Temper Embrittlement of Pressure Vessel Steels

BY R. A. SWIFT A N D J . A. GULYA

Pressure vessel steels and weld metals operating up to 1100 F vary in tendency to temper embrittlement depending on composition, structure, heat treatment and, in submerged arc welding, the choice of flux

R. A. SWIFT is Research Engineer and J. A. GULYA is Supervisor, Alloy Research, Lukens Steel Company, Coatesvi/le, Pa.

ABSTRACT. Temper embrit t lement is a particularly important consideration for steels used in high pressure con­tainment vessels operating in the temperature range of 600 to 1100 F. On this basis, a comprehensive evaluation of commonly used pres­sure vessel steels and weld metals was undertaken. The steels investi­

gated include C-Mo, Mn-Mo, Mn-Mo-Ni and Cr-Mo steels. Weld metals in­clude C-Mo, Mn-Mo and Cr-Mo com­positions. A total of 95 composit ion-treatment combinations, including 20 composition-treatment combinations of weld metal, have been evaluated.

Of the materials tested, suscepti­bility to temper embritt lement varied with composition, structure, and heat treatment. The susceptibility was low­er for the C-Mo and Mn-Mo steels than for the Mn-Mo-Ni and Cr-Mo steels. The addition of Ni to Mn-Mo

apparently increases susceptibility to embritt lement. With in the scope of the tests reported herein, no f i rm trends in the effects of As, Sb, P, and Sn on the susceptibility to temper em­britt lement are observed.

The results of embritt lement tests on the weld metal were similar to those obtained for the corresponding plate steels. However, flux is a critical factor in determining susceptibility. It was shown that the susceptibility of weld metal is generally more severe than that of the base metal.

Table 1—Specifications and Chemical Analysis of Materials Investigated

Type

C-Mo

ASTM Specs (a)

A204C

Mn-Mo A302B

A302B, A533A A302C, A533B

1 Cr-Mo A387B

2V4Cr-Mo A387D, A542

3Cr-Mo

None

None A387E

Form

Plate Weld™

Plate Weld , b |

Plate

Plate Plate

Plate Weld |b |

Weld |b)

Plate Plate Plate Weld |b)

Weld ,b)

Plate . Weld(b» Plate Plate

Plate Plate Plate

Heat Mn

A2948 .36 .73 (RACO) .06 1.02

B5217 .19 1.34 (Linde) .07 1.32 A4069 .20 1.20

C4689 .21 1.27 C4694 .24 1.32

15756 .19 (RACO) .06 (L&R) .11

.13 12

.12

.13

.12

.12

B6730 A2766 A3588 (L&R-1) (L&R-2) C6669 (L&R-3) -(L-105)(C'11 A9367 .12

(L-104)(c)13 B3683 .12 A2966 .14

.57

.80

.43

.46 48 .50 53 .44 .46 .47 .49 .39

.51

.40

.43

.023 .009

.023 .015

.023 .008

.016 .012

.018 .009

.016

.015

.024

.012

.022

.017

.018

.015

.012

.017

.024

.011

.020

.009

.025

.021

.011

.009

.014

.015

.010

.011

.017

.017

.015

.010

.020

.008

.008

.009

.010

.010

Cu

24 .29

.25

.12

.11

.12

.11

.22

.24

.20

.17

.18 ,16 .16 .15 .17 .16 .19 .15

.19 23 .15

.23

.45

26 .43 .24

.25

.15

.30

.57

.35

.21

.26

.26

.27

.22

.21

.20

.29

.24

.31

.26

.29

Ni

.13 08

.13

.04

.08

.56

.56

.13

.19

.16

.17 ,18 .20 .17 .18 .48

Cr

.07

.05

.10

.08

2.06 2.27 2.28 2.25

.17

.31

.22 30 .11

Mo

.03

.54

,14 .49 .05 .47 .07 .47

.60

.55

.21 1.11 .51

.14 1.34 .54

.20 1.34 .49

.96

.88

.95

.95

.96

.95

.92

.97

.98

.18 3.12 .48

.17 3.00 .89

.22 2.90 .93

Al

.015

.013

.042

.008

.031

.014

.019

.023

.013

As Sb Sn

.018

.016

.008

.007

.008

.005

.007

.009

.015

.027 .0026

.024 .0023

.0034 .024

.0032

.0029

.012 .0025

21.0(l

18.6"

.017 18.7(( — 20.9

18.5 «J)

.012 .00015 .00 9.3

.012 .011 .0001 .00 10.2

.023 .018 .0040 .015 19.8

.010 .025 .0025 .013 18.9

(a) Related plate specifications (b) Sub-arc welds

(c) Laboratory melts (d) "X calculated assuming 0 015% Sn

W E L D I N G R E S E A R C H S U P P L E M E N T ! 57-s

+ 240

+ 200

+ 160

+ 120

+ 80

< +40

S 0

- 4 0

- 8 0

- 1 2 0

-160 SERIES TREATMENT MATERIAL

- Fig- f — Fracture appearance transition temperature data for C-Mo, Mn-Mo andMn-Mo-N

A—As Received or As Welded T—Stress Relieved S -S tep Cooled

—

o

< o CM

| r i < o

CO

o

< o «* -* 1 o < o CM

| r i < n

cn

i ) CO

1 O < o •>t —i 1

o < o CM

| ri < o m co -H '< <

< o <* I

•i <

o CO

1

< o «tf

1

< <

< o cv

I r i < o LC CO

—

O < 1 ) • = *

•-< 1 O

< n CM

o < o in CO

-"*

CJ CO

1 O < o vr

-—< u < o CM

I C)

< o i n CO

—• 5 >*

< n <* _̂, 1

5

<

O C/5 | < n «*

• i

<

o CO

CJ

< i n

< i

< o . -1 CO

i ; C/) 1

o < o 1(1

1 < ) < o i i CO

O < O i n

~ i

c > 1

o

: : CO

o UJ

1 o < n i n

—* t )

< o

: ! IO

o o 5 5

< < o o 2 9

1 L > si cp I 1 co

-o

& o

! CO

o

< LO r* CM

I O

<* o CO

t 1 u> 1

o < i n CV'

1 o < o

co

CJ

< i n CM

I O

< o in CO

C ) CO

1 O <

CM

| o < o m CO

CJ CJ CO CO

1 1 CJ CJ O CJ U_ U- LL. LL. o o o o in in un m

1

1 O

<-o lf> I ! )

| o o

1 O ' s i o n

cn

| O 3 o <* - * » * -

1 o 5

n̂ o 5 o

1 O £ o

1 2 3 A T S A204C

4 5 6 A T S WELD

7 8 9 A T S A302B

10 11 12 A T S WELD

13 14 A S

A302B

15 16 A S

A302B

17 18 A S

A533A

19 20 A S

A302C

21 22 A S

A302C

23 24 A S

A533B

25 26 A S

A533B

In t roduct ion

The development of economies in costs of material, fabrication, trans­portation and operation has promoted the use of alloy plate steels in in­creasingly close proximity to max­imum capability. Typically, steels are being used at the higher strength levels afforded by heat treatment and at more extreme temperatures, both for reasons for process economy. This trend is apparent even in the re­latively conservative approach used in the field of pressure vessel con­struction for power generation, LNG transport, and petroleum refining.

Under these conditions, properties or characteristics that were previous­ly considered secondary may become prime or even limiting. Fatigue, creep resistance, hydrogen res is tance , hardenability, t empe rab i l i t y and temper embrit t lement are among the features that fit this classification. Creep resistance and temper em­britt lement are particularly important considerations for steels used in con­tainment vessels operating in the temperature range of 600 to 1100 F. An evaluation of the creep properties of various pressure vessel steels has been presented.1 Subsequently, a pre­l im ina ry assessment of the propensity of these same pressure vessel steels to temper embrit t lement was undertaken.

Temper embrit t lement can be de­fined as a shift in the Charpy V-notch

energy absorption and fracture ap­pearance curves to higher tempera­tures when a heat treated steel is either slow cooled through the temperature range of 1100 to 600 F or held for long periods of t ime wi th in this range. Britt le fractures in temper embritt led steels fol low intergranular paths along prior austenitic grain boundaries. The embrit t lement is re­versible. Short t ime heating above 1100 F fol lowed by rapid cooling through the embritt l ing temperature range results in the el iminat ion of temper embritt lement. Britt le frac­tures in a de-embritt led steel exhibit transgranular cleavage.

Temper embrit t lement has been studied extensively over the past twenty years in plate steels2 - 5 and more recently in weld metals.6 - 8

Present theories as to cause focus on the role of impurities such as As, Sb, P and Sn, combined w i th their com­plex reactions in the presence of cer­tain alloying elements, particularly Ni, Cr, M n a n d S i .

Bruscato6 has developed an embrit­t l ing factor X such that the impact energy absorbed at +50 F by step-cooled 2V4Cr-1Mo steel can be qual­itatively predicted. The factor, given by

_ 1 0 P + 5 S b + 4 S n + A s X= PPm

M 1

100 0 )

can be used to show there is a s_trong dependence of susceptibility to X and

temper embrit t lement on (Mn + Si) %. The energy absorbed at +50 F is plotted as a function of X and (Mn + Si) and a trend is seen. When the (Mj i + Si) is high, as is weld metal, X must be low. However, plate material has a relatively low (Mn + Si), typicah_ ly 0.60 to 0.80, so that much larger X can be tolerated. This is detailed in Bruscato's paper.

Whi le the end-user, designer and fabricator have an academic interest in the theories, they are primarily con­cerned wi th the extent of degradation of standard materials expressed in fa­miliar parameters. Al though no single experimental approach is likely to simulate the broadly varying condi­tions of application, this study covers a range of commercial materials and provides comparisons in terms of the familiar Charpy test.

Exper imenta l Procedure General

Step-cooling treatments were uti l­ized to induce temper embrit t lement and cannot be i n te rp re ted quantitatively in terms of actual process conditions. Step-cooling is a sequence of programmed isothermal aging treatments at progressively lower temperatures in the range of 1100 to 600 F that provides an effec­tive, accelerated screening technique to determine the relative suscepti­bility of a material to temper em­britt lement. The propensity to embrit-

58-s I F E B R U A R Y 1 9 7 3

Fig. 2 — Fracture appearance transition temperature

data for 1 Cr-'AMo steel

Table 2 — Typical Step-Cool Cycle

Heat to 1100 F Hold 1 hour Furnace cool to 1000 F Hold 16 hours Furnace cool to 975 F Hold 24 hours Furnace cool to 925 F Hold 72 hours Furnace cool to 875 F Hold 96 hours Furnace cool to 850 F Hold 120 hours Furnace cool to 650 F Air cool

+ 240

+ 200

+ 160

+ 120

- + 8 0 C±J or _J

<: + 4 0 or UJ OL

E o

- 4 0

- 8 0

- 1 2 0

160

A—As Received or T—Stress Relieved

< o

- cn

1

- < o LO

( ) < «* 1

o < o LI)

| < o LO

o </) 1 < ) o r̂

CJ

< Q u>

| CJ

< o ?

«S

O < o < * T ? •3

o CO

1 o < o .—i

I S: <

S-Step Cooled

? <

o < o U)

7 ? •a

o CO

o < o

7 <: <

SERIES 27 28 29 30 31 32 33 34 35 TREATMENT A T S A T S A T S MATERIAL A387B WELD WELD

As Welded

?. <

o <. o CJ CN

7 ? •a

O CO

1 o < o CJ CM

7 3 <

33 36 37 A T S WELD

Fig. 3 — Fracture appearance transition temperature

data for 2'ACr- 1Mo steel

cr

<

LJ h-

+ 120

+ 80

+ 40

-40

-80

- 1 2 0

- 1 6 0 SERIES TREATMENT MATERIAL

38 39 40 A S R A387D

<

CJ CJ CO CO

I I o o o < < <

o o o o o

A—As Received or As Welded S-S tep Cooled R—Retempered

41 42 43 A S R

A542

44 45 46 A S R

A542

47 48 49 A S R

A542

50 51 52 A S R

A542

53 54 A S A542

£

o

CM CM

55 56 A S A 542

57 58 A S A542

W E L D I N G R E S E A R C H S U P P L E M E N T ! 59 -s

Table 3—C-Mo Tensile Data

ASTM Specif.

A204C

Form

Plate

Weld

Heat

A2948

(RACO)

Test No.

1 2 3 4 5 6

Treatment ( a |

1650 AC + 1250 AC Test 1 + 1140 AC Test 2 + Step-cool As-welded Test 4 + 1140 AC Test 5 + Step-cool

.2% YS, ksi

66.3 57.4 50.8 65.2 58.2 53.5

UTS, ksi

79.8 79.8 73.8 81 .1 ( b )

81.6 (b )

76.9 , b )

Elong. 2 in.,

% 30.0 29.5 32.0 16.0 21.0 17.5

Red area.

% 64.5 64.8 67.6 62.7 62.6 65.2

Rock­wel l

A

49.8 52.3

— 51.1 50.6

—

(a) AC = air cooled (b) Fracture occurred in base metal

Table 4 — C - M o Impact Data

ASTM Specif.

A204C

Form

Plate

Weld

Heat no.

A2948

(RACO)

Test no

1

3 4 5 6

Treatment (a)

1650 AC + 1250 AC Test 1 + 11 40 AC Test 2 + Step-cool As-welded Test 4 + 11 40 AC Test 5 + Step-cool

F.ATT|bl

F

+30 +80 +60 +60 +80 +80

Energy af FATT,

ft- lb

72 85 82 39 53 42

Energy at 50 F,

ft-lb

82 70 82 34 45 38

40 ft- lb transit ion

temp, F

- 1 0 - 5 - 1 0 +75 +25 +60

Upper shelf

energy, ft-lb

110 102 120

52 63 55

(a) AC = air cooled (b) FATT = f racture appearance t rans i t ion tempera ture

Table 5—Mn-Mo Tensile Data

ASTM Specif.

A302B

A533A

A302C

A533B

Form

Plate

Weld

Plate

Plate

Plate

Plate

Plate

Plate

Heat no.

B5217

(Linde)

A4069

A4069

C4689

C4694

C4689

C4694

Test no.

7 8 9

10 11 12 13 14 15 16 17

18 19 20 21 22 23

24 25

26

Treatment

1 650 AC + 1 250 AC Test 7 + 11 40 AC Test 8 + Step-cool As-welded Test 1 0 + 1140 AC Test 11 + Step-cool 1 600 AC + 1 1 50 AC Test 13 + Step-cool 1 600 AC + 1 100 AC Test 1 5 + Step-cool 1 6 0 0 W Q + 1 1 0 0 A C + 1 150-18-WQ Test 17 + Step-cool 1650 AC + 1275 AC Test 19 + Step-cool 1650 AC + 1275 AC Test 21 + Step-cool 1 6 5 0 W Q + 1 2 4 0 W Q + 1150-50-FC Test 23 + Step-cool 1 650 W Q + 1 240 WQ + 1150-50-FC Test 25 + Step-cool

.2% YS, ksi

61.6 60.1 56.2 60.6 59.4 60.1 73.5 70.8 79.5 74.3 96.0

87.3 70.2 67.0 75.2 72.5 81.9

81.7 88.7

87.5

UTS, ksi

80.1 78.4 77.6 82.4 |C )

80 .0 ( c ' 82 .2 l c )

93.4 92.7 99.0 95.4

104.9

102.1 89.6 89.1 96.7 96.1 96.7

96.2 100.3

100.3

Elong. 1 in..

% 31.0 (b)

30.5 |b)

31.0 (b)

17.0 l b» 22.5 (b)

19 .0 ( b )

26.3 26.8 23.4 27.3 23.8

25.0 28.7 28.7 26.9 28.8 28.8

28.2 27.5

26.2

Red. area,

% 69.6 71.4 70.7 69.7 69.8 68.0 63.2 63.3 61.5 63.5 69.1

67.0 72.9 71.0 68.0 68.8 74.9

72.1 69.8

67.6

Rock­wel l

A

52.5 50.4

— 54.5 49.8

— 55.3 54.6 56.0 55.0 59.2

56.1 53.8 53.2 55 3 55.2 56.0

55.8 57.1

57.1

(a) AC = air cooled; W Q - wa te r quenched; FC = furnace cooled (b) 0 .505 in. d iam test spec imen; o thers 0 . 2 5 2 in. diam (c) Fracture occurred in base metal

6 0 - s I F E B R U A R Y 1 9 7 3

+ 240

+ 200

+ 160

+ 120

U_

L J + 8 0

or z> < +40 or UJ Q_

S o I -

- 4 0

- 8 0

-120

- 1 6 0 SERIES TREATMENT MATERIAL

-

, ,

-

o CO

o o CO CO

o o o

< < < - o o o i n i n i n CM CM CM

- o < IT)

_ o 1

o . g Q

IT) 1 ^

o < LO

o 1

O g o LO r̂

O < ID r~. o

7 O g o LO .—i

g <

O < o LO CM

1

Q UJ Z

or UJ H CLi Q (-O z

g <

o < o LO CM

1 g <

A—Received or As Welded T—Stress Relieved S—Step Cooled SS—Double Step Cooled

o CO

o CO

o o < r - < *

Q UJ Z

QC UJ 1 -UJ Q 1-Q Z

O

< i n r-v o .-H

1 o o LO 1 ^

1-1 LO CM

o < in r-. o

7 O g o LO | \ <—(

o LO <M

O < CO

o

7 o g o LO

o CO

o < o in CM

o < LO

o 1

o g o in > — 1 l

o < o i n CM

1 g <

Q LU

Z

cr LU h-LU Q

1-O z

o CO

O <

% o o o < < < i n t ^ CM >—i

1

8 r̂

o m CM

rh <y g 8 r*.

LO

r « 4 -

7 o g o in K

8

o < LO r̂ CM

7 O g o IO r̂

59 60 61 T S SS

A542

62 63 64 A T S

WELD

65 66 67 A T S

WELD

68 69 70 71 A T S SS

A542

72 73 74 A T S

WELD

75 76 A S A542

77 78 A S A542 + Ni

Fig. 4 — Fracture appearance transition temperature data for 21ACr-1Mo steel

tie was measured mainly in terms of the shift in the fracture appearance transit ion temperature (FATT) ob­served in Charpy V-notch (CVN) tests. The FATT in CVN tests is generally ac­cepted as one of the more reliable and convenient standards of tough­ness. Other criteria such as upper shelf energy and energy absorbed or fracture appearance at a given temperature were convenient out­growths of the data and are included for completeness.

Base data on unembritt led plate and some weld metals were initially generated. The data included chem­istry, microstructure, impact, tensile and hardness test results. Plates were tested in their normal condit ion of heat treatment and welds were tested in the as-deposited condit ion. Al l welds and some plate composi­tions were then subjected to a temper or stress-relief prior to the next evaluation. Al l were subjected to a step-cooling treatment and retested. Finally, selected step-cooled series were given a temper or stress-relief prior to the concluding tests. Tensile and/or hardness tests were used to monitor strength changes. Metal­lography and fractography were em­ployed to establish the nature of ob­served embritt lement. General struc­ture, fracture path and grain size are among the metallographic features examined.

Materials

The plate materials used in this in­vestigation were primarily f rom production heats of commercial low alloy steel plate compositions. In one case where a laboratory melt was used (3Cr-1/2Mo steel) a companion laboratory heat of 2V4Cr-1Mo steel was also tested. The similarity of em­britt lement data from both production and laboratory 2ViCr-1Mo heats sup­ported the validity of the data on the 3Cr-V4Mo analysis.

Weld metals appropriate for the plate steel were all deposited by the submerged arc process using a single head DC reverse polarity technique wi th a nominal heat input of 55 kJ / i n . The variations in flux, fi l ler, and procedure were all w i th in the scope of good practice. The welding parameters are not detailed because of the l imited number of welds produced.

Table 1 lists the chemical composi­tion of the weld deposits and the plate materials, along wi th the asso­ciated ASTM specifications.

Heat Treatment

Initial plate tests were conducted on material subjected to the standard treatment appropriate for the grade. Nonstandard treatments were ap­plied only in those cases where it was desired to develop special char­

acteristics for study, such as coarse grain or very high strength. Weld met­als were initially tested in the as-de­posited condition. Selected plates and all of the weld metals were given a tempering or Stress-relieving treat­ment. Al l materials were subjected to the step-cooling treatment described in Table 2 to induce embritt lement. Some materials were tested to deter­mine the reversibility of embritt le­ment by subsequent retempering for one hour at 1100 F. The treatments for each test are reported in conjunc­tion wi th the mechanical property re­sults in Tables 3 through 12.

Test Procedures

Al l tests were performed in accord­ance wi th pertinent ASTM specifica­tions. Standard 0.252-in.-diam ten­sile specimens of plate material were oriented longitudinal to the f inal rol l­ing direction. These tests were con­ducted to confirm adherence to ma­terial specifications and to disclose any effects of subsequent tempering or step-cooling treatments included in the experimental program. For those stages of plate evaluation where sufficient material for tensile specimens w a s not ava i lab le , strength was monitored by hardness testing. Al l welds were monitored by hardness testing after each stage of processing.

W E L D I N G R E S E A R C H S U P P L E M E N T ! 61-s

Table 6 — M n - M o Impact Data

ASTM Specif.

A302B

A533A

A302C

A533B

Form

Plate

Weld

Plate

Plate

Plate

Plate

Heat no.

B5217

(Linde)

A4069

A4069

C4689

C4694

C4689

C4694

Test no.

7 8 9

10 11 12 13 14 15 16 17

18 19 20 21 22 23

24 25

26

Treatment 8

1650 AC + 1250 AC Test 7 + 11 40 AC Test 8 + Step-cool As-welded Test 10 + 1140 AC Test 11 + Step-cool 1 600 AC + 11 50 AC Test 13 + Step-cool 1600 AC + 1100 AC Test 15 + Step-cool 1 600 W Q + 1100 AC + 1150-18-WQ Test 17 + Step-cool 1650 AC + 1275 AC Test 19 + Step-cool 1650 AC + 1275 AC Test 21 + Step-cool 1650 W Q + 1240 W Q f 1150-50-FC Test 23 + Step-cool 1 650 W Q + 1240 W Q + 1150-50-FC Test 25 + Step-cool

FATT" " F

+40 +75 +65 +45 +60 +80 +10 +30 +15 + 15 -95

- 7 5 +40 +60 +75 +85 - 6 5

- 2 0 - 8 0

- 2 0

Energy at FATT,

ft- lb

82 95 88 46 55 47 72 70 66 68 73

65 48 62 50 46 46

42 39

40

Energy at 50 F,

f t- lb

95 90 85 48 56 65 83 83 85 85

130

123 52 48 4 0 35 82

77 67

62

4 0 ft-lb t ransit ion

temp, F

< - 4 0 < - 1 0

- 5 0 +20

< - 1 0 +40 - 7 5 - 4 0 - 4 0 - 5 0

- 1 7 5

- 1 3 5 +10 +20 +45 +70 - 8 0

-25 -75

- 2 0

Upper shelf

energy, ft- lb

130 120 120

68 80 65

130 120 100 120 145

140 90 82 80 80 82

85 67

70

(a) AC - air cooled; W Q = water quenched; FC = furnace cooled {b} FATT = f racture appearance t rans i t ion tempera tu re

Table 7—1 Cr-Mo Tensile Data

ASTM Specif.

A387B

Form

Plate

Weld

Weld

Heat no.

15756

(RACO)

(L&R)

Test no.

27 28 29 30 31 32 33 34 35 36 37

Treatment

1750 AC + 1250 AC Test 27 + 1140 AC Test 28 + Step-cool As-welded Test 30 + 1140 AC Test 31 + Step-cool As-welded Test 33 + 11 50 AC Test 34 + Step-cool Test 33 + 1 200 AC Test 36 + Step-cool

.2% YS, ksi

48.3 46.0 44.1 51.0 47.5 49.7

UTS, ksi

68.6 68.7 6 7 - 7 i b . 73.3<b)

69.6 ( b )

73.3 ( b )

Hardness Tests

Elong. 2 in.,

% 32.5 32.5 34.5 15.0 19.0 15.5

only

Red. area

% 70.8 71.4 70.5 67.9 70.7 67.4

Rock­wel l

A

49.5 48.6 47.0 56.7 57,6 —

56.5 54.7 58.7

— 54.6

(a) AC = air cooled (b) Fracture occurred in base metal

Table 8 — Cr-Mo Impact Data

ASTM Specif

A387B Form Heat no.

Plate 15756

Weld (RACO)

Weld (L&R)

Test no.

27 28 29 30 31 32 33 34 35 36 37

(a) Treatment

1750 AC Test 27 + Test 28 + As-weldei Test 30 + Test 31 + As-weldei Test 33 + Test 34 + Test 33 + Test 36 +

+ 1250 AC 1140 AC Step-cool i 1140 AC Step-cool i 11 50 AC Step-cool 1 200 AC Step-cool

FATT,(b)

F

+40 +35 +45 +85

+115 +140

+75 + 85 + 50 + 75 + 30

Energy at FATT,

ft-lb

67 93 87 35 40 31 36 37 50 42

130

Energy at 50 F,

ft- lb

75 120

95 30 20 11 28 28 40 33

145

40 ft- lb transit ion

temp, F

+ 15 - 1 0 - 5

+115 +115 + 170

+70 +95 +50 +70 +10

Upper shelf

energy, ft-lb.

105 130 125

50 50 45 60 60 70 64

145

(a) AC - air cooled (b) FATT = f racture appearance t rans i t ion tempera tu re

6 2 - s I F E B R U A R Y 1 9 7 3

CC

z>

< or

+ 240 -

+ 200 -

+ 160

+ 120

+ 80

+ 40

0

- 4 0

- 8 0

- 1 2 0

- 1 6 0

o CO

I O O < LO

CM

3 8 r-s

< rtrr r̂ CM ,—i

O g 8 T— i

SERIES 79 80 TREATMENT A S MATERIAL 3Cr '/i Mo

A—As Received S—Step Cooled

SS—Double Step Cooled

o < 1/5

o g 8 r̂

C J CO

o < LO CM

Cv

g 8 r̂ -

(,) U -

un r̂ CM

C) < i n ( N (T) T—t

cr S

8 r v

*-'

O CO

O u_ LO

rv OJ > — 1

O

< LO l \ l ro .-< o 5 8 l>s

—>

CJ

< LO CM i—1

Cv

5 8

O CO

o < i n CM

O g 8

o < 1*1

<

CO

o < i n i n CM CM

rtrr co ,~4 —t

o

8

O g 8

81 82 A S

A387E

83 84 85 A S SS A387E

86 87 A S

QT A387E

88 89 A S

QT A387-E

90 91 A S

QT A387E

92 93 A S

QT A387E

94 95 A S

QT A387E

Fig. 5 — Fracture appearance transition temperature data for 3Cr-Mo steel

Impact tests of the plate were fu l l -size, transversely-oriented Charpy V-notch type. The weld metal impact specimens were full size and cut f rom the quarterline of the weld and trans­verse to the weld axis. The energy ab­sorption was measured over the range of temperatures necessary to provide 0 to 100% fibrous fracture and to define the upper shelf energy.

For both tensile and impact testing, specimens from heavy gage plate (1 % in. and over) were removed from the quarter-thickness location. For light plate, the specimens were removed from the surface to conserve ma­terial. In these light gages the two locations are proximate both physical­ly and metallurgically. Specimen loca­t ion and plate gage is not signif icant to embrit t lement evaluation as long as both are consistent throughout a series of tests on a specific material. Thus, these specimen details are not included in the tabulations.

Metallographic examinations were conducted on impact specimens of all materials in all conditions. Fracture surfaces were nickel plated and ex­amined w i th the light microscope for general structure, prior austenitic grain size and fracture path. Standard two stage replication techniques were used for electron microscope fracture edge examination. In addi­t ion, fractography was performed using plastic replicas and electron microscopic techniques.

R e s u l t s

C-Mo Steel Plate c o n f o r m i n g to A S T M

specification A204C was tested in the normalized and tempered condi­t ion, after an additional tempering cycle, and finally after step-cooling. A submerged arc weld deposited in this plate was tested as-deposited, after tempering, and after step-cooling. Re­sults are identified as Tests 1 through 6 in Tables 3 and 4 and Fig. 1.

A reduction in yield strength is noted (Table 3) after each thermal treatment and the weld appears to be slightly more temper resistant. The tensile strength of neither the weld metal nor the plate was affected by the additional temper. The step cool cycle decreased the tensile strength of both plate and weld by equivalent amounts. The trends in tensile elongation and reduction of area are not clear from this minimal data, a l ­though each subsequent treatment appears to cause an increase in ductility.

The trends in the various tough­ness criteria (Table 4) are not readily discerned from the available quantity of data. In general, the plate material offers slightly more favorable values. Relatively little propensity to em­britt lement is observed in either case.

Mn-Mo Steels Plate c o n f o r m i n g to A S T M

specification A302B and an asso­ciated submerged arc weld metal were subjected to the same sequence of treatment and testing as the A204C previously described. The ten­sile tests (Table 5, Tests 7 through 1 2) showed little change w i th additional-treatments, except for a slight drop in plate tensile and yield strength. Both plate and weld impact data suggest little deterioration from the step-cool­ing as show in Figure 1 and Table 6.

A second A302B plate initially sub­jected to two different tempering temperatures (1100 and 1150 F) was step-cooled without an intervening temper. The tensile results (Table 5, Tests 13 through 16) show a slight drop in yield and tensile wi th a cor­responding minor increase in ductility. Figure 1 and Table 6 show little effect of step-cooling on the im­pact properties; the largest change being a 35 F (-75 to -40) increase in the 40 ft-lb transition temperature.

Tests (Nos. 17 and 18) were also conducted on this same plate in the quenched and tempered condition (ASTM A533A), both as-treated and after a step-cool. The various notch toughness criteria indicate minor em­britt lement.

ASTM A302C and A533B are s im­ilar to A302B and A533A respectively except for the addition of about 0.5% nickel. Plates from two different heats of these Mn-Mo-Ni steels were evaluated. Again, the tensile data

W E L D I N G R E S E A R C H S U P P L E M E N T ! 63-s

, ^ y \

*i • •-

& A

'.-,

.-.0X

• x - . v . ' :. . • •

Fig. 6 — Fracture edge of as tempered (a) and step cooled (b) Charpy V-notch impact specimens. Nital-picral etch, X250, reduced 20%

- f ... • . * * > . > " \ - •Jf •;< ,Jcr " I

i«?x rt *-,r„>*,v%>

Vii*

v? ,'•>#.<?

b * »• • \

,& ** -,€

f/gr. 7 — Electron micrographs of nickel plated fracture edge of as tempered (a) and step cooled (b) Charpy V-notch impact specimens. Nital-picral etch. Two stage plastic replica. Au-Pd shadowed from fracture edge, X3.200, not reduced

_ »

NC

,!:-' ,'*^%.-r*m|->-i. l f'-%v

• a r

• % , . , .

i b

F/c;. 8 — Electron fractographs ofas tempered (a) and step cooled (b) Charpy V-notch impact specimens. Au-Pd shadowed from notch, X4.500, not reduced

6 4 - s I F E B R U A R Y 1 9 7 3

Table 9—21/4Cr-MoTensile Data

ASTM Specif.

A387D

A542

Form

Plate

Plate

Heat no.

B6730

B6730

A2766

Weld

Weld

Plate

A3588

(L&R-1)

(L&R-2)

C6669

Weld (L&R-3)

None

Plate

Plate

LI 05

A9367

Treatment (a)

Test no.

38 1750 FC + 1325 AC 39 Test 38 + Step-cool 40 Test 39 + 1100 AC 41 1750 WQ + 1325 AC 42 Test 41 + Step-cool 43 Test 42 + 1100 AC 44 1750 WQ + 1125 AC 45 Test 44 + Step-cool 46 Test 45 + 1100 AC 47 1750-8-WQ + 1325-8-WQ 48 Test 47 + Step-cool 49 Test 48 + 1100 AC 50 1750-8-WQ + 1125-8-AC 51 Test 50 + Step-cool 52 Test 51 + 1100 AC 53 1700 WQ + 1275 AC 54 Test 53 + Step-cool 55 2000 WQ + 1275 AC 56 Test 55 + Step-cool 57 2300 WQ+1275 AC 58 Test 57 + Step-cool 59 1750 WQ+1075-15-AC +

1250 AC 60 Test 59 + Step-cool 61 Test 60 + Step-cool 62 As-Welded 63 Test 62 + 1250 AC 64 Test 63 + Step-cool 65 As-welded 66 Test 65 +1250 AC 67 Test 66 + Step-cool 68 1750 WQ+ 1075-15-AC 69 Test 68 + 1250 AC 70 Test 69 + Step-cool 71 Test 70 + Step-cool 72 As-welded 73 Test 72 + 1250 AC 74 Test 73 + Step-cool 75 1 700 WQ + 1275 AC 76 Test 75 + Step-cool 77 1750 WQ+1275 AC 78 Test 77 + 1100 AC +

Step-cool

2% YS, ksi

42.7

— 41,0 63.6 68.1 67.4

140.4 123.5 118.3

68.4 68.1 67.0

117.3 118.3 115.8

99.9

81.6 80.6

UTS, ksi

68.2 No

68.8 84.3 89.8 88.6

155.8 137.5 131.7

86.9 86.9 85.5

133.3 134.6 132.1

Elong. 1 in.,

% 33.5

tTested 31..0 30.5 26.5 26.0 16.5 18.0 19.0 26.5 25.0 25.0 20.0 20.0 20.0

Hardness Tests Or

115.2

iy

23.5 Hardness Test Only

99.4 99.6

21.5 22.0

Red. area,

% 73.5

— 73.3 77.4 75.0 77.4 64.8 68.2 69.6 75.7 74.4 76.3 66.9 67.0 68.5

70.7

62.7 64.6

Rock­wel l

A

46.6 47.5 47.3 54.8 53.6 52.9 63.6 6 2 9 64.1 55.1 53.3 53.5 64.7 64.5 63.8 53.6 52.6 54.1 54.7 54.2 52.1 56.0

55.4 57.4 63.7 55.4 56.8 62.9 54.9 56.3 61.8 55.2 55.1 56.8 62.2 54.9 56.9 55.0 55.5 54.5

—

I AC - air cooled; WQ - water quenched; FC = furnace cooled

(Table 5, Tests No. 19 through 26) show step cooling caused a slight yield strength loss wi th minimal duc­ti l i ty loss restricted to A 5 3 3 B (quenched and tempered) . The deterioration of toughness (Table 6, Fig. 1) from step cooling is not major as evidenced by a maximum shift of 20 F in FATT for A302C and 60 F for A533B. It should be noted that despite the larger shift for A533B, its FATT after step cooling as - 2 0 F for both plates, whi le A302C plates were + 1 0 a n d + 2 0 F .

Detailed chemical analyses of these two heats were performed and are listed in Table 1. The analyses for embritt l ing constituents As, Sb and P are similar as a re j he Mn and Si con­centrations. An X factor6 was calcu­lated assuming 0.01 5% Sn (1 50 ppm) which is a typical Sn concentration. Because of the similarit ies in com­

position, no large differences in sus­ceptibility to temper embrit t lement was expected. This is substantiated by the data.

1 Cr-Mo Steel

One plate of ASTM A387B steel was investigated in the usual normalized and tempered condit ion, after an additional temper, and after a subsequent step cool. Tests 27, 28 and 29 in Tables 7 and 8 and Fig. 2 detail the results. The later treat­ments reduced both yield strength and toughness by a small degree only.

The two welds made in this plate and similarly treated (Tests 30 through 35) responded differently to step-cooling; one decreased in tough­ness (Test 30-32) whi le the other in­creased (Tests 33-35). The weld which increased in toughness ex­

hibited an additional increase when the tempering temperature prior to step cooling was raised (Tests 33 through 37). The difference between the two welds was in the welding flux used indicating the importance of this variable.

2V4Cr-Mo Steels

A one-inch gage plate of 21/4Cr-1Mo (ASTM A387D) was furnace cooled after austenitizing to simulate a normalization of an 11 in. gage plate. After tempering, it was step-cooled and subsequently tempered at 1100 F, (Tests 38, 39, 40). The em­brittl ing treatment appeared to have little effect on strength or notch toughness. The largest shift in FATT was +10 F. A l l data for th is composi­t ion are given in Tables 9 and 10 and Figs. 3 and 4.

W E L D I N G R E S E A R C H S U P P L E M E N T ! 65-s

Table 10—21/4Cr-Mo Impact Data

ASTM Specif. Form

Heat no.

A387D Plate B6730

A542 Plate B6730

A2766

A3588

Weld (L&R-1)

None

(L&R-2)

Plate C6669

Weld (L&R-3)

Plate (L105)

Plate A9367

Test no.

38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59

60 61 62 63 64 65 66 67 68 69 70 71 72 73 74 75 76 77 78

Treatment (a)

1750 FC + 1325 AC Test 38 + Step-cool Test 39 + 1100 AC 1750 WQ + 1325 AC Test 41 + Step-cool Test 42 + 1100 AC 1750WQ+ 1125 AC Test 44 + Step-cool Test 45 + 1100 AC 1750-8-WQ + 1325-8-WQ Test 47 + Step-cool Test 48 + 1100 AC 1750-8-WQ + 1125-8-AC Test 50 + Step-cool Test 51 + 1100 AC 1700 WQ + 1 275 AC Test 53 + Step-cool 2000 WQ + 1 275 AC Test 55 + Step-cool 2300 WQ+ 1275 AC Test 57 + Step-cool 1750 WQ + 1075-1 5-AC + 1250 AC

Test 59 + Step-cool Test 60 + Step-cool As-welded Test 65 + 1250 AC Test 66 + Step-cool As-welded Test 65 + 1 250 AC Test 66 + Step-cool 1750 WQ+ 1075-15-AC Test 68 + 1250 AC Test 69 + Step-cool Test 70 + Step-cool As-welded Test 72 + 1250 AC Test 73 + Step-cool 1700 WQ+ 1275 AC Test 75 + Step-cool 1750 WQ+ 1275 AC Test 77 + 1100 AC + Step-cool

FATT,,b)

F

+ 75 + 90 + 85 -100 - 80 -100 - 25 + 35 - 50 - 35 + 20 - 20 +230 +240 + 180 - 60 - 10 - 70 + 5 - 50 - 40 -110

- 55 - 45 + 170 + 30 —

+ 45 - 10 —

- 60 - 25 - 15 +120 - 50 —

-130 -115 -100 - 40

Energy at FATT, ft-lb

100 102 127 80 100 100 56 48 58 100 94 87 57 64 66 80 86 70 70 90

101 85

75 105 34 30 — 34 44 —

85 72 70 52 80 — 47 51 44 33

Energy, at 50 F, ft-lb

45 62 100 150 150 125 80 55 75 140 120 135 26 30 34 140 118 128 86 142 143 140

140 140 9

65 80 35 80 20

120 115 120 38 115 61 90 90 69 62

40 ft-lb transition temp, F

+ 25 - 5 + 15 -185 -155 -200 - 65 0

-110 - 45 - 55 -105 +150 + 165 +100 -150 - 90 -110 - 60 -140 - 90 -150

-105 -100 +200 + 20 + 10 + 90 - 20 + 95

-130 - 95 - 80 + 70 - 80 + 10 -165 -155 -100 - 45

Upper Shelf Energy

ft-lb

150 150 150 150 150 140 80 60 80 140 140 140 70 75 80 140 118 126 117 144 146 140

140 140 45 105 105 42 95 88

120 125 130 85 115 85 105 105 68 62

(a) FC - furnace cooled; AC - air cooled; WQ - water quenched (b) Fracture appearance transition temperature

Material f rom this same plate was also quenched and tempered to two tensile strength levels, 84 ksi (A542) and 156 ksi. This was fol lowed by a step-cool and a final retemper at 1100 F. Results of these tests (Nos. 41-46) show that step-cooling the low strength material increased the yield and ultimate by approximately 5 ksi w i th a corresponding decrease in elongation and reduction of area. The trend was reversed in the high strength material. After the f inal temper, the low strength material showed little further change, whi le the high strength plate continued its trend. Step cooling of the low strength plate caused a A FATT of +20 F but the original FATT o f - 1 0 0 F was restored by the f inal temper. The FATT of the 156 ksi plate shifted +60 F after step cooling. After the f inal

temper it shifted -25 deg F. A similar series of tests were per­

formed on another A542 steel heat treated to 87 and 1 33 ksi UTS levels. In this series, (Nos. 47 through 52) tensile property changes w e r e insignificant. The toughness of the 87 ksi plate was more affected by step cooling shift ing the FATT +55 F than the high strength condit ion wh ich shifted only +10 F. An improvement in the toughness of the 133 ksi plate wi th the 1100 F temper was ob­served. The apparent low sus­ceptibility of the plate at the high strength level to temper britt leness may be due to the poor toughness to begin w i th . As reported by Mc-Mahon9 , there is an upper l imit em­britt lement for a particular steel and the loss in toughness wi l l approach this l imit by either temper brittleness

or an inherently poor toughness. This same p late w a s t h e n

austenitized at three temperatures to ascertain grain size effects (Tests 53-58). The grain sizes varied from ASTM -0.4 to 7,5. The strength was monitored by hardness tests wh ich showed a range in Rockwell A hard­ness of 52.1 to 54.7. This series of tests indicated that there is a critical grain size for maximum embrit t le­ment. The f ine grain material had a shift in FATT of +50 F and the coarse grained material showed a shift in FATT of +10 F, However, the material w i th the intermediate grain size (ASTM 5.5) experiences a shift in FATT of +75 F. See Fig. 3. This same trend is evident in the other measures of toughness.

Plate from a third heat of 21/4Cr-1Mo in the quenched and tempered

66-s | F E B R U A R Y 1 9 7 3

Table 11—3Cr-Mo Tensile Data

ASTM Specif.

None

A387E

A387E

Form

Plate

Plate

Plate

Plate

Heat no.

L104

B3683

B3683

A2966

Test no.

79 80 81 82 83 84 85 86 87 88

89 90 91 92 93 94 95

ta)

Treatment

1700 WQ + 1275 AC Test 79 + Step-cool 1750 AC + 1325 AC Test 81 + Step-cool 1 700 AC + 1 275 AC Test 83 + Step-cool Test 84 + Step-cool 1 700 WQ + 11 25 AC Test 86 + Step-cool 1700 AC + 1325 AC + 1275 FC Test 88 + Step-cool 1700 WQ + 1125 AC Test 90 + Step-cool 1700WQ+ 1225 AC Test 92 + Step-cool 1700 WQ + 1325 AC Test 94 + Step-cool

.2% YS ksi

88.6

70.2

92.3

102.8

62.2

115.3

90.9

63.7

i, UTS, ksi

106.2 Hardness Test Only

92.5 Hardness Test Only

110.7 Hardness Test Only Hardness Test Only

119.0 Hardness Test Only

87.8

Hardness Test Only 131.9

Hardness Test Only 111.3

Hardness Test Only 90.7

Hardness Test Only

Elong. 1 in..

% 25.9

28.9

23.9

24.0

31.0

20.5

24.0

30.0

Red. area

% 73.8

76.8

70.8

74.2

74.6

71.6

73.8

76.9

Rock­we l l

A

51.8 53.4 53.0 53.9 56.5 57.1 58.0 59.9 61.0 51.0

51.3 63.6 62.9 59.5 59.7 54.0 54.7

(a) W Q - wa te r quenched; AC = air cooled; FC = fu rnace cooled

condition was subjected to two step-cooling cycles (Tests 59, 60, 61). Af­ter the first step-cool cycle, the FATT increased from - 1 1 0 to - 5 5 F. The second step cool caused a further in­crease to -45 F. This indicates that most of the embrit t lement occurs dur­ing the first step cooling. See Fig. 4.

Two submerged arc welds each w i th a different flux were made in this third plate, tempered and step-cooled. Tests 62-67 detail the results. Hardness tests on both indicated a de­crease in strength after temper at 1250 F, and a slight increase after step-cooling. Both welds showed significant toughness improvement after the 1 250 F treatment. One weld exhibited litt le change in toughness despite step-cooling, whi le the sec­ond weld clearly decreased. As w i th the welds in A387B, these data show the importance of the flux used in welding.

The fourth heat of quenched and tempered 21/4Cr-1Mo steel showed a loss in toughness as a result of step-cooling. A second step-cool cycle caused a further but lesser drop in toughness (Tests 68 -71 , Fig. 4). The submerged arc weld in this plate was more responsive to both tempering and step-cooling (Tests 72-74). Upon step-cooling, the weld loss in notch toughness was more than that of the plate (a shift in the 40 ft- lb transit ion temperature of +90 F for the weld vs +35 F for the plate). The fracture ap­pearance of this weld and the two previous welds could not be deter­mined due to the wide variation in grain size concomitant w i th a mult i ­pass weld.

A f i f th heat was a laboratory melt made for comparison w i th a similar melt of another analysis discussed later. The tests (75-76) indicated little

effect f rom step-cooling (AFATT of +15 F) corresponding to production plate results.

The last heat of this group was a non-standard chemistry in that it con­tained about 0.5% nickel. The results (Tests 77 and 78) show some reduc­t ion in toughness incurred by step-cooling in that the FATT increased from - 1 0 0 to - 4 0 F: a shift of +60 F.

As w i th the Mn-Mo-Ni steels re­ported earlier, a detailed chemical analysis fai led to show the interac­t ion of the embritt l ing constituents and (Mn + Si). This is due to the l im-ited_chemical data and the fact that the X factors range from 18.4 to 20.8 wi th the (Mn + Si) factors ranging f rom 0.67% to 0.74%. Wi th l imited data, a tight grouping of data points such as experienced here makes cor­relations difficult. However, conf i­dence in the data can be increased by comparing the production heats w i th the laboratory heat (Tests 75 and 76). The X factor for Test 75 is 9.3 ppm and (Mn + Si) is .78%. The A FATT for the lab heat was only +1 5 F whereas shifts as high as +70 F were obtained w i th production heats. One produc­t ion heat (Test 41) has a A FATT of +20 F but generally, the susceptibility to embrit t lement increases as X and (Mn + Si) increases.

3Cr-Mo Steels

The laboratory melt of 3Cr-1/2Mo steel was quenched and tempered to 106 ksi UTS prior to step cooling. The data for this material (Tests 79, 80) are contained in Tables 11 and 12 and the fracture appearance data are plotted in Fig. 5. The validity of tests on lab heats was indicated by Tests 75 and 76. This heat showed only a minor effect of step cooling, being

wel l w i th in experimental scatter. The shift in FATT amounted to only 5 F and can be attributed to low X~ and (Mn + Si) factors.

Plate from one heat conforming to the requirements of A387E (3Cr-1 Mo) was evaluated after normalizing and tempering to both 93 and 110 ksi UTS (Tests 81-85). Both exhibited minimal susceptibility to embrit t le­ment from step-cooling. The second step-cool cycle applied in one case produced another small increment of embritt lement. This is reported in Ta­bles 11 and 1 2 and Fig. 5.

Tests Nos. 86-89 are on the same plate quenched and tempered to 88 and 119 ksi ultimate. Minimal step cooling effects were apparent, except for the low strength plate showing a 60 F FATT shift.

A plate from another production heat of A387E chemical composition was quenched and tempered to 9 1 , 111 , and 132 ksi ultimate prior to step cooling (Tests 90-95). Although some transition temperature shifts occurred, all values were in the - 6 0 to +20 F range. The shift in FATT in­creased wi th decreasing strength.

The two 3Cr-1Mo heats have s im­ilar susceptibilities to temper em­b r i t t l emen t . C o m p a n n g the i r compositions, especially X and (Mn + Si) shows these factors to be close. Therefore, similar responses to step cooling are expected.

Metallography

Optical microscopy was employed to confirm that the general structure, grain size, and cleanliness of the experimental materials were typical of commercial product. The only ob­served deviations f rom microstruc­tures typical to plates produced by standard practices occurred in the

W E L D I N G R E S E A R C H S U P P L E M E N T 67-s

Table 1 2 ^ 3 C r - M o

ASTM Specif.

None

A387E

A387E

Impact

Form

Plate

Plate

Plate

Plate

Data

Heat no.

L104

B3683

B3683

A2966

Test no.

79 80 81 82 83 84 85 86 87 88

89 90 91 92 93 94 95

Treatment (a)

1700 W Q + 1275 AC Test 79 + Step-cool 1750 AC + 1325 AC Test 81 + Step-cool 1700 AC + 1275 AC Test 83 + Step-cool Test 84 + Step-cool 1700 WQ + 1125 AC Test 86 + Step-cool 1700 AC + 1325 AC + 1275 FC Test 88 + Step-cool 1700 W Q + 1125 AC Test 90 + Step-cool 1700 W Q + 1225 AC Test 92 + Step-cool 1700 WQ + 1325 AC Test 94 + Step-cool

(bl FATT

--+

+ + +

-+ 4-

+

------

F

90 85 0 5

40 55 75 10 20

5

65 50 20 90 10

100 40

Energy at FATT,

ft- lb

65 60 46 37 32 42 45 20 27 39

46 25 29 30 28 38 42

Energy at 50 F,

ft- lb

100 105

70 63 40 40 30 40 46 56

42 45

— 64 54 82 75

40 ft- lb transit ion

temp, F

_• -• -+ + + + + + +

+

+ + -+

--

I 50 I 40

10 25 50 50 70 50 40 10

40 10 30 60 20 95 60

Upper shelf

energy, ft-lb

102 106 71 64 51 61 65 50 55 72

70 48

— 64 65 82 90

(a) W Q - water quenched; AC - air cooled; FC = furnace cooled (b) FATT = f racture appearance t rans i t ion temperature

study of the effects of austenitizing temperature where the anticipated coarse grain was developed after high temperature treatments.

Optical examination of fracture edges of impact specimens permitted determination of fracture type and path. Fracture edges were nickel plated to preserve the fracture during sectioning. The nonembritt led speci­men (Fig. 6a) has an intergranular fracture path. These can be seen in finer detail in the electron micro­graphs of Fig. 7.

Electron f ractographs of the fracture surface are shown in Fig. 8. The nonembritt led specimen (Fig. 8a) has a cleavage fracture. The em­brittled specimen shows cleavage and intergranular fractures. Repli­cated precipitates can be seen on the grain facets in Fig. 8b.

Conclusions

Temper embrit t lement is charac­terized by a displacement of the Charpy V-notch energy absorption and fracture appearance curves to higher temperatures. Using these criteria, all of the materials evaluated by step-cooling showed some degree of propensity to temper embritt le­ment. Despite a variety of materials and conditions, the maximum shift was about 80 F wi th the majority of the displacements 30 F or less. Em­brittled materials exhibited the inter­granular fracture associated wi th temper embritt lement. Step-cooling did not affect the upper shelf energy absorption values, thereby conf irm­ing the transit ion temperature shift to be caused by temper embritt lement.

The susceptibility to temper em­britt lement varies wi th composition,

structure and heat treatment. The susceptibility was lower for the C-Mo and Mn-Mo steels than for the M n -Mo-Ni and Cr-Mo steels. The addition of Ni to Mn-Mo apparently increases susceptibility to embritt lement. W i th ­out Ni, the A FATT for quenched and tempered material is +20 F and it is +45 to +60 F when Ni is present. There are no apparent differences be­tween the susceptibilities of 21/4Cr-1Mo and 3Cr-1Mo to embritt lement. With in the scope of the tests reported herein, no f i rm trends in the effects of As, Sb, P and Sn are observed.

Weld metal tests showed the same trends as did the tests on plate steel. However, the susceptibility of weld metal to temper embritt lement in­duced by step cooling is generally more severe than that of the base metal. This is attributed to the effect of flux on the susceptibility. Proper selection of flux can inhibit temper embritt lement greatly. For example, in the 1Cr-ViMo weld metal, the shift in FATT ranged from an increase of 25 F to a decrease of 45 F depending upon the flux and stress relieving treatment.

Tempering affects the susceptibility to temper embritt lement through its effect on strength. Generally, the loss in notch toughness after step cooling is more severe for low strength levels than for higher strengths. This is attributed to the fact that there is a minimum toughness for any one com­position and the high strength condi­tions are closer to this minimum than are the low strength conditions. Therefore, step cooling wi l l have a greater effect on a high toughness condition than on a condition wi th lower toughness. However, the ab­solute toughness for a temper em­

brittled low strength steel is often bet­ter than that of a nonembritt led high strength condition of the same steel.

The susceptibility of a normalized and tempered structure appears to be independent of composition. A max­imum shift in FATT of 20 F was ob­tained. However, the toughness of a temper embritt led quenched and tem­pered material may exceed that of the nonembritt led normalized and tem­pered structure of the same material.

References

1. Gulya, J. A., Swift, R. A. "Steels for High Temperature Petroleum Pressure Vessels" ASME Paper 70 — Pet — 26, Petroleum Mechanical Engineers Con­ference, Denver, Colorado, 1 970.

2. Jaffe, L. D., Buffum, D. C, "Isother­mal Temper Embrittlement," Trans ASM 42 1950, pp. 604-618.

3. Low, J. R., Stein, D. F., Turkalo, A. M., Laforce, R. P. "Alloy and Impurity Ef­fects on Temper Brittleness of Steel." Trans Met. Soc. AIME 242 (1) 1968, pp. 14-24.

4. Restaino, P. A., McMahon, C. J. Jr., "The Role of Antimony in Temper Brittle­ness," Trans ASM 60 (4) 1967, pp. 699-706.

5. Swift, R. A. "Temper Embrittlement of 21/4Cr-1Mo Steel," Lukens Steel Com­pany RDR 68-28 December 1 968.

6. Bruscato, Robert, "Temper Embrittle­ment and Creep Embrittlement of 2ViCr-1Mo Shielded Metal Arc Weld Deposits," Welding Journal 49 (4) Research Suppl. pp. 148-S to156-S, 1970

7. Simmers, A. M. "Temper Embrittle­ment Susceptibility of General Low Alloy Steels and Their Weld Metals," Lukens Steel Company, RDR 69-20, September 1969.

8. Swift, R. A. — Unpublished Research. 9. McMahon, C. J. Jr. "Temper Brittle­

ness — An Interpretive Review" ASTM STP 407, Philadelphia, Pa. 1 968.

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