ORNL-TM-2483.txt

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P rsartcey—of

 

OAK RIDGE NATIONAL LABORATORY
operated by
UNION CARBIDE CORPORATION %
NUCLEAR DIVISION

 

for the _
U.S. ATOMIC ENERGY COMMISSION (}4 c
ORNL- TM- 2483
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BT T MAR 2

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PREIRRADIATION AND POSTIRRADIATION MECHANICAL

PROPERTIES OF HASTELLOY N WELDS

H. E. McCoy
D. A. Canonico

 

 
 

 

—_———————— LEGAL NOTICE ——————— —

This report was prepored os an gccount of Government sponsored work. Majther the United States,

nor the Commission, nor any person acting on behalf of the Commission:

A. Makes any warranty or representation, expressed or implisd, with respect to the occuracy,
completeness, or usefulness of the information ceontoined in this report, or that the use of
any information, opparctus, method, or process disclosed in this report may not infringe
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B. Assumas any liabilities with respect 1o the use of, or for domages resulting from the use of
any information, apporatus, methed, or process disclosed in this report.

As used in the ohove, "'person acting on behalf of the Commission” includes any employe= or

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or contractor of the Commission, or employee of such contractor prepores, disseminates, or
pravides occess to, any information pursuont to his employment or contract with the Commissien,
or his employment with sueh contractor,

 

 
ORNL-TM-2483

Contract No. W-7405-eng-26

METALS AND CERAMICS DIVISION

PREIRRADIATION AND POSTIRRADIATION MECHANICAL
PROPERTIES OF HASTELLOY N WELDS

H. E. McCoy and D. A. Canonico

Submitted to the Welding Journal without the Appendix.

 

MARCH 1969

OAK RIDGE NATIONAL LABORATORY
Oak Ridge, Tennessee
operated by
UNION CARBIDE CORPORATTION
for the
U.S. ATOMIC ENERGY COMMISSION
 

A 0 BN A e et

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R R 1
iii

CONTENTS

Abstract .

Introduction .

Experimental Details .

Experimental Results .
Tensile Properties
Creep-Rupture Properties
Postweld Heat Treatments
Metallography . . « « « « + « « &

SUMMAYY « « o« e e e e e e e e e

Acknowledgments

References .

Appendix .
PREIRRADIATION AND POSTIRRADIATION MECHANICAL
PROPERTIES OF HASTELLOY N WELDS

 

 

H. E. McCoy and D. A. Canonico

ABSTRACT

Welds were made by the TIG process in several heats
of Hastelloy N. The mechanical properties of transverse
weld samples and the base metal were compared in tensile
tests over the range of 75 to 1600°F and in creep tests at
1200°F. The as-fabricated welds exhibited lower fracture
strains than the base metal under all test conditions, but
the properties of the welds were improved markedly by post-
weld heat treatments. The postirradiation tensile and
creep properties of the welds and base metal at elevated
temperatures were about the same, although the properties

were widely different before irradiation.

 

INTRODUCTION

Hastelloy N is a trade name given the solid-solution-strengthened

nickel-base alloy developed at the Oak Ridge National Laboratory specif-

ically for use in molten fluoride salts up to 1500°F (ref. 1). The

alloy was originally designated INOR-8 and has a nominal composition of

Ni—17% Mo—7% Cr—5% Fe. This material is the primary metallic structural

material in the Molten Salt Reactor Experiment (MSRE) which achieved

criticality on June 1, 1965, at Oak Ridge, Tennesse.? This alloy has
been used for numerous other applications, and we anticipate that a
slightly modified composition will be used in a future molten salt
breeder reactor experiment.

The weldability of this material has received considerable atten-

3,4

tion in our program. We found it necessary to establish a weldability

test to determine whether specific heats of material could be joined

satisfactorily. Through this screening process, it was possible to
select heats for use in fabricating the MSRE components.

During the past few years we have found that Hastelloy N as well
as other nickel- and iron-base alloys are subject to a type of neutron
irradiation damage that decreases the creep-rupture strength and fracture

556 Thisg damage 1is rather general for

strain at elevated temperatures.
all austenitic iron- and nickel-base alloys and is attributed to the
production of helium within the material due to the interaction of a
neutron of thermal energy with '“B to produce “Li and “He (refs. 7-14).
The property changes due to the helium must be evaluated for base metal
and welds.

In the present paper, we shall show how the mechanical properties
of transverse welds differ from those of the base metal and what types
of annealing treatments can be used to improve the properties of welds.
We shall also show how neutron irradiation alters the mechanical proper-
ties of the welds and the base metal. The mechanical properties measured

in this study were tensile properties over the temperature range of 75

to 1600°F and creep properties at 1200°F.

ok b e R A R PR B R R e vt SR et ke R e ok e G e 1 e b R AR 20, Rttt L e L
EXPERIMENTAL DETAILS

Several heats of Hastelloy N were used in this study and their
chemical compositions and other pertinent details are given in Table 1.
Heats 65-552 and 2477 were wvacuum melted; all others were air melted.

All base metal was annealed for 1 hr at 2150°F prior to testing unless
otherwise specified.

Seven weldments were involved in this study and the details are
given in Table 2. The welds were all highly restrained and made by the
manual tungsten-arc welding process. A standard welding procedure was
used on all the welds, although some deviation was necessary in welds 7
and & because of the type of filler metal being used. The joint configu-
ration and pass sequence are shown in Fig. 1. The primary working
direction in the plates was perpendicular to the weld axis.

We used small mechanical property samples throughout the study,
since this geometry was required for the irradiation experiments (Fig. 1).
The samples were cut perpendicular to the weld axis and parallel to the
stringers. Three layers of specimens were cut from all the welds except
No. 3; we could find no systematic variation in properties from top to
bottom. However, the gage portions of the transverse weld specimens did
contain different amounts of weld metal since the joint tapered (Fig. 1).

The tensile tests were run in an Instron Universal testing machine
1

(10,000-1b capacity) at strain rates of 0.05 or 0.002 min~ The strain

measurements were taken from the crosshead travel. A similar machine
was located in a hot cell for testing the irradiated samples. The labo-
ratory creep-rupture tests were run in standard lever-arm creep machines.

The strain measurements were taken from the pull-rod displacement after
Table 1. Chemical Analysis of Test Materials

 

 

 

 

Heat . Composition
Designa- Meltl?g

etiin Fractice Cr Fe Mo c si Co W (w;n%j v P S AL Ti Cu B (pgm) N
5065 Air 7.2 3.9 16.5 0.065 0.e0 0.08 0.04 0.55 0,22  0.004 0.007 0.01 0.01 0.01 10-37 16 110
5067 Air 7.4 4.0 17.3 0.060 0.43 0.08 0.06 0.50  0.30 0.013 0.007 0.01L 0.01 0.015 10-24 12 100
5101 Air 6.9 3.9 16.4 0.05 0.63 0.05 0.44  0.34 0,001 0.009 - 0.022 — 3.5
5055 Air 7.9 3.8 16.2 0.06 0.6l 0.03 0.69  0.21  0.006 0.008 0.06 0.02 50
2477 Vacuum 7.05 4.25 16.3 0.057 0.015 0.14 0.47 0.04 0.008 0.003 0.055 0.10 0.10 8 5 7
65-552  Vacuum 6.89 4.06 16.2 0.045 0.16 0.050 0.006 0.45 < 0.0005 0.002 0.006 0.25 0.5 1.3 9

 

aCom.bined amount of eluminum and titanium.
Table 2. Identification of Welds Made in This Study

 

 

Weld Thickness
Nurber Base Metal Filler Metal of.Weld
{in.)
1 5065, 5067 5101 1
3 2477 2477 1/2
4 5065, 5067 5101 1
5 5065, 5067 5055 1
6 5065 65-552 11/8
7 5065 5055 + Al,052 11/8
5065 5055 + WC? 11/8

 

8piller material was 1/8 in. wire of heat 5055 plasma
sprayed with 0.002 in. of indicated material.

the load was applied. The postirradiation creep-rupture tests were run
in 1ever-arm creep machines located in the ORNL hot cells. The strain
was measured by an extensometer with rods attached to the upper and lower
specimen grips. The relative movement of these two rods was measured by
a linear differential transformer. All tests were run in an air environ-
ment. We used a standard equilibrating time of 1/2 hr for each test
prior to loading.

The irradiations were conducted in two facilities: the ETR,
Idaho Falls, Idaho, and the ORR, Oak Ridge, Tennessee. A core facility
was used in the ETR where the thermal and fast (> 1 Mev) fluxes were
each 3.2 X 10%* neutrons em™? sec™!. The fluence obtained there was
5 x 10°° neutrons/cmz. The ETR experiments were uninstrumented and the

design temperatures were either less than 300°F or 1112 + 180°F. A

core facility was also used in the ORR where the thermal flux was
ORNL-DWG 68-12878R
DIMENSIONS IN INCHES

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FILLET WELDED TO STRONG BACK TO PRODUCE HIGH RESTRAINT WELD

29

 

 

 

 

 

 

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JOINT DESIGN AND WELDING SEQUENCE

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Q.0008 in.
Q.0000 in.
0.0005 in.
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MECHANICAL PROPERTIES SPECIMEN.

Fig. 1. The INOR-8 High-Restraint Weldability Test Specimen Used
to Provide Samples for the Mechanical Properties Study.
2.5 x 10'% neutrons cm™? sec™l. The fluences obtained were

8.5 x 102° neutrons/em® thermal and 7.0 x 102° neutrons/cm? fast

(> 1 Mev). The temperature in this facility was 110°F. We could not
observe any effect of fluence over the small range involved here, so

the data are presented without reference to a particular experiment.

Irradiation temperature was important in some cases and will be indicated.
EXPERIMENTAL RESULTS

Tensile Properties¥
The variation of the fracture strain in a tensile test with test
temperature is shown in Fig. 2 for both transverse welds and base metal.
The behavior of this particular heat of base metal is typical for
Hastelloy N. The decrease in fracture strain above 1100°F is due to

ORNL-DWG 6812875
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T T T T
HEAT 5065 WELD1 WELD 4
AS PROCESSED ° A &

 

AFTER IRRADIATION . .
80 POST WELD ANNEAL OF 8hr
AT 1600°F
UNIRRADIATED 0
IRRADIATED -

 

 

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UNIRRADIATED BASE METALX
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FRACTURE STRAIN (%)
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UNIRRADIATED WEL[;?\\\ /
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o 200 400 600 800 1000 1200 1400 1600 1800

TEST TEMPERATURE (°F)

Fig. 2. Ductility of Hastelloy N Welds and Base Metal in Tensile
Tests at a Strain Rate of 0.05 min™?!.

 

*See Tables A-1 and A-2, Appendix, for the tabulated tensile data.
the transition from transgranular to intergranular fracture. The
recovery above 1400°F is associated with increasing grain boundary
mobility. The fracture strain of the welds is much lower. Most of the
welds fractured in the weld metal, but several samples tested at 75 and
392°F did fail in the base metal.

The fracture strain after irradiation is also shown in Fig. 2.
Since these samples were irradiated at less than 300°F, there is some
displacement damage in the base metal and the fracture strain is lower
at test temperatures of 75 and 392°F. At test temperatures of 1200 and
1600°F, where intergranular fracture predominates, the fracture strain
of the base metal is reduced dramatically. The postirradiation fracture
ductility of the irradiated welds is also shown in Fig. 2. The strain
at fracture of the welds is reduced slightly by irradiation, but the
changes are much less than those observed for the base metal. Thus,
although the ductility of the welds is much lower than that of the base
metal before irradiation, the ductilities of welds and base metal at
elevated temperatures are quite similar after irradiation.

The yield strengths of typical welds and base metal are compared
in Fig. 3. The yieldlstrengths of the transverse weld samples are con-
sistently higher than those of the base metal. This same observation
was also made by Gilliland and Venard.’ After a postweld anneal of & hr
at 1600°F, the welded and base metal samples have very similar yield
strengths.

Irradiation increases the yield strength of the base metal and
welded samples at low temperatures. This damage anneals out as the test

temperature is increased, and the welds and base metal show similar

 
ORNL - DWG 6B-12876

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

120 l .
| ! ‘
HEAT 5065 WELD1 WELD 4
s L,R,fsf‘ED'GE%EL AS PROCESSED o 0 a
0o | ¥ AFTER IRRADIATION o s
R N A POST WELD ANNEAL
7 \\\\\\ OF 8hr AT 1600°F
o 80 —— % UNIRRADIATED o
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8 a ‘-~::::§\ IRRADIATED | | oo
= . —— —~— |, UNIRRADIATED WELDS
; — o \ T
3 40 O S —O
u UNIRRADIATED BASE METAL — &
N
20
0
0 200 400 600  BOO {000 {200 1400 1600 1800

TEST TEMPERATURE (°F)

Fig. 3. Yield Strength of Hastelloy N Welds and Base Metal in Ten-
sile Tests at a Strain Rate of 0.05 min~1.
recovery. Thus, above 1200°F the yield strengths of base metal and
welds annealed for & hr at 1600°F are equivalent and are unaffected by
irradiation.

Figure 4 shows that the ultimate tensile strength is about the same
for welded samples and for base metal. This property is not affected
significantly by postweld heat treatment at 1600°F. Irradiation causes
a slight (approx 10%) increase in the ultimate tensile strength at the
iower test temperatures and a decrease at test temperatures of 1200°F
and greater. This reduction at higher temperatures is due to the reduced
ability of the material to deform plastically after irradiation (i.e.,
fracture occurs before the stress increases to the higher wvalues noted
for unirradiated materials).

Thus, the effects of welding on the strength parameters measured by
standard tensile tests are relatively small and the most significant
factor is the reduction in the fracture strain. The fracture strains

observed for the welds involved in this study are summarized in Fig. 5.
FRACTURE STRAIN (%)

140

120

100

80

60

40

ULTIMATE TENSILE STRESS (1000 psi)

20

Fig. 4.

10

ORNL-DWG 68-12877

 

 

 

 

 

 

 

 

 

 

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AFTER IRRADIATION . a

 

[ POST WELD ANNEAL
OF 8 hr AT 1600°F

 

 

 

 

 

 

 

 

UNIRRADIATED o
I IRRADIATED .
0 200 400 600 800 1000 1200 1400 1600 1800

TEST TEMPERATURE {°F)

Ultimate Tensile Strength of Hastelloy N Welds and Base

Metal in Tensile Tests at a Strain Rate of 0.05 min~?!.

ORNL -DWG 68- 12888

 

 

 

 

   

  
  

 

   
     

 

 

 

 

 

 

  

 

 

 

80
T B As wELDED - — = .
72 7?,, ANNEALED 8hr AT 1600°F o ]
o 8 | AS WELDED - IRRADIATED
R e - T TTTT 7T T ANNEALED 8hr AT 1600°F - IRRADIATED 13 T T
64 Fu—O—F — e = —0
- E T ®©
I <L_ E . - b —— _ -a
w Y < _J
T w o
56 " 7?-7 - =
~ & 33
11 R | -
48 - x ***2 ® *'&*nn *ul':
l " o g w 8 o
T .88 755
x = oW
40 = ] - ! 75 I -
| Y
: = =I
ST
\ J \ J -
Y
7°F 1200°F
Fig. 5. Comparison of the Fracture Strains of Welds and Base Metal
in Tensile Tests at a Strain Rate of 0.05 min~>.

ek e A G S e R Tk IS L e o e
11

Welds 1, 4, and 5 involve air-melted materials with relatively high
boron levels. Heat 2477, a vacuum-melted heat containing & ppm B, was
used as the base and filler metal in weld 3. Weld 6 utilized filler
metal from heat ©65-552, a vacuum-melted heat containing 1 ppm B, and
base metal from heat 5065, an air-melted heat. Welds 7 and 8 utilized
air-melted base and filler materials. The filler rod for weld 7 was
plasma-spray-coated with Al,03 in an effort to reduce the interdendritic
spacing in the cast weld metal and to provide additional sites for helium
collection. However, most of the Al,03 floated on the weld and had to
be removed mechanically before making the next pass; the aluminum content
of the deposited weld metal was only 0.06% compared with 0.84% for a
typical cross section from the filler metal before deposition. Weld 8
involved a filler material coated with WC. The deposited weld metal
contained 2% W and 0.1% C, so the WC coating dissolved in the melt
altered the properties of the weld appreciably.

The fracture strains shown in Fig. 5 are compared for
1. base metal given a standard solution anneal of 1 hr at 2150°F,
2. base metal given the 1 hr at 2150°F treatment followed by an anneal

of 8 hr at 1600°F,
3. transverse weld samples in the as-welded condition, and
4. transverse weld samples annealed for & hr at 1600°F.
These properties are considered for unirradiated and irradiated condi-
tions where data are available.

At a test temperature of 75°F the following points are illustrated

in Fig. 5.
12

1. The fracture strains of all welds are lower than those of the
base metal.

2. Welds 1, 4, and 5 all involved air-melted alloys and exhibited
the lowest fracture strains.

3. Welds 3 and 6 involved vacuum-melted weld metal and have the
best properties.

4. Welds 7 and 8, which involved air-melted alloys with the filler
metal modified with Al,05; and WC, exhibited intermediate properties.

5. Postweld heat treatments generally improved the properties.

6. Irradiation decreased the fracture strain of all welds and
base metal. The welds had lower postirradiation ductilities than the
base metals, but the welds were affected less by irradiation.

7. The postirradiation ductilities of welds did not depend
appreciably on whether they had received a postweld heat treatment prior
to irradiation.

8. The postirradiation properties of all welds except No. 8 were
equivalent.

Figure 5 also illustrates several important points at a tensile
test temperature of 1200°F.

1. The fracture strains of all welds were lower than those of the
base metal.

2. The as-welded ductilities of all welds except No. 8 were about
equivalent, ranging from 12 to 18%. Weld & had 25% strain at fracture

under these conditions.
i3

3. A postweld heat treatment of 8 hr at 1600°F markedly improved
the fracture strain, with values of 22 to 30% being observed. Weld &
again was better.

4. After irradiation, all welds were less ductile than the base
metal. There were some differences in the ductilities of the welds, but
no consistent trends are apparent.

5. Postweld annealing for 8 hr at 1600°F had little effect on the
postirradiation properties.

At 1600°F, Fig. 5 shows the following trends.

1. The fracture strains were drastically lower for welds than for
base metal. Welds 3 and 8 had properties superior to those of the other
welds. DPostweld heat treating had a beneficial effect on the ductility.

2. Irradiation reduced the fracture strain to about 2% independent
of whether the test sample was base metal or a transverse weld.

All of the postirradiation results presented thus far have been
for materials irradiated at less than 300°F. We have irradiated several
samples of these heats and others at 1200 to 1400°F and find that there
are at least two significant differences in the results obtained. The
displacement damage anneals at these irradiation temperatures and the
tensile properties up to about 1000°F are the same for irradiated and
unirradiated materials. At higher test temperatures the fracture strain
is reduced even Turther if the irradiation temperature is in the range
of 1200 to 1400°F. For example, heat 5065 was found to have a fracture
strain of 11.3% (irradiated at 1200°F) at 1200°F compared with the
value of 22.2% (irradiated at less than 300°F) shown in Fig. 5. Weld 7

was observed to have fracture strains at 1200°F of 7.2 and 10.8% after
14

irradiation at 1200°F and less than 300°F, respectively. Thus, the
trends shown in Fig. 5 for samples irradiated at 300°F seem to hold for
irradiation temperatures in the 1200 to 1400°F range, although the actual

fracture strains may be lower for the higher irradiation temperature.

Creep-Rupture Properties*

The stress-rupture properties at 1200°F of several heats of base
metal are shown in Fig. 6. The data are described reasonably well by
a single line, although both air- and vacuum-melted materials are
involved. The line for the base materials is shown in Fig. 7 where a
comparison is made between the stress-rupture properties of the base
metal and the various welds involved in this study. All of the welds
had lower creep-rupture strength in the as-welded condition than the
base metal. Weld 3, which involved only vacuum-melted materials, had
significantly higher strength than the other welds. Welds 7 and 8, which
involved A1,0; and WC additions, also showed notably better performance
than welds 1, 4, and 6. The rupture lives of the welds approached those
of the base metal as the stress level was reduced; this is likely due to
thermal recovery of the weld metal during the long time at 1200°F. The
strength was improved greatly by a postweld anneal of 8 hr at 1600°F.

The fracture strains of the welds and several heats of base metal
are compared in Fig. 8. The fracture strain of the base metal was about
25% for short rupture lives of a few hours and decreased to values of
10 to 15% when the rupture life was 1000 hr. There is no appreciable

difference between air- and vacuum-melted alloys. The welds had fracture

 

*See Tables A-3 and A-4, Appendix, for the tabulated creep-rupture
data.
15

ORNL-CWG 68-12879

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

70
TR Il
° 5065
\ a 5067
60 o\} o 2477
b\\“
50 N
0 \
N
o 40
Q
o N
l&‘ 30 D\\ o
@ N
\\
20 N
10
0
o 2 5 10 2 5 10% 2 5 100 2 5 10°

RUPTURE TIME (hr)

Fig. 6. Stress-Rupture Properties of Hastelloy N Base Metal at
1200°F.

ORNL-DWG 68-12880

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

70
N n
BASE METAL
WELD 3
60 ~ 1
\ N N \\\
N\ 0\\2 Y’ \\\ i
50 — ] /,)N\ \\ ™ \\ 1:
0N NG RR |
= N
2 \ NG TN \\ )
8 a0 &{ Sh, ; ,
Q N AN ‘
z \\ ™ \\A\ 1
h A\ \\ \\\N !
) WELDS 1,4,6 \3 NG -\\\ |
~ 30 O ;\\ 2
* NG TN NLWELD 8
NG TN
NG TN
O WELD1 N N
AWELD3 \fiL______ BN
20 O WELD 4
vVWELD®SB )
0 WELD 7 j
<WELD B
10 OPEN POINTS -AS WELDED T g
CLOSED POINTS—ANNEALED ;
8hr AT1600°F | *
. DL 1L |
10° To) 102 103 104

RUPTURE LIFE (hr)

Fig. 7. Stress-Rupture Properties of Several Hastelloy N Welds at
1200°F.
16

ORNL-DWG 68-12884

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

28 *3
26
24
A
22 ' ~
20
o
; 5
18 ¢
.<?.
3 & HEAT 5065 &
~ 16 A HEAT 2477
< : A
= o WELD <1 A
x A WELD 3
2 va 0o WELD 4 A
o v WELD 6 4
D
E 0 WELD 7
3 < WELD 8 N '
w 12 CLOSED SYMBOLS-ANNEALED &S
8 hr AT 1600 °F Al
BEFORE TESTING >
10
O
8 A
Fa
6 v A A
. T
G
4 T F
& o O
O ’:]JP VC‘ S |G <> 0
0
2 %* v
J
0
107! 10° 101 102 103 104

RUPTURE TIME (hr)

Fig. 8. Variation of Fracture Strain with Rupture Life for Welds and
Base Metal Tested at 1200°F.
17

strains of only about 3% for rupture lives up to a few hundred hours.
Samples having rupture lives over 1000 hr showed some improvement with
strains in the range of 3 to 5%. Weld 3 had much superior fracture
straing of about 6% for all test conditions. All welds and even the
base metal exhibited much higher fracture strains after an anneal of

8 hr at 16C0°F.

As shown in Fig. 9, irradiation produces some rather dramatic
changes in the stress-rupture properties. The lines shown in Fig. 9
were obtained from Fig. 6 for the unirradiated base metal, Fig. & for
the minimum strength of unirradiated welds, and from Ref. 6 for the
strength of irradiated base metal. Several points are included for
base metals irradiated under exactly the same conditions as the welds.
The properties of heat 2477 (low boron, vacuum-melted) are superior to

ORNL —DWG 68 — {2882

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

70
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"N\ UNIRRADIATED
BASE METAL
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NNIRRADIATED N
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* ANNEALED 8hr AT 1600 °F
o L1 1L oLl
0% 2 5 0! 2 5 t0° 2 5 10> 2 5 0%

RUPTURE TIME (hr)

Fig. 9. Influence of Irradiation on the Stress-Rupture Properties
of Hastelloy N Welds and Base Metal at 1200°F. All welds in the as-
welded condition unless otherwise indicated.
18

those of heat 5065 for the low irradiation temperature of less than
300°F, but further work has shown that these materials have comparable

properties when irradiated at 1200°F or higher.®

The rupture lives of
the welds seem quite comparable with those of the base metals and all
welds seem to have about the same properties. A single specimen was
given a postweld heat treatment of 8 hr at 1600°F and had superior post-
irradiation strength, but this observation should be repeated before con-
cluding that this anneal is effective. Moteff and Smith!® also found
that the postirradiation creep properties of Hastelloy N base metal and
welds were equivalent at 1200°F. They also found that a postweld anneal
of 4 hr at 1600°F did not improve the postirradiation properties of
welds.

The fracture strain is shown in Fig. 10 as a function of rupture
life. The fracture strain of the base metal had a low value of about
0.5% for rupture lives of only a few hours and increased to aboutVB% for
rupture lives of 1000 hr. The welds fall into this same pattern, but
the strains were consistently slightly lower. Again the single test on
a sample having a postweld anneal exhibited better properties.

We also compared the minimum creep rates for these same materials
to obtain a measure of strength. The creep properties of several heats
of base metal in the irradiated and unirradiated conditions are shown in
Fig. 11. Most of the data are contained within the band in Fig. 11,
with heat 5065 favoring the right side and heat 2477 the left side.

This same band is transferred to Fig. 12, where the data for the welds
are shown for comparison. The data do not deviate significantly from
the band obtained for the base metal. Thus, the creep rate of Hastelloy N

is not affected appreciably by welding or by irradiation.
19

ORNL-DWG 68—12883

 

]

HEAT 5065
HEAT 2477
WELD 3
WELD 4
WELD 6
WELD 7
+ WELD 8
* ANNEALED 8nr AT {600°F

 

 

 

&« 4 B P DO

 

 

 

 

STRAIN AT FRACTURE (%)
n

 

 

 

 

o
a®
X
%

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

o 2 s w0 2 5 102 2 5 03 2 s 0%
RUPTURE TIME {hr)

Fig. 10. Influence of Irradiation on the Fracture Strain of
Hastelloy N Welds and Base Metal in Creep at 1200°F. All welds in the
as-welded condition unless otherwise indicated.

ORNL—-DWG 68-12884

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

70 AT
/
/1
60 AL
Y |
A a /
A
| /]
Y
= / //fi
8 40 { "...
A
_9; ,/ . /
fi 30 .:‘_I’-.Qn . o HEAT 5065
& .,/ F/ 4 HEAT 5067
" ° / o HEAT 2477
20 3 OPEN SYMBOLS — UNIRRADIATED
1 A CLOSED SYMBOLS - IRRADIATED
10
0

10°% 2 5 073 2 5 02 2 5 ' 2 5 10° 2 5 10
MINIMUM CREEP RATE (%/hr}

Fig. 11. Influence of Irradiation on the Creep Rate of Hastelloy N
Base Metal at 1200°F.
20

ORNL-DWG 68-12885

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

ST T ’
/ g
O WELD 1 s/ 4
A WELD 3 s .
60} o WELD 4 // .
v WELD 6 p 4
0 WELD 7 v,<raoo,,’
< WELD B q 7/
50 [OPEN SYMBOLS - UNIRRADIATED /,/ #
CLOSED SYMBOLS-IRRADIATED o, £
% q
2 / nd
S 40 Lot —oim »
o 4 1
O LA 7
= A /Y
A e, MY | S
& 30 ApSwL—
Wy 7 /‘
y ¥
/ /’
20 dF<fg /4
7/ s
10
0
10°% 2 5 4073 2 5 1072 » 5 101 2 5 40°

MINIMUM CREEP RATE (%/hr)

Fig. 12. Influence of Irradiation on the Creep Rate of Hastelloy N
Welds at 1200°F.

Postweld Heat Treatments*

We have mentioned previously the improvement of the properties of
welds by a postweld heat treatment of 8 hr at 1600°F. We chose this
treatment on the basis of our work with weld 1 in which we tested samples
at 40,000 psi and 1200°F after annealing samples for various periods of
time at temperatures of 1200, 1400, and 1600°F. The influence of these
treatments on the rupture life is shown in Fig. 13. The as-welded sample
had a rupture life of only 19 hr compared with a rupture life of about
350 hr for the base metal (Fig. 6). Long annealing times at 1200°F were

required to produce much improvement, and the rupture life was only 60 hr

 

*See Table A-5, Appendix, for the tabulated results for creep-
rupture tests.
21

RNL-DW -
1000 ORNL-DWG 65-10798R

ANNEAL ING

ENVIRONMENT
500 Hp Ar

4+ & 1600°F
A $400°F
200 e 1200°F

100

50

RUPTURE LIFE (hr)

20

 

10
01 02 05 1 2 5 10 20 50 t00 200 500 1000 2000

TIME AT TEMPERATURE (hr)

Fig. 13. Influence of Postweld Heat Treatments on the Rupture Life
of Weld 1 Tested at 1200°F and 40,000 psi.
after a 1000-hr postweld heat treatment at 1200°F. Annealing at 1400°F
brought about more rapid recovery with the rupture life of the weld being
comparable with that of the base metal after a 50-hr anneal. Only about
8 hr at 1600°F was required to restore the rupture life. We have no
explanation for the apparent reduction in the rupture life due to a post-
weld anneal at 1600°F of 50 hr compared with anneals of 8- and 100-hr
duration.

The change in the fracture strain with postweld heat treatment is
shown in Fig. 14. The fracture strain is about 12% for the base metal
(Fig. 8) and the welds can be annealed to obtain equivalent or even
better fracture strains. There is an apparent inversion in the recovery
curve at 1600°F; this corresponds to the conditions producing short
rupture lives in Fig. 13.

The reduction in area for these same samples is shown in Fig. 15.

Values for the welds as high as the 15% obtained for base metal can be
22

 

ORNL-OWG 68-12886
28 T I

ANNEALING
ENVIRONMENT
Hp Ar
24 —— 4 0 1600°F
& 1400°F \
0 1200°F /

20 /

 

 

 

¥ o .

FRACTURE STRAIN (%)
™~
\
)
el
7
f
0
N
\‘
1
1
1
1
=

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

i
0 s 0% 2 5 10 2 5 0% 2 5 403
TIME AT TEMPERATURE (hr)

 

 

bt
o‘r M IJ L

 

Fig. l4. Influence of Postweld Heat Treatments on the Fracture
Strain of Weld 1 Tested at 1200°F and 40,000 psi.

o ORNL-DWG 68-12887

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

28 Ir T
24
C

20 p “)
= ANNE ALING V
& ENVIRONMENT /
EJ H Ar / IT fnaat
&6 —— 2 pomt
a & o 1600°F / l
> A 1400°F _ 2
s o 1200°F
& M
E 12 :
> /] /
o // /
& A /

8 A /?

/J /
LA L/
4 = ,.—/
/1 et i
41 T ? T

 

 

 

 

 

 

 

 

 

 

5 109 2 s w0 2 5 102 2 5 103
TIME AT TEMPERATURE (hr)

Fig. 15. Influence of Postweld Heat Treatments on the Reduction in
Area of Weld 1 Tested at 1200°F and 650°C.
23

obtained by postweld anneals at 1400 and 16CO°F. The unusual behavior
due to annealing for 50 hr at 1600°F again appears for two of the
samples. The reduction in area is not affected appreciably by annealing
st 1200°F, indicating that the improved elongation at fracture (Fig. 14)
is due primarily to the formation of cracks during testing.

We examined the influence of whether the postweld annealing environ-
ment was argon or hydrogen at 16C0°F, since Gilliland and Venard® had
reported better properties for samples annealed in hydrogen than for
those annealed in argon. Comparative points are shown in Figs. 13, 14,
and 15 for these annealing environments, and no appreciable differences

are apparent.

Metallography

Many of the samples were examined metallographically and a few
typical microstructures will be presented. Figure 16 shows the fusion
zone of weld 4. The structure of the base metal is characterized by
stringers of Mg(C~type precipitates and a relatively fine grain size.

The weld metal contains a fine cellular structure within coarse grains.
The fusion line is quite sharp and the transition of some of the large
M¢C precipitates to a lamellar phase is apparent. The exact identifica-
tion of the lamellar phase has eluded us for some time, but our evidence
indicates that it is a Mo-Ni-Si compound.!® Microprobe studies have
shown that the precipitates are enriched in silicon and that the dark
etching portions of the dendritic cast structure are also higher in

silicon.
24

A o =R - - - - 3 3 s 5 L
Re - : A=
R TR R R
- FONr. ] 2 . : 2 - = el ¥, 1
5 SR > L SRR TR
- -~ N oy - > j T ;‘\}‘-:1—: - \ "-'} "
: < S St .
oy e R AN v .fi\_-
- ¢ T N L 2 R it B fflqf:’ :‘.P'& Ok \l.:-t \ - ~
) ; - : S SRt A
SEvte et N —] Rt \.““in
& ¥ ARG A A O R B, 8
o = % Ny SR NN
‘ s = = a e e e,
s : DR R S S,
y A :'.: - = Ji : I".\ """'-_._;’:1 l'Lh ““T‘.-vh\ H'::‘ Il\"hq. e
g S e R USSR
\ \'-,_\‘ ¢ 'F'I ‘i\ N < -_\“‘-
NG T s A ) -
l\r ,‘).--\_"‘\‘:‘ 'l---J L’; F : A
% A . . £ L “‘xfk"\_l ‘iw 1
. 3 E‘_\ - . F‘?.*\ .'5_-\'. ; A \___L
- Sy .
= c T A LTk e Jitt\ Wi ™
_ TN e R S W BSRN TR T . 4
« sl v R NI B K oy ; HEETES
= N . g TR AT N ' . < iy
. _4_".-%.‘__‘ ; v ; _'.. *_"' j? 7 L o 7 - -4 2
Sl war PRI TR £ X% ¢ w\ ol ol ; ; =t
- " L I o '\;J‘t—fif,‘.']'.- :“"-‘. f - "l \ . \' £ ’ . %, % ‘
(a) » f b 13 00 Jr\ 3 \:J,' X \; ",-'Ir N ¥ t e g < -
- Wty Y-63005
- B i y oy
1 » &
v o -
\ e -
~—g™ o
b a .l "‘. - h - . \_h
- H - ' é.
'w \l"“ - -
. - b -
. . S L - ,l'"\‘__‘
- - - 3 § o
" W i g vy .
- e
i3
- . p /J # I‘1..
— - “* 2 ?‘h
3 & - h-'k_ -
r-fl?-'— o A “ ;"'
- n *a ‘*"-14 ‘ Y
Y - -
- g Sk Y <
"
3 - ¥ YR e
.-

(b) b —
Fig. 16. Photomicrographs of the Fusion Zone of Weld 4. (a) Base
metal on left and weld metal on right. 100x. (b) Transformation of
stringer precipitates to lamellar product at the fusion line. 500x.
Etchant: glyceria regia.
This same general description applies to the fusion zones of the
other welds except 3 and 8. The base metals used in weld 3 were vacuum-
melted and the stringers are minimal. The bulk silicon content was low

(0.015%) and the precipitates were primarily of the MosC type and were

found to dissolve easily. Thus, not much of the lamellar transformation

 
25

product was present at the fusion line. Weld 8 involved the filler mate-
rial that was coated with WC, and the structure of the deposited weld
metal is much finer (Fig. 17).

A typical microstructure of a creep-rupture fracture in Hastelloy N
base metal is shown in Fig. 18. This sample was tested at 40,000 psi and
1200°F and failed after 312 hr with 16.6% strain. There is considerable
grain boundary cracking and the fracture is predominantly intergranular.

Typical microstructures of a tested sample from weld 1 are shown
in Fig. 19. This sample was tested at 40,000 psi and 1200°F in the as-
welded condition; failure occurred in 18.7 hr with 3.1% strain and only
0.77% reduction in area. The fracture occurred in the weld metal and
followed the grain boundaries; however, little grain boundary cracking
adjacent to the fracture is evident.

Postweld heat treatments generally improved the tensile and creep
properties. Figure 20 shows the fracture of a sample from weld 1 that
was annealed for 2 hr at 1600°F and tested at 40,000 psi and 1200°F.

This sample failed in 200 hr with 9.4% strain and 6.4% reduction in area.
The failure still occurred in the weld metal; however, intergranular
cracks did form throughout the weld zone, indicating that deformation

occurred generally throughout this zone.

SUMMARY

This study has shown that the mechanical properties of welds in
Hastelloy N are generally inferior to those of the base metal. The
strain at fracture is the property affected most severely, although

appreciable reductions in the high-temperature creep-rupture strength
e -~ : o 20s
- AT N N s i Ntk y b} Y-662

g- X _\::r:::‘ - -‘--:_‘fl\n\-'-l\g F‘ 'Qtfi*- fiI"“% fik}r“:’\r‘fi | ". 1'11 Rk o .

'-\"h. gt gt e { ki n‘._. o T

2 e e e TR ? . -1\,-.-“-11. ”_‘1.\“ 3 Pt .\ ]

SN SCTEANGR AR T b L

& T:';;"' "\7\1{" Ffi Y .}(\F"fl‘:'“x.-\\# Py """ 0 N ~

:fi.dfl?_ E:E o - ;\‘ o \ h o 'li aj ‘u;\:',lu‘ ANy [\‘H-'Air" \. I. :

N = Vol e ol YW L | AL Ry

N < C 1 INEEL o :fi"“‘n' At

 

E.':_‘;'ffi‘ A .: ¥Lli'-l.1;{t I’l“. .!1‘_ in' ;.l;‘ wfi :‘ ‘:,‘" y -'. .\
W b [ I -JL “”“”; "'\}1\ \\ 4 s

 

 

1_ & b ‘l i
F " l\1I|'ul - \:i: Y‘L_‘_?‘!\Q!""fi ?-
v X : Ly
\ 1‘ I;l. J ‘;‘xl"-l'.": )‘;l._ .I.-'?.'I'-‘l ‘I t‘.‘ .-{ ‘.\lh P v .
\ e NS g _'.. = % A . %
e AN MEIRL B, A - r.¥
.‘.i‘fi..‘}i;\%‘:“ I g \ g 3
1IN ety &,
g W A -‘.'].'il
o fi o W {
j-" % "‘,J' ™, iy A ¥
e N i ¥t he ‘ol
: Vg i 0
- ' o
fol 1 ::I b 3 |
- - % 'I * =
c f i .n. i {
oy iy W iy
. ' L | 6
. I w1
g s " : I
v 3 . = L
i I - ¢ I','_ . y . wis
. i AT A Y o ¥ § ‘l., i
A . LT 2 ¥ ¢ ' |
R Nt TN % = ¥ fx. N
* o el 1 A '
#E I I—| |'i.; 1 ol I,{ “I'l B A
'y \ g . LT Al
i | Vi 4 g

Fig. 17. Photomicrograph of the Fusion Line of Weld 8. Weld metal
in the upper left and base metal in the lower right. 100X. Etchant:
glyceria regia.

 

‘ : e ,.f | : o T ’ A b g Y-6L729
Y~ ¥ J,li 1 \l . j / E ?\ E (.\. o
{ X QYT R S ¥
Py Y, PSRk e
7 CEF Ly el e A
- ( ) ; I { P NS :4 = K
\ y Eal? L i M-
} ] e 5 : { -,.1/ "I,_f‘ : ;}:w e
) - ot Py il J: 1s Yoy r L
E‘r=J" 4 > 3 l / Y ol ‘-( i - " A e i /
>’ 1 : . ‘ \ \ f s = ~.r = LS

A |_. , s i \ i j . ' o .r
: BN =l T Y G 7 -
i I i 3 " - e )( -

Fig. 18. Photomicrograph of the Fracture of a Hastelloy N
(Heat 5065) Sample Tested in Creep at 40,000 psi and 1200°F Following
a Pretest Anneal of 1 hr at 2150°F. Failed in 312 hr with 16.6% strain.
100x. Etchant: glyceria regia.

 
 

- - -
- t .
- - - - = § i &
44 N L ¥
v -l
v - '
- A 1 L Pl
? .
i - 4 i 8 g
o 1 g
f ol 5
= \ L - N e
. i
# . ’
‘ 2 } .
iy I
i \
5 b '
& -
- .
Ly -~
- £
- J
. S _II
¥
. g - 1 ¥
& 2 1
- -
; ; > Bl
A f F gk
¥ = L
-
i .
1 ry
& o r r ™
e i
¢ =
= P
. -

Y-66857

 

Fig. 19. Photomicrographs of Sample from Weld 1 Tested at
40,000 psi and 1200°F in the As-Welded Condition. Failed in 18.7 hr
with 3.1% strain. 100x. (a) Fracture in the weld metal. Etchant:
glyceria regia. (b) An unetched photomicrograph of a cracked area
away from the fracture.

 
 

(o) L

Fig. 20. Photomicrographs of Tested Creep Sample from Weld 1,
Sample given a postweld anneal of 2 hr at 1600°F in argon and tested
at 40,000 psi and 1200°F. Failed in 200 hr with 9.4% strain. 100x.
(a) Fracture. Etchant: glyceria regia. (b) An unetched photomicro-
graph of deformed area in the weld metal.

 
29

also occurred. The properties of welds can be improved by suitable post-
weld heat treatments to almost equal those of the base metal.

Several welds were studied which involved air- and vacuum-melted
materials. The weld involving vacuum-melted base and filler metal
(weld 3) had superior creep-rupture properties in the as-welded condition.
Weld 8 involved air-melted materials and the filler metal was coated with
WC. The properties of this weld were generally superior to those of the
other welds, but the fracture strain under creep conditions was not
improved.

Neutron irradiation caused a decrease in the fracture strain of
both welds and base metal. The postirradiation fracture strains in
tensile tests of the welds and the base metal became closer as the test
temperature was increased; they were identical at 1600°F. Although post-
weld heat treatments improved the fracture strain in preirradiation
tensile tests, the postirradiation fracture strains of welds seemed
independent of postweld anneal. The postirradiation creep-rupture proper-
ties at 1200°F were about the same for welds and base metal, even though
the preirradiation properties differed greatly. Although some of the
irradiated welds showed improved performance in tensile tests, they all

had about the same creep-rupture properties.

ACKNOWLEDGMENTS

The welds were made under the supervision of T. R. Housley and the
mechanical property tests were run by B. C. Williams, E. Bolling,
N. O. Pleasant, J. T. Feltner, and V. G. Lane. H. R. Tinch was respon-

sible for the metallographic work. We are also indebted to
30

G. M. Slaughter and J. R. Weir for their interest in this work and for
their careful review of the manuscript. The drawings were prepared by
the Graphic Arts Department and the manuscript was prepared by the

Metals and Ceramics Division Reports Office.

REFERENCES

1. W. D. Manly et al., "Metallurgical Problems in Molten Fluoride
Systems," Progr. Nucl. Energy Ser. IV 2, 164~179 (1960).

2. H. G. MacPherson, "Molten Salt Reactor Shows Most Promise to Congerve
Nuclear Fuels, Part 2," Power Eng. Zé(Z), 56-58 (1967).

3. R. G. Gilliland and J. T. Venard, "Elevated Temperature Mechanical

Properties of Welds in Ni-Mo—Cr—Fe Alloy," Welding J. (N.Y.) 45(3),

 

103-5-110-s (1966).

4. R. G. Gilliland and G. M. Slaughter, private communication.

°. W. R. Martin and J. R. Weir, "Postirradiation Creep and Stress
Rupture of Hastelloy N," Nucl. Appl. 3, 167 (1967).

6. H. E. McCoy, "Variation of the Mechanical Properties of Irradiated

Hastelloy N with Strain Rate," submitted to Journal of Nuclear

 

Materials.
7. D. R. Harries, "Neutron Irradiation Embrittlement of Austenitic

Stainless Steels and Nickel Base Alloys," J. Brit. Nucl. Energy Soc.

 

5, 74 (1966).
8. G. H. Broomfield, D. R. Harries, and A. C. Roberts, "Neutron Irradia-
tion Effects in Austenitic Stainless Steels and a Nimonic Alloy,"

J. Iron Steel Inst. (London) 203, 502 (1965).

1 e g~ gD A B A R e S N ast L, L L
31

9. F. C. Robertshaw et al., "Neutron Irradiation Effects in A-286
Hastelloy Y and René 41 Alloys," Spec. Tech. Publ. No. 341, p. 372,
American Society for Testing and Materials, Philadelphia, Pa., 1963.

10. N. E. Hinkle, "Effect of Neutron Bombardment on Stress-Rupture
Properties of Some Structural Alloys," Spec. Tech. Publ. No. 341,
p. 344, American Society for Testing and Materials, Philadelphia, Pa.,

1963.

11. P.C.L. Pfeil afid D. R. Harries, "Effects of Irradiation in Austenitic
Steels and Other High-Temperature Alloys," p. 202 in Flow and
Fracture of Metals and Alloys in Nuclear Environments Spec. Tech.
Publ. No. 380, American Society for Testing and Materials,
Philadelphia, Pa., 1965.

12. J. T. Venard and J. R. Weir, "In-Reactor Stress-Rupture Properties
of a 20 Cr—25 Ni Columbium-Stabilized Stainless Steel,” p. 269 in
Flow and Fracture of Metals and Alloys in Nuclear Environments Spec.

Tech. Publ. No. 380, American Society for Testing and Materials,

 

Philadelphia, Pa., 1965.

13. P.C.L. Pfeil, P. J. Barton, and D. R. Arkell, "Effects of Irradiation
on the Elevated Temperature Mechanical Propertles of Austenitic
Steels,” Trans. Am. Nucl. Soc. 8, 120 (1965).

14. P.R.B. Higgins and A. C. Roberts, "Reduction in Ductility of Austenitic
Stainless Steel after Irradiation,” Nature 206, 1249 (1965).

15. J. Moteff and J. P. Smith, Sixth Annual Report of High Temperature

Materials Program, Part A, March 31, 1967, GEMP-475A, p. 185 (1967).
32

16. R. E. Gehlbach and H. E. McCoy, "Phase Instability in Hastelloy N,"
paper presented at the International Symposium on Structural
Stability in Superalloys, Seven Springs, Pa., Sept. 4—6, 1968;

to be published in the proceedings.
APPENDIX
35

Table A-1l. Tensile Properties of Welds in the Unirradiated Condition

 

 

 

- Test . - Reduction
Specimen Weld  Postweld S;:::n Tenpera- Stress, psi Elongation, % in Area
Number Number  Anneal . =1 ture Yield Ultimate Uniform Total (%)
( min ) ( o F)

13 1 None 0.05 75 g1,100 113,100 22.0 22.3 37.7

14 1 None 0.05 390 59,100 81,200 11.3 12.3 24 4

15 1 None 0.05 1200 50,000 64,800 6.0 7.8 13.1

16 1 None 0.05 1200 40,800 42,000 3.0 5.8 6.9

17 1 None 0.05 1600 23,200 23,400 1.8 21.8 25.5

18 1 a 0.05 1600 41,600 42,100 3.9 19.2 18.6
3000 1 None 0.05 75 76,300 111,900 27.1 28.3 29.5
3001 1 None 0.05 1200 59,300 79,800 10.2 11.7 26.5
3005 3 None 0.05 75 71,700 122,800 35.2 37.2 42,4
3006 3 None 0.05 1200 53,200 77,300 12.6 14.4 23.6
10285 3 None 0.05 1600 41,300 44,900 T 13.8 14.8
10287 3 a 0.05 75 61,000 123,700 44,0 46.1 41.02
10286 3 a 0.05 1200 43,700 82,400 21.4 22.5 26.93
10300 3 a 0.05 1600 40,900 44,200 5.3 21.7 18.15
3007 4 None 0.05 75 74,300 108,900 24.9 25.9 47.8
3008 4 None 0.05 1200 55,300 72,300 13.5 15.3 31.7
10288 4 None 0.05 1600 40,900 44,700 3.1 5.9 7.66
10289 4 a 0.05 75 57,700 107,500 31.4 31.9 38.4
10290 A a 0.05 1200 39,300 76,900 29.6 30.4 35.4
5 None 0.05 75 73,300 119,800 33.0 34.2 45.1

5 None 0.05 1200 51,100 70,100 14.2 15.9 42.8

10299 5 None 0.05 1600 43,300 45,400 1.7 11.7 18.0
10296 5 a 0.05 75 63,800 119,500 32.7 33.1 41.1
10297 5 a 0.05 1200 49,200 91,300 25.6 25.9 38.8
10295 5 a 0.05 1600 39,000 39,400 1.4 36.1 62.6
1144 6 None 0.05 75 68,200 120,900 38.7 40,7 38.3
1145 6 None 0.05 1200 49,800 81,300 16.6 18.6 20.9
1147 6 None 0.05 1600 39,000 44,400 2.2 2.9 3.04
1140 6 a 0.05 75 63,100 119,100 37.0 38.5 47.3
1141 6 a 0.05 1200 44,400 85,400 29.0 29.9 36.8
1142 6 a 0.05 1600 39,900 44,500 4.5 13.9 16.7
1164 7 None 0.05 75 71,900 125,700 32.3 33.8 45.0
1165 7 None 0.05 1200 52,000 79,100 13.7 14 .4 11.9
777 7 None 0.002 46,900 65,700 7.3 8.6 18.7
1166 7 None 0.05 1600 41,500 45,100 2.0 3.9 7.8
10284 7 a 0.05 75 64,100 119,400 33.8 34.5 34.9
10283 7 a 0.05 1200 45,100 86,100 24.0 24.6 37.6
10295 7 a 0.05 1600 39,000 39,400 1.4 36.1 62.6
3011 8 None 0.05 75 72,300 124,900 35.8 38.8 49.4
g None 0.05 1200 54,000 91,300 22.9 24.9 29.1

778 8 None 0.002 1200 39,000 66,600 11.8 16.7 28.2
10280 8 None 0.05 1600 38,400 44,600 4.5 13.0 8.1
10282 8 a 0.05 75 58,900 120,100 33.9 34.0 28.4
10281 8 a 0.05 1200 41,700 91,200 35.7 37.0 37.7
10294 8 a 0.05 1600 39,800 43,800 4.3 53.5 6.4

 

29 hr at 1600°F.
Table A-2. Tensile Properties of Welds in the Irradiated Condition

 

 

 

Test
Temper- ) ] .
Specimen Weld Postweld Experiment S;::;n ature Stress, psi Elongation, % R?iuZEQZn
Number Number Anneal Number (min-1) (°F) Yield Ultimate Uniform Total (%)

469 3 None  ETR-41-30  0.05 1200 57,600 70,800 5.9 6.8 18.6
470 3 None ETR-41-30 0.002 1200 56,400 62,200 3.0 3.6 12.6
699 4 None ORR-149 0.05 75 109,900 120,500 16.8 19.2 48.7
700 4 None  ORR-149 0.05 392 92,900 113,300  18.9  19.1 25.3
701 4 None ORR-~149 0.05 1200 53,300 69, 300 8.6 8.7 12.0
703 4 None  ORR-149 0.002 1200 51,900 63,400 4.7 5.2 12.8
702 4 None  ORR-149 0.05 1600 41,200 41,600 1.4 2.0 1.3
704 A None ORR=-149 0.002 1600 26,000 26,000 0.9 2.0 9.2
708 4 a ORR-149 0.05 75 104,500 122,800 17.2 18.7 14.5
709 4 a ORR-149 0.05 392 82,700 107,600 21.0 22 .2 27.3
710 4 a ORR~149 0.05 1200 42,200 63,300 9.8 10.3 21.9
712 4 a ORR-149 0.002 1200 45,400 58,300 5.6 6.6 28.1
711 4 a ORR-149 0.05 1600 38,700 39,600 1.4 2.1 1.8
713 4 a ORR-149 0.002 1600 25,300 25,300 1.0 2.0 5.9
457 4 None  ETR-41-30  0.05 1200 57,200 69,800 5.9 7.3 20.1
458 4 None ETR-41-30 0.002 1200 57,900 64,900 2.5 3.3 11.3
461 4 a ETR-41-30 0.05 1200 44,600 70,400 16.2 16.6 32.6
462 4 a ETR-41-30  0.002 1200 43,200 59,700 8.9  11.0 15.6
714 6 None  ORR-149 0.05 75 111,000 130,700 22.6  25.4 47.0
715 6 None  ORR-149 0.05 392 88,800 112,100  20.4  22.4 27.4
716 6 None  ORR-149 0.05 1200 45,100 72,000  14.2  14.6 13.2
718 6 None ORR-149 0.002 1200 51,700 63,300 6.1 7.0 30.4
717 6 None  ORR-149 0.05 1600 38,900 40,500 1.4 2.5 3.2
719 6 None ORR-149 0.002 1600 24,900 24,900 0.8 1.4 6.0
723 6 a ORR-149 0.05 75 106,500 127,100 21.1 22."7 33.4
724 6 a ORR-149 0.05 392 83,900 111,200 25.0 25.9 34.9
725 6 a ORR- 149 0.05 1200 44,700 60,300 8.8 9.5 6.1

9¢
 

 

 

Table A-2. (continued)
Test . .
. Stress, psi Elongation, % uctio
Specimen Weld Postweld Experiment S;:i:n Temper- Yield U;_:') . — g ,t ] Rii A;c':a.n
Number Number Anneal Number (min- atlgre 1€ imate niform Tota (%)
(°F)
727 6 a ORR-149 0.002 1200 43,500 53,600 4.6 5.4 9.2
726 6 a ORR-149 0.05 1600 36,400 40,100 2.0 5.3 11.9
728 6 a ORR-149 0.002 1600 18,600 18,900 0.9 1.1 5.1
529 7 None ORR-149 0.05 75 112,400 135,100 18.4 21.4 25.7
530 7 None ORR-149 0.05 392 89,700 114,000 18.8 20.1 34.0
531 7 None ORR-149 0.05 1200 45,900 75,000 10.6 10.8 5.6
533 7 None ORR-149 0.002 1200 53,400 65,200 5.0 6.4 5.6
532 7 None ORR-149 0.05 1600 40,300 41,300 1.5 1.5 5.6
534 7 None ORR-149 0.002 1600 25,100 25,100 1.0 1.6 5,9
538 7 a ORR-149 0.05 75 107,700 122,900 14.6 16.7 54.2
539 7 a ORR-149 0.05 392 88,100 110,500 19.1 20.3 31.8
540 7 a ORR-149 0.05 1200 44,300 66,400 10.2 15,2 7.4
542 7 a ORR-149 0.002 1200 45,100 56,600 4.2 4.6 10.4
541 7 a ORR-149 0.05 1600 39,500 40,200 1.3 2.4 10.5
543 7 a ORR-149 0.002 1600 24,800 24,800 1.0 1.4 2.4
1162 7 None  ETR-41-31  0.05 1200 50,000 65,600 7.2 7.2
1163 7 None ETR-41-31 0.002 1200 48,300 56,300 3.4 3.5
729 8 None ORR-149 0.05 75 110,300 137,100 25.0 29.0 48.91
730 8 None ORR-149 0.05 392 86,700 121,200 30.0 33.8 47.7
731 8 None ORR~149 0.05 1200 45,900 75,000 14.6 14.9 16.7
733 8 None ORR-149 0.002 1200 48,200 64,100 6.6 6.8 11.3
732 8 None ORR=149 0.05 1562 38,700 41,400 1.7 2.0 2.7
734 8 None ORR- 149 0.002 1562 27,500 27,500 1.0 1.3 8.9
738 8 a ORR~149 0.05 75 108,900 134,800 27.5 30.5 11.7
739 8 8 ORR-149 0.05 392 88,400 117,900 25.4 26 .4 31.9
740 8 a ORR-149 0.05 1200 41,800 70,500 13.5 13.8 14.5
742 8 a ORR-149 0.002 1200 42,700 61,400 9.0 9.2 9.2
741 8 a ORR~149 0.05 1562 2.8 3.0 20.4
743 8 a ORR- 149 0.002 1562 27,900 28,000 1.1 1.6 6.0

 

&2 nhr at 1600°F.

LE
38

Table A-3. Creep-Rupture Properties of Welds at 1200°F
in the Unirradiated Condition

 

Minimum

 

Sample  Test Weld Postweld Stress Rupgure Creep F;act?re R?ductlon
Number Number Number  Anneal (psi) Life Rate Strain in Area
(nr) (%) (%)
(%/br)
35 399¢ 1 None 55,000 4.6 0.15 3.10 2.71
1 3780 1 None 40,000 18.7 0.013 13.1 0.77
33 5476 1 None 20,000 3864.6 0.0013 4.7 5.1
3 4080 1 None 20,000 1215.6 0.001 9.4 0.32
4éy 5583 1 None 20,000 3184.3 0.00080 3.6 1.9
2 3779 1 None 30,000 80.9 0.005 4.7 0.79
31 3996 1 a 40,000 308.8 0.029 19.4 13.2
30 5202 1 a 40,000 551.9 0.028 28.0 9.4
3345 5889 3 None 55,000 26.5 0.11 6.0 6.4
3345 5890 3 None 47,000 77.9 0.046 6.2 1.8
3346 5891 3 None 35,000 468.9 0.0072 7.4 1.8
4016 3 None 40,000 273.7 0.0163 7.8 g.2
5110 3 None 30,000 1228.0 0.0028 5.2 3.6
3395 5893 4 None 47,000 6.2 0.024 3.0 1.5
2 3997 4 None 40,000 25.2 0.0265 3.1 1.6
3 4013 4 None 30,000 80.2 0.0098 4.7 0
5894 4 None 21,500 495.9 0.0012 1.7 0.7
7060 5140 6 None 55,000 2.8 0.584 6.3 6.4
5881 6 None 47,000 6.3 0.0388 2.0 0.8
7061 5141 6 None 40,000 14.5 0.0326 3.2 3.2
1158 5882 6 None 30,000 158.2 0.0046 1.9 0.8
7127 5462 6 a 40,000 423.7 0.028 17.0 4.5
7071 5195 7 None 55,000 7.3 0.11 3.1 7.1
1168 5883 7 None 47,000 27.9 0.0406 3.6 1.0
7069 5192 7 None 40,000 47.8 0.016 3.1 0.8
7129 5464 7 None 32,400 179.0 0.070 2.5 0.6
7164 5312 7 None 30,000 194.8 0.0050 2.9 2.4
7135 5472 7 None 21,500 2173.2 0.0030 3.4 2.1
7126 5464 7 a 40,000 372.0 0.030 18.3 3.2
7073 5221 g None 55,000 129.1 0.080 4.4 4.8
3340 5884 8 None 47,000 26.5 0.048 3.0 0.8
7070 5194 g None 40,000 82.4 0.014 3.1 1.6
7125 5460 8 None 30,000 520.9 0.0035 3.5 4.0
7190 5576 g8 None 21,500 3172.2 0.0012 5.2 4.0
7128 5463 8 a 40,000 692.9 0.026 29.9 23.1

 

%2 nr at 1600°F.

 
Table A-4. Creep-Rupture Properties of Welds at 1200°F in the Irradiated Condition

 

 

Rupture Minlmum Fract Reducti
Sample Test Weld Postweld Experiment  Stress L?f Creep gtc gre ? uirlon
Number Number Number Anneal Number (psi) (1 S Rate raln in Area
hr) ey () (%)
471 R-90 3 None ETR-41-30 32,400 175.5 0.0064 1.59
706 R-36"7 4 None ORR-149 39,800 7.6 0,037 0.38 3.5
705 R-363 4 None ORR- 149 32,400 47.8 0.0037 0.20 4,3
459 R-83 4 None ETR-41-30 32,400 111.3 0.0031 0.59
463 R-71 4 a ETR-41-30 32,400 585.2 0.0056 3.52
720 R~ 107 6 None ORR-149 35,000 9.6 0.020 0.37
721 R-370 6 None ORR-149 30,000 51.4 0.0045 0.39 12.7
535 R-37"7 7 None ORR-149 39,800 17.7 0.0085 0.29
536 R-116 7 None ORR-149 32,400 45,1 0.0084 0.38
735 R-108 8 None ORR- 149 35,000 3.8 0.0605 0.37
736 R-365 g None ORR-149 30,000 210.8 0.0032 0.87 4.5

6¢

 

%% hr at 1600°F.
Table A-5. 1Influence of Postweld Heat Treatment on the Creep Properties
of Weld 1 at 1200°F and 40,000 psi

 

 

 

Postweld Treatment Rupture Minimum Fracture Reduction
Sample Test : . Creep . .
Numb er Number Tempfrature Time Environment Life Rate Strain in Area
(°F) (hr) (hr) (%) (%)
(%/hr)

1 3780 18.7 0.0125 3.1 0.77
22 3979 1200 100 Ar 20.3 0.0265 3.1 0.78
24 3958 1200 300 Ar 25.3 0.0557 9.4 2.4
27 5028 1200 1000 Ar 79.7 0.027 6.3 1.6
26 5112 1200 1000 Ar 39.8 0.027 9.4 1.6
42 5111 1400 2 Ar 46.6 0.014 4.7 2.4
21 3984 1400 8 Ar 97.3 0.020 4.7 3.6
19 3976 1400 20 Ar 77.6 0.032 6.3 5.1
32 3972 1400 50 Ar 320.9 0.023 12.5 7.9
34 4024 1400 200 Ar 296.6 0.030 15.0 17.2
45 5315 1400 1000 Ar 434.3 0.024 15.6 14.6

3996 1600 0.5 Ar 54,5 0.023 3.1 4.0
3989 1600 2 Ho 200.0 0.023 9.4 6.4

4 3793 1600 2 Ar 193.9 0.029 9.4 8.6

7 3817 1600 8 Ar 416.8 0.033 21.9 11.0
31 3966 16C0 8 Ho 308.8 0.029 19.4 13.2
30 5202 1600 8 Hp 551.9 0.028 28.0 9.5
37 3995 1600 50 Ar 151.7 0.032 6.3 7.3
38 4065 1600 50 Ar 193.6 0.046 12.5 9.4
43 5249 1600 50 Ar 266.3 0.033 15.6 21.3

0%

 
1-3.
4=5.

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