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Evaluation of necessary delay before inspection for hydrogen cracks

Richard Pargeter

Paper presented at 82 nd Annual AWS Convention, Cleveland, Ohio, 6 - 10 May 2001 and published in Welding Research Supplement of Welding Journal, vol.82, no.11, Nov.2003, pp.321s-329s.

(N.B. The AWS published version contains an error in Table 5, where the > and < signs are not shown. This version contains the corrected table.)

Introduction

In the welding of ferritic steels, the most common form of fabrication cracking is that caused by hydrogen embrittlement. It is well known that there can be some delay between the completion of welding and the formation of hydrogencracks in ferritic steels. Therefore, if inspection is carried out too soon after welding, these cracks may not be detected, with potentially catastrophic consequences. On the other hand, excessive delays after welding prior toinspection can have serious financial implications due to, for example, hold ups in production.

Currently there are recommendations for delays of between 16 and 48 hours in various national and industry standards (Refs. [1-5] ), but there is no firm basis for these times. Furthermore,there is generally no discrimination between different materials, joint geometries or welding conditions, with just one delay time recommended for all circumstances.

It was against this background that a programme of experimental work was initiated in which crack development was monitored using ultrasonic techniques in a variety of test welds.

Background

It is well established that hydrogen cracking in ferritic steels only occurs when a critical combination of the four basic factors involved is exceeded. These factors are:

  1. Hydrogen
  2. Susceptible microstructure
  3. Tensile stress
  4. Temperature

The likelihood of cracking increases with increasing hydrogen level, microstructural susceptibility, tensile stress, and as the temperature approaches about 20°C (68°F). In the final stages of cooling following welding,three of these parameters are slowly changing, as depicted schematically in Fig.1. Even when cooling has stopped, however, and the developed contraction stresses have stopped increasing, the hydrogen content will continue tochange. For the weld as a whole, the hydrogen level will decrease as a result of diffusion out of the weld metal and heat affected zone (HAZ) and the adjacent base material. However, at local points within the weld metal and HAZ,particularly those of high triaxial stress, the hydrogen content will increase for a period of time as a result of stress assisted diffusion. Hydrogen will diffuse up stress gradients from regions of lower to higherconcentration. Thus, it could be some time after completion of welding and cooling that, locally, the hydrogen concentration first reaches a critical value and cracking commences. This effect of hydrogen diffusion was noted as early as1961 when Beacham et al (Ref. [6] ) demonstrated that cracking in Lehigh restraint tests could be delayed by storing test panels at low temperatures. In their tests, they found that crackingoccurred between ¼ and ½ hour after welding under normal conditions. For similar welds quenched to -110°F immediately after welding, cracking was suppressed during storage at -110°F but did occur approximately¼ hour after reaching room temperature again. Hydrogen diffusion is therefore one feature that can contribute to the observed delayed nature of hydrogen cracking.

≤0.003%S, 0.0009%O). Carbon equivalent levels (0.38-0.45%) were selected to help with generation of cracking. One low alloy steel with SMYS of 690Mpa (100 ksi)(grade HY100) was used, and a limited number of tests were carried out on some 565MPa (82 ksi) yield HSLA steel (grade Q1N).

Table 1 Chemical compositions of parent materials

  Element %wt
C S P Si Mn Ni Cr Mo V Cu Nb Ti Al O N CE IIW Ca
Low sulphur C-Mn steel 0.19 <0.002 0.021 0.28 1.38 0.01 0.02 <0.005 <0.002 0.005 0.024 0.002 0.047 0.0004 0.0041 0.43 -
High sulphur C-Mn steel 0.18 0.037 0.018 0.45 1.57 0.04 0.02 0.01 0.002 0.06 0.047 <0.002 0.045 0.0035 0.0076 0.45 -
450N/mm 2 QT steel 0.09 <0.002 0.010 0.41 1.23 0.50 0.02 0.17 0.05 0.01 <0.002 0.004 0.030 0.0009 0.0055 0.38 -
450N/mm 2 yield steel 0.06 0.003 0.010 0.18 1.47 0.76 0.02 <0.005 0.002 0.23 0.015 0.014 0.058 - - 0.38 0.0012
HY100 0.17 0.002 0.008 0.28 0.28 2.89 1.59 0.51 <0.002 0.13 <0.002 0.004 0.024 0.014 0.0096 - -
Q1N steel 0.13 <0.002 0.009 0.23 0.29 2.88 1.25 0.4 <0.002 0.02 <0.002 0.003 0.019 - - 0.70 <0.0003


CE IIW = C + Mn/6 + (Cr + Mo + V)/5 + (Ni + Cu)/15

Shielded metal arc consumables used were E7018 for 350MPa yield steel, E8018G for 450MPa yield steel, E9016G for the Q1N steel, and E12018MM for the HY100 steel. An SD3 wire was used for submerged arc welding 350MPa yield steel, andan SD3 1Ni 1/4Mo wire for 450MPa yield steel.

Shielded metal arc consumables, with the exception of E9016G and E12018MM, were supplied in part-dried condition, thus allowing them to be dried to a required hydrogen level, between about 4 and 12 ml/100g deposited metal. A partdried basic agglomerated flux was used for submerged arc welding, dried as required to give between about 8 & 13ml/100g deposited metal. Typical weld metal chemical compositions are given in Table 2.

Table 2 Chemical composition of weld metals

  Element wt%
C S P Si Mn Ni Cr Mo Cu V Nb Ti Al
E7018 0.11 0.008 0.011 0.44 1.65 0.02 0.03 0.005 0.01 0.01 0.002 0.008 <0.003
SD3 Sub Arc 0.08 0.003 0.018 0.33 1.49 0.04 0.03 <0.005 0.20 0.002 0.005 0.002 0.012
SD3/Ni¼Mo/OP121TT 0.06 0.005 0.016 0.24 1.34 0.99 0.02 0.48 0.19 0.002 <0.002 0.004 0.014
E8018G 0.10 0.006 0.008 0.41 1.62 0.97 0.02 <0.005 0.03 0.01 <0.002 0.005 0.003
E12018 M 0.11 0.008 0.010 0.43 1.58 2.25 0.55 0.49 0.03 0.01 <0.002 0.006 <0.003
E9016 G 0.05 0.005 0.011 0.35 1.44 0.82 0.03 0.20 0.01 0.002 <0.002 0.029 0.016


The dimensions of welded groove and butt joint test panels are shown in Figs.2 & 3. In both cases, a good surface was required to facilitate automated ultrasonic inspection from the weld root side. The groove samples weremachined flat at the same time the groove was machined, and for the welded butt joint panels, the backing bar was machined off after completion of the root pass. Variations in the groove panel geometry to provide reduced restraint(through end and side slits), different groove depths, and different groove widths were incorporated in the programme. Heat inputs of between about 0.6 and 2.4kJ/mm were used for shielded metal arc welds in C-Mn steels and between 3.5and 5kJ/mm for submerged arc welds. For the tests on the low alloy (690MPa yield) steel, all welds were made at a heat input of about 1.1kJ/mm.

A summary of all the test series is presented in Table 3.

Table 3 summary of test series

Hydrogen: High Low Submerged Arc
Restraint: High Low High        
Groove Depth: 20mm 20mm 20mm   40mm    
Groove angle: 2x10° 2x10° 2x10° 2x30° 2x10°    
350 MPa yield, clean 0.8-1.1 0.8 0.56-0.8 0.8 1.1 5.0  
350 MPa yield, dirty     0.7-0.8        
450 MPa yield     0.7-0.8, 2.0-2.4     3.5  
HSLA (Q1N)     2.0        
Low alloy (HY100) 1.1 1.1 1.1   1.1    


Range of heat input (kJ/mm) employed given where tests were carried out.

Non destructive examination

Details of inspection equipment and techniques developed to some extent throughout the project, but the principals, which aimed to achieve constant and reproducible coupling, remained the same. For the groove welds, coupling wasachieved by immersing the specimen in an oil bath ( Fig.4). A pan of fixed probes, providing a combination of longitudinal and transverse pulse echo and time of flight diffraction examinations, was automatically traversed overthe reverse sides of the specimens. Test samples were immersed in the oil bath when they had cooled to about 40°C adjacent to the weld. The time of the first inspection varied, due to differences in weld cooling times and otherexperimental variables, between about 30 minutes and eight hours after completion of welding, generally being between 1 and 3 hours.

≤ 2 inch) thickness. Until proven otherwise by experiment, it should be anticipated that for significantly greater material thickness, longer delay time could berequired. The importance of hydrogen diffusion indicates that a further variable, ambient temperature, should be considered. Although the question of ambient temperature has not been addressed in this work, it is very likely thatlowering ambient temperature will increase delay times (Ref. [6] ). This should be borne in mind if welding is to be performed at temperatures below 20°C (68°F), and the possibilityof longer delay times should be allowed for.

Conclusions

  1. For C-Mn type steels of up to 450MPa (65 ksi) yield strength, delay times were found to increase with increasing heat input over the range studied of 0.7 to 5kJ/mm.

  2. Delay times for weld metal cracking when welding Q1N steel at a heat input of 2.0kJ/mm with the E9016G used, or similarly alloyed consumables, were similar to those for C-Mn steels.

  3. Significant delay times of up to 64 hours before the last initiation of a new crack and of up to 140 hours before all subsequent crack growth ceases have been recorded for welds in the 690MPa (100 ksi) yield strength steel. The longest time before detectable cracks were first produced was 21 hours.

  4. Delay times before cracking in both the 350MPa (51 ksi) yield C-Mn steel 1B789 and the 690MPa (100 ksi) yield HSLA steel 1B778 appear to be largely insensitive to the variables restraint, hydrogen level and weld thickness, within the ranges studied.

  5. Increasing the weld volume (cross section area) by changing the angle of preparation had no significant influence on delay times in a 350MPa (51 ksi) yield strength steel.

  6. Guidelines for delay times before inspection for hydrogen cracks have been produced. Seventy-two hours is recommend for 690MPa (100 ksi) yield low alloy steel, welded at <1.1kJ/mm heat input at 20°C (68°F) ambient temperature. Recommended delays for C-Mn steels are tabulated in Table 5.

Table 5: Guidelines for delay time before inspection for C-Mn steels of yield strength of up to and including 450N/mm 2 and up to 50mm thick.

Arc Energy, kJ/mm (kJ/inch) Heat Input, kJ/mm (kJ/inch) Delay time before inspection (at an ambient temperature of 20°C (68°F))
Observed
Greatest delay time for crack initiation (hrs)		 Proposed
Ultrasonic inspection (hrs)
≤3 ( ≤76)* ≤2.4 ( ≤61)* 4.7 12
≤3.5 ( ≤89)** ≤3.5 ( ≤89)** 12.3 24
3.5-5**
(89-127)		 3.5-5**
(89-127)		 16.5 36


*   For SMAW only
** For SAW only

References

  1. British Standards Institution. Welding - Recommendations for Welding of Metallic Materials - Part 2: Arc Welding of Ferritic Steels. BS EN 1011-2:2001

  2. American Welding Society. AWS D1-1:2000, Structural Welding Code - Steel, Miami Fla.

  3. DNV Rules for Classification of Fixed Offshore Installations, Part 3.

  4. Construction Specification for Fixed Offshore Structures in the North Sea. EEMUA 158, 1994 revision.

  5. National Structural Steelwork Specification for Building Construction, 3 rd ed. BCSA and SCI publication No.203/94, July 1994.

  6. Beacham, E. P; Johnson, H. H.; and Stout, R. D. 1961. Hydrogen and delayed cracking in steel weldments. Welding Journal 40 (4): 155s-159s.

  7. Andersson, B.A.B. 1982. Hydrogen induced crack propagation in a QT steel weldment. Journal of Engineering Materials and Technology 104(4) (October): 249-256.

  8. Troiano, A. R: The role of hydrogen and other interstitials in the mechanical behaviour of metals. Transactions of the ASM 52: 54-80.

  9. Bailey, N.; Coe, F. R.; Gooch, T. G.; Hart, P. H. M.; Jenkins, N.; and Pargeter, R. J. 1993. Welding steels without Hydrogen Cracking, 2 nd ed. Abington Publishing, 1993.

  10. Interrante, C. G.; and Stout, R. D. 1964. Delayed cracking in steel weldments. Welding Journal 43 (4): 145s-160s.

  11. Alcantara, N. G.; Oliveras, J. and Rogerson, J. H. 1984. Non-destructive testing in the fitness-for-purpose assessment of welded constructions. Proceedings International Conference, London Nov. 20-22. The Welding Institute.

  12. Hart, P. H. M. 1978. Low sulphur levels in C-Mn steels and their effect on HAZ hardenability and hydrogen cracking. International Conference on Trends in Consumables for Welding. London, Nov. 14-16. The Welding Institute.

  13. Böhme, D.; and Eisenbeis, C. 1980. Untersuchungen über die verzögerte Rissbildung am beispiel von querrissen im einlagigen unterpulverschweissgut von feinkorn baustählen. Schweissen und Schneiden 32 (10): 409-413.

  14. Juers, R. H. 1982. Determination of intra and post-weld hydrogen removal thermal soaking treatments for HY-130/MIL-14018 SMAW weldments. First International Conference on Current Solutions to Hydrogen Problems in Steels. Washington, D.C., Nov.1-5. ASM.