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|| 0F STAN0AR0S 4 tech R.I.C. ELD HEAT TREATMENT

aiSwia. IA FOR REPAIR WELDS IN

2-1/4Cr-1Mo SUPERHEATER HEADERS

AN EXPERIMENTAL STUDY

D. T. Read

H. I. McHenry

National Bureau of Standards

U.S. Department of Commerce

Boulder, Colorado 80303-3328

August 1988

. U 5 6

#87-3075

g America's Progress 1913-1988
NATIONAL INSTITUTE OF STANDARDS &

TECHNOLOGY

Research Information Center

Gaithersburg, MD 20899
NBSIR 87-3075

O-Cloo

/lO. 3 7

' POSTWELD HEAT TREATMENT mz

CRITERIA FOR REPAIR WELDS IN

2-1/4Cr-1Mo SUPERHEATER HEADERS

AN EXPERIMENTAL STUDY

D. T. Read

H. I. McHenry

Fracture and Deformation Division

Institute for Materials Science and Engineering

National Bureau of Standards

Boulder, Colorado 80303-3328

August 1988

Supported in part by

Naval Sea Systems Command

Washington, DC 20362-5101

U.S. DEPARTMENT OF COMMERCE, C. William Verity, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director
IMPORTANT NOTICE

Identification of commercial bibliographic services by

trade name does not imply recommendation or endorsement

by the National Bureau of Standards, nor does it imply

that the specific services identified are the best

available

.
CONTENTS

ABSTRACT v

1. STATEMENT OF THE PROBLEM:

DOES POSTWELD HEAT TREATMENT OF REPAIR WELDS IN SUPERHEATER

HEADERS CONTRIBUTE TO THE PREVENTION OF FRACTURE? 1

1 . 1 Literature Review 2

2. APPROACH: ELASTIC-PLASTIC FRACTURE MECHANICS 3

3. WELD METAL: 2^Cr-lMo 4

4. EXPERIMENTAL SPECIFICS:

WELD AND SPECIMEN GEOMETRIES AND TEST PROCEDURES 5

4.1 Wide Plates 7 4.2 Standard Specimens 8

5. EXPERIMENTAL RESULTS 9

5.1 Weldment Characterization 9 5.2 Wide-Plate Results 10

5.3 J- Integral Fracture-Toughness Tests 14

5.4 Tensile Tests 16

6. RESULTS AND DISCUSSION:

DIRECT MEASUREMENT OF THE J- INTEGRAL IN THE WIDE PLATE 17

7. IMPLICATIONS 21

8. NOTE ON THE EXPERIMENTAL APPROACH 22

9. ACKNOWLEDGMENTS 22

10. REFERENCES 23

APPENDIX A. DESCRIPTION OF THE WELDING PROCEDURE 25

APPENDIX B. ULTRASONIC MEASUREMENTS OF RESIDUAL STRESS 27

APPENDIX C. MECHANICAL MEASUREMENTS OF RESIDUAL STRESS 29

APPENDIX D. MEASUREMENT OF APPLIED J- INTEGRAL

PRODUCED BY RESIDUAL STRESS

ABSTRACT

Wide-plate and standard- size specimens cut from repair welds

in 2^Cr—IMo plate were tested as -welded and after postweld heat

treatment (PWHT) . Wide plates were prepared with surface cracks

and tube -hole -ligament cracks in simulated repair welds made with

the shielded-metal - arc process. The plates were instrumented with

strain and displacement gages for direct measurement of the applied J- integral. After notching and fatigue precracking, the

plates were tested in bending. Pop-ins occurred in the as-welded

plates, but not in the PWHT plates.

Three -point-bend specimens with cracks oriented in the TS

direction were used to measure weld-metal and heat-affected-zone

(HAZ) toughness values. Results of direct measurements of the

applied J- integral on the wide plates were compared with critical

J-value measurements of three -point-bend specimens. The comparison indicated that PWHT was highly beneficial, because it reduced

the crack-driving force from residual stresses and increased the

weld-metal and HAZ toughness. In the as-welded condition, very

low toughness values were measured at the HAZ. These low tough- ness values, together with the measured crack-driving forces,

indicated critical crack depths of a few millimeters.

To extend the usefulness of these results, a new approach

to the problem of the applied J- integral produced by residual

stresses is being explored: strain-gage measurements made during

notching are analyzed to obtain an applied J- integral as a function of crack depth. The preliminary results are encouraging.

The residual - stress -produced J-value is roughly equivalent to that

produced by a remote elastic loading to the same stress level.

Key words: Fracture; hydrotest; J- integral; pressure vessel;

residual stress; safety; stress relief.

POSTWELD HEAT TREATMENT CRITERIA FOR REPAIR WELDS

IN 2kCr—IMo SUPERHEATER HEADERS

AN EXPERIMENTAL STUDY

1. STATEMENT OF THE PROBLEM: DOES POSTWELD HEAT TREATMENT OF REPAIR WELDS

IN SUPERHEATER HEADERS CONTRIBUTE TO THE PREVENTION OF FRACTURE?

Postweld heat treatment (PWHT) of repair welds made in superheater

headers is currently mandated by documents, including MIL-STD-278E(SH) [1].

Omission of the PWHT is allowed in certain circumstances listed in the standard; these include several classes of welds in grooves 6 mm (1/4 in) deep

or less. The standard indicates that NAVSEA may specifically approve

omission of the PWHT in "other welds."

The PWHT of repair welds can entail logistical problems above and beyond the consumption of time and energy. Specialized equipment and extensive setup procedures may be needed because the superheater headers are located deep within the hull. There are no known reports of fractures or failures caused by omission of the PWHT in the allowed circumstances.

Specific justification for the original inclusion of the PWHT requirement in

the MIL- STD is unavailable. The purpose of the PWHT requirement is now

assumed to be fracture prevention in the boiler components.

Fracture of boiler components is unacceptable because of danger to

personnel, loss of ship function, and cost of repair. The actual benefits

of PWHT should be carefully evaluated so that they can be balanced against

the associated costs.

Risks of superheater header fracture occur during hydrostatic testing

(at room temperature) and during operation (at superheated steam temperature) . The three causes of fracture are high stresses, brittle material,

and large defects. The stresses are highest and the material is most

1
D. T. Read and H. I. McHenry

brittle during hydrotest. Inspection and maintenance procedures are intended to eliminate large defects before hydrotest; however, the boilers

have many welds that can contain undetected and potentially dangerous

cracks

.

This program addresses fracture at weld cracks during a hydrotest after

weld repair. It seeks to identify the difference in fracture safety between

a weld-repaired component that is given a PWHT and a component that receives

no PWHT. The technical information obtained in this program will support

Navy administrative decisions on stress relief -criteria for repair welds.

1.1 Literature Review

Two approaches were adopted to take advantage of previous research

on postweld treatment of repair welds in 2^Cr—lMo steel pressure vessels:

computerized searches of abstracts on file with abstract services, and

tracking of significant references starting with recent high-quality work.

The computer index searches were not productive; however, several useful

references were found by reference tracking.

The key word combination {weld fracture .and. (pressure vessel .or. stress relief) } was sought in abstracts in Compendex (which contains engineering information) . Two hundred and forty-one resulting abstracts were

examined. The combination (weld .and. stress relief .and. pressure vessel) was sought in the Weldasearch file. Ninety-eight resulting abstracts were

examined. These efforts confirmed the lack of a good data base on stress

relief of repair welds in pressure vessels. The most useful references

(tracked manually) are listed below.

An excellent review of the metallurgy of 2*z;Cr— IMo was given by C . D. Lundin, B. J. Kruse, and M. R. Pendley [2]. The effects of PWHT time and

temperature on the toughness and tensile properties of a multipass submerged

arc weld in 2^Cr—IMo were reported by R. A. Swift and H. C. Rogers [3].

This report confirmed the large differences between as -welded (AW) and PWHT

weld metal found in the present study.

2

-
Postweld Heat Treatment Criteria

The first report, and still one of the most accessible and unambiguous

reports, of the effect of residual stress on the fracture of wide plates was given by T. W. Greene [4]. It provides the precedent for the use of wide

plates in bending in the present study. Unfortunately, this study was conducted well before fracture mechanics reached its current state of

development

.

The effect of PWHT on 2*4^—IMo base -metal Charpy- impact toughness

was reported by J . P. Gudas , C. A. Zanis , E. J. Czyryca, G. L. Franke, E. M.

Hackett, and W. A. Palko [5]. In their study, ductile tearing occurred at room temperature in 2kCr—lMo base metal, a behavior not observed in the

present study, possibly because a different specimen orientation was used.

A recent literature review of the weldability and properties of 2kCr—

IMo steel by J . M. Blackburn [6] concentrates on the controllable aspects

of the welding process, especially preheat, and their effects on the weld

metallurgy

.

2. APPROACH: ELASTIC-PLASTIC FRACTURE MECHANICS

Fracture mechanics is the branch of engineering science that seeks

to measure the resistance of materials to fracture and to balance applied

stresses, material resistance to fracture, and defect size, so that the risk

of fracture in service is appropriately controlled. The present approach

centers on a fracture mechanics evaluation of severe artificial defects in

simulated repair welds in large, thick plates of superheater header material, 2*sCr—IMo steel. Large plates are needed to retain the residual

stresses and to simulate the actual structural behavior. Plates with PWHT

are compared with AW plates. Room- temperature tests conveniently match the

hydrotest temperature.

Included in the approach are evaluations of the residual stress present

in the PWHT and AW plates and conventional small-specimen tests of weld
D. T. Read and H. I. McHenry

metal in both PWHT and AW conditions. The small -specimen tests include

weld-metal fracture - toughness , Charpy- impact- toughness , and tensile -property

tests, as well as metallographic examination.

Analysis and interpretation of the test data will contribute to the

development of guidelines for stress relief of boiler components following

weld repair.

3. WELD METAL: 2kCr-lMo

Steel plate manufactured to ASTM Standard A387 Grade 22 Class 1 (1984),

in the annealed condition, was obtained from a commercial source in the form

of two plates, each 5 cm (2 in) thick by 2 m (80 in) by 3.7 m (144 in). The

plate chemistry supplied by the manufacturer is given in table 1, along with

a check analysis performed by a commercial laboratory.

The welding electrodes used were the same type, E9018-B3L, and from

the same manufacturer as those used by the Navy [1]. In the electrode

designation, L indicates low carbon, which is now standard. A preheat of

200°C (400°F) was used. The welding procedures are detailed in Appendix A

of this report, and the chemical compositions of the weld and base metal are given in table 1. The differences between the AW and PWHT specimens are thought to reflect the level of accuracy in the analysis rather than actual

differences caused by PWHT.

According to [1], the holding temperature for PWHT of 2kCr-lMo vessels

should be between 730 and 775 °C (1350 to 1425 °F)

.

The holding time should

be 1 hour per 2.5 cm of thickness (1 hour per inch of thickness). Following

[5], the PWHT was performed as follows:

a. Heat from 200 to 745 °C (400 to 1375 °F) at a rate less

than 110 °C (200 °F) per hour;

b. Hold at 745 °C (1375 °F) for 2 hours;

c. Cool to 200 °C (400 °F) at a rate less than 110 °C (200

°F) per hour;

d. Air cool.

4

-
Postweld Heat Treatment Criteria

Table 1. Chemical compositions of materials, in mass percent.

Element Base Metal Weld

Mfr

.

Check AW PWHT

Cr 2.24 2.14 2.43 2.41

Mo 0.98 0.86 1.05 1.02

C 0.11 0.097 0.042 0.035

Mn 0.47 0.45 0.76 0.70

Ni NA 0.14 0.05 0.05

P 0.009 0.010 0.024 0.024

S 0.015 0.012 0.018 0.016

Si 0.20 0.18 0.048 0.043

Al NA 0.011 0.012 0.010

Ti NA 0.002 0.007 0.006

Co NA 0.011 0.017 0.016

V NA 0.008 0.010 0.009

Cu NA 0.083 0.056 0.054

Fe bal

.

bal. bal. bal.

4. EXPERIMENTAL SPECIFICS:

WELD AND SPECIMEN GEOMETRIES AND TEST PROCEDURES

The test matrix, including wide-plate and small-specimen tests of

repair welds AW and after PWHT , is illustrated in figure 1. Two types of

repair welds were used: trough welds and simulated tube -hole -ligament

repairs

.

The Welding Group at David Taylor Naval Ship Research and Development

Center (DTNSRDC) , Annapolis, has developed a standard repair-weld test

geometry consisting of a trough in a plate with specified dimensions

(fig- 2). The trough is 20 cm (8 in) long by 4 cm (1-1/2 in) deep in a plate that is 5 cm (2 in) thick, 45 cm (18 in) long, and 30 cm (12 in) wide.

The geometries of six of the eight troughs welded in this study were based

on the DTNSRDC geometry—two for wide-plate tests and four for standardspecimen tests (fig. 1). The other two were designed to simulate a repair

to a ligament between tube holes (discussed below); see figure 3.
D. T. Read and H. I. McHenry

AS WELDED

180 cm

(72 in)

POST-WELD HEAT TREATED

65 cm

(26 in)

(] Standard Navy Trough Weld

30 cm 013 Tube-Hole-Ligament Repair-Weld Simulation

(12 in)

Figure 1 . Weld and plate geometries used to test benefits of PWHT

in 2kCr—lMo repair welds. Summary of simulated A387,

Grade 22, Class 1 (E9018-B3L) repair welds.

r«— 5 cm —

(2 in)

Figure 2. Geometry of the standard Navy trough weld designed by

David Taylor Naval Ship Research and Development Center.
Postweld Heat Treatment Criteria

5 cm

(2 in)

Figure 3. Geometry of a simulated tube-hole- ligament repair weld.

4.1 Wide Plates

The wide-plate geometries are indicated in figure 1. Trough welds

(fig. 2) were made in two wide plates, one weld in each plate. One of these

plates received PWHT , the other did not; both were tested by the same procedure. Ultrasonic scans were taken before and after welding and after PWHT

to determine residual stresses. This technique was successful in the case

of the AW plate but inconclusive after PWHT, because the micros tructura

1

changes produced by PWHT altered the ultrasound velocities by amounts

comparable to the residual stress effects. See Appendix B.
D. T. Read and H. I. McHenry

Simulated tube -hole- ligament repairs were made in two plates. Details

of the weld geometry are given in figure 3. One plate was given a PWHT , and

the other was not. Each test plate was instrumented with strain gages; the

strain- gage layouts are given in Appendix C. A part- through surface notch

(an artificial crack) was cut into the wide plate with a special saw made

for this program. The special saw was needed to produce a notch of the

desired geometry without interfering with the strain gages. It used a circular blade 7 cm (2-3/4 in) in diameter by 0.5 mm (0.02 in) thick. The

notches were approximately semi-elliptical, 5 cm (2 in) long at the plate

surface and 1 cm (0.4 in) deep at the notch center. The notching process

also affects residual stresses in these plates. Cutting relieves residual

stresses adjacent to the cut; the amount of strain relaxation is related to

the magnitude of the residual stress. Residual stress values are derived

from measured strains in Appendix C.

After application of additional strain gages, each plate was fatigued

in bending in a 1-MN (222-klbf) servohydraulic testing machine, to provide

an ideally sharp crack. Each specimen was then tested quasi-statically in

four-point bending until failure or until the displacement capacity of the

testing facility was exhausted.

4.2 Standard Specimens

Four trough welds, two in each of two plates, were made and then

machined into conventional small specimens. These plates were 45 cm (18 in)

long by 60 cm (24 in) wide. The troughs were positioned as shown in fig- ure 1. Two of the trough welds received PWHT and two did not, The plan for

specimen extraction from each of these welds is shown in figure 4. These

plates yielded a total of sixteen bend-bar specimens (5 cm by 2.5 cm; 2 in

by 1 in) for J- integral fracture toughness measurements, six round-bar tensile specimens, and fifteen Charpy- impact specimens. The AW weld metal was much harder, more difficult to machine, and wore out the tools much faster

than the PWHT weld metal.
Postweld Heat Treatment Criteria

Charpy

Figure 4. Plan for extraction of standard three-point-bend fracture toughness, Charpy, and round tensile specimens from trough welds.

5. EXPERIMENTAL RESULTS

Experimental results indicate clear and significant differences between

the properties of simulated repair welds with and without PWHT . The welds

with PWHT have a much higher toughness and a lower yield strength. Furthermore, the crack- dr iving force produced by the residual stresses in the AW

plates is comparable to the base-metal and HAZ toughness values. The PWHT

relieves the residual stresses and reduces the crack- dr iving force to negligible levels.

5.1 Weldment Characterization

Sections cut through the trough welds of fractured AW and PWHT specimens were examined metallographically , and their hardness was tested.

Figure 5 shows a metallographic section of a trough weld in the AW condition. The Knoop microhardness and the Rockwell hardness of the AW weld

metal were considerably higher than that of the PWHT weld metal. Repeated

microhardness traverses revealed no hardness peak at the HAZ in either case

(fig. 6).

9
D. T. Read and H. I. McHenry

Figure 5. Microstructure of an AW trough weld, near the end.

Note that Charpy impact tests were conducted to determine the transition temperatures for the AW and PWHT weld metal. The results (fig. 7) show

that the 54-J (40-ft-lbf) transition temperature of the AW specimens is

about 10°C (50°F) , and that of the PWHT specimens is about —68°C (—90°F)

.

5.2 Wide-Plate Results

Two wide -plate tests were run on trough-welded wide plates. Brittle

fracture in the form of an extremely large pop -in was observed in the AW

plate at a high applied J -value. As is typical, pop- in was marked by a sharp noise, a sudden drop in the measured load, and a sudden increase in

the crack-mouth-opening displacement (CMOD) , indicating that the crack had

suddenly advanced into the specimen. The records of load versus CMOD are reproduced in figure 8. The sizes of the starter crack and the pop- in are shown in figure 9. Examination of the fatigue -marked crack in the AW plate

indicated no stable crack growth.

10-
KNOOP

HARDNESS

Postweld Heat Treatment Criteria

350

300

250

200

150

100

1 1

—

1

1 o o

_

O

(-f~> o o _

o O Metallographic^ c HAZ o

CO

_ < 1 As Welded

l

o o

o o

o

Base

1 l

Weld

l 1

1 1 1 1

Metallographic

HAZ

PWHT

o o

o

—1— o o

o o /°°o - Base

1 1 Weld

J l 0 5 10 15 20 25 5 10 15 20 25

POSITION, mm POSITION , mm

Figure 6. Knoop microhardness as a function of location

near the fusion line, AW and PWHT.

>•O

oc

LUz

HIo

<

CL>

>-

0.cc<

I

o

Figure 7. Charpy-V impact-energy transition curves

for AW and PWHT weld metal.

11
LOAD

D. T. Read and H. I. McHenry

CRACK MOUTH OPENING DISPLACEMENT

Figure 8. Load—displacement trace for a wide-plate specimen with a semielliptical surface crack in an AW trough weld, showing a large

pop- in, compared with the load—displacement trace for a wideplate specimen with a crack in a PWHT trough weld, showing no brittle fracture.

Fatigue

Figure 9

.

12

Schematic diagram of the wide-plate fracture surface with the

crack in the AW trough weld, showing the extent of the pop -in.

The weld was tested in four-point bending at room temperature.
Postweld Heat Treatment Criteria

Fatigue

Figure 10. Schematic diagram of the wide -plate fracture surface with the

crack in the PWHT trough weld, showing fracturing in the base

metal. The weld was tested in four-point bending at room

temperature

.

A schematic drawing of the fracture surface of the PWHT plate is shown

for comparison in figure 10. No macroscopic brittle fracture was observed

during the test. The fracture surface contained a small region of apparently brittle fracture in the base metal adjacent to the weld, but no fracture in the weld metal. This was consistent with the standard-specimen

tests, because the crack-driving force in the wide-plate tests did not reach

the high- toughness values observed for the weld metal in the standardspecimen tests. It also showed that the PWHT weld metal was more fracture

resistant than the base metal, consistent with results of standard bend-bar

tests

.

Two tests were run on wide plates with simulated tube-hole- ligament

repairs. The load—displacement traces are compared in figure 11. Brittle

fracture in the form of four audible pop -ins was observed in the AW plate.

The first pop -in occurred at a J -value consistent with the toughness of the

standard specimens. However, the ligament did not fracture through, and the

test was continued. Three more pop- ins were observed at higher J -values.

Post-test examination showed that the four pop-ins resulted in a crack

extension of 11 mm (7/16 in) through the ligament thickness.
D. T. Read and H. I. McHenry

CRACK MOUTH OPENING DISPLACEMENT

Figure 11. Comparison of the load-displacement trace for a wide-plate

specimen with a crack in a simulated AW tube -hole - ligament

repair weld (with audible pop- ins noted) with the loaddisplacement trace for a wide-plate specimen with a crack

in a simulated PWHT tube -hole - ligament repair weld, where

no pop- ins occurred.

5.3 J -Integral Fracture-Toughness Tests

The results of 31 fracture - toughness tests are listed in table 2. High

initiation toughness values followed by stable crack growth were observed in

the two tests of PWHT weld metal. The average initiation toughness level was 536 kN/m (3060 lbf/in) . After initiation, the crack growth resistance

of the PWHT weld metal was very high, practically indistinguishable from the

resistance curve associated with crack blunting. In four tests of AW weld

metal, the average initiation toughness value was 87 kN/m (478 lbf/in).

Unstable crack propagation, in the form of repeated pop- ins followed by

brittle fracture of the whole specimen, was observed in all the AW weldmetal specimens tested.

Eight tests were performed on specimens with cracks located in the

heat-affected zone (HAZ) , four with PWHT welds and four with AW welds. The

results of these tests indicate that the HAZs in both PWHT and AW welds

14-
Postweld Heat Treatment Criteria

Table 2. Fracture Toughness Values for Standard Trough Welds

in 2kCr—lMo Plate

J - Integral CTOD

K(J r ) or K(6,~) Specimen

Number

Notch

Location

Condition

a/W J J 6 S MPaVm ksi^in

Criterion kN/m lbf/in Criterion

mm in 10 -3

1NJ

1

WM AW 0.51 Jc 68.3 397 0.077 3.0 123 112 1NJ2 WM AW 0.52 Jc 77.9 445 *c 0.086 3.4 131 119

2NJ

1

WM AW 0.51 Jc 138 789 «c 0.132 5.2 174 158 2NJ2 WM AW 0.51 Jc 122 694 0.104 4.1 164 149

1SJ1 WM PWHT 0.51 J Ic 415 2370 «i 0.521 20.5 302 275

1SJ2 WM PWHT 0.53 J Ic 245 1400 *i 0.307 12.1 232 211 1SJ4 WM PWHT 0.53 869 4960 1 . 10 43.2 437 398 2SJ2 WM PWHT 0.48 J Ic 569 3250 «i 0.673 26.5 354 322

lNJ3t HAZ AW 0.45 Jc 38.0 217 0.0549 2.16 91 83 2NJ3 HAZ AW 0.44 Jc 5.25 30.0 0.0053 0.21 34 31 CMALC3 HAZ AW 0.54 Jc 14.4 82.4 «c 0.0272 1.07 55 50 CMALC4 HAZ AW 0.43 Jc 8.84 50.5 6 C 0.0130 0.51 44 40 CMALC5 HAZ AW 0.50 Jc 5.67 32.4 *c 0.0068 0.27 35 32 2SJ

1

HAZ PWHT 0.50 Jc 39.8 227 «c 0.082 3.2 94 85 2SJ3 HAZ PWHT 0.48 Jc 50.8 290 «c 0.090 3.6 106 96 2SJ4 HAZ PWHT 0.52 Jc 25.7 147 0.051 2.0 75 68 1SJ3 HAZ PWHT 0.46 Jc 32.0 183 S c 0.075 2.9 84 76 CMPB5 BM PWHT 0.54 Jc 65.7 375 &C 0.150 5.9 120 109 CMPB6 BM PWHT 0.54 Jc 167 952 0.384 15.1 192 174 CMPB7 BM PWHT 0.42 Jc 114 650 «c 0.259 10.2 158 144 CMPB8 BM PWHT 0.32 Jc 96.8 553 0.244 9.6 146 133 CMPB9 BM PWHT 0.44 Jc 98.9 565 «c 0.253 10.0 147 134 CMPB10 BM PWHT 0.25 Jc 67.1 383 «c 0.220 8.6 121 111 CMPB11 BM PWHT 0.27 Jc 101 579 «c 0.279 11.0 149 136 CMPB13 BM PWHT 0.14 Ju 496 2830 «u 1.32 52.1 330 300 CMPB14 BM PWHT 0.65 — — — «c 0.094 3.7 80 73 CMPB15 BM PWHT 0.34 — — — «c 0.43 16.9 171 156 CMPB16 BM PWHT 0.14 Ju 427 2440 «u 1 . 12 44.0 306 279 CMPB17 BM PWHT 0.10 — — — «u 0.49 19.3 183 166 1NJ4 BM AW 0.51 Jc 41.2 235 «c 0.089 3.5 95 87 2NJ4 BM AW 0.50 Jc 40.1 229 «c 0.087 3.4 94 85

t Out-of-plane crack

WM - weld metal; HAZ growth

.

- heat- affected zone

;

BM - base metal; AW - as welded; PWHT- post-weld heat treated; CTOD — crack tip opening displacement;

Jc “ J-integral measured at a "pop-in"; Jjc — fracture toughness for slow,

stable crack growth; Jm - J-integral measured at maximum load; J u - J-integral

measured at brittle fracture following stable crack growth (J—R curve not available)

.

15
D. T. Read and H. I. McHenry

are brittle. The PWHT HAZ had an average critical J -value of 47 kN/m

(270 lbf/in) . The AW HAZ had an average critical J-value of 15 kN/m

(87 lbf/in). No stable crack growth was observed in these tests, and

the crack in one AW specimen appeared to follow the HAZ instead of going

straight

.

Two of the base-metal specimens had very low a/W ratios, which caused

the final J-value to be very high at failure. These two tests were not

included in the calculation of the average toughness values.

5.4 Tensile Tests

Triplicate tensile tests of round-bar tensile specimens, 6.2 mm

(0.25 in) in diameter, indicated that the weld metal before PWHT is significantly stronger than both the base metal and the PWHT weld metal. The

results are listed in table 3. For AW weld metal, the yield strength was

767 MPa (111 ksi) and the ultimate strength was 843 MPa (122 ksi) ; for the

PWHT weld metal, the yield strength was 454 MPa (66 ksi) and the ultimate

strength was 579 MPa (84 ksi). According to the manufacturer, the yield

strength of the base plate was 270 MPa (39 ksi) , and the ultimate strength

was 499 MPa (72 ksi). The weld metal in the AW and PWHT conditions had

similar ductilities.

Table 3. Tensile Properties for Standard Trough Welds in 2hCx—IMo Plate.

Material No.

Obs .

Yield

Strength

,

MPa (ksi)

Ultimate

Strength

,

MPa (ksi)

Elongation,

percent

Reduction

of Area,

percent

WM—PWHT 2 454 (65.8) 579 (84.0) 28 74

WM-AW 3 767 (111) 843 (122) 23 66

WM = weld metal; PWHT = postweld heat treated; AW = as welded;

No. Obs. = number of observations

16
Postweld Heat Treatment Criteria

6. RESULTS AND DISCUSSION:

DIRECT MEASUREMENT OF THE J- INTEGRAL IN THE WIDE PLATES

The applied J- integral of the four wide-plate specimens as a function

of applied bending stress is shown in figures 12 and 13, the primary experimental results of this series of tests. These figures portray the effect of

residual stress and applied bending stress on the crack-driving force, as characterized by the J- integral. The stress values plotted are the elastically calculated outer-fiber tensile stresses for the wide-plate specimen

BENDING STRESS, ksi

<

cco

LU

D

UJ

_J

Q.

CL

<

300

10 20 30 40 50

2-1/4 Cr-IMo Repair Welds

Surface Cracks

18 mm Deep x 50 mm Long

A

AA

A

200

100

Condition: o AW

a PWHT

Hydrotest

Stress

Fracture Toughness:

L

A A&

A

AA

O

°o

O

o

o WM-PWHT

WM-AW

oBM

A

A C A

A O

A O

AAo

A O A

A

A

AA

O

o

100 200

HAZ-PWHTl

HAZ-AW

i 300

1.0

Q.

<

CC

(D

LU

hQ

LU

_J

Q.

CL

<

BENDING STRESS, MPa

Figure 12. Applied J- integral as a function of imposed bending stress for

the AW and PWHT wide-plate specimens with surface cracks in

trough welds. The vertical line marks a convenient stress level

approximating the stress near a tube hole during a maximumpressure hydrostatic test. Relevant toughness levels are indicated by horizontal lines.

17
D. T. Read and H. I. McHenry

E

z

<

DC

O

yj

i ”3

o

UJ

_l

Q_

0_

<

oOO

200

100

BENDING STRESS, ksi

10 20 30 40 50

1 -H 1 1-7 o

2

- 1 /

A o 4 Cr-IMo Repair Welds a

Condition:

Tube Hole Cracks * ° 19 mm Deep

A °

AO

o AW Fracture Toughness:

a PWHT WM-PWHT o

Hydrotest

^

<

« o

o

o

1 Stress U ^ WM-AW ° — Pop-in A

o

° o a

A O

° -BM a HAZ-AW

HAZ-PWHT A

A — A A

A

» A A A £_ A A

^

i i 100 200 300

BENDING STRESS, MPa

1.5

1.0

0.5

a

<

DC

0

LU

I —>Q

UJ

_J

0.

CL

<

Figure 13. Applied J- integral as a function of imposed bending stress for

the AW and PWHT wide -plate specimens with surface cracks in

simulated tube -hole- ligament repair welds. The vertical line

marks a convenient stress level approximating the stress near a tube hole during a maximum-pressure hydrostatic test. Relevant

toughness levels are indicated by horizontal lines.

loaded in four-point bending. For the PWHT specimens, J -value was close to zero for zero bending stress. As the stress was increased, J increased

gradually. When the stress reached approximately the yield strength of the

base plate, plastic strain became significant. The applied J- integral

increased along with the plastic strains, whereas the stress stayed nearly

constant

.

18
Postweld Heat Treatment Criteria

For the AW specimen, residual stresses caused a significant level

of crack- driving force even before the bending stress was applied. The

J- integral increased gradually with the bending stress. When the stress

reached an effective yield strength, plastic strain became significant,

and the J- integral increased rapidly with stress. The effective yield

strength of the AW plate was higher than that of the PWHT plate because

of the higher yield strength of the AW weld metal.

For the AW plate with the surface crack in the trough weld, the pop- in

indicated in figure 8 occurred at a J-level of about 275 kN/m (1570 lbf/in)

.

This is much higher than the critical J- levels observed in the standard

specimen tests, about 87 kN/m (478 lbf/in), but the orientation of the crack

relative to the weld was different. In the wide plate, the crack was perpendicular to the weld, whereas in the standard specimen, the crack was parallel to the weld. The observed results indicate the possibility of an orientation effect on the weld toughness as well as an effect due to size

and geometry.

The effect of the residual stresses on applied J for both PWHT and AW

plates was calculated by including in the applied analysis residual strains

equal to the measured strain relaxation when the notch was cut. This is the

first known use of this technique. The results (figs. 12 and 13) show that

the residual stress—strains raise the applied J- levels by 20 to 30 kN/m

(114 to 171 lbf/in) for the cracks tested. This increase is especially

significant when considered in conjunction with the low toughness values of

the base metal and HAZ , which are noted in the figures.

The rather low values of the AW-HAZ toughness compared with the moderate applied J -values produced by the residual stresses stimulated interest

in the applied J -values as a function of notch depth. Because strains and

crack depths had been recorded continuously during the cutting process, the

data needed were available. Applied J was calculated at each crack depth.

The results are shown in figure 14.

1
D. T. Read and H. I. McHenry

E

"v. z

-J

<

CC

O

LU

I “D

Q

LLj

_J

CL

CL

<

40

30

20

10

CRACK DEPTH, in

0.1 0.2 0.3 0.4 0.5

2-1/4 Cr-IMo Repair Welds

Residual Stress J

o SESC-PWHT

a THLC-PWHT

+ SESC-AW

x THLC-AW x

x

xx

Fracture

Toughness ^ HAZ-PWHTl

++

+ +r

X +

ix#8 00 0 0 ©0 0 c0 © e

lAHAZ-AW

o

200

100

5 10

CRACK DEPTH, mm

15

Figure 14. Applied J- integral as a function of cut depth for the AW and

PWHT wide-plate specimens with semielliptical surface cracks

(SESC) in trough welds and in simulated tube -hole- ligament

crack (THLC) repair welds. For comparison, the mean toughness,

15 kN/m (187 lbf/in) , of the AW-HAZ is indicated by the horizontal line labeled Fracture Toughness : AW-HAZ near the left

side of the plot.

20-

APPLIED

J-INTEGRAL,

Ib/in
Postweld Heat Treatment Criteria

7 . IMPLICATIONS

These experimental results of 2kCr—IMo repair welds indicate that the

PWHT of the current weld consumable is definitely beneficial. The major

benefit is a much higher fracture resistance in PWHT weld metal than in the

AW weld metal. Reduction of residual stresses is another advantage.

Figures 12 and 13 compare the measured toughness values to the applied

J- integral at a selected stress value corresponding to an estimated maximum

hydrotest stress near a tube hole. The hydrotest stress is indicated by the

vertical line in each figure; the toughness levels are indicated by horizontal bars. If the toughness is less than the applied J, fracture is predicted. The toughness of the PWHT weld metal is above the range of these

plots. These figures indicate that the PWHT and AW weld metals are tough

enough to prevent fracture at the crack sizes used in these tests. The

base-metal and PWHT-HAZ toughness values are above the PWHT driving- force

curves, indicating a safe condition. However, the AW-HAZ and base-metal

toughness values are below the AW driving- force curve, indicating the possibility of fracture. Linear scaling of the test crack sizes indicated that

a safe condition could be reached by limiting the crack sizes to about 5 mm

(0.2 in). This estimated critical crack size was refined by using the data

of figure 14, which plots applied J- integral against crack depth for PWHT

and AW wide plates. The crack depths correspond to the depths of the saw

cut when cutting was interrupted to record strains from strain gages for

J- integral calculation. The figure shows that the J- integral increased much

more rapidly with crack depth in the AW plates; it reached the toughness of

the AW-HAZ at depths of about 5 mm. However, the toughness was measured

using standard bend bars.

If it is necessary to omit the PWHT, compensating action should be

taken to ensure the safety of the superheater header, such as restricting

the omission of PWHT to very shallow welds, so that the possible crack size

is limited. Extra- stringent inspection could be required for the AW welds,

to ensure that all cracks have been removed. Omission of PWHT could be

21
D . T. Read and H. I. McHenry

restricted to welds in low-stress regions. A special local stress relief of some sort could be developed, which would give sufficient toughness to the

weld deposit, but which might not relieve the residual stresses as well as

a conventional PWHT. Selection of a different weld-metal consumable for

repair welds could be considered as a metallurgical remedy to the nonoptimum

properties of the current consumable in the non- PWHT state.

8. NOTE ON THE EXPERIMENTAL APPROACH

The experimental technique of direct measurement of the applied

J- integral was developed previously [7]. Some applications to surface flaws

have been reported, for example, in reference 8. The behavior of the

J- integral in three dimensions was discussed in reference 9, and its behavior in welds is a topic of current research [10].

This is the first report known to the authors of the application of

this technique to repair welds and to an evaluation of PWHT. Appendix D, a companion report, gives the details of this technique and its basis.

Apparently, this technique can provide valuable data on the applied

J- integral in welds. In particular, the data given in figure 14 are unique.

Previously, only engineering estimates of the applied J- integral produced by

residual stresses have been available. Because this approach is relatively

new, the results should be applied with caution.

9 . ACKNOWLEDGMENTS

The financial support provided by the Naval Sea Systems Command through

project sponsor Commander L. R. Easterling is gratefully acknowledged.

Helpful discussions with W. C. Fraser and M. A. E. Natishan of the David

Taylor Naval Ship Research and Development Center are appreciated. Technical assistance by J . D. McColskey, M. Moyer, and J. J. Broz of the National

Bureau of Standards is also appreciated.

22 —
Postweld Heat Treatment Criteria

10. REFERENCES

[1] Naval Sea Systems Command. Military standard: fabrication welding

and inspection; and casting inspection and repair for machinery,

piping, and pressure vessels in ships of the United States Navy. MILSTD-278E(SH) , Department of the Navy, Washington, D.C.; 1976.

[2] Lundin, C.D., Kruse, B.J., and Pendley, M.R. High temperature properties of 2-1/4 Cr 1 Mo weld metal. Weld. Res. Counc . Bull. 277, May

1976, pp. 1-27.

[3] Swift, R.A. and Rogers, H.C. Embrittlement of 2-1/4 Cr—1 Mo steel

weld metal by postweld heat treatment. Weld. J. — Res. Suppl., April

1973, pp . 145s—153s.

[4] Greene, T.W. Evaluation of the effect of residual stresses. Weld.

J. — Res. Suppl., May 1949, pp . 193s—204s.

[5] Gudas , J.P., Zanis, C.A., Czyryea, E.J., Franke , G.L., Hackett, E.M.,

and Palko, W.A. Fitness for service analysis of defects in the USS

Saratoga (CV-60) superheater headers. DTNSRDC/SME- 83/62 , David Taylor

Naval Ship Research and Development Center, Annapolis, Maryland,

August 1983.

[6] Blackburn, J.M. The weldability and properties of 2-1/4% chromium —

1% molybdenum steel used for Navy ship boiler components — A

literature review. (Distribution limited to U.S. Government agencies

only.) DTNSRDC/SME- 84/146 , David Taylor Naval Ship Research and

Development Center, Annapolis, Maryland, January 1985, 58 pp

.

[7] Read, D.T. Experimental method for direct evaluation of the J-contour

integral, in Fracture Mechanics: Fourteenth Symposium — Volume II:

Testing and Applications. J.C. Lewis and G. Sines, Eds. ASTM STP

791, American Society for Testing and Materials, Philadelphia, 1983,

pp. II- 199—11-213

.

[8] King, R.B., Cheng, Y.W.

,

Read, D.T., and McHenry, H.I. J- integral

analysis of surface cracks in steel plates, in Elastic—Plastic

Fracture: Second Symposium, Volume I — Inelastic Crack Analysis.

C.F. Shih and J.P. Gudas, Eds. ASTM STP 803, American Society for

Testing and Materials, Philadelphia, 1983, pp . I—444—1-457.

[9] Carpenter, W.C., Read, D.T., and Dodds, R.H., Jr. Comparison of

several path independent integrals including plasticity effects. Int . J. Fract. 31, 1986, pp . 303-323.

2
D. T. Read and H. I. McHenry

[10] Read, D.T. Experimental results for fitness-for-service assessment

of HY130 weldments. NBSIR 84-1699, National Bureau of Standards,

Boulder, Colorado, March 1984.

24-
APPENDIX A. DESCRIPTION OF THE WELDING PROCEDURE

T. A. Siewert

All elements of the welding procedure were chosen in consultation with

the David Taylor Naval Ship Research and Development Laboratories (DTNSRDL)

in an attempt to reproduce the field welding conditions.

The weld joints had been machined; therefore, no grinding was needed to remove surface oxides or surface irregularities. However, the joint regions

were carefully degreased to remove residual machine lubricants. The final

step before welding was preheating the plate to 205°C (400°F) . The plate

was maintained between 205° and 260°C (400° and 500°F) until the weld was completed and then allowed to cool slowly.

Welding was performed using a 2.4-mm (3/32-in) diameter MIL 9018-B3L

electrode conforming to MIL- E- 22200/8B . At 80 A and with a weave of about

one electrode diameter, each electrode produced about 10 cm (4 in) of weld

bead and a heat input of 1 kJ/mm (24 kJ/in). A new 4.5-kg (10-lb) hermetically sealed can of electrodes was used for each weld to keep hydrogen

potential low.

All these procedural details were kept constant for all welds. The

pass -layer sequence was different for the two types of weld joint, but was kept constant for each type.

The DTNSRDC standard joint geometry was a boat-shaped weld about 20 cm

(8 in) long, 5 cm (2 in) wide, 3.3 cm (1.3 in) deep, and with a rounded

bottom. Since each pass required two electrodes, there was a possibility

of differing microstructure or weld flaws in the intermediate start—stop

region. To minimize the effects of this variable on the property measurements, all the intermediate starts and stops occurred within 1 cm of the

centerline, and this region was avoided in the mechanical testing program.

Each layer near the bottom of the trough contained two passes; the top layer

had five or six passes. These welds contained between 16 and 18 layers for an average number of passes near 85 (170 electrodes).
Appendix A: Description of the Welding Procedure

The second joint geometry simulated the repair of a through-wall crack

in the ligament between two boiler tube holes. The joint was about 2.5 cm

(1 in) wide near the top, about 0.3 cm (0.12 in) wide near the bottom, and was about 5 cm (2 in) in length. Here, each pass could be completed with a single electrode. Start and stop regions were inside the tube bore so they

would be removed during the subsequent remachining of the tube bore. Each

layer near the bottom contained one pass; layers near the top had four

passes. These welds contained 19 to 22 layers for an average number of

beads near 58.

26
APPENDIX B. ULTRASONIC MEASUREMENTS OF RESIDUAL STRESS

A. V. Clark

Ultrasonic measurements of residual stresses were made using electromagnetic-acoustic transducers (EMATs) . Since the EMATs are noncontacting,

we can measure the arrival times of ultrasonic waves at locations of

interest before the plates are welded and then remeasure arrival times at

the same locations after welding. If (a) liftoff effects are negligible,

(b) plate thickness changes are due solely to elastic strains, and (c) no microstructural changes occur, then changes in arrival times are due solely

to stress.

Mathematically, we have

At

= “ ( aCT XX + kCTyy)

At

c y

(auyy + baxx )

(Bla)

(Bib)

where a and b are constant and Atx

is the difference of arrival times before and after welding, with EMATs polarized in the ax -direction. (We here

assume that the principal stresses ax and Oy act along material symmetry

axes.) If a

2 / b2

, then ct x and Oy can be determined by inverting the two

equations above.

Because of the cavity geometry and the welding pass sequence, we expect

the stress to have a complicated variation in the plate thickness direction.

We make the simplifying assumption that stresses ct xx and o yy

are caused

mainly by transverse and longitudinal shrinkage of weld metal, respectively.

Consequently, we expect that ct xx will be composed of both a bending and a membrane stress, because there will be more shrinkage of weld metal at the

top of the cavity than at the bottom. We also expect Oyy to be primarily a membrane stress (constant through the plate thickness). Our ultrasonic

technique measures the thickness - averaged stresses.
Appendix B. Ultrasonic Measurements of Residual Stress

Arrival times were measured along the plate centerline before and after

welding (with temperature compensation) . Equation B1 was used to calculate

axx anc* °yy outside the weld. These values are shown in figure B1

.

(a) (b)

Figure Bl . Residual stress, measured ultrasonically , as a function of position near a trough weld in a wide plate,

(a) along the weld and (b) across the weld.

To obtain the longitudinal stress inside the weld, we use the selfequilibrating nature of residual stress to give

L

Oyy dx = 0. (B2)

-L

where 2L is the plate width. We calculated Jayy dx outside the weld, and

assumed that <7yy is a constant inside the weld. We found that in order for

equation B2 to be satisfied, we have CTyy = 380 MPa inside the weld.

28
APPENDIX C. MECHANICAL MEASUREMENTS OF RESIDUAL STRESS

D. T. Read

Residual stresses are eliminated by mechanical relaxation when the

associated force paths are cut. This principle is the basis of the standard

method of determining residual stress by drilling a hole near strain gages.

The strain gages measure the relaxation that occurs during drilling; the

residual stress level is calculated from these relaxations. This same principle allows calculation of residual stress perpendicular to a cut from

strain relaxations measured during cutting.

Figure Cl shows the strain-gage layout used on the front side and

back side of the plate. The gages along the weld axis provide checks on

the paired gages near the cut and are used in J- integral calculations. Fig- ure C2 displays a plot of strain relaxation 7 mm (0.28 in) from the cut as a function of cut depth. The strain relaxation is nearly complete at a cut

depth of 8 mm (0.31). Gages nearer to the cut saturate at lower depths, and

vice versa. Figure C3 shows that strain relaxation decreases as distance

from the cut increases (circles) . The magnitude of the relaxation of

strains parallel to the cut is also plotted (squares)

.

The complex dependence of strain on distance from the cut in the AW

plate is shown in figure C4 . Instead of a uniform exponential decay with

distance from the cut, which would appear as a straight line when plotted in

semilogarithmic coordinates, figure C4 shows that the relaxation strains are nearly constant near the cut and then decay further away. Figure C5 shows

the behavior near the cut in more detail. Figures C6 and C7 show the behavior of the PWHT plate, where an exponential decay of strain with distance

occurred at distances up to 20 mm. The specific reason for this difference

between the AW and PWHT plates is not known.

The same sort of difference between AW and PWHT welds was observed in

the simulated tube-hole- ligament repair welds (figs. C8 through CIO).

Figure Cll shows strain relaxations measured on the bottom side of one PWHT

and one AW plate. These were small, so only the strain directly opposite

the cut was measured in subsequent tests.
Appendix C. Mechanical Measurements of Residual Stress

Figure Cl. Strain-gage layout for measurement of residual

strain relaxation during the cutting of the notch;

(a) front-side strain, (b) back-side strain.

30-
STRAIN

RELAXATION,

Postweld Heat Treatment Criteria

CUT DEPTH, mm

Figure C2 . Strain relaxation measured by a gage centered

7 mm (0.28 in) from the semicircular cut in an AW trough weld, as a function of cut depth.

- magnitude of relaxation of strains

perpendicular to the cut

magnitude of relaxation of

strains parallel to the cut

Figure C3. Strain relaxation in specimen 2N as a function of

distance from the cut at a depth of 11 mm (0.43 in).

1

1
Appendix C. Mechanical Measurements of Residual Stress

Figure C4 . Strain relaxation as a function of distance from the semicircular cut in trough weld in AW specimen 3N for several cut

depths. Each of the symbols at 4 mm (0.16 in) from the cut

represents an approximate average, excluding outliers, of the

maximum strains from an array of small strain gages near the

coO

<

X

<

LLJ ccz

<

DCI-co

3000

2000

1000

700

500

300

cut

.

The actual data from the

1 A387-3

l i (as welded)

1 i

A A

O A ° o

A Cut Depth o O

O 3.0 mm

6.4

a 10.7

1

mm

mm

l l i

O

1 2 4 6 8 10

DISTANCE FROM CUT, mm

12

Figure C5.

Strain relaxation as a function of distance near

the cut in an AW tough

weld of specimen 3N,

for several cut depths.

32
Postweld Heat Treatment Criteria

Figure C6

.

Strain relaxation as a function of

distance from the semicircular

cut in a trough weld in PWHT

specimen 4 for two cut depths.

Figure C7

.

Strain relaxation as a function

of distance near the cut in a trough weld in PWHT Specimen 4

for three cut depths

.
Appendix C. Mechanical Measurements of Residual Stress

DISTANCE FROM CUT, cm

Figure C9

.

Strain relaxation as a function of

distance from the cut in a simulated

tube-hole-ligament repair weld

of PWHT specimen 5S.

Figure C8

.

Strain relaxation as a function of

distance from the cut in a simulated

tube -hole - ligament repair weld

of AW specimen 6N.

DISTANCE FROM CUT, cm

34-
STRAIN

RELAXATION,

x

10°

STRAIN

RELAXATION,

Postweld Heat Treatment Criteria

CD

DISTANCE FROM CUT, cm

Strain relaxation as a function of distance near the cut in a simulated tube-hole- ligament repair weld of PWHT specimen 5S

.

DISTANCE FROM CUT, in

100

-100 -

-200 _L

Back Side Relaxation

o As Welded

a Post Weld Heat Treated

_L JL 5 10 15 20

DISTANCE FROM CUT, cm

25

Figure Cll. Strain relaxation on the back side as a function of distance

from the semicircular cut in two trough welds, AW and PWHT.
Appendix C. Mechanical Measurements of Residual Stress

The residual strains reported in table Cl represent the extrapolation

to the cut position of the strain values measured 2 cm (3/4 in) and farther

away from the cut. This extrapolation ignores the saturation effect noted

above. The residual stresses represent the average saturated strain value

measured near the cuts, multiplied by Young's modulus. The Poisson correc- tion was determined to be insignificant by using data from transversely

oriented strain gages.

The residual strains are affected by the PWHT more than the residual

stresses (see table Cl) . The mechanism of this difference and of the noted

strain saturation near the cut are unknown.

In the J- integral calculations, the saturated strain values were used,

so that the reported values of applied J- integral produced by residual

stresses are lower bound values.

Table Cl. Residual Stresses at Simulated Repair Welds

in Plates of 2kCr—IMo Steel.

Standard Troughs

Condition Me thod Stress, Strain

MPa (ksi) (post yield)

PWHT

AW

AW

Strain gage (cutting)

Strain gage (cutting)

Ultrasonic

325 (47.1)

372 (54.0)

380 (55.1)

1.6 x 10’ 3

2.6 x 10* 3

Simulated Tube! -Hole -Ligament Repairs

Condition Method

Stress

MPa (ksi)

Strain

(post yield)

PWHT

AW

Strain gage (cutting)

Strain gage (cutting)

55 (8.0)

383 (55.5)

0.3 x 10‘ 3

2.9 x 10‘ 3

36
APPENDIX D. MEASUREMENT OF APPLIED J- INTEGRAL PRODUCED BY RESIDUAL STRESS

D. T. Read

INTRODUCTION

A quantitative evaluation of the applied J- integral produced by

residual stresses is needed to make decisions on post-weld heat- treatment

(PWHT) criteria and inspection criteria based on fitness - for- service (FFS)

assessment in structures where fracture must be avoided. Existing FFS

assessment procedures typically include a procedure such as adding the

residual strain to the mechanical strain to obtain the total effective

strain [ D 1 ] . This strain becomes the starting point for a calculation of

the crack- driving force. In some procedures, residual strains, being selfequilibrating , are treated differently from mechanical strains [D2], These

analytical treatments have been justified on the basis of experimental

evidence of the adverse effect of residual stresses on fracture loads,

failure analyses that list residual stresses as a contributing cause of

failure, and engineering judgment.

However, few straightforward calculations of the effect of residual

stresses on crack- driving forces are available. The state of the art is

indicated by a recently published finite-element analysis of a centercracked tensile panel subjected to thermal and mechanical loading [ D 3

]

. Therefore, the opportunity exists to utilize experimentally derived information on the effect of residual stresses on fracture driving forces to verify

and improve FFS assessment methods and to provide information for specific

cases not covered by these methods

.

This paper reports the first use known to the author of the technique

of measuring contour strains to evaluate the applied J- integral produced by

residual stresses in simulated repair welds.

The technique of direct measurement of the J-integral has been applied

to base-metal plates in tension and bending for various crack sizes and up
Appendix D. Measurement of Applied J - Integral Produced by Residual Stress

to applied strain levels well beyond yield [D4,D5], The uncertainty of

the results was estimated at 10 percent in one base-metal study [D4]. The

presence of boundaries between weld metal and base metal in the specimen

interior considerably complicates the conceptual basis of the measurement

procedure [D6] and requires an approximation in the present technique.

MATERIALS AND SPECIMENS

Steel plate made to ASTM specification A-387 Grade 22 was used as base

plate. This is a 2kCr—IMo pressure -vessel steel in the annealed condition.

The plate thickness was 5 cm (2 in) . Troughs 20 cm (8 in) long by 5 cm

(2 in) wide by 3.2 cm (1-3/8 In) deep were filled with E-9018-B3L electrodes

2.3 mm (3/32 in) in diameter by the shielded-metal-arc (SMA) procedure.

The yield and tensile strengths of the as-deposited weld metal were 767 and

843 MPa (111 and 122 ksi) , and the base-plate yield and ultimate strengths were 270 and 499 MPa (39.2 and 72.4 ksi). Of four plate specimens tested,

two were tested as -welded (AW) and two matching specimens were tested after

PWHT . In the PWHT condition, the yield and ultimate strengths were 454 and

579 MPa (65.8 and 84.0 ksi). The weld properties were obtained from triplicate tensile specimens oriented along the weld direction extracted from

duplicate welds. The base-metal properties were supplied by the steel

maker

.

The J- integral measurements were made in simulated repair welds in

large plates 70 cm (27 in) wide by 180 cm (72 in) long. Two weld geometries

were used: a simple trough and a simulated boiler tube -hole - ligament repair;

these are shown in figure D1 . The simple troughs were 20 cm x 5 cm x 3 . 2 cm

(8 in x 2 in x 1-3/8 in), as described above. The notch was cut perpendicular to the weld direction, starting at the surface and cutting toward the

back face with a circular saw 7.0 cm (2-3/4 in) in diameter. This diameter

imposed a relation between cut depth and cut length such that at a depth of

3 mm (1/8 in), the length was 28 mm (1-1/8 in); at a depth of 6 mm (1/4 in),

the length was 39 mm (1-1/2 in); and at a depth of 10 mm (3/8 in), the

length was 49 mm (1-15/16 in).

38
Postweld Heat Treatment Criteria

0.7 m *

Semi-Elliptical Tube Hole

Surface Crack Ligament Crack

.Weld -Weld

/

1.8 m - *

>h

oro

ii \

Notch “Notch

2 plates: 2 plates

1 as welded 1 as welded

1 post weld 1 post weld

heat treated heat treated

Thickness = 5 cm

Figure Dl . Weldments of the wide-plate 2kCr—IMo steel used in this study.

The two simulated tube holes were 3.2 cm (1-1/4 in) in diameter, separated by a ligament of width 2 cm (3/4 in). In the ligaments, a deep groove

was machined extending from one hole to the other; it was filled with weld

metal. The machined notch extended from one hole to the other, starting at

the front surface and penetrating into the ligament thickness.

TECHNIQUE AND INSTRUMENTATION

Strain gages were located on a contour extending along the weld both on the front face where the notch opened and on the back face (opposite). In

the first specimen 34 gages were used, but in the others only 8 or 9 gages were needed. The strain-gage layouts are shown in figures D2 and D3 . Aft'

' »
Appendix D. Measurement of Applied J - Integral Produced by Residual Stress

Figure D2

Figure D3

.

Strain-gage layout on the front side of the plate for

measurement of the applied J- integral in a trough weld.

Strain-gage layout on the front side of the plate for mea- surement of the applied J- integral in a tube -hole ligament.

40-
Postweld Heat Treatment Criteria

the first test showed that the relaxations on the back face were minimal,

only one gage was used on the back face; it was positioned directly opposite

the notch location. A beam instrumented with an LVDT was used to indicate

the overall relaxation of the plate during the cutting.

The strain gages were wired to a microcomputer-based amplification and

recording system that provided a reproducibility of about 5 x 10' 6 strain

for all except one gage, which was excluded from the analysis because of

drift. Near the cut, an array of small gages separated by 1.5 mm was used.

The center of the gage closest to the mouth of the cut was about 2 mm from

the cut. The width of the saw blade was 1.5 mm.

After setup, checkout, equilibration, and recording of the zero point

data, the procedure was to lower the circular saw blade into the cut until

the depth was increased by about 1 mm (0.04 in), then to stop cutting to

measure the new cut depth and to record the strain gage strains and the

plate deflection indicator reading. The test itself required approximately

a half day for a final cut depth of about 14 mm (1/2 in).

ANALYSIS

Measured Strain Relaxations and Residual Strains

The measured strain relaxations of the front face gages were com- pressive, indicating the relaxation of a state of residual tension. The

relaxation strains decreased with distance from the cut, as required by

St. Venants's principle. In the PWHT plates, the straight-line dependence

of the strain relaxation on the distance from the cut in a semilog plot

(figs. D4 and D5) indicated an exponential decay. The strains were extrapolated to the cut location. Converting to stress using the elastic modulus

gave residual stresses of 325 MPa (47.1 ksi) at the semielliptical surface

crack and 55 MPa (8.0 ksi) at the tube-hole- ligament crack. Transverse

gages, not shown in figures D2 and D3 , relaxed only slightly, so that no contribution from transverse strains was used in the residual stress

calculation.
Appendix D. Measurement of Applied J - Integral Produced by Residual Stress

Figure D4 . Strain relaxation versus distance from cut in a PWHT trough

weld. Cut length increases along with the depth, as noted

in the text.

Figure D5 . Strain relaxation versus distance from cut in

a PWHT simulated tube -hole ligament weld.

42'
Postweld Heat Treatment Criteria

In the AW plates, the dependence of the strain relaxation on the distance from the cut was more difficult to understand (figs. D6 through D8)

.

Far from the cut, the dependence was exponential, as before; but near the

cut, a saturation occurred, so that the measured strains were below the

extrapolated value. Figure D7 illustrates this behavior in detail. The

saturation strain level is roughly equal to the base-metal yield strain.

For the J- integral calculations, a value of the residual strain was

needed. For the PWHT plates, the extrapolated value at the cut location

seemed appropriate. However, for the AW plates, because the measured strain

relaxations near the cut were below the extrapolated value, the strain value

at saturation was used.

These observations seem to indicate that the severity of a residual

stress field depends not only on its peak magnitude but also on its spatial

distribution.

Crack-mouth-opening displacement (CMOD) was obtained by integrating the

measured strains. The total y-direction displacement over the instrumented

contour was assumed to be zero, because of the restraint provided by the

large specimen plate. So

(Dl)

0

using the assumption of symmetry. CMOD was obtained by using equation Dl

.
Appendix D. Measurement of Applied J -Integral Produced by Residual Stress

Figure D6 . Strain relaxation versus distance from cut in an AW trough weld.

The cut length increases along with the depth, as noted in the

text. Each of the symbols at 4 mm from the cut represents an approximate average, excluding outliers, of the maximum strains

from an array of small gages near the cut. The actual data from

these are shown in figure D7

.

44.
Postweld Heat Treatment Criteria

coO

X

Z

O

l-<

x

<

LUQC

<

oc

I

CO

3000

2000

1000

700

500

300

0 2 4 6 8 10 12

DISTANCE FROM CUT, mm

1 1 [ A387-3 (as welded)

I

A A

O A

=30 __L A Cut Depth o O

-

o 3.0 mm

° 6.4 mm

^ 10.7 mm

_l i i

i

O

1

Figure D7 . Strain relaxation near the cut in an AW trough weld.

Figure D8 . Strain relaxation versus distance from cut

in an AW simulated tube -hole - ligament weld.
Appendix D. Measurement of Applied J - Integral Produced by Residual Stress

Accounting for the Weld in Calculating the J - Integral

Although an expression for the J- integral that includes thermal

stresses and strains has been derived [D7], it seemed preferable here to

treat the weld as a bimaterial composite of base metal and weld metal.

Within each, the J- integral is path- independent , as usual. The definition

is

J W dy — T • du/dx ds

f

(D2)

where, as usual, x and y are Cartesian spatial coordinates with their origin

at the crack mouth, with x the direction of self-similar crack propagation

and y the direction perpendicular to the crack plane; T is a contour enclos-

— ing the crack tip, W is strain work density, T is traction normal to the

— contour, u is displacement, and ds is increment of arc length along the con- tour. Usually, the combination of strain measurements on half of the con- tour (say, y > 0 only), the measured CMOD , and the assumption of symmetry

about the crack plane enable both terms of equation Dl to be evaluated, with

some cancellation of errors between the two terms.

However, because of the yield strength mismatch between the present

specimen base and weld metal, the J- integral is not path- independent for

contours that lie partially in the base metal and partially in the weld

metal. In the present case, the experimental contour cannot be restricted

to the weld metal, because the weld is in a groove. Therefore, an additional term in the J- integral relating to the interface between the weld

metal and the base metal must be considered. Mathematically,

J ^usual ^interface (D3)

The second term in equation D2 , the traction-bending term, could not be

measured in the present case because the CMOD was not measurable directly.

Equation Dl implies that this term is zero. The contour was chosen so that

46
Postweld Heat Treatment Criteria

the end lay as far as possible from the crack plane, and thus the tractionbending term would be small. The severe restraint provided by the largeplate specimen contributes to minimizing the traction-bending term. The

assumption is:

^usual ^front face ^back face

where

J front face

*L W dy,

o front face

(traction-bending neglected) (D4)

(D5)

and Jback face is defined similarly.

In strain- gage measurements of the J- integral in base metal with no crack (a = 0) , the magnitudes of the front-face and back-face terms would

both increase with strain, but they would cancel each other. That is,

J = _J front face( a=°) + Jback face( a=0 ) = 0 (base plate) (D6)

where a is crack depth.

After the notch is cut, the strains near the crack mouth relax, causing

W to decrease and thus reducing the magnitude of the front- face term and

increasing the total J -value. Under normal conditions, that is, when the

crack depth is less than half the plate thickness and when there is no net

section yielding, the back- face term remains about the same as it would have

been if there were no crack. Therefore, for a given strain level,

Jback face ( a ) - l J front face( a=0 )| (normal conditions) (D7

)

and

J “ “ I dfront face( a )| + | J front fac e ( a=0 )| (normal conditions) (D8)

where the absolute value signs have their usual meaning.

In a weldment, equations D3 and D4 become

4 7
Appendix D. Measurement of Applied J - Integral Produced by Residual Stress

D " ^front face ^back face ^interface (D9)

The difficulty in evaluating J in equation D9 is the interface term. Equation D9 must apply both before and during the cutting of the notch. Before

cutting, the full residual strains are present, but the total J- integral

must be zero, because, as yet, there is no notch:

0 = -J front face (a=0) + Jback face (a=0) + J,interface (a=0)

where a is crack length.

(DIO)

Therefore, an expression for J can be formed from equations D9 and DIO:

J - Jfront face

(

a ) + Jback face( a ) + ^interface

(

a ) (Dll)

^front face ( a_^) ^back face( a_^^ ^ inter fac e ( a—^ )

Equation Dll shows that what is needed is the change in the interface

term, rather than in its absolute value. This term cannot be measured

experimentally because the interface is buried within the specimen. Therefore, the behavior of the interface term can only be assumed. The known

behavior of the back- face strain relaxations provides a clue. These relaxations were measured and found to be small; the interface is much closer to

the back face than to the front face, so the strain relaxations at the

interface should also be small. Therefore, the interface term of the

J- integral was assumed to remain constant during cutting. That is,

^interface ( a ) = constant (assumption) (D12a)

^interface ( a ) ~ ^inter f ace ( a=^ ) = ^ (Dl2b)

Various arguments in support of equation D12 were considered, but none seemed definitive. Therefore, equation D12 is treated here as a needed

and plausible assumption.

From equations Dll and D12,

J = -J front + Jback

face ( a ) ^front face( a

face ( a ^ ^back face( a—

(D13)

Because the changes in Jback face during cutting were small, equation

48-
Postweld Heat Treatment Criteria

D13 for welds is essentially the same as equation D8 for base plate. Equa

tion D13 was used to estimate the applied J-integral in the present weld

specimen experimentally. The uncertainty results from the assumptions in

equations D4 and D12. The uncertainty in the J-integral was estimated to

be plus or minus 20 percent.

RESULTS AND DISCUSSION

The measured J-integral values are plotted against crack depth in fig

ure D9 . The clear differences between the AW and PWHT plates result from

the higher residual strains in the AW plates. The initial linear increase

of J-integral with depth and the saturation at the latter stages are as anticipated from results reported in the literature for semielliptical

surface flaws in tension and bending.

E

Jx'0

cn

CD jz

1 ~>

~o

_<dQ_

CL <

0.0 0.1 Crack Depth, in 0.2 0.3 0.4 0.5 40

30

20

10

2 1/4 Cr 1 Mo Repair Welds

Residual Stress J

° = SESC-PWHT

a = THLC-PWHT

< = SESC-AW

x = THLC-AW x

x X X

.X*

44

200 c

cn

(D

100 _Lxs

"o. CL <

i48 00 8

A A A . h- ,

O0 o§ 0§o°

A A_

10 Crack Depth, mm

15

Figure D9 . Measured applied J-integral values plotted against cut depth.

SESC = surface crack in trough weld

THLC = surface crack in tube -hole - ligament repair weld
Appendix D. Measurement of Applied J - Integral Produced by Residual Stress

Stress intensity, K, values for a surface - cracked plate in bending,

calculated by Newman and Raju [D8], were converted to units of J-integral

using the standard formula K = (E-J) , where E is Young's modulus, for comparison to the present results (fig. DIO) . The approximate agreement of the

present results with the Newman—Raju values was obtained by choosing a stress level of 519 MPa (75.2 ksi) , which is equal to the average of the

yield strengths of the weld and base metal. This stress is higher than the

measured residual stress of 372 MPa (54 ksi) . The extrapolated residual

strain at the cut location was 2.6 x 10' 3

. When converted to stress this

gives 538 MPa (78 ksi) , which would also give approximate agreement between

predicted and measured applied J -values. The agreement between the present

results and the Newman—Raju results supports the validity of the present

results

.

0.0 Crack Depth, in0.1 0.2 0.3 0.4 0.5

100

0

cr>

_0> j=

1 —>

"a

QJa_a. <

Figure D10. Comparison of calculated and measured J-integral

values as a function of cut depth.

50-
Postweld Heat Treatment Criteria

The calculation of CMOD from measured strains is straightforward.

Therefore, it would be attractive to correlate measured CTOD and J- integral

with calculated CTOD and J- integral to verify the measured J -values. But

only approximate calculated CMOD values are available for surface cracks

in tension [D9] and none are available for surface cracks in bending.

Therefore, ratios of K to CMOD from calculations for single-edge cracks in

tension and bending and for surface cracks in tension were compared with the

experimental ratio of K to CMOD. The ratios are listed in table 1 for the

final crack depth of the surface-cracked specimen. These data indicate that

the present results for the ratio of K to CMOD are consistent with the trend

of the calculated values for single-edge cracks. The handbook ratio for

bending is about 10 percent less than for tension; the present value for

a semielliptical surface crack is about 10 percent less than that for such

a crack in tension.

Table 1. Normalized ratio of K to CMOD,

at a = 10.9 mm (0.428 in).

K

for various crack types

E • CMOD

Type K/ (E • CMOD) , cm

~h K/ (E • CMOD) , in

-h

Center crack, tension

Single-edge crack, tension

Single -edge crack, bending

Semielliptical surface crack, tension

Present, semielliptical surface crack

0.425

0.379

0.324

0.283

0.338

(0.677)

(0.516)

(0.451)

(0.604)

(0.539)
Appendix D. Measurement of Applied J- Integral Produced by Residual Stress

SUMMARY

A destructive technique has been demonstrated for measurement of the

applied J- integral produced by residual strains from welding. This technique clearly distinguishes as-welded from stress-relieved plates. The

applied J- integral was much higher in two AW plates than in two postweld

heat treated plates. The technique shows the increase of J- integral with

crack depth. The measured J -values, converted to units of stress intensity,

approximately matched those calculated for a yield- level bending stress

applied to an equivalent plate, when the effective yield stress was taken as the average of base-metal and weld-metal yield stresses. The J-values for a crack in a simulated boiler tube -hole - ligament repair were slightly lower

than those for a semielliptical surface crack in a repair weld in an intact

plate

.

ACKNOWLEDGMENTS

The financial support provided by the Naval Sea Systems Command through

its project sponsor Commander L. R. Easterling is gratefully acknowledged.

Helpful discussions with W. C. Fraser and M. A. E. Natishan of the David

Taylor Naval Ship Research and Development Center are appreciated. Technical assistance by J . D. McColskey, M. Moyer, and J. J. Broz of the

National Bureau of Standards is also appreciated.

REFERENCES

[Dl] British Standards Institution. Guidance on some methods for the

derivation of acceptance levels for defects in fusion welded joints.

PD 6493: 1980, British Standards Institution, London, 1980.

[D2] Harrison, R.P., Loosemore, K. , Milne, I., and Dowling, A.R.

Assessment of the integrity of structures containing defects,

R/H/R6-Rev. 2. Central Electricity Generating Board, London, 1980.

52'
Postweld Heat Treatment Criteria

[D3] Hellen, T.K. and Blackburn, W.S. Post-yield fracture mechanics

analysis of the combined thermal and mechanical loading of a centrecracked plate. International Journal of Fracture 32, 1987, pp

.

185-199.

[D4] Read, D.T. Experimental method for direct evaluation of the

J-contour integral, in Fracture Mechanics: Fourteenth Symposium — Volume II: Testing and Applications. J.C. Lewis and

G. Sines, Eds. ASTM STP 791, American Society for Testing and

Materials, Philadelphia, 1983, pp. 11-199—11-213

.

[D5] Dodds, R.H. and Read, D.T. Experimental and analytical estimates of

the J- integral for tensile panels containing short center cracks.

Int. J. Fract. 28, 1985, pp . 39—54.

[D6] Read, D.T. Experimental results for fitness-for-service assessment

of HY130 weldments. NBSIR 84-1699, National Bureau of Standards,

Boulder, Colorado, March 1984, 90 pp

.

[D7] Chell, G.G. Elastic—plastic fracture mechanics, in Developments in

Fracture Mechanics — I. G.G. Chell, Ed. Applied Science Publishers, London, 1979, p. 97.

[D8] Newman, J.C., Jr. and Raju, I.S. Analyses of surface cracks in

finite plates under tension or bending loads. NASA Technical Paper

1578, National Aeronautics and Space Administration, October 1979,

43 pp.

[D9] Mattheck, C., Morawietz , P., and Munz , D. Stress intensity factor

at the surface and at the deepest point of a semielliptical surface

crack in plates under stress gradients. Int. J. Fract. 23, 1983,

pp. 201-212.
NBS-114A (REV. 2-8C ) U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET (See instructions)

4. TITLE AND SUBTITLE

PUBLICATION OR

REPORT NO. NBSIR 87-3075

2. Performing Organ. Report No. 3. Publication Date

August 1988

POSTWELD HEAT TREATMENT CRITERIA FOR REPAIR WELDS

in 2-l/4Cr-lMo SUPERHEATER HEADERS

AN EXPERIMENTAL STUDY

5. AUTHOR(S)

D.T. Read and H.I. McHenry

6. PERFORMING ORGANIZATION (If joint or other than NBS. see instructions) 7. Contract/Grant No. NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE 8. Type of Report & Period Covered

WASHINGTON, D.C. 20234

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City. State. ZIP)

Department of the Navy

Naval Sea Systems Command

Washington, DC 20362

10. SUPPLEMENTARY NOTES

Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here) Wide-plate and standard-size specimens cut from repair welds in 2-l/4Cr-lMo plate

were tested as-welded and after post-weld heat treatment (PWHT) . Wide plates were prepared with surface cracks and tube-hole-ligament cracks in simulated repair welds made

with the shielded-metal-arc process. The plates were instrumented with strain and

displacement gages for direct measurement of the applied J- integral. After notching and

fatigue precracking, the plates were tested in bending. Pop-ins occurred in the as- welded plates, but not in the PWHT plates.

Three-point-bend specimens with cracks oriented in the TS direction were used to measure weld-metal and heat-affected-zone (HAZ) toughness values. Results of direct

measurements of the applied J-integral on the wide plates were compared with critical

J-value measurements of three-point-bend specimens. The comparison indicated that PWHT

was highly beneficial, because it reduced the crack-driving force from residual stresses

and increased the weld-metal and HAZ toughness. In the as-welded condition, very low

toughness values were measured at the HAZ. These low toughness values, together with

the measured crack-driving forces, indicated critical crack depths of a few millimeters.

To extend the usefulness of these results, a new approach to the problem of the

applied J-integral produced by residual stresses is being explored: strain-gage measure- ments made during notching are analyzed to obtain an applied J-integral as a function

of crack depth. The preliminary results are encouraging. The residual-stress-produced

J-value^is

e joughly equivalent to that produced by a remote elastic loading to the same

12. KEY WORDS (Six to twelve entries ; alphabetical order; capitalize only proper names; and separate key words by semicolon s) fracture; hydrotest; J-integral; pressure vessel; residual stress; safety; stress relief

13. AVAILABILITY 14. NO. OF

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