Effect of welding and post-weld heat treatment on QT 35 steel

1 ^ JAM. 7971

CRANFIELD

INSTITUTE OF TECHNOLOGY

EFFECT OF WELDING AND POST-WELD HEAT

TREATMENT ON QT 35 STEEL

BY

CRANFIELD INSTITUTE OF TECHNOLOGY

E F F E C T OF WELDING AND POST-WELD HEAT TREATMENT ON QT 35 STEEL b y E. Smith. Ph. D . , B. Sc. . A. I. M. and R. L. Apps, P h . D . , B. S c , F . I. M . , M. Weid. I. SUMMARY

A simulation technique has been employed to exannine the s t r u c t u r e s and properties of the heat affected zone in single and multipass submerged a r c welds in QT 35 steel for a range of heat inputs i . e . 1.2, 2 . 1 , and 4.2 k J / m m (30, 54, and 108 k J / i n ) .

Mechanical properties were a s s e s s e d using Charpy V-notch, tensile and hardness tests and related to m i c r o s t r u c t u r e as determined by electron microscopy. Notch impact values deteriorated over the range of t h e r m a l cycles examined and were most severely affected by high peak tem.peratures and by slow cooling r a t e s which gave the c o a r s e s t metallurgical s t r u c t u r e s . Additional t h e r m a l cycles corresponding to those applied by tempering weld beads were shown to improve Impact performance in some parts of the heat affected zone, but not in o t h e r s .

The effects of post-weld heat t r e a t m e n t s at 45G°C, 550°C, 650 C and 700°C for 1 hour on the c o a r s e grained heat affected zone s t r u c t u r e s were examined. F u r t h e r embrittlement occurred after treatment at 450 C but the notch toughness improved at higher t e m p e r a t u r e s . Treatment at 650 C was the most effective and reliable method of improving notch toughness without adversely affecting the parent plate strength.

INTRODUCTION 1 EXPERIMENTAL

2. 1. M a t e r i a l s 2 2.2 P r o c e d u r e 2

2. 2 . 1 Simulation of weld HAZ s t r u c t u r e s 2 2. 2. 2 Simulation of post-weld heat t r e a t m e n t 3

2. 2. 3 Mechanical testing 3 2. 2. 4 Metallographic examination 3

RESULTS

3.1 Single cycle simulation 3 3. 1. 1 Mechanical testing 3 3. 1. 2 Metallographic examination 4

3.2 Double cycle simulation 5 3. 2 . 1 Mechanical testing 5 3. 2. 2 Metallographic examination 6 3 . 3 P o s t - w e l d heat t r e a t m e n t 6 3. 3. 1 Mechanical testing 6 3. 3. 2 Metallographic examination 6 DISCUSSION 7 4. 1 F r a c t u r e toughness r e q u i r e m e n t for parent m a t e r i a l 7

4. 2 F r a c t u r e toughness r e q u i r e m e n t for the weld HAZ 9

4. 3 Post weld heat t r e a t m e n t 10

CONCLUSIONS 11 REFERENCES 12 TABLES

QT 35 is a low carbon, low alloy, Mn - Ni - C r - Mo - V s t e e l which was developed for use in s u b m a r i n e hulls and highly s t r e s s e d p a r t s of surface ships where ductile behaviour under impact loading at low t e m p e r a t u r e s is d e s i r a b l e . It is used in a quenched and t e m p e r e d condition to the Ministry of Defence (Ship Department)

specification DG S h i p s / P S / 9 0 1 4 (Ref. 1). The s t e e l has a yield strength (0.2% proof s t r e s s ) in the range 36 - 44 ton f/in (Ref 2) and a minimum Charpy V-notch toughness

(longitudinal) of 81J (60ft. Ibf) at -40 C with a fracture appearance of not g r e a t e r than 75% c r y s t a l l i n i t y . The m a t e r i a l specification i s contained in DG Sliips/5135 B (Ref. 2). Weldability is claimed to be adequate if modest p r e h e a t s (120 - 150 C) and suitable matching e l e c t r o d e s a r e used (Ref. 3)

Post-weld heat t r e a t m e n t of QT 35 is not recommended because of p r a c t i c a l difficulties but the specification provides for a s t r e s s relieving t r e a t m e n t in special cases, such a t r e a t m e n t to be c a r r i e d out at 625 C + 25 C for 1 hour p e r inch of thickness of the thickest m e m b e r of the weldment. In c a s e s where excessive r e s t r a i n t and cooling effects a r e p r e s e n t the preheat t e m p e r a t u r e may be maintained for 1 hour after completion of welding to reduce the cooling r a t e of the deposited weld and p a r e n t plate. The use of a post-weld heat t r e a t m e n t aimed at t e m p e r i n g of the weld heat affected zone (HAZ) has apparently not been considered n e c e s s a r y . The development of simulation techniques, however, has enabled notch toughness m e a s u r e m e n t s of the weld HAZ s t r u c t u r e s to be made and has shown that v e r y low notch toughness can occur in QT 35 s t e e l (Ref. 4). This is due to the production of virtually quenched s t r u c t u r e s in the HAZ and it would be logical to expect that t e m p e r i n g would be v e r y effective in improving notch toughness.

P r e v i o u s work on QT 35 and other low alloy s t e e l s (Refs 4-6) has shown that the structures produced by the weld t h e r m a l cycle s i m u l a t o r a r e comparable to actual weld HAZ s t r u c t u r e s . It was also shown that s e v e r e degradation of notch toughness o c c u r r e d in the HAZ of s u b m e r g e d a r c welds in QT 35 using a heat input of 4. 2 k J / m m (108 k J / i n ) in 38mm (1.5 in) plate (Ref. 4). The lowest notch toughness was associated with a p r e -dominantly upper bainitic s t r u c t u r e in the region of g r a i n c o a r s e n i n g where an impact value of 28J (20 ft. Ibf) at 50 C was r e c o r d e d . However, this heat input is above the maximum of 2 . 5 k J / m m (65 k J / i n ) specified for submerged a r c welding of QT 36 s t e e l (Ref. 1). In addition Savage and Owczarski (Ref. 7) and Nippes et a l (Ref. 8) have shown that in this c l a s s of s t e e l a distinct i m p a i r m e n t of notch toughness in the weld HAZ r e s u l t s from high heat input. With lower heat inputs cooling r a t e s a r e i n c r e a s e d and r e s u l t in the formation of s t r u c t u r e s containing a g r e a t e r proportion of lower bainite and autotempered m a r t e n s i t e and consequent improvennent in notch toughness.

The p r e s e n t work had two main o b j e c t i v e s :

1. To define m o r e closely the i m p a i r m e n t of notch toughness, as m e a s u r e d by the Charpy V-notch t e s t , in the HAZ of s u b m e r g e d a r c welds in QT 35 s t e e l at heat inputs of 1.2, 2 . 1 . and 4 . 2 k J / m m (30, 54, and 108 k J / i n ) . A simulation technique, employing both single and double t h e r m a l c y c l e s , was used to evaluate the HAZ of single and m u l t i p a s s welds. Additional information was obtained from h a r d n e s s and tensile t e s t s and the mechanical p r o p e r t i e s w e r e r e l a t e d to m i c r o s t r u c t u r e using the carbon extraction replica technique.

2. To evaluate the effectiveness of post-weld heat t r e a t m e n t s for 1 hour in the range 450 - 700 C in overcoming s e c o n d a r y hardening p r o b l e m s and

2. 1. M a t e r i a l s

The 38 m m . (1. 5 in. ) thick QT 35 s t e e l plate had the following composition which conformed to specification DG. Ships/5135 B. (Ref. 2).

c

Tensile and Charpy V-notch impact t e s t s conducted on specimens taken from the plate m i d - t h i c k n e s s yielded the following r e s u l t s which conformed to specification DG Ships/ 5135 B (Ref. 2). Specimen orientation Longitudinal Charpy Energy*at-40°C J 133 ft. Ibf. 98 0. 2% proof s t r e s s M N m ' ^ 617 tonf. in 40. 1 U . T . S . M N m ' ^ 713 -2 tonf. in 46. 3

* notched in the through thickness direction 2 . 2 . P r o c e d u r e

2. 2, 1. Simulation of weld HAZ s t r u c t u r e s

Weld t h e r m a l cycle simulation was c a r r i e d out on an apparatus designed and built at Cranfield in which a m a t e r i a l blank. 10. 7 m m x 10. 7 mm x 83 m m (0. 42 in x 0. 42 in X 3. 25 in) i s heated by virtue of its own r e s i s t a n c e to the passage of an e l e c t r i c c u r r e n t and cooled by the flow of w a t e r through hollow b r a s s clamping blocks. The Cranfield s i m u l a t o r has been d e s c r i b e d in detail by Clifton and George (Ref. 9). The blanks w e r e machined from the m i d - t h i c k n e s s of the plate with their length p a r a l l e l to the rolling direction. They w e r e held rigidly in the clamping blocks during simulation in an attempt to induce some r e s t r a i n t although it was accepted that this would be far l e s s than the r e s t r a i n t o c c u r r i n g in p r a c t i c e .

T h e r m a l cycles with peak t e m p e r a t u r e s of 1300°C. 950°C, and 830 C were chosen to give s p e c i m e n s with s t r u c t u r e s corresponding to c o a r s e grained, fine grained, and i n t e r c r i t i c a l regions of the HAZ. Actual t h e r m a l cycles for heat inputs of 1, 2, 2.1, and 4. 2 k J / m m (30, 54 and 108 k j / l n ) were used to give a wide variation in cooling r a t e s , i . e . 18, 1 0 . 5 , and 5.5 C / s e c . through the t e m p e r a t u r e range 700 - 300 C, corresponding to both r e c o m m e n d e d and unacceptable welding conditions. The nine t h e r m a l cycles used a r e shown in figs. 1 - 3 and c o r r e s p o n d to those experienced under the following

conditions:-Heat Input k J / m m 1.2 2. 1 4 . 2 k J / i n 30 54 108 P r e h e a t Temp. none 120°C none A r c c u r r e n t (amps) 130 500 390 A r c Voltage 23 30 30 T r a v e l speed m m / s e c . 2.48 7.05 2. 75 i n . / min. 5 . 8 16. 7 6. 5 Welding p r o c e s s automatic m e t a l a r c submerged a r c submerged a r c Weld preparation single Vee bead-on-plate bead-on-plate Reference Nippes et al (Ref. 10) Smith and Kellock (Refs 11 and 12) Coward (Ref.l3)

The following t h e r m a l cycles and combinations of t h e r m a l cycles w e r e used to examine the HAZ s t r u c t u r e s of single and m u l t i p a s s weldments for each heat

input:-1 Peak Temperature, °C | First cycle 1300 950 830 1300 1300 1300 Second cycle

1

1

2. 2. 2. Simulation of post-weld heat t r e a t m e n t

Some of the s p e c i m e n s cycled once to the peak t e m p e r a t u r e of 1300 C for each heat input w e r e given furnace heat t r e a t m e n t at 450 C, 550 C, 650 C, and 700 C for 1 hour. After 1 hour a t t e m p e r a t u r e the s p e c i m e n s w e r e removed from the furnace and allowed to cool in s t i l l a i r .

2. 2. 3, Mechanical testing

Ten s t a n d a r d Charpy V--notch impact specimens w e r e p r e p a r e d from the simulated blanks for each condition studied with the notch machined in the control thermocouple position and in the through thickness direction of the original plate. Impact t r a n s i t i o n curves w e r e determined by testing over the t e m p e r a t u r e range - 196 C to 180 C.

T h r e e No. 13 Hounsfield tensile t e s t specimens with a modified gauge length of 7. 6 m m (0. 3 in. ) w e r e p r e p a r e d from the simulated blanks for each condition with the gauge length a c c u r a t e l y positioned within the heat t r e a t e d zone at the c e n t r e of the blanks. The s p e c i m e n s w e r e tested on a s t a n d a r d Instron tensile machine at a s t r a i n r a t e of approx-imately 3 X lO" / s e c . The values of 0. 2% proof s t r e s s , tensile strength, and reduction of a r e a w e r e r e c o r d e d .

One simulated blank from each condition was sectioned through the centre of the heat t r e a t e d zone and the h a r d n e s s d e t e r m i n e d using the Zwick h a r d n e s s t e s t e r and a load of 5 kg.

2 . 2 . 4 . Metallographic examination

The sections used for h a r d n e s s determinations w e r e reground and p r e p a r e d for optical metallographic examination using 2% nital for etching. E s t i m a t e s w e r e made of the p r i o r austenite grain size in specimens cycled once to the 1300 C peak t e m p e r a t u r e for each heat input. P r i o r to the p r e p a r a t i o n of carbon extraction r e p l i c a s the s p e c i m e n s w e r e given a d e e p e r etch in 2% nital in o r d e r to e n s u r e the production of s a t i s f a c t o r y r e p l i c a s . The carbon r e p l i c a s w e r e then p r e p a r e d in the s t a n d a r d m a n n e r and examined in an electron m i c r o s c o p e . In some of the specimens e s t i m a t e s of the proportions of upper and lower bainite and autotempered m a r t e n s i t e w e r e made using a s y s t e m a t i c point count method. Approximately 2000 point counts w e r e used for each estimation.

3. RESULTS 3 . 1 Single cycle simulation 3. 1. 1. Mechanical testing

and i n t e r c r i t i c a l regions of the weld HAZ produced by t h e r m a l cycli.ig to pes.x t e m p e r a t u r e s of 1300°C, 950°C, and 830°C respectively, a r e shown in figs. 4-6 for the three hsat inputs. The tensile and h a r d n e s s r e s u l t s a r e shown in Table 1 and various Charpy transition c r i t e r i a a r e shown in Table 2. The r e s u l t s show a marked drop in notch-toughness of the c o a r s e grained HAZ for each heat input. With the low and medium heat inputs the impact value at 40 C i s 19 21 J (14 15 ft. Ibf) with a fracture appearance of 85% crystallinity c o m -p a r e d with 10 J (7 ft. Ibf) and 95% c r y s t a l l i n i t y for the high heat in-put. The u-p-per sheK energy, however, is higher for the high heat input, 102 J (75 ft. Ibf) compared with about 68 J (50 ft. Ibf) for the medium and low heat inputs. The h a r d n e s s d e c r e a s e d from 384 HV5 for the low heat input to 346 HV5 for the high heat input with a corresponding drop in the s t r e n g t h p r o p e r t i e s and i n c r e a s e in ductility a s m e a s u r e d by reduction of a r e a . This i s consistent with the formation of higher t e m p e r a t u r e transformation products and reductions in dislocation density and solid solution strengthening, p a r t i c u l a r l y by carbon, accompanying i n c r e a s e s in heat input.

The impact values for the fine grained and i n t e r c r i t i c a l regions of the HAZ w e r e s i m i l a r for each heat input and significantly higher than in the c o a r s e grained HAZ. The fine grained HAZ had values in the range 43 - 52J (32 - 38 ft. Ibf) at -40°C with a fracture appearance of about 75% crystallinity, while the i n t e r c r i t i c a l HAZ had values in the range 57 - 67 J (42 - 49 ft. Ibf) at -40°C with a fracture appearance of about 65% crystallinity. In these s t r u c t u r e s the h a r d n e s s d e c r e a s e d with i n c r e a s i n g heat input although no distinct t r e n d s w e r e apparent in the tensile p r o p e r t i e s . Table 1.

3. 1. 2. Metallographic examination

The parent plate s t r u c t u r e , fig. 7, consisted of t e m p e r - c a r b i d e s , sited mainly at high angle boundaries, in a f e r r i t e m a t r i x . The c a r b i d e s were mainly c o a r s e with an elongated or spheroidised morphology and have been identified as mainly cenientite by Dolby, (Ref. 14). Some a r e a s showed comparatively little precipitate and it is probable that these w e r e originally upper bainite in the quenched condition.

The c o a r s e grained HAZ s t r u c t u r e s produced by t h e r m a l cycling to a peak t e m p e r a t u r e of 1300 C a r e shown in fig. 8. Complete austenitisation o c c u r r e d during the weld t h e r m a l cycle with a considerable degree of g r a i n growth, as shown by the large p r i o r austenite g r a i n s i z e of each of these s t r u c t u r e s . With the low heat input the s t r u c t u r e was predominantly autotempered m a r t e n s i t e , fig. 8a, but containing s m a l l amounts of upper and lower bainite. The formation of a u t o t e m p e r e d m a r t e n s i t e can be attributed to the r e l a t i v e l y high Ms t e m p e r a t u r e of QT 35 s t e e l . Nisbet and Kelly (Ref. 15) have determined an Ms t e m p e r a t u r e of 430 C for QT 35 s t e e l and this a g r e e s with values calculated using the formulae of Steven and Haynes (Ref. 161t (415°C) and Andrews (Ref 17), (420°C). The significance of a high Ms t e m p e r a t u r e i s that any m a r t e n s i t e formed during cooling will be t e m p e r e d a s the t e m p e r a t u r e falls to ambient and this will be manifest by the precipitation of c a r b i d e s in a Widmanstatten distribution within the m a r t e n s i t e laths. With the medium heat input t h e r e was a g r e a t e r proportion of upper and lower bainite, fig. 8b, although

autotempered m a r t e n s i t e was still the predominant constituent. At the high heat input the s t r u c t u r e consisted mainly of upper bainite, fig. 8c, but with some lower bainite and autotempered m a r t e n s i t e . Grain s i z e m e a s u r e m e n t s , using the l i n e a r intercept method and m i c r o s t r u c t u r a l point counts, yielded the following r e s u l t s ;

-Heat Input k J / m m 1. 2 2. 1 4. 2 k J / i n 30 54 108 Average grain d i a m e t e r 0. 054 mm 0. 060 mm 0. 071 nmi M i c r o s t r u c t u r e (%) M a r t e n s i t e 85 70 35 Bainite 15 30 65

Upper and lower bainite w e r e counted a s one constituent because of the difficulty in distinguishing between them in a r e a s containing a mixed bainitic s t r u c t u r e . The r e s u l t s show an i n c r e a s i n g proportion of bainite, mainly upper bainite, and an i n c r e a s i n g p r i o r austenite grain size with i n c r e a s i n g heat input. This can be attributed to the slower cooling r a t e and the g r e a t e r time spent above the g r a i n c o a r s e n i n g t e m p e r a t u r e with i n c r e a s i n g heat input.

The fine grained HAZ s t r u c t u r e s produced by the t h e r m a l cycling to a peak t e m p e r a t u r e of 950 C a r e shown in fig. 9. Austenitisation was virtually complete during the t h e r m a l cycle indicating a peak t e m p e r a t u r e close to the Ac , but the fine p r i o r austenite g r a i n s i z e showed that the g r a i n c o a r s e n i n g t e m p e r a t u r e had not been r e a c h e d . On cooling the austenite t r a n s f o r m e d to a g r a n u l a r type of s t r u c t u r e which was neither a u t o t e m p e r e d m a r t e n s i t e nor upper or lower bainite. This t r a n s f o r m a t i o n product will be called g r a n u l a r bainite in the p r e s e n t r e p o r t . A few fine globular c a r b i d e s w e r e d i s p e r s e d throughout the s t r u c t u r e s .

The i n t e r c r i t i c a l HAZ s t r u c t u r e s produced by t h e r m a l cycling to a peak

t e m p e r a t u r e of 830 C a r e shown in fig. 10. These w e r e somewhat s i m i l a r to those of the fine grained HAZ but with l a r g e a r e a s of f e r r i t e containing little p r e c i p i t a t e , and t r a n s f o r m e d regions s t r u c t u r a l l y s i m i l a r to the g r a n u l a r bainite d e s c r i b e d above but containing s m a l l a r e a s of a u t o t e m p e r e d m a r t e n s i t e . The absence of c a r b i d e s in the f e r r i t e regions indicates that the c a r b i d e s initially p r e s e n t in the p a r e n t plate were r e a d i l y dissolved during the t h e r m a l cycle and the carbon atomis w e r e sufficiently mobile to diffuse to the austenitised r e g i o n s . The c a r b o n - e n r i c h e d austenite then t r a n s f o r m e d to a m i x t u r e of a u t o t e m p e r e d m a r t e n s i t e and g r a n u l a r bainite on subsequent cooling.

3. 2. Double cycle simulation

In this section simulated t h e r m a l cycles a r e designated solely by the peak t e m p e r a t u r e attained e . g . 1300 C r e f e r s to a single cycle having a peak t e m p e r a t u r e of

1300 C w h e r e a s 1300/950 C r e f e r s to a double cycle with first and second cycle peak t e m p e r a t u r e s of 1300 C and 950 C r e s p e c t i v e l y .

3 . 2 . 1 . Mechanical testing

The Charpy t r a n s i t i o n c u r v e s for the s p e c i m e n s given double t h e r m a l cycles in the s i m u l a t o r a r e shown in figs. 11 - 13 and the h a r d n e s s , tensile p r o p e r t i e s and Charpy t r a n s i t i o n c r i t e r i a a r e included in tables 1 and 2. At the low and medium heat inputs the 1300/1300 C t r e a t m e n t produced s t r u c t u r e s with lower notch-toughness than after the 1300°C t r e a t m e n t with impact values of 15 - 16J(11 - 12 ft. Ibf) at -40°C and a fracture a p p e a r a n c e of 85 - 90% c r y s t a l l i n i t y . figs. 11 and 12. With the high heat input the 1300/1300 C t r e a t m e n t produced a s t r u c t u r e of s i m i l a r notch-toughness to that produced by the 1300 C t r e a t m e n t , fig. 13. In each c a s e the h a r d n e s s was slightly lower for the double cycled condition while the s t r e n g t h was i n c r e a s e d with the low heat input and d e c r e a s e d with the medium and high heat inputs.

The 1300/950 C t r e a t m e n t r e s u l t e d in i n c r e a s e d notch-toughness over the 1300 C t r e a t m e n t at each heat input. The low and medium heat inputs were s i m i l a r with impact values of 25 - 26 J ( 1 8 - 19 ft. Ibf.) at - 40°C and a f r a c t u r e a p p e a r a n c e of 75 - 80 % c r y s t a l l i n i t y , figs. 11 and 12. The notch-touchness was significantly lower with the high heat input with an impact value of 15 J (11 ft. Ibf) at 40 C and a f r a c t u r e a p p e a r -ance of 95% c r y s t a l l i n i t y , fig. 13. The hai'dness and strength w e r e i n c r e a s e d with the low heat input, but d e c r e a s e d with the medium and high heat inputs, table 1.

The 1300/830 C t r e a t m e n t produced a furtlier improvement in notch-toughness at each input. The low and medium heat inputs had impact values of 41 55 J (30 -40 ft. Ibf) at - -40 C and a fracture appearance of 65 - 70 % c y r s t a l l i n i t y , figs. 11 and 12, while the high heat input was significantly lower in notch-toughness with

c o r r e s p o n d i n g values of 24 J (17 ft. Ibf) and 90% c r y s t a l l i n i t y at - 40°C, fig. 13. At each heat input the h a r d n e s s and s t r e n g t h w e r e significantly reduced by the second t h e r m a l cycle.

3 . 2 . 2 . Metallographic examination

The s t r u c t u r e s produced by the double cycle simulation t r e a t m e n t s a r e shown in figs. 14 - 16. The 1300/1300 C t r e a t m e n t produced s t r u c t u r e s s i m i l a r to the 1300°C t r e a t m e n t , fig. 14 but with a r a t h e r c o a r s e r p r i o r austenite grain s i z e . The low heat input s t r u c t u r e , fig. 14a, was predominantly autotempered m a r t e n s i t e with v e r y little upper and lower bainite. With the medium heat input the s t r u c t u r e contained m o r e upper bainite but autotempered m a r t e n s i t e was s t i l l the main constituent, fig. 14b. At the high heat input the s t r u c t u r e was predominantly upper bainite with few a r e a s of autotempered m a r t e n s i t e , fig. 14c.

The s t r u c t u r e s produced by the 1300/950 C t r e a t m e n t a r e shown in fig. 15. The g r a i n s i z e was much finer than after the 1300/1300 C t r e a t m e n t and some

proeutectoid f e r r i t e was p r e s e n t in each s t r u c t u r e . The predominant microconstituent v a r i e d from a u t o t e m p e r e d m a r t e n s i t e at low heat input to upper bainite at high heat input.

The s t r u c t u r e s produced by the 1300/830 C t r e a t m e n t a r e shown in fig. 16. L a r g e a r e a s of f e r r i t e containing little precipitate w e r e p r e s e n t in each of these s t r u c t u i e s a s a r e s u l t of t h e r m a l e x c u r s i o n into the i n t e r c r i t i c a l r a n g e . The r e s t of the s t r u c t u r e s consisted mainly of autotempered m a r t e n s i t e for the low and medium heat input, figs. 16a and 16b, and upper bainite with some lower bainite and autotempered m a r t e n s i t e for the high heat input, fig. 16c.

3. 3 P o s t - w e l d heat t r e a t m e n t 3 , 3 . 1 . Mechanical testing

The effects of 1 hour post weld heat t r e a t m e n t s in the range 450 - 700 C on the impact p r o p e r t i e s of the c o a r s e grained HAZ a r e shown in figs. 17 - 19 for each heat input. The tensile and h a r d n e s s p r o p e r t i e s a r e included in Table 1. The r e s u l t s for each heat input follow the s a m e pattern. After 1 hour at 450 C t h e r e was a further e m b r i t t l e m e n t of the c o a r s e grained HAZ with impact values at - 40 C in the range

8 - 15 J (6 - 11 ft. Ibf) with a f r a c t u r e appearance of 90 - 100 % crystallinity. T h e r e was an a p p r e c i a b l e drop in h a r d n e s s and a slight reduction in strength.

With i n c r e a s i n g post-weld heat t r e a t m e n t t e m p e r a t u r e above 450 C the notch-toughness improved. After 1 hour at 550 C the impact values at - 40 C w e r e in the range 20 - 25 J (15 - 18 ft. Ibf) with a f r a c t u r e a p p e a r a n c e of 80 - 90 % crystallinity. After 1 hour at 650°C the impact values at - 40°C w e r e in the range 57 - 64J (42 - 47 ft. Ibf) for the low and medium heat inputs with a f r a c t u r e appearance of 50 - 55 % c r y s t a l l i n i t y . The improvement was l e s s m a r k e d with the high heat input after this t r e a t m e n t with c o r r e s p o n d i n g Charpy values at - 40 C of 37 J (27 ft. Ibf) and 80 % c r y s t a l l i n i t y . 1 hour at 700°C produced a c o n s i d e r a b l e improvement in notch-toughness with values r e s t o r e d close to parent plate l e v e l s .

The effects of t h e s e post-weld heat t r e a t m e n t s on the s t r e n g t h of the parent plate a r e shown in Table 1. T h e r e was little change in strength with heat t r e a t m e n t up to 650°C2 but after 1 hour at 700°C the strength of the parent plate dropped by about 3 tonf/in . The 0. 2 % proof s t r e s s was s t i l l , however, within the specified range. 3. 3. 2 Metallographic examination

The m e t a l l u r g i c a l s t r u c t u r e s produced by the post-weld heat t r e a t m e n t s were s i m i l a r for each heat input and so one s e r i e s of m i c r o g r a p h s only is shown, fig. 20, for the medium heat input. At 450°C and 550°C considerable precipitation of fine

Widmanstatten carbides occurred inside the martensite laths and coarse, mainly elongated, carbides precipitated at the lath boundaries, as shown in figs. 20a and 20b. The

embrittlement observed at 450 C coincided with the early stages of precipitation Above this temperature the improvement in notch toughness was associated with an advanced stage of precipitation. At 650 C the coarse lath boundary carbides grew at the expense of the fine Widmanstatten carbides and this was accompanied by a marked improvement in notch toughness although the strength level was still high compared to the parent plate due, presumably, to the continued existence of some fine carbides, fig. 20c. At 700 C most of the fine carbides had disappeared and the resulting structure consisted of coarse, mainly spheroidised. carbides in a ferrite matrix, fig. 20d, similar to the parent plate. This was accompanied by a further marked improvement in notch toughness with strength levels slightly above those of the parent plate. Table. 1.

4. DISCUSSION

The interrelation of microstructure and properties is an important factor in an investigation of the weld HAZ of structural steels. The simulation technique provides a means by which this can be assessed and the success of the technique will depend, in the first instance, upon reproducing accurately the structures present in the actual weld HAZ. Previous work, (Refs. 4 - 6) on QT 35 and other low allow steels has shown that the structures produced by the weld thermal cycle simulator are comparable to those occurring in the weld HAZ in terms of optical metallography and hardness tests.

The Charpy V-notch test provides a comparatively cheap and quick method for evaluating fracture toughness, and hence the resistance to brittle fracture, of the simulated structures under impact loading. The fracture toughness of different structures present in the weld HAZ can then be compared and the effect of welding variables and post-weld heat treatment assessed. The results of such investigations can then be used to formulate large scale test procedures on full plate thickness weldments, which are more representative of the practical situation and this may finally lead to improvements in welding procedures on full plate thickness weldments, which are more representative of the practical situation and this may finally lead to improvements in welding procedures. An equivalent programme based on the crack opening displacement test could be used to evaluate the behaviour under static loading where the notch toughness requirement will be quite different.

The successful performance of a welded steel structure in avoiding brittle fracture will depend upon the fracture toughness characteristics of the parent material and weld metal in addition to the HAZ. Low fracture toughness in any of these regions may result in complete failure of the structure by catastrophic brittle fracture, since each of these regions provides a continuous path along which such fractures may propagate. Two questions need to be answered; namely what level of fracture toughness is required for safe operation in a particular application and how is this fracture toughness to be measured. 4. 1 Fracture Toughness requirement for parent material

Pellini and Puzak (Refs. 18 - 22) have developed a successful procedure for the fracture - safe engineering design of steel structures based on a fracture analysis diagram. Fig. 21. This approach defines four critical transition temperature range reference points which also serve as design points

:-1. NDT (nil ductility transition) reference point

Restricting the service temperature to slightly above the NDT ensures that fractures do not initiate, even when small flaws are present, unless gross plastic over-loading occurs. In the event of a fracture initiating, resistance to propagation is low.

2. NDT to F T E (fracture t r a n s i t i o n elastic) m i d - r a n g e r e f e r e n c e point R e s t r i c t i n g the s e r v i c e t e m p e r a t u r e to above NDT + 1 7 C i. e. the m i d -range of the NDT and F T E region, provides fracture a r r e s t protection if the nominal s t r e s s level does not exceed half the yield s t r e s s .

3. F T E r e f e r e n c e point

R e s t r i c t i n g the s e r v i c e t e m p e r a t u r e to above F T E provides fracture a r r e s t protection if the nominal s t r e s s level does not exceed the yield s t r e s s .

4. F T P (fracture t r a n s i t i o n plastic) r e f e r e n c e point

R e s t r i c t i n g the s e r v i c e t e m p e r a t u r e to above the F T P e n s u r e s that only fully ductile f r a c t u r e i s possible.

The successful application of this approach to f r a c t u r e - s a f e design depends upon the c o r r e c t choice of c r i t e r i o n and will need to take account of flaw s i z e , s t r e s s level, and s e r v i c e t e m p e r a t u r e . Pellini and Sprawley (Ref. 20) have discussed the fracture toughness r e q u i r e m e n t s for s u b m a r i n e s t e e l s . F o r non-combatant s u b m a r i n e s the p r i m a r y r e q u i r e m e n t is r e s i s t a n c e to f r a c t u r e propagation at regions of tensile s t r e s s which o r d i n a r i l y involve points of i n t e r s e c t i o n . At such points the tensile s t r e s s e s may be considered to potentially attain, but not g r e a t l y exceed, yield point l e v e l s . T h e r e f o r e operation above the F T E t e m p e r a t u r e i s r e q u i r e d . F o r the c a s e of combatant s u b m a r i n e s the f r a c t u r e toughness r e q u i r e m e n t is much higher because the s t r u c t u r e m u s t be capable of withstanding underwater explosion attack i. e. operation above the F T P t e m p e r a t u r e i s r e q u i r e d .

The s i m p l i c i t y of this approach i s that only one r e f e r e n c e t e m p e r a t u r e , the NDT, has to be d e t e r m i n e d experimentally, since the F T E and F T P r e f e r e n c e points occur at a l m o s t constant t e m p e r a t u r e i n t e r v a l s above NDT. The NDT t e m p e r a t u r e can be determined d i r e c t l y from the drop weight t e s t or the explosion bulge t e s t and is highly reproducible. It has a l s o been shown that a good c o r r e l a t i o n e x i s t s between Charpy test r e s u l t s and the NDT t e m p e r a t u r e d e t e r m i n e d from drop weight or explosion bulge t e s t s . This c o r r e l a t i o n v a r i e s a c c o r d i n g to the type of s t e e l so that it is n e c e s s a r y to establish the c o r r e l a t i o n before it can be applied to a new type of s t e e l . Pellini and Srawley (Ref. 20) have established the c o r r e l a t i o n for HY-80 s t e e l , which is a s i m i l a r quenched and t e m p e r e d s t e e l to QT35, so that application of this c o r r e l a t i o n to QT35 m a y be considered to be justified. The c o r r e l a t i o n established was that at the NDT t e m p e r a t u r e the Charpy test gave values in the r a n g e 20 - 45 ft. Ibf. with an a v e r a g e value of 25 ft. Ibf. Since the use of QT35 s t e e l in s u b m a r i n e s t r u c t u r e s i s a c r i t i c a l one, in the s e n s e that failure would have s e r i o u s con-s e q u e n c e con-s , it i con-s e con-s con-s e n t i a l , a con-s far a con-s p o con-s con-s i b l e , to e l i m i n a t e any r i con-s k of b r i t t l e fracture and so a c o n s e r v a t i v e e s t i m a t e of the NDT t e m p e r a t u r e is n e c e s s a r y . This m a y be related to the 50 ft. Ibf. t e m p e r a t u r e in the Charpy t e s t . In any c a s e the choice between a 25 ft. Ibf. and a 50 ft, Ibf. t e m p e r a t u r e is unlikely to lead to any s e r i o u s differences in f r a c t u r e toughness r e q u i r e m e n t because of the s t e e p r i s e in the Charpy e n e r g y curve in this region. It was a l s o shown that for HY-80 s t e e l the F T E and F T P t e m p e r a t u r e s w e r e a p p r o x i m a t e l y 25 C and 50 C r e s p e c t i v e l y above the NDT t e m p e r a t u r e .

T h u s , the f r a c t u r e toughness r e q u i r e m e n t s for the two types of s u b m a r i n e d e s c r i b e d above can be stated in t e r m s of the Charpy values n e c e s s a r y to e n s u r e operation above the F T E o r the F T P at -10 C which is below the lowest t e m p e r a t u r e encountered during s u b m e r g e d operation. These a r e ;

-(a) 50 ft. Ibf. minimum at -35 C for non-combatant s u b m a r i n e s . (b) 50 ft. Ibf. m i n i m u m at -60 C for combatant s u b m a r i n e s .

The QT35 used in the p r e s e n t work satisfied both of these r e q u i r e m e n t s with a notch impact value of about Go ft. Ibf. at -60 C. However, c a s t s of QT35 s t e e l which just m e e t the m i n i m u m specified Charpy r e q u i r e m e n t óf 60 ft. Ibf. at -40 C would fail to meet the r e q u i r e m e n t for combatant s u b m a r i n e s , although the r e q u i r e m e n t for non-combatant

s u b m a r i n e s would be satisfied.

4. 2 F r a c t u r e toughness r e q u i r e m e n t for the weld HAZ

A s i m i l a r a t t e m p t to evaluate the behaviour of the weld HAZ for s e r v i c e conditions is much m o r e difficult. The f r a c t u r e a r r e s t c h a r a c t e r i s t i c s considered n e c e s s a r y for the p a r e n t m a t e r i a l will be r e q u i r e d only if f r a c t u r e propagation o c c u r s along the weld HAZ. On the b a s i s of s e r v i c e experience it s e e m s m o r e likely. however, that f r a c t u r e s initiating in the weld HAZ will propagate e i t h e r into p a r e n t m a t e r i a l or into weld m e t a l since the welds a r e n o r m a l l y of V o r double V g e o m e t r y and thus produce a HAZ slanted 45 to 60 with r e s p e c t to the usual s t r e s s vector d i r e c t i o n s . In c a s e s w h e r e f r a c t u r e s propagate into the p a r e n t m a t e r i a l it has a l r e a d y been shown that f r a c t u r e a r r e s t is likely to occur. The m o r e s e r i o u s mode of p r o -pagation would be into weld m e t a l . Although no a t t e m p t was made in the p r e s e n t work to evaluate the f r a c t u r e toughness c h a r a c t e r i s t i c s of the weld m e t a l it s e e m s unlikely that this would be sufficient to a r r e s t a moving f r a c t u r e so that it becomes n e c e s s a r y to p r o -vide protection against f r a c t u r e s propagating into the weld m e t a l . This m e a n s that the f r a c t u r e toughness r e q u i r e m e n t for the HAZ will have to be based on preventing f r a c t u r e initiation. Since s t r e s s - r e l i e v i n g is not n o r m a l p r a c t i c e for s u b m a r i n e s t r u c t u r e s , high yield level r e s i d u a l s t r e s s e s will be p r e s e n t in the weld HAZ and m a y initiate f r a c t u r e s due to s m a l l flaws in the weld regions for a wide range of nominal s t r e s s e s if the s e r v i c e t e m p e r a t u r e falls below NDT. Thus, a f r a c t u r e toughness r e q u i r e m e n t for the HAZ can be postulated based on operation above NDT since it i s u n r e a l i s t i c to c o n s i d e r a total absence of s m a l l flaws in this region.

With the p r e s e n t r e s u l t s it is not possible to state c a t e g o r i c a l l y whether the HAZ i s likely to be exposed to t e m p e r a t u r e s below its NDT although this s e e m s a distinct possibility. An i n c r e a s e in the NDT t e m p e r a t u r e of the p a r e n t m a t e r i a l of about 50 C would be sufficient to move the NDT t e m p e r a t u r e above the lowest s e r v i c e t e m p e r a t u r e . The single cycle simulation r e s u l t s showed an i n c r e a s e in the 20 ft. Ibf. Charpy

t e m p e r a t u r e of about 75 C for the c o a r s e grained region of the HAZ of single p a s s

s u b m e r g e d a r c welds made within the r e c o m m e n d e d heat input range and a further i n c r e a s e of about 20 C for a heat input above this r a n g e . More d r a s t i c i n c r e a s e s in the 50 ft. Ibf. Charpy t e m p e r a t u r e s o c c u r r e d , well o v e r 100 C for all heat inputs, due to the slow r i s e in Charpy energy with i n c r e a s i n g t e m p e r a t u r e for these s t r u c t u r e s .

The double cycle simulation technique used in the p r e s e n t work enables p r e -dictions to be made r e g a r d i n g the HAZ of m u l t i p a s s welds since the HAZ produced by a single weld pass will be expected to be t e m p e r e d by subsequent weld p a s s e s o r , in the c a s e of the final edge bead HAZ, by the tenaper-bead. This is m o r e r e p r e s e n t a t i v e of the p r a c t i c a l situation where m u l t i p a s s welding i s used. A multiplicity of combinations of t h e r m a l c y c l e s m a y be envisaged in the HAZ of m u l t i p a s s welds. In the p r e s e n t instance work was confined to an examination of the effect of a second t h e r m a l cycle on the c o a r s e grained HAZ s t r u c t u r e since this had the lowest notch toughness in s i n g l e -p a s s w e l d s .

The p r e s e n t r e s u l t s showed that in regions of the c o a r s e g r a i n e d HAZ w h e r e effective t e m p e r i n g from subsequent weld p a s s e s o r the t e m p e r - b e a d does not occur, the 20 ft. Ibf. Charpy t e m p e r a t u r e s a r e r a i s e d by m o r e than 80 C above p a r e n t m a t e r i a l levels at each heat input and the 50 ft. Ibf. Charpy t e m p e r a t u r e s a r e r a i s e d by well over 100 C. In s o m e c a s e s , i. e. for heat inputs within the r e c o m m e n d e d r a n g e , the upper shelf e n e r g y does not r e a c h 50 ft. Ibf. Such a situation will a r i s e in regions w h e r e the fusion boundary of a weld pass a p p r o a c h e s close to an existing c o a r s e grained HAZ and subsequent deposits of weld m e t a l may then be too far r e m o v e d from this region to c a u s e effective t e m p e r i n g . The final edge bead HAZ d e s e r v e s s p e c i a l attention since this r e g i o n contains the toe of the weld which may be a point of a p p r e c i a b l e s t r e s s concentration. Effective t e m p e r i n g in this region r e l i e s upon the a c c u r a t e positioning of the t e m p e r - b e a d . If the t e m p e r - b e a d

is positioned too close to the final edge bead HAZ no improvement in notch toughness is indicated, and if the temper-bead is positioned too far away little or no tempering will occur. Oldridge (Ref. 23) has shown that there a r e immense practical difficulties involved in maintaining accurate positioning of the temper-bead during deposition so that its efficiency in providing effective tempering in the final edge bead HAZ is open to conjecture.

In regions where effective tempering from subsequent weld passes and the temper-bead does occur a considerable improvement in notch toughness is indicated by the present results. In general the HAZ of welds made within the specified heat input range had notch toughness properties which indicated that the NDT requirement is satisfied whereas with heat input above the specified maximum of 2. 5 k J / m m (65 kj/in) there is some doubt whether this requirement is satisfied.

4. 3 Post-weld heat treatment

The philosophy underlying the post-weld heat treatment programme was to examine the possibility of improving fracture toughness in the coarse grained region of the HAZ without adversely affecting the strength of the parent material. Such a treatment would also reduce the high yield level residual s t r e s s e s present after welding and so improve resistance to brittle fracture initiation provided notch toughness is not adversely affected.

The results showed that treatment at 450 C for 1 hour reduced notch toughness in the coarse grained region of the HAZ at all heat inputs examined, due to the precipitation of fine Widmanstatten carbides within the martensite and bainite laths and coarser, mainly elongated carbides at the lath boundaries. This effect can be attributed to the secondary hardening characteristics of QT 35 steel due to the precipitation of molybdenum and vanadium carbides taken into solution during the weld thermal cycle. This secondary hardening

phenomenon has been reported in other steels containing similar quantities of molybdenum and vanadium. (Refs 6, 24 and 25).

With heat treatment temperatures above 450 C notch toughness improved as precipitation became more advanced with the coarse lath boundary carbides growing at the expense of the fine Widmanstatten carbides. The improvement was more marked in the HAZ of welds made within the specified heat input range and this was attributed to the superior notch toughness of tempered martensite compared with tempered upper bainite. Treatment at 650 C was necessary to give 20 ft. Ibf. and 50 ft. Ibf. Charpy temperatures less than 50 C above parent material levels and so indicate an NDT temperature below the lowest service temperature. This treatment also had no detrimental effect on the strength of the parent material. Raising the heat treatment temperature of 700 C produced a

further marked improvement in notch toughness with Charpy values restored close to o r ^ b o v e those of the parent material, but this was accompanied by a reduction of about 3 tonf/in in the 0. 2% proof s t r e s s of the parent material. J n the present instance this still met the minimum specification requirement of 36 tonf/in but this would probably not be true for casts of QT35 which fall near the bottom end of the strength range.

It is suggested, therefore, that adequate protection against brittle fracture initiation from small flaws in the HAZ of submerged arc welds in QT35 steels can be achieved by post-weld heat treatment of 650 C without adversely affecting the strength of the parent material. The same effect can not be guaranteed by relying on the uncertainty of tempering produced in multipass weldments or by the application of the temper-bead. There may be some practical difficulties in applying such a greatment to a large submarine fabrication, but this could be overcome, perhaps, by using the electric strip contact heaters used for preheating and maintaining a slowly decreasing temperature gradient away from the HAZ and into the unaffected parent material.

and since this m u s t be affected by post-weld heat t r e a t m e n t it is n e c e s s a r y to establish that this p a r t of the weldment i s not a d v e r s e l y affected by such heat t r e a t m e n t before it can be r e c o m m e n d e d a s a p r a c t i c a l p r o c e d u r e .

5. CONCLUSIONS

1. T h e r m a l c y c l e s experienced by the c o a r s e grained, fine grained, and i n t e r c r i t i c a l r e g i o n s of the HAZ of s i n g l e - p a s s s u b m e r g e d a r c welds in QT35 s t e e l c a u s e a m a r k e d drop in f r a c t u r e toughness a s m e a s u r e d by the Charpy V-notch impact t e s t . The g r e a t e s t l o s s o c c u r s in the c o a r s e grained region.

2. A significant i m p r o v e m e n t in f r a c t u r e toughness of the c o a r s e grained

region of the HAZ of s u b m e r g e d a r c welds in QT35 s t e e l c a n r e s u l t from t e m p e r i n g occuring in m u l t i p a s s welds. This i s shown, however, to be an unreliable method of achieving good f r a c t u r e toughness in a l l p a r t s of the HAZ.

3. The f r a c t u r e toughness of the c o a r s e grained region of the HAZ in both single and m u l t i p a s s welds in QT35 s t e e l d e t e r i o r a t e s with heat inputs above the 2. 5 k J / m n (65 k J / i n ) m a x i m u m specified for s u b m e r g e d a i c welding of this s t e e l due to the formation of s t r u c t u r e s containing a g r e a t e r proportion of upper bainite and with an i n c r e a s i n g p r i o r austenite g r a i n s i z e .

4. An e s t i m a t e of the f r a c t u r e toughness r e q u i r e m e n t s for s e r v i c e , based on the Pellini - Puzak f r a c t u r e a n a l y s i s d i a g r a m , indicates that the QT35 used in this work has adequate f r a c t u r e toughness in the unwelded condition for the most s e v e r e s e r v i c e conditions i. e. u n d e r w a t e r explosion attack at the lowest s e r v i c e t e m p e r a t u r e . C a s t s of QT35 s t e e l which just meet the specified Charpy r e q u i r e m e n t of 60 ft. Ibf. minimum at

-40 C fail to m e e t the f r a c t u r e toughness r e q u i r e m e n t for the most s e v e r e s e r v i c e conditions. 5. It is suggested that b r i t t l e f r a c t u r e s m a y be initiated from s m a l l flaws in the c o a r s e grained r e g i o n of the HAZ in QT35 s t e e l due to low f r a c t u r e toughness. This could lead to total collapse of the s t r u c t u r e if propagation o c c u r s along the HAZ or into weld m e t a l .

o

6. P o s t - w e l d heat t r e a t m e n t at 650 C for 1 hour is a m o r e effective and r e l i a b l e method of achieving good f r a c t u r e toughness in the c o a r s e g r a i n e d . r e g i o n of the HAZ of s u b m e r g e d a r c welds in QT35 s t e e l . Additional benefit would be derived from s t r e s s - r e l i e f ,

7. The HAZ of s u b m e r g e d a r c welds in QT35 s t e e l made within the specified heat input r a n g e (1. 2 - 2. 5 k J / m m ) has s u p e r i o r f r a c t u r e toughness after post-weld heat t r e a t m e n t at 650 C for 1 hour than those produced by higher heat inputs.

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14. DOLBY, R . E .

15. NISBET, T . H . M. , and KELLY, P . M .

Welding and F a b r i c a t i o n of QT 35 Steel, P r o c e s s Specification No. DG. S h i p s / P S / 9 0 1 4 , O c t . , 1964 QT35 Quality Steel P l a t e s , DG. Ships/5153B.

Production of high-strength tough s t e e l s at Consett, ISI/BISRA Joint Conference on Strong Tough S t r u c t u r a l Steels, ISI Publication 104, p . 9 3 Scarborough, A p . , 1967.

The weld heat affected zone s t r u c t u r e and

p r o p e r t i e s , o f QT 35 s t e e l , CoA Report Mat. No. 5. Cranfield Institute of Technology. Cranfield, M a r c h , 1968.

The weld heat affected zone s t r u c t u r e and p r o p e r t i e s of two mild s t e e l s . , CoA Report Mat. No. 2, Cranfield Institute of Technology, Cranfield, J a n . , 1968.

The weld heat affected zone s t r u c t u r e and

p r o p e r t i e s of two low carbon M n - C r - M o - V Steels, CoA Report Mat. No. 3, Cranfield Institute of Technology, Cranfield, F e b . , 1968.

The m i c r o s t r u c t u r e and notch impact behaviour of a welded s t r u c t u r a l s t e e l ; Weld. Jnl. , 45 (2), p . 5 5 - s , 1966.

Studies of the weld heat affected zone of T - 1 steel, Weld. Jnl. , 36, (12), p. 531s 1957.

Design and construction of a weld heat affected zone s i m u l a t o r . Met. Constr. , 1, (9), p. 427. 1969.

The heat-affected zone in a r c - w e l d e d type 347 s t a i n l e s s s t e e l . Weld. J n l . , 34, (6) p. 169s, 1955. Unpublished work, Cranfield Institute of Technology.

Cranfield.

A study of simulated weld heat-affected zone s t r u c t u r e s and p r o p e r t i e s of HY-80 low alloy s t e e l , D.A. E. T h e s i s , Cranfield Institute of Technology, Cranfield, 1969.

M e a s u r e m e n t of t h e r m a l cycles in the weld h e a t -affected zone of mild s t e e l , CoA Note Mat. No. 13, Cranfield Institute of Technology, Cranfield, Sept, 1967.

The effects of welding on the f r a c t u r e toughness of quenched and t e m p e r e d low alloy s t r u c t u r a l s t e e l s , Weld. Inst. Rep. C 2 0 1 / 1 0 / 6 8 , 1968.

E l e c t r o n Microscopy of QT 3 5 s t e e l in relation to its weldability. Report No. N152, Naval

Construction R e s e a r c h E s t a b l i s h m e n t , Dunfermline. Fife. May, 1962.

STEVEN, W . . and HAYNES. A. G. The t e m p e r a t u r e of formation of m a r t e n s i t e and bainite in low alloy s t e e l s , - Some effects of c h e m i c a l composition. J . I . S . I , . 183. p. 344. 1956 ANDREWS. K , W . E m p i r i c a l formulae for the calculation of s o m e

t r a n s f o r n i a t i o n t e m p e r a t u r e s , J. I . S . I . , 203, p. 721. 1965.

PUZAK. P . O . . ESCHBACHER. E , W . , and PELLINI. W.S.

PUZAK. P . O . . and PELLINI. W . S .

PELLINI. W . S . and SRAWLEY. J. E.

PELLINI, W. S. and PUZAK, P . P .

PELLINI. W . S .

OLDRIDGE, D.

WATKINS, B . , VAUGHAN, H. G. and LEES, G.M.

SAUNDERS. G . G . . and DOLBY. R . E .

Initiation and propagation of b r i t t l e fracture in s t r u c t u r a l s t e e l s . Weld. J n L . 31 (12), p. 561s 1952.

Evaluation of the significance of Charpy t e s t s for quenched and t e m p e r e d s t e e l s . Weld. J n l . , 35 (6). p. 275s 1956.

P r o c e d u r e s for the evaluation of f r a c t u r e toughness of p r e s s u r e v e s s e l m a t e r i a l s , N. R. L. Report 5609. June. 1961.

F r a c t u r e analysis d i a g r a m p r o c e d u r e s for the fracture - safe engineering design of s t e e l s t r u c t u r e s . N. R. L. Report 5920. March. 1963. Evolution of engineering p r i n c i p l e s for f r a c t u r e

-safe design of s t e e l s t r u c t u r e s , N. R. L. Report 6957. S e p t . . 1969.

The effect of a t e m p e r i n g bead on the s t r u c t u r e and p r o p e r t i e s of the weld HAZ in QT35 s t e e l . D. A. E. T h e s i s . Cranfield Institute of Technology, Cranfield, 1964.

E m b r i t t l e m e n t of simulated heat-affected zones in low alloy s t e e l s . Brit. Weld. J n l . , 1_3. (6), 1966, p. 350.

Effects of welding and post-weld heat t r e a t m e n t on the f r a c t u r e toughness of M n - C r - M o - V s t e e l s . Brit. Weld, J n l . , 1 5 , (5), 1968. p. 230.

( k J / m m ) P a r e n t t i t i f t TI 1.2 ?» Tt I t t l I I t l t t t t t l I t 2 . 1 " f t I t t l I t t l Tl I I f t 4 . 2 " t l t l f t t t ' I t l Tl (°C) p l a t e M Tt I t t t none IT Tl I T t t I t I t I t T l I t Tt 120 t l t t I t t t I t TT t l t t t l „ none I t I I I I I t 11 I t I t t l I t P e a k t e m p . (°C) _ -1.300 950 830 1,300 1,300 1.300 1.300 1,300 1,300 1,300 1,300 1,300 950 830 1,300 1.300 1.300 1.300 1,300 1,300 1,300 1.300 950 830 1.300 1,300 1,300 1.300 1, .300 1.300 1.300 P e a k t e m p . (°C) _ -_ -1,300 950 830 -_ _ . 1.300 950 830 -_ - -_ 1.300 950 830 _ _ . -h e a t t e m p . (°C) _ 450 550 650 700 _ _ -450 550 600 650 700 -_ _ _ _ 450 550 650 700 -Z -450 550 650 700 (MN/m^) 617 620 620 614 572 1.035 813 640 1,100 1.100 840 1,015 960 952 848 675 1,048 _ ..875 735 980 980 846 990 880 845 633 920 724 668 890 880 748 886 866 798 623 (tonf/in'') 4 0 . 1 4 0 . 3 4 0 . 3 3 9 . 9 3 7 . 2 6 7 . 2 5 2 . 8 4 1 . 6 7 1 . 3 7 1 . 5 5 4 . 5 6 6 . 0 6 2 . 4 6 1 . 8 5 5 . 1 4 3 . 9 6 8 . 1 5 6 . 9 . 4 7 . 8 6 3 . 8 6 3 . 8 5 5 . 0 6 4 . 2 5 7 . 2 5 4 . 9 4 1 . 1 5 9 . 9 4 7 . 0 4 3 . 4 5 7 . 9 5 7 . 2 4 8 . 6 5 7 . 6 5 6 . 2 5 1 . 9 4 0 . 5 (MN/m'^) 713 710 706 705 670 1.194 951 756 1.227 1.250 1,008 1.060 1,038 1,010 918 760 1,135 996 890 1,100 1,122 1,018 1,040 944 962 722 1.012 853 788 992 1,020 897 954 944 868 715 (tonf/in^) 4 6 . 3 4 6 . 1 4 5 . 9 4 5 . 8 4 3 . 5 7 7 . 6 6 1 . 7 4 9 . 1 7 9 . 5 8 1 . 4 6 5 . 5 6 8 . 8 6 7 . 3 6 5 . 8 5 9 . 6 4 9 . 4 7 3 . 8 6 4 . 7 5 7 . 8 7 1 . 5 7 3 . 0 6 6 . 1 6 7 . 6 6 1 . 2 6 2 . 5 4 6 . 8 6 5 . 9 5 5 . 4 5 1 . 2 6 4 . 4 6 6 . 3 5 8 . 2 6 2 . 0 6 1 . 3 5 6 . 4 4 6 . 4 area (%) 73 70 73 73 75 57 48 65 54 55 58 61 61 62 66 71 61 63 66 61 58 61 60 62 63 71 61 69 68 64 66 69 63 63 65 73 (HV5) 234 _ _ 232 224 384 349 304 378 397 323 346 341 335 299 242 378 323 291 35i3 352 318 340 331 293 243 346 292 256 327 320 287 319 313 291 248

( k J / m m ) P a r e n t " t l " I t 1.2 11 11 " " H " »t " " t t 2 . 1 " " I t t t " • 1 t l " " 4 . 2 t t •• " •• •• " ' 1 • • 1 1 p l a t e I t 11 I t IT none I t 11 • t n IT 11 " M M 1 1 120 TI 11 I t I t t l Tl I I t l none I I M Tl t l I t I t ' I " " P e a k t e m p . ( C) -1,300 950 830 1.300 1,300 1,300 1,300 1,300 1,300 1,300 1,300 1,300 950 830 1,300 1,300 1,300 1,300 1.300 1,300 1,300 1,300 950 830 1,300 1,300 1,300 1,300 1,300 1,300 1,300 P e a k t e m p . ( C) -1.300 950 830 -1,300 950 830 -— -_ -_ 1,300 950 830 -t e m p . ( C) -450 550 650 700 -_ -450 550 600 650 700 -450 550 650 700 _ _ -450 550 (J) 133 -133 162 19 52 60 16 24 54 15 24 35 57 150 19 44 67 15 26 39 11 22 65 141 11 50 57 14 15 23 8 20 650 1 35 700 117 t e m p . ( C) - 76 -- 76 - 87 + 20 - 52 - 58 + 56 - 14 - 52 + 54 - 15 - 30 - 56 - 134 + 12 - 42 - 65 + 22 - 10 - 38 + 32 - 6 - 62 - 112 + 30 - 52 - 58 + 32 + 8 - 10 + 40 - 2 - 36 - 88 temp. (°C) - 52 -- 52 - 62 + 28 - 20 - 32 + 64 - 10 - 32 + 60 - 2 - 15 - 35 - 112 + 27 - 12 - 32 + 32 + 2 - 15 + 4 0 + 6 - 40 - 82 + 42 - 10 - 22 + 38 + 34 + 20 + 52 + 22 - 14 - 60 energy (J) 163 -• 163 163 68 150 139 60 72 106 73 90 102 120 152 69 133 163 64 86 98 72 92 121 150 101 155 151 99 109 126 92 98 115 154

input 1.2 k J / m m (30 k J / i n ) . no preheat, (after Nippes et a l , Ref. 10).

2. Weld t h e r m a l c y c l e s adjacent to submerged a r c welds in 38 mm (1. 5 in) s t e e l plate. Heat input 2 . 1 k J / m m (54kJ/in). P r e h e a t 120°C. (after Smith. Ref. 11 and

Kellock. Ref, 12).

3. Weld t h e r m a l cycle adjacent to s u b m e r g e d a r c welds in 38 m m (1. 5 in) s t e e l plate. Heat input 4. 2 k J / m m (108 k J / i n ) , no p r e h e a t . (after Coward. Ref. 13),

4. Weld HAZ impact p r o p e r t i e s of QT35 after single t h e r m a l c y c l e s . Heat input 1. 2 k J / m m (30 k J / i n ) . no preheat.

5. Weld HAZ impact p r o p e r t i e s of QT35 after single t h e r m a l c y c l e s . Heat input 2. 1 k J / m m (54 k J / i n ) . preheat 120°C.

6. Weld HAZ impact p r o p e r t i e s of QT35 after single t h e r m a l c y c l e s . Heat input 4. 2 k J / m m (108 k J / i n ) . no p r e h e a t ,

7. QT 35 parent p l a t e - c a r b o n extraction r e p l i c a s .

8. C o a r s e grained HAZ s t r u c t u r e s produced by single t h e r m a l cycles with peak t e m p e r a t u r e 1300 C - carbon extraction r e p l i c a s .

9. Fine grained HAZ s t r u c t u r e s produced by single t h e r m a l cycles with peak t e m p e r a t u r e 950 C - carbon extraction r e p l i c a s .

10. I n t e r c r i t i c a l HAZ s t r u c t u r e s produced by single t h e r m a l cycles with peak t e m p e r a t u r e 830 C - carbon extraction r e p l i c a s .

11. Weld HAZ impact p r o p e r t i e s of QT35 s t e e l after double t h e r m a l c y c l e s . Heat input 1. 2 k J / m m (30 k J / i n ) , no preheat.

12. Weld HAZ impact p r o p e r t i e s of QT35 s t e e l after double t h e r m a l c y c l e s . Heat input 2 . 1 k J / m m (54 k J / i n ) , p r e h e a t 12G°C.

13. Weld HAZ impact p r o p e r t i e s of QT35 s t e e l after double t h e r m a l c y c l e s . Heat input 4. 2 k J / m m (108 k J / i n ) , no p r e h e a t .

14. S t r u c t u r e s produced by double t h e r m a l c y c l e s . Peak t e m p e r a t u r e s : F i r s t cycle 1300 C. second cycle 1300 C - carbon extraction r e p l i c a s .

15. S t r u c t u r e s produced by double t h e r m a l c y c l e s . Peak t e m p e r a t u r e s : F i r s t cycle 1300 C, second cycle 950 C - carbon e x t r a c t i o n r e p l i c a s .

16. S t r u c t u r e s produced by double t h e r m a l c y c l e s . Peak t e m p e r a t u r e s : F i r s t cycle 1300°C. second cycle 830°C.

17. Effect of 1 h r . post-v/eld t r e a t m e n t s on the impact p r o p e r t i e s of the c o a r s e grained HAZ in QT35 s t e e L Heat input 1. 2 k J / m m ( 3 0 k J / i n ) . no p r e h e a t .

18. Effect of 1 hr. post-weld heat t r e a t m e n t s on the impact p r o p e r t i e s of the c o a r s e grained HAZ in QT35 s t e e l . Heat input 2 . 1 k J / m m (54 kJ/in). p r e h e a t 120°C.

19. Effect of 1 h r . post-weld heat t r e a t m e n t s on the impact p r o p e r t i e s of the c o a r s e g r a i n e d HAZ in QT35 s t e e l . Heat input 4 . 2 k J / m m (108 k j / i n ) , no p r e h e a t .

Fig 1. Weld thermal cycles adjacent to arc welds in 38mm.(V5in) steel plate.Heat input V2KJ./mm(30KJ./in),no preheat.

{after Nippesetalio)

KOO

10 20 30 40

_so

TIME, SEC

Fig, 2 Weld thermal cycles adjacent to submerged arc welds in 38mm. (l-Sin) steel plate. Heat input 2=1 KJ/mm l54KJ/in). Preheat 120'C (after Smith'Oand Kellock")

1000

10 20 30 40 TIME, SEC.

Fig.3 Weld thermal cycles adjacent to submerged arc welds in 38mm. I1-5in) steel plate. Heat imput 4-2 KJ/mm.(108KJ/in)

No preiieat (after Coward'^)

la)

uo

11

_L

Fig.A Weld HAZ impact properties of Q735 steel after single thermal cycles. Heat input 1-2 KJ/mm.|30KJ/in). Preheat-none

100 IA « eo

I 60

Fig 5. Weld HAZ impact properties of QT 35 after single thermal cycles. Heat input 2-1 KJ/mm (5ilKJ/in).Preheat 120''C

FINE HAZ INTERCRITICAL HAZ COARSE HAZ _ 140 1 2 0 -M U .J W O ^ <9 80S 60 Ó. a. < X t o " 20 SO 60 ï^ tos >• 20 -140 -60 -20 20 TEMPERATURE C O

Fig. 6 Weld HAZ impact properties of QT35 steel after thermal cycles. Heat input 4-2 KJ/mm (lOSKJ/in)

Preheat-single none.

(b)

(a) Heat input 1.2 kJ/mm, no preheat x 4500

7 .t»,

(b) Heat input 2.1 kJ/mm, 120 C preheat x 4500

(c) Heat input 4.2 kj/mra, no preheat x 4500

Fig. 8 Coarse grained HAZ structures produced by single thermal cycles with peak temperature 1.300 C -carbon extraction replicas.

(b) Heat input 2.1 kJ/mm, 120°C preheat x 4500

(C) Heat input 4.2 kJ/mm, no preheat x 4500

Fig. 9 Fine grained HAZ structures produced by single thermal cycles with peak temperature 950 C -carbon extraction replicas.

\ *,

(b) 11<'.U iiipiit ;.'. I k-I/nini, 120 C.' preheat x 4500

(c) Heat input 4 .2 kj/mm, no preheat x 4500

Fig. 10 Intercritical HAZ structures produced by single thermal cycle with peak temperature 830 C -carbon extraction replicas.

100 8 0 6 0 -9 U i > • o. < 40 2 0 (b) 80 z _ l < 40 > u 201 O00/t>B*C 1300/950*C 1300'C 1300/1300*0 i l 4 0 1 1 2 0 | •UUQ» oe UJ 80 5 a.

«1

Fig. 11 Weld HAZ impact properties of QT35 steel after double thermal cycles.Heat input 1-2KJ/mm. (30KJ/in) Preheat-none.

140 —Hoo/eso-C) «oo/sso-c -1300*C ••«óó/isoo'c " 140 120!2 - I 1 0 0 -> o et 80 ë UI >• « o | u ' 4 0 PARENT HATERIAL 1300/1300'C 1300'C 1300/950*C 1300/830*C 20 80 60 > to 5 20 100 60 20 20 60 TEMPERATURE ( ' O 100 140

Fig.12 Weid HAZ impact properties of QT35 steel after double t h e r m a l cycles. Heat input 2-1KJ/mm.{5^KJ/in.) Preheat 120 C

100 — 80 0

>-e

<

/

/

/

I

i

Fig, 13 Weld HAZ Impact properties of QT35 steel after double thermal cycles.Heat input A-2 KJ/mm (108KJ/in) Preheat - none.

(b) Heat input 2.1 kJ/mm, 120°C preheat x 4500

(c) Heat input 4.2 kj/mm, no preheat x 4500

Fig. 14 Structures produced by double thermal cycles.

Peak temperatures ; first cycle 1300 C, second cycle 1,300 C carbon e>ctraction replicas.

;H-^Av|?; (a) Heat input l.!:! kJ/'nim, no p r e h e a t x 4500

(b) Heat input 2 . 1 k j / m m , 120 C p r e h e a t x 4500

^/j

(^) Heat input 4.2 k J / ' m m , no prelical., .\ 4300

F i g . 15 S t r u c t u r e s produced by double Lhermnl^^yclcs. ^^ P e a k t e m p e r a t u r e s : f i r s t cycle inOO C, second cycle 950 C.

Pr ^''^^' '

i > /

V

-(b) Heat input 2.1 kJ/mm, 120 C preheat x 4500

l^f^ss^

(c) Heat input 4.2 kJ/mm, no preheat x 4500

Fig. 16 Structures produced by double thermal cycles.

Peak temperatures ; first cycle 1300 C, second cycle 830 C. Carbon extraction replicas.

60 20 20 TEMPERATURE C C )

Fig 17 Effect of Ihour post heat treatment on the impact properties of the coarse grained HAZ in QT 35 steel. Heat input 1-2KJ./mm l30KJ/inJ Preheat-none. POST HEAT AT 650*C POST HEAT AT 550*C ' ' A S W E L D E D ^ .-••'TOST HEAT AT 450'C PARENT MATERIAL POST HEAT AT 700*C POST HEAT AT 6 5 0 ' C POST HEAT AT 550*C AS WELDED » POST HEAT AT 450*C Ui 100 q (9 UJ z •-|60 >-a: 40 5 u 20 SO - 6 0 z 40 2 u it) UO 100 60 20 20 40 » 0 140 TEMPERATURE C O

F i g . 18 Effect of 1 hour post-weld heat t r e a t m e n t s on the properties of the coarse grained H A Z in QT35 steel. Heat 2-1 KJ/mm (54KJ/in.) Preheat 120°C.

impact input

140 , POST HEAT AT 700'C POST HEAT AT 650*C POST HEAT AT 5 5 0 % POST HEAT AT 450" C AS WELDED — PARENT MATERIAL —POST HEAT AT 7 0 0 ' C POST HEAT AT 6 5 0 ' C POST HEAT AT 5 5 0 ' C » POST HEAT AT 450*C AS WELDED 140 1 2 0 M 111 . J O 100 3 >-o 80 g z 60 > 0. < 40 5 20 80 460 5 z 40 3 < > at 20 " 100 60 20 20 140 TEMPERATURE («C)

Fig 19 Effect of 1 hour post-weld heat treatments on the properties of the coarse grained HAZ inQT35steel. Heat i 4-2KJ/mm.(108KJ/in.) Preheat -none.

impact nput

Fig. 20

Effect of post-weld heat treatment on the coarse

o

grained HAZ - hoat input 2.1 k j / m m , 120 C preheat.

Carbon extraction replicas.

YIELD «o STRESS UI Ui (A V4 Y.S. V2 Y.S. Vt Y.S. FLAW SI2ES1 m i l m i l 1 1 1 1 1 1 I I n i l I 6 - 8 K S I 11 (STRESS LIMITATIONJV ELASTIC LOADS FRACTURES . / >r__fs 00 NOT ^^ /A ^ PROPAGATE • y ^ (TEMPERATURE LIMITATION) / / • y^ ^ C A T CURVE J I L J L J I L NDT N D T * 3 0 » F N 0 T * 6 0 * F TEMP ^ NDT»120'F