting from the deformation. Following this, the specimen was

subsequently chemically thinned and then electropolished to produce

46
UJ

0T

I-O

<

or

UJ

Z>CD

100

80

60

40

20

uniaxial

I ICJ<

0_

orUJ

i-<

80

60

40

20

uiaxiai

COz

UJ 60

or 40

< triaxial

100 200

TEMPERATURE,

K

300

Figure 42. Temperature dependence of martensite formed after

impact -fatigue fracture in AISI 304.

47
foils suitable for examination in the electron microscope.

Examination of the tensile specimens) using optical microscopy)

revealed the grain size and slip traces. When a grain is deformed,

slip will occur initially on the slip system where the resolved shear

stress exceeds the critical resolved shear stress, t c . In facecentered cubic metals and alloys> there are 1Z possible slip systems>

defined by the four (ill} slip planes which each contain three ( 1 1 )

slip directions. In polycrystalline specimens) the primary slip system

(the slip system on which t c is first exceeded) is dependent on the

orientation of a grain in respect to the applied load; therefore, for

grains of different orientations, slip will produce slip traces in different directions relative to one another, delineating individual grains.

Furthermore, initiation of conjugate slip, slip on a second slip system,

is dependent on the orientation of a grain. Figure 43 shows an area of

a specimen that was strained ~ 3% where there are examples of grains

having slip on the primary slip system only (grain A) and grains showing cross slip (grain B), slip on both the primary and conjugate slip

systems (this is representative of both 304 and 310). The frequency of

grains containing cross slip is an indication of the amount of plastic deformation in a specimen, since a random distribution of orientations for

the grains in each specimen can be assumed. For both alloys, it was

observed that there were substantially more cross slipped grains at 9%

strain than at 3% strain and that the rate of loading did not affect the occurence of cross slip. The temperature of deformation likewise did

not affect the frequency of observed cross slip. This is unexpected

since tc is temperature dependent. It, therefore, appears

that the application of the tensile load, either impact or slow, and the

temperature of deformation had little influence on the slip behavior of

these alloys.

48
Face-centered cubic alloys undergo three distinct stages of deformation characterized by three dislocation structures. These are as follows

:

Stage I: low strain

(a) Long straight dislocations on the primary slip system.

(b) A few dislocation interactions, the beginning of conjugate

slip.

(c) More dislocation interactions.

(d) Dislocation density at end of Stage I ~ 1 cm . Stage II: intermediate strain

(a) Rapid appearance of rough dislocation cell structure;

not possible to distinguish individual slip planes.

(b) High density of loops in cell walls.

11 — 2 (c) Dislocation density at end of Stage II ~ 1 cm . Stage III: high strain to rupture

(a) Misorientation between cells increases.

(b) Boundaries of cells show a tendency to lie along

primary or conjugate slip traces.

Evidence of Stage I is observed for both 304 and 310 stainless steel independent of deformation temperature (the deformation of 304 is complicated by a stress-induced martensitic reaction which will be discussed

later). Figure 44 shows an example of deformation characteristics of

the later part of Stage I for 310 stainless steel, which corresponds to ~ 3% strain. The figure shows, as indicated by the arrows, the exist- ence of two active slip systems, as well as dislocations which have interacted after piling up at (a) where two slip planes intersect; at (b)

there is a stacking fault separating two partial dislocations. For ~ 9%

strain in the same alloy, which corresponds to early Stage II, a typical

dislocation structure is given in Figure 45, which shows the initiation

49
"•if ',.

>':.• \

N

if ' *jK2< «>

Figure 43. 304 specimen strained ~ 3%; "A" grain with slip on primaryslip system only, "B" grain showing slip on two systems. 200X

-• A~r -^V*'- -! • I V -*

\ *t ft

! *\

\ * *

* "•r V / >

4 r i

•

? P

*M

e «. r ? tV • C < c j

1 J^1 v •< k' it f w * f

>¥* ''^%

Figure 44. 310 stainless steel strained ~3%. Arrows indicate slip

systems; (a) are dislocation interactions at dislocation pile-up where two

slip planes intersect and (b) is a stacking fault separating two partial

dislocations. 40, 000X

50
of a cell structure and a substantially higher dislocation density. The

dislocation structures of 310 seemed to be the same, independent of

loading and temperature; whereas the 304, while having the same general dislocation arrangements as the 310 at a corresponding strain,

contained deformation twins in some grains when impacted to ~ 9%

strain (Figure 46).

Furthermore, since 304 stainless steel undergoes a stress induced martensitic transformation, the 304 specimens both impacted

and slowly strained ~ 9% at 297K* and all specimens strained at 76K

contained hexagonal close-packed, -, and body-centered cubic, a , phases associated with the transformation. Figure 47 shows an electron micrograph (a) and its selected area diffraction pattern (b) for a 304 stainless steel which has transformed after deformation. The diffraction pattern contains spots corresponding to the y, the parent facecentered cubic phase, e , and a . The long parallel bands in the micrograph are £ and the dark lenticular rods in the £ are the crosssections of a lathes. In this micrograph it is evident that a has transformed on two different variants, (a) and (b), of its habit plane. At low

strain and 76K, there is evidence that £ is not always associated with the

a , as can be seen in Figure 48. When subjected to high strains at room temperature or even low strains at 76K, independent of the

method of loading, 304 stainless steel undergoes a martensitic transformation producing a and e , which is associated with dislocation

structures similar to those .of 310 under similar conditions of deformation.

Optical microscopy observations of fractured specimens re- vealed that, in general,the fracture path is random, following neither

grain boundaries nor planar (martensitic) features within the grains.

Other microcracks were not detected. Since all specimens exhibited

51
>^t

Figure 45. Cell dislocation structure of 310 stainless steel after

strain of -9%. 40, 000X

Figure 46. Deformation twin in 304 stainless steel after ~9%

strain. 100, 000X

52
Figure 47. 304 stainless steel strained ~9%. (a) Electron micrograph at 100, 000X; (b) selected area diffraction, e(hcp), a ' (bcc), and y(fcc).

53
Figure 48. 304 stainless steel strained ~3% at 76K; a and y present but no e. 16, 000X

54
some necking prior to fracture, it is not surprising that other cracks

were not observed; the necking served to concentrate the stress within

a very limited region of the specimen.

Table 5 summarizes the deformation structures and the phases

present for the alloys investigated under different conditions of testing.

55
ao

i-H 3 i ai—i

0) o

CO

•rH

Q

w

co

Q

>>o ? I—1

r—1

<!

a

"O o

<dg

•rH

4-> COa

•rH o !h £

CD

H

Qu P om

+->a

CD .

s co ^ u

CD

d o

ft ' '

co<D

a co IT) ri ^ i—i

ft ,Q .—-,

rtf 'C Oh H

0)

+->U3 w CJ

Vh

a

CO *£

•rH

a I h-> CO

•iH

+->

ctf au

(1) Q

a

•rH

ctf O

m

J

>>u

cti s a si

u

01o

CDCO 0) CD <D CD CD CD

CD

a

aa

a

aa

a

h o o o o o o

Ph

a £ aa £ a

a

a

a

a £

CD CD CD CD

en CO en co CD CD CD CD

CD <D CD CD

a

aa

a

U

Um

U o o o

a, C^ Ph & a jjj

aa

aaaa

CD CD CD CD

co co CO co CD CD CD

CD CD CD CD

a £ aa

UUU

U o o o

Oh iX a. On

a

a

ag

CT- On

O

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CJ

CTi Oh

O

Oha

CO

H

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r~- r~ vO ^D o Is- [> r~ o o r- r- o o o^ o

ro 00 ro 00 00 00 o.—I < ^ 1* ^ i1 o o o o

CO CO CO CO

O

CO

56
6. Discussion

6. 1 Impact- Fatigue Equipment

A new apparatus, designed to simulate the application of repeated

variable loads at high strain rates on controlled specimens, has been

employed in this study. The test results, using this "impact-fatigue"

machine, have proven very consistent. Applied stress versus number

of cycles to failure (S-N) curves blend very well with conventional

"push-pull" or flexure fatigue curves (Figure 15). Since only ductile

specimens(AISI 310, 304) were tested, it is of considerable interest

to employ this equipment in the study of the characteristics of higher

strength aerospace alloys. Presumably, crack propagation behavior

under impact loading would vary considerably- from high cycle fatigue

loading conditions.

6. 2 Fatigue Strength

Apparently, no previous fatigue studies of austenitic stainless

steels have included AISI 310; this investigation represents the first

documentation of AISI 310 in the temperature range 76 to 297K. As

Figures 4-6 suggest, the fatigue strength increases with decreasing

test temperature and is stronger for uniaxially- applied impact loads.

While the ratio of the yield strength at 7 6K to the yield strength at 297K is about 2. 6, the ratio of the increase of fatigue strength between

the two temperatures is considerably less, only about 1.6. This ratio

decreases more if the impact stresses are applied biaxially or triaxially.

Triaxial AISI 310 endurance limits are less than the corresponding yield strengths at each temperature. In fact, all triaxial S-N curves

appear to merge at longer fatigue lives and to have endurance strengths

of about 20, 000 psi at over 10 4 cycles. This strength represents the

lowest design stress level found for AISI 310 and is independent of

57
operating temperature in the interval 76 to 297K.

Fatigue strengths of AISI 304, (Figure 7-9) exhibit the same

trends as AISI 310. In general, however, at given fatigue life, temperature, and stress conditions, the fatigue strength of AISI 304 is

slightly higher . Also, AISI 304 endurance limits are better, with the

minimum endurance strength of about 25 to 30, 000 psi under triaxial

loading conditions at all temperatures. Again, under triaxial conditions

at longer fatigue life, the effect of temperature is insignificant.

Compared to tensile yield strengths, the fatigue strength in

uniaxial and biaxial conditions increases more as the test temperature

is lowered (yield strength ratio, 76 to 297K, of about 1. 3; fatigue

strength ratio, 76 to 297K, of about 1. 7). This is opposite to the effect

of temperature on AISI 310 yield strength and fatigue strength ratios.

But, the ratio of fatigue strength increase with decreasing temperature

for the two alloys is nearly identical (AISI 310= 1.6, AISI 304 =- 1.7).

This implies that the effect of martensite formation on fatigue life is relatively insignificant; in AISI 304 significantly more ma rtensite

forms in the 76K specimens than forms in the 297K. This conclusion

is startling since purer martensitic alloys (Fe-Ni alloys, AISI 400

series) are well known to exhibit more brittle behavior at lower

temperatures . Even under triaxial loading conditions, when the effect of

temperature becomes insignificant, (implying some embrittlement

at lower temperatures) the argument cannot be made that martensitic

transformation is responsible; the same trend is found in AISI 310,

where no martensite was observed (see Table V)' As AISI 310 yield

strengths are equivalent or better than AISI 304 tensile yield strengths,

the slightly better fatigue strengths obtained in AISI 304 may be

attributed to martensite formation. The same explanation is usually
presented for the higher ultimate strengths obtained in AISI 304, compared to AISI 310 at 76K (see Table IV). Therefore, after examination

of AISI 310 (which has no stress -induced transformation) and AISI 304

(which has a stress -induced transformation) impact-fatigue data, no

deleterious behavior can be attributed to the martensitic transformation. Reduction of endurance strengths under triaxial loading conditions

is almost identical in both alloys, indicating that face-centered cubic

alloy dislocation interactions or deformation twins are responsible,

not martensitic transformations.

6. 3 Welding Effects

In the course of this program, other variables such as weld

properties, the effects of cold work, and the effects of thermal cycling

between 76 and 297K (to produce about 5% martensite in AISI 304) were

examined. Some of the results are briefly discussed below.

If AISI 304 is carefully welded, using the heliarc process,

impact-fatigue strengths of the weld area are approximately 15% lower

than the parent material strengths (see Figures 10 and 11). Weld

tensile strengths (Table IV) are also lower than parent material

strengths, as fractures occurred in the weld portion of specimens.

Fatigue strengths at 10 4 cycles to failure are compared in Figure 14

for AISI 304 (welded), AISI 304 (annealed) and AISI 310 (annealed).

Under biaxial loading conditions, the welded AISI 304 is about 15%

weaker than the annealed materials. Unfortunately, it was not possible

to manufacture triaxial specimens (round, notched) from the 1/8 inch

thick sheet specimens. But the data of Figure 14 do imply that the

weakest portion of a welded, AISI 304 structure would be that section

which is welded and exposed to triaxially oriented loads under impact

conditions. At 1 4 cycles, the impact-fatigue strength of such sections

would be about 20, 000 psi or greater. The endurance limit of welded,

59
triaxially loaded specimens, however, would likely be less, 10, 000

psi being a reasonable, though conservative estimate. One can arrive

at this minimum stress estimate by extrapolating the welded uniaxial and

biaxial fatigue strengths to triaxial strengths, using the trend of the

annealed AISI 304 fatigue strengths as a guideline.

6. 4 Thermal Cycling Effects

It was thought that thermal cycling to induce martensite formation spontaneously would, perhaps, alter the impact-fatigue behavior.

Earlier, we had shown that prior transformation in polycrystalline

AISI 304 reduces the "easy glide" portion of the stress-strain curve

5 at 76K. This "easy glide" region was shown to correspond to formation of hexagonal close packed martensite during plastic straining.

Perusal of Figures 7 and 8 indicate that previous formation of small

amounts of martensite has no detectable effect on subsequent impactfatigue life.

6. 5 Cold Working Effects

Cold working of AISI 310 or 304 at 76K significantly increases

the subsequent impact-fatigue properties at 76 and 297K (see Figure

17). It is interesting to compare the ratios of the fatigue strengths

(at 10 3 cycles) of the cold worked to annealed conditions. Both AISI

304 and 310 at 76 and 297K have ratios (fatigue strength, cold worked/

fatigue strength, annealed) of between 1.4 and 1.5. This again strongly

implies that martensite formation has little effect on fatigue behavior,

since cold working of AISI 310 produced no martensite, while cold

working of AISI 304 produces about 12% hexagonal close packed

martensite and 50% body-centered cubic martensite (see Reed, Guntner

for martensitic product dependence on applied stress or plastic strain).

60
The data contained in Figures 15 and 16 indicate the high degree

of correlation between our results and the conventional fatigue test

r , 15-Z2

results of others. It proves informative to compare Figure 15

with the tensile results of Figure 40. At one cycle, equivalent to a high strain rate tensile test, the "fatigue" strength corresponds very

well to the tensile strength. At lower applied stresses (higher cycles

to failure) the endurance strength falls to approximately the minimum

yield strength (20, 000 psi).

6. 6 Fatigue Fracture Stress

If the data is converted to fatigue fracture stress (applied load/

reduced area at fracture), the temperature dependence of the data is considerably less. In particular the biaxial and triaxiai results for

both AISI 304 and 310 are essentially independent of temperature.

Slight temperature dependence is shown by the uniaxial data.

6. 7 Deformation Energy Density

Consideration of the energy density curves in Figures 26-33

leads to some interesting speculation. It is clear that as the impact

load is decreased, providing longer fatigue life, the energy absorbed

by the specimen prior to fracture is reduced toward zero. For this

to occur, localized deformation must replace more uniform, "tensiletype" deformation. That is, it is very likely that crack initiation

processes are changing as a function of applied load. Although for

simplicity the energy density-cycles to failure curves are characterized as a linear function in Figures 26-33, it is very probable that

the function becomes asymptotic to zero load as the number of cycles

to failure continue to increase. Following this, it is reasonable to speculate that zero energy density represents the material endurance

limit; impact loading,below the stress necessary to produce some

finite specimen strain, will never produce specimen fracture. Of

61
further interest is the tendency of all assumed linear functions (energydensity versus cycles to failure), if extrapolated, to converge to zero energy density at the same number of cycles to failure (see Figures 27,

28 and 31, for example).

6. 8 Martensite Formation in AISI 304

Martensite formation was monitored as a function of number

of cycles, and the results are presented in Figure 36. The increase

of martensite (a ferromagnetic phase) with increasing number of

fatigue cycles promotes the feasibility of monitoring fatigue life.

The undeformed material, austenite, is paramagnetic; therefore,

magnetometric monitoring of specimens or components exposed to repeated loads should very easily indicate impending fracture.

The monitored specimen, tested at room temperature, exhibits an unusual increase in uniform reduction in area, elongation, and

percent martensite transformed just prior to fracture (Figure 36).

To our knowledge,the rather sudden increase in uniform specimen

strain (not specimen necking) has not been reported in the literature

and is quite unexpected. Further consideration and documentation of

this is needed. Arguments cannot be advanced that fracture propagation is solely responsible, since little uniform martensite transformation would be expected to occur except along fracture paths.

6. 9 Tensile Data

Compiled tensile data for both alloys, shown in Figure 39 and

40, indicate the increase of AISI 310 yield strength, compared to AISI

304, at lower temperatures. Again, the lower 304 yield strength can

be attributed to hexagonal martensite formation. Higher 304 tensile

strengths have been attributed to body-centered martensite

strengthening. Further discussions of these contributions are contained

5,29,30

in many references*

62
Figure 41 contains data on martensite (b.c.c. ) formation as a function of specimen strain for comparison of impact-loaded and

tensile-loaded specimens. Room temperature data of Alloys A and B

do show significant difference. Both of these alloys are commercially

classified and sold as AISI 304, but it is easy to discern their different

transformation characteristics. This difference is also apparent at 76K. Comparison of the effect of strain rate on transformation tendency is not as conclusive. From our available data,no definitive

statement can be made related to this question, although many others

claim that increased strain rate produces lower amounts of martensite

. 3,9,24,25

per unit strain.

6.10 MicroscopyDiscussion of optical and electron microscopy results is presented in the Experimental Results section. Several observations

provide points for additional discussion. All specimens examined

had necked prior to crack propagation and fracture. In this sense,

therefore, fracture was very similiar to conventional tensile fracture, as contrasted to fatigue crack propagation and fracture. Martensite played no detectable role in crack propagation, nor were premature cracks initiated by the martensitic structures.

It is well documented that high energy rate forming of facecentered cubic metals produces an abundance of deformation at ambient temperatures or higher. Similarly, impact-loading of AISI

304 produced deformation twins (see Table V), however, none were

31

found in 310. As suggested by many (see Venables for review), this

evidence supports the view that lower stacking fault energy promotes

twinning. At lower temperatures in 304, martensitic transformations

(hep and bec) evidently replace twinning as deformation products.

63
6. 11 Possible Additional Programs

(32) There are indications " that the fatigue strengths of AISI 3 00

Series steels which undergo partial martensitic transformations are

less at 20K in the large amplitude range than at 76K. Therefore,

it should be clearly understood that the conclusions drawn in the re- port only apply in the temperatures range 7 6 to 294K. Additional

data should be acquired below 7 6K to enable accurate prediction of

the long-time life of components exposed to periodic sudden loads

in this temperature regime.

64
7. References

1. P. P. Benham, Metallurgical Reviews 3, 203(1958).

2. J. D. Campbell and J. Duby, Proc. Roy. Soc. (London) A236,

24 (1956).

3. H. G. Baron, J. Iron Steel Inst. 1_82, 354 (1956).

4. J. D. Campbell, J. Duby and K. E. Puttick, Phil. Mag. Z, 548 (1957).

5. C. J. Guntner and R. P. Reed, Trans. ASM 55, 399 (1962).

6. T. Williams, Sheet Metal Industries _39, 487 (1962).

7. S. W. Hollingum, High Energy Rate Working of Metals, Proc.

NATO Advanced Study Institute, Central Institute for Industrial

Research, Oslo, Norway, 278 (1964).

8. T. Kvam and B. Augland, source cited in ref. 7, p. 264.

9. J. P. Bressanelli and A. Moskowitz, Trans. ASM _59, 223(1966).

10. L. E. Murr and T. I. Grace, Expt. Mech. _9, 145 (1969).

11. D. V. Neff, T. E. Mitchell, and A. R. Troiano, Trans. ASM

62, 858 (1969).

12. A. Kumar and R. G. Kumble, J. Appl. Phys. 40, 3475 (1969).

13. A. Kumar, High Speed Testing Vol. VII; The Rheology of

Solids, Interscience Publishers, N. Y. , 67 (1969).

14. R. N. Orava and H. E. Otto, J. Metals _22 #2, 17(1970).

15. R. M. Brick and A. Phillips, Trans. ASM _2_9 , 435 (1945).

16. D. W. McDowell, "Low Temperature Properties of Type 304

Stainless Steels", Technical Report, Development and Research Division, The International Nickel Co. , New York (1959).

17. E. Krempl, "Notch High-Strain Fatigue Behavior of Three LowStrength Structural Steels, General Electric Co., Atomic

Power Equipment Dept. , San Jose, Calif. (1969), NTIS GEAP5714.

65
18. K. Fushimi, H. Okamato, E. Fukishima, and H. Yonemitsu,

Low Temperature Engineering _5_, 105 (1970).

19. J. T. Berling and J. B. Conway, Met. Trans. 1, 805 (1970).

20. M. G. Fontana, "Investigation of Mechanical Properties and

Physical Metallurgy of Aircraft Alloys at Very Low Temperatures, Pt. II", Tech. Rept. 5662, Ohio State Univ. Research

Foundation (1948).

21. J. L. Zambrow and M„ G. Fontana, Trans. ASM 41 , 480(1949).

22. J. W. Spretnak, M. G. Fontana, and H. F. Brooks, Trans.

ASM 43, 547 (1951).

23. A. Kramer and W. M. Baldwin, Jr., Trans. ASM 50, 803 (1958).

24. B. Eftest«Sl, B. F. Haegland, P. Storvik, and A. Windfeldt, source cited in ref. 7, p. 296.

25. G. W. Powell, E. R. Marshall, and W. A. Backofen, Trans.

ASM 50_, 478 (1958).

26. J. F. Watson and J. L. Christian, Report ERR-AN-057 of

Third Quarterly Progress Report - Phase I, Contract

AF 33(6l6)-7984, Dec., 1961, General Dynamics/Astronautics,

San Diego, California, NTIS AD 273221.

27. C. J. Guntner and R. P. Reed, Advances in Cryogenic Engineering, Vol. 6, K. D. Timmerhaus, Ed. , Plenum Press, Inc. , New York (1961), p. 565.

28. K. A. Warren and R. P. Reed, "Tensile and Impact Properties

of Selected Materials from 20 to 300°K", National Bureau of

Standards Monograph 63 (1963).

66
29. A. J. Nachtigall, "Comparison of Tensile Properties of 304 L, and 310 S Stainless in Liquid Helium", NTIS NASA-TM -X- 5270

(1969).

30. R. P. Reed and C, J. Guntner, Trans. Met. Soc. AIME 230 , 1713 (1964).

31. J. A. Venables, Deformation Twinning , Met. Soc. Conf. Vol

25, (Gordon and Breach Science Publ. , N. Y. 1964) p. 77.

32. F. Schwartzberg, "Cryogenic Materials Data Handbook",

ML-TDR-64-280 (PB171809), Clearinghouse for Tech, and

Sci. Inform., Washington D. C. (1968).

67
Appendix

The following paper, "Apparatus for Impact- Fatigue Testing, "

describes the modifications made to automate a standard impact

machine. It has been accepted for publication in Section C, Journal

of Research, NBS (1971).

Another paper summarizing the main findings in this report

is presently in progress.

68
Apparatus for Impact-Fatigue Testing

R. E. Schramm, R. L. Durcholz and R. P. Reed

Cryogenics Division, NBS-Institute for

Basic Standards, Boulder, Colorado, 8030Z

A standard impact machine was extensively modified to allow the

measurement of the response of specimens to repeated, controlled impact

pulses. This equipment enables one to vary the temperature (76 - 297K),

specimen geometry (uniaxial, biaxial, triaxial stress systems), and load

levels. At stress levels in the neighborhood of the yield stress on the order

of 10, 000 impact cycles are needed to fatigue specimens to fracture. Strain

rates achieved are moderately high, of the order of 1000 min -1

, which conveniently form intermediate data between tensile (max. of about 100 min _1

) and explosive straining data (about 6000 min -1

). Contrasted to standard

fatigue tests, no constraint is placed on specimen elongation and only unidirectional stresses are imposed. Typical impact-fatigue results for

AISI 310 stainless steel are presented.

Keywords: Cryostat; fatigue; impact; low temperature; mechanical property

equipment; stainless steel.

* This work was carried out at the National Bureau of Standards for the

Naval Air Engineering Center, Philadelphia, Pennsylvania.

69
1. INTRODUCTION

In many low temperature applications such as in the transportation of dewars or in the transfer of cryogenic liquids through pipe, system components

experience periodic sudden jolts or loads. Prediction of expected life of these

components can best be accomplished by testing controlled specimens under

repeated impact loads which approach up to the yield load. Such a test, then, can be described as impact-fatigue, as the specimen eventually fractures in

a manner similar to conventional fatigue fractures.

To perform such tests a standard commercial impact machine was extensively modified. These modifications resulted in an apparatus that could

repeatedly apply sudden loads in the range to 1800 kilograms for the many

repetitions sometimes required for specimen fracture. Additionally, the specimen chamber was insulated to allow temperature environments from 76 to 297 K. Our paper describes this new equipment.

A standard Riehle model PI-2'~ impact testing machine was available.

This machine has three hammer sizes, 15, 30, and 60 lbs. (6.8, 13. 6 and

27. 3 kg) and the hammer drop is continuously adjustable from to 48 inches

(0 to 1.3m). Our impact-fatigue life determinations, however, would re- quire greater than 2 0, 000 cycles in some cases so it became obvious that the single cycle operation for which this machine was designed was

not feasible. We designed major modifications to automate the repetition, to allow tests to be conducted at cryogenic temperatures, and to permit the use

of specimens in a great variety of configurations. Stress pulses are applied

to the specimen in one direction, while the conventional fatigue tests are

typically either "push-pull" with alternating tensile and compressive forces

or flexure with alternating applied bending forces. Another feature of the

impact-fatigue test is that the specimen is not constrained with respect to specimen strain.

The use of trade names in this paper in no way implies endorsement or approval by NBS and is included only to define the experimental procedure.

70
2. MODIFICATIONS

Figure 1 a is a schematic view of the impact- fatigue tester. A 1/6 hp,

1725 rpm electric motor is geared to rotate a 24 inch diameter wheel at 7. 3 rpm. The shaft and bearings which support this wheel are clamped to the upright part of the impact machine frame by U-bolts. A brass hammer release

arm protruding through the outer edge of this moving wheel serves to pick up

the hammer and raise it to a predetermined height. At this point, the trip

arm (Fig. 2) encounters a block which is supported behind the rotating wheel.

The trip arm rotation causes a cam motion which swings a locking bar out of

position allowing the hammer release to simply fall back under the weight of

the hammer. A second trip block is at the top of the wheel's revolution to in- sure that the system does not bind there as it passes the pendulum shaft. After

each trip, springs pull the release bar back into a locked position to repeat the

cycle.

The standard anvil on the base of the impact machine has been closed

off on the right end to serve as the fixed point to which the right specimen pull

rod is anchored, either directly or through a load cell (Fig. 3). The left pull

rod passes through a three piece impact yoke assembly held together by eight

i inch stainless steel bolts. The hammer impact is absorbed by this yoke

which transmits it to the specimen, since it is free to move in a right-left

direction on four rollers. When the sample breaks, the yoke moves to the

left and allows the hammer to strike a switch on the base which removes electrical power from the whole system.

A solenoid system prevents secondary impact by the recoiling hammer

(Fig. 1). This stop is a 3/4 inch round bar which rides in a brass bushing inserted into a drilled hole in the impact machine. When this bar is in the fully

inserted position, the impact hammer cannot swing past it to strike the yoke.

A portion of the bar has a flat depression ground into it so that when the solenoid is activated and the bar partially withdrawn from the anvil the hammer

may pass by freely. A switch on the stationary wheel near the block that trips

the hammer release mechanism controls the power to the solenoid. As the

71
u

a,

> <u

s §

<U GO

&0 h

72
HAMMER RELEASE

MICROSWITCH TRIP TABLE

(for secondary impact stop)

Figure 2. Impact hammer release mechanism. The trip arm is rotated by a stationary block fixed at the desired position behind the 24 inch wheel. This allows the hammer

release to fall back under the hammer weight. Springs

then return the mechanism to a locked position.

LN 2 TILL SPECIMEN

LN 2 VENT

TEFLON CRYOSTAT

Figure 3. Specimen impact system.

73
trip table (Fig. 2) passes by this switch, the secondary stop is removed from

the impact hammer path. By the time the hammer has recoiled, power has

been removed from the solenoid and springs have pushed the stop back into

position. A pin through this bar operates another switch to advance a cycle

counter.

Wherever possible, the springs in the hammer-release mechanism were

doubled, so if one fails during a test the second will continue to function until

repairs can be made. It is important to use high quality steel springs to minimize difficulty due to breaks; soft, tempered springs were found more reliable

than high strength units. Since this modification results in a much higher repetition rate than was originally intended, the welds between the pendulum shaft

and fittings on both ends were reinforced.

3. CRYOSTAT

The cryostat used for 76 K tests consisted simply of a 2 inch diameter

Teflon rod, 7 inches long, with a 1 inch bore through the center. The specimen

pull rods are supported and centered by Teflon end caps screwed into the main

chamber (Fig. 3). Rubber O-rings in both of these caps served as an initial seal around the pull rods to minimize the loss of liquid nitrogen. However, in

cooling to 76K, the Teflon contracts considerably and then serves as the main

seal. Because of this great contraction, it is necessary to check for free

movement of the pull rods when they are at test temperatures. Two small

holes in the side wall serve for filling with liquid nitrogen and venting. A

liquid flow of about GA/hr was generally used to maintain temperature. At room

temperature this cryostat was still used to assure the proper positioning of the

sample.

For 195 K tests, a slight modification of this system was employed.

Instead of the small holes in the side wall for filling and venting nitrogen, a 1X4 inch panel was removed. A styrofoam box was attached to act as a reservoir for the powdered dry ice used to fill the cryostat. Consumption in this case was about 0.51b/hr.

74
4. DISCUSSION

We used a load cell specially constructed to withstand sudden impacts

and a high speed recorder to measure impact loading. When using the 301b.

(13.6kg) hammer, the impulse time is about 2msec. Load cell response

(Fig. 4) indicates a damped ringing, but this is a characteristic of the load

cell. While the specimen probably does vibrate in the uniaxial direction, our system is incapable of measuring the compressive forces it may experience,

but they are believed to be very small.

During the progress of a test the specimen undergoes plastic strain.

This allows the yoke to move in the direction of motion of the impact hammer.

Thus, by measuring the position of the yoke frequently during a test, it is possible to determine specimen strain as a function of the number of cycles.

Naturally, this function depends on variables such as material, temperature,

and applied pulse. Figure 5 illustrates a typical curve of strain versus number of cycles for an AISI 310 stainless steel specimen. Initially, the specimen

undergoes considerable strain but, after sufficient work hardening (several

hundred cycles), its length becomes nearly constant. Since each cycle corresponds to a 2 ms impulse time, the coordinate

giving the number of cycles is also a time coordinate and its slope gives the

strain rate. With our experimental conditions, the maximum initial strain

rate was 2100 min" . This is conveniently intermediate between the maximum

strain rate in tensile tests of about 100 min"1 and the approximate lower limit

in explosive straining of 6000 min"1

. From the strain readings and load calibration, a typical stress-strain curve maybe constructed. Figure 6 shows this curve for AISI 310 at 297 K.

The stress used to construct this curve is the engineering stress, i.e., load

divided by original area. This stress level is constant throughout the test.

75
500 A

400

f\ 300

200

100

- 100 —

-200 —

-300

-400 r

-500 — I

200

160

120

80

40 o> »5

o

-40 J

-80

-120

-160

-200

.01 02 .03 .04 .05 .06 .07 TIME, second

.08 .09 0.1

Figure 4. Time response carve recorded by load cell and high speed

recorder.

< 4

iAISI 310 STAINLESS STEEL

297 K

Applied Stress = 70,000 psi (49.2 kg/mm2

)

400 800 1200 1600 2000 NUMBER OF IMPACT CYCLES

2400 2800

Figure 5. Specimen strain during impact-fatigue life.

76
By varying the height of the hammer (stress amplitude), a typical

fatigue (S-N) curve of stress level versus number of cycles to failure may be

obtained. The shape of such curves for AISI 310 at 297, 195 and 76K, as shown

in Fig. 7, is very similar to those obtained by conventional fatigue tests in which the high strain rates are not achieved and either the load or specimen

deflection amplitude is held constant. Any specimen configuration, limited only by length (about 5 inches), may

be measured. In practice, we have measured both uniaxial (tensile), biaxial,

and triaxial (notch tensile) specimens of both AISI 304 and 310 stainless steels

at 76, 195 and 297K. These test results will be published later.

5. REFERENCES

E. K. Henriksen, I. Lieberman, J. F. Wilkin, and W. B. McPherson,

"Metallurgical Effects of Explosive Straining", in Symposium on Dyna - mic Behavior of Materials (ASTM Spec. Tech. Pub. No. 336, 1963),

p 104.

77
80 x lO^i

.02 .04 .06 .08 .10 STRAIN, €

Figure 6. Typical stress-strain curve for AISI 310 stainless steel

showing repeated impact cycles at 297 K (1 kg/mm^ = 9.8 x 106 N/m2

). Compressive stresses due to recoil of

specimen upon itself are very small and not shown.

200 x!0 3 -

180

Q- 160

CO

Ld 140

120

100

B0

60

1

I !

I III 140

130

120 - 110 - 100 90

80

7060

50

1 10" 10° CYCLES TO FAILURE

Figure 7. Fatigue life curve for AISI 310 at 297, 195, and 76 K.

78
FORM NBS-114A (1-71) U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET

1. PUBLICATION OR REPORT NO. NBS-TN-609

2. Gov't Accession No.3. Recipient's Accession No.

4. TITLE AND SUBTITLE Study of Cryogenic Storage Tank Fatigue Life (Low Temperatur5. Publication Date August 1971

Mechanical Testing of AISI 304 and 310 Stainless Steels)

6. Performing Organization Code

7. AUTHOR(s) r. p. Reed, R. L. Durcholz, R. E. Schramm,

and T. J. Patrician

8. Petfotming Organization

10. Ptoject/Task/Wotk Unit No. 2750533

9. PERFORMING ORGANIZATION NAME AND ADDRESS NATIONAL BUREAU OF STANDARDS DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

11. Conttact/Gtant No. 9-8029

12. Sponsoring Organization Name and Address Naval Air Engineering Center

Philadelphia, Pennsylvania

13. Type of Report & Period Covered N/A

14. Sponsoring Agency Code

15. SUPPLEMENTARY NOTES

16. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant bibliography or literature survey, mention it here.) Two 300 series stainless steels were tested for impact-fatigue life and tensile

properties as a function of temperature, percent transformed phase (martensite),

stress level, and specimen geometry. These alloys were studied to predict the

fracture characteristics of parent material and weld joints in large cryogenic dewars

which are subjected to periodic stresses. Normally, AISI 304 is employed in the

construction of such dewars. Under fatigue loading conditions at cryogenic temperatures, the behavior of AISI 304 is complicated by the gradual transformation to a martensitic structure. The influence of this transformation on the fracture

characteristics was studied. For this purpose, a new impact-fatigue test and

necessary equipment were developed.

The fatigue strength of AISI 304 and 310 exhibit similar trends, but at a given

fatigue life and temperature, the fatigue strength of AISI 304 is slightly superior.

AISI 310 displays an endurance limit of about 20, 000 psi *; the 304 limit is 25-30, 000

psi. However, the lowest endurance limit (weakest) specimens appear to be the triaxially-loaded AISI 304 weld specimens which have an endurance limit near 10, 000

psi. Little temperature-dependence of the impact-fatigue properties was observed.

* For ease in interpretation all stress values in the text are expressed in units of psi.

Conversion to other stress units are

10 3 psi = 0. 703 kg/ mm3 = 6.90x10 s N/m 3

17. KEY WORDS (Alphabetical order, separated by semicolons) r . . r . . , fatigue; fracture; impact; low temperature;

mechanical property equipment; stainless steel.

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the world's literature and critically evaluated. Product Standards. Provide requirements for sizes, types, quality and methods for testing various industrial products. These standards are developed cooperatively with interested Government and industry groups and provide the basis for common understanding of product characteristics for both buyers and sellers. Their use is voluntary.

Technical Notes. This series consists of communications and reports (covering both other agency and

NBS-sponsored work) of limited or transitory interest. Federal Information Processing Standards Publications. This series is the official publication within

the Federal Government for information on standards adopted and promulgated under the Public Law

89-306, and Bureau of the Budget Circular A-86

entitled, Standardization of Data Elements and Codes

in Data Systems.

Consumer Information Series. Practical information, based on NBS research and experience, covering

areas of interest to the consumer. Easily understandable language and illustrations provide useful back- ground knowledge for shopping in today's technological marketplace.

NBS Special Publication 305, Supplement 1, Publications of the NBS, 1968-1969. When ordering, include Catalog No. C13.10:305. Price $4.50;

foreign, $5.75.

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