Weldability of High Strength Line Pipe Steels

Weldability of High Strength Line Pipe Steels 90 to 100° C preheat eliminates cracking at moderate levels of applied stress when Grade 483 line pipe is misaligned during welding

BY T. H. NORTH, A. B. ROTHWELL, A. G. GLOVER AND R. J. PICK

ABSTRACT. Full-scale weldability tests showed that Grade 483 line pipe material was resistant to cracking up to very high levels of general stress, in the absence of misalignment. When pipe misalignment was introduced a preheat of 90 to 100°C (194 to 212°F) was necessary to eliminate cracking at moderate levels of applied stress. Full-scale weldability and laboratory weldability test results correlated well when using slot testing and WIC restraint cracking tests. The implant test can give good indications of the susceptibility of a material to HAZ cracking, but its relevance must be questioned in situations where cracking occurs mainly in the weld metal. Introduction The field welding of high strength line pipe steels has been studied for many years, and numerous recommendations have been advanced concerning the determination of weldability and the specification of materials to ensure adequate field performance. Much of the recent steel development activity in the line pipe area has been specifically directed towards the improvement of hydrogen-assisted cold cracking resistance, a property upon which heavy demands are made by the still-prevalent stovepipe welding technique. After nearly t w o decades of development in this area, it may appear unlikely that anything new can be contributed to the subject. Those close to the pipeline industry, however, will be aware that the problem of cold cracking has far from disappeared; a number of recent, major projects have experienced serious difficulties of this kind. It is legitimate to ask

why such a situation still exists, and whether, despite the vastly increased fund of knowledge related to hydrogenassisted cold cracking, we are still approaching practical problems in the wrong way. One of the difficulties which arises in addressing the specification of suitably crack-resistant materials is the plethora of formulae which are proposed for the assessment of a material's cracking tendency. While many authorities agree that the IIW carbon equivalent formula is inadequate for modern, low-carbon steels (Ref. 1, 2), most codes and standards and —indeed —most company specifications still depend primarily on this expression. The additional consideration of different ranges of carbon content in some of the more progressive standards and specifications represents a major improvement. Again, a vast array of laboratory tests have been advanced as providing an assessment of the "weldability" of a material. While some such tests do combine the virtues of simplicity, economy and reproducibility, they are of very little use in the present context if it is not possible to relate the experimental variables and outcome to the way in which

cracking arises (and may be prevented) in the real world. Weldability tests that are able to simulate, in full-scale, the exact sequence of welding and manipulation which a pipe may experience in the field are the only ones in which this relationship is direct. Such tests are expensive and cumbersome, however, and only one pipeline company is known to use them on a production basis (Ref. 3, 4). Lastly, it is important to realize that the one "full-scale" test which is habitually carried out prior to production welding—the procedure qualification test — provides no protection against any but the grossest of inadequacies relative to cold cracking behavior. It is quite unreasonable to assume that the welding of t w o 500 mm (19.7 in.) long pups, under carefully-controlled shop (or at worst, yard) conditions will give any indication of the likely behavior of 24 m (78.7 ft) long double joints weighing, perhaps, over 10 tons, when welded and manipulated on a roller-coaster right-of-way at subzero temperatures.

It was with some of these perplexities in mind that the present program of work was begun. Initially, a theoretical stress analysis was undertaken in order to determine the sources and relative importance of stresses which could lead to cracking. In particular, a finite element analysis was used to relate local stresses Paper presented on the 63rd Annual AWS in different regions of the root pass to Convention in Kansas City, Missouri, during general bending stress arising from lifting April 26-30, 1982. and to assess the effects of weld geomeT. H. NORTH is Senior Research Associate, try. Stelco Inc.; A. B. ROTHWELL is ManagerThe second phase of the program Welding Technology, NOVA, An Alberta Corporation; A. G GLOVER is Manager-Industrial involved cold cracking tests on full-size Applications, Welding Institute of Canada; and samples of line pipe, using a purposeR. J. PICK is Professor, Mechanical Engineeringbuilt, four-point bending frame. Finally, a number of laboratory-scale tests were Department. University of Waterloo, Canada. examined to determine whether any log-

WELDING RESEARCH SUPPLEMENT | 243-s

1 ]

1 ST ROOTBEAD COMPLETED

3 RD

ROOTRCOT- I z BEAD BEAD ! o | y3 ICOMPLETED]! - JCCMPL£~Q 1 i/i I u~ 1 Z I

5 !

1

13 1

z 5

5 TH

4 TH ROOTBEAD

tCMPLETED

Q

0 z

ROOT- ! BEAD |

~c

ROOT-BEAD

COMPLETE!

O (min)

to cool into the cracking range also contributes to the diffusion of hydrogen out of the root-pass region; other results are the softening of hardened microstructures and the reduction of any local stresses resulting from pipe manipulation. Stress Considerations

5 MINUTE

HOT-PASS| I ICOMPLETEDI

I COMPLETE

HOT-PASS I I

[COMPLETE)

! 0 (min)

26

29

z

i

O I

" I

il l

35

38

Fig. 1—Schematic representation of root and hot pass welding operations

ical framework existed within which they could provide useful indications of field behavior. The steels examined, in the experimental part of the work, were typical of those currently being supplied by Canadian pipe makers for use in the construction of demanding pipeline projects such as the Alaska Highway Gas Pipeline System. Background Before turning to the analytical and experimental work, it is in order to review the basic factors which affect hydrogen-assisted cold cracking and the way in which they relate to pipeline field welding practices. There is a general understanding of the contributions which the material susceptibility, hydrogen and applied tensile stress make to the probability of cracking (Ref. 5), although the incorporation of this knowledge into quantitative formulations (analogous to those of fracture mechanics) has not yet been accomplished. It is also clear that hydrogen cracking will not normally occur at temperatures in excess of about 100°C (212°F). Role of Hardness There has been a widespread tendency to consider hardness as a reasonable indicator of susceptibility to cold cracking. In general terms, the prediction of hardness can be a relatively complex exercise, involving the detailed consideration of heat flow, of hardenability (the relationship between microstructure and cooling rate) and of the hardness of the different microstructures which can develop (Ref. 1, 6, 7). For the root-pass welding of relatively thick-walled line pipe by the stovepipe technique, the problem is somewhat simplified. This is because the relationship between heat input and wall thickness is such that very high cooling rates through the transformation range can be guaranteed. Consequently, weldability considerations are thus most reasonably based on

244-s | AUGUST 1982

the maximum hardness which can be developed in a given material. It must also be realized that cracking can take place in either the HAZ or the weld metal. As a result, pipe materials, which are themselves extremely resistant to cracking, are not always sufficient to ensure freedom from field welding problems. For this reason, laboratory tests which do not force cracking to occur in one region rather than the other are those which should relate most closely to field experience. Hydrogen Diffusion Large quantities of hydrogen ( > 4 0 mL/100 g deposited metal) are introduced into the weld region by the cellulosic-coated electrodes used for stovepipe welding. The extent to which this hydrogen can diffuse away from the weld is largely determined by the cooling rate through the low temperature range; this, in turn, is mainly determined by preheat temperature. Duren et ai. have shown that an increase in preheat from 20 to 100°C (68 to 212°F) leads to a decrease in diffusible hydrogen content from over 35 to « 1 0 mL/100 g (Ref. 8). The addition of a second pass, normally with a significantly higher heat input than the first, before the weld region has time

Stresses acting on the root pass before deposition of the hot pass could, in principle, arise from a number of sources, and the relative importance of these has not been well understood until quite recently. Bending stresses acting on a pipe due to lifting were considered by Mercer and Needham (Ref. 9), by Lumb and Fearnehough (Ref. 10), and, more recently, by Bufalini et al. (Ref. 11). In the first part of the present program, Higdon et al. (Ref. 12, 13, 15), using simple beam theory, derived the general bending stresses resulting from lifting, in relation to the lifting geometry. They then calculated the local stress concentrations in the root area, using a finite element analysis. They also considered other sources of stress, such as thermally-induced residual stresses and the restrained recovery of ovality when the line-up clamp is released. The main features of this work are discussed elsewhere in this paper. The important conclusion was that lifting stresses are likely to make by far the most important contribution to the maximum tensile stress acting on the weld; the sequence of welding and manipulation which takes place in a typical field-welding operation must be analyzed in this light. Stovepipe Welding The operations associated with the deposition of the first t w o passes in stovepipe welding, and their timing, have been considered in a recent publication (Ref. 14). They can be summarized as follows: • A pipe joint is supported at its center by a side-boom and brought up to the end of the already-welded string, where

POINT FORCE

\.\\

. \ n r ^ \

.

a

a

a

|

J SOLID FOUNDATION

Fig. 2 —Lifting of pipe

CRITICAL A R E A 'A'

800 | 700

Q.

2

600

LIFTING

500

LIFTING * OVALITY LIFTINGRESIDUAL LIFTING* RESIDUAL' OVALITY

m 400 UO

UJ

cc

300

I— t/1 _l




3A

20 105 120 150

Cracked Cracked Uncracked Uncracked

303 285 296 281

45

20 100

Cracked Uncracked

289 275

Pipe material

°5 kg load during hardness testing.

(Ref. 25-28). The reproducibility is, in general, not very satisfactory except at very high or very low levels of cracking. The discrimination capability of this criterion is thus poor in the intermediate range. In fact, in our o w n tests (Table 9), there was no clear-cut effect of chemical composition on cracking percentage. This difficulty is overcome by using the critical preheat to avoid cracking as the weldability criterion. As discussed later, this parameter relates directly to something that can be controlled in the field. As can be seen from Table 10 (and Fig. 10), there was a general increase in critical preheat temperature as PCM increased. Yurioka (Ref. 1) observed a similar trend when testing Japanese steels

and suggested the following relation: Critical Preheat Temperature (°C) = 718 PCM - 88 This relationship was derived when testing a range of steels with varying PCM values (from 0.124 to 0.358) using E7010 electrodes. In our results the relationship between preheat temperature and PCM was: Critical Preheat Temperature (°C) = 747 PCM - 58 This equation applies to slot testing using E8010 electrodes and incorporates a delay time of 5 min after welding. Within the accuracy of the present experiments ( = 3 0 ° C or 86°F steps in

The tests carried out to determine the critical preheat temperature to raise vcr to ffY for the three steels used in the fullscale investigation gave the results shown in Table 12. Critical preheating temperatures avoiding complete specimen rupture in 1000 min were 150, 120 and 150°C (302, 248 and 302°F) for steels 3A, 45 and 57. Although preheating temperatures up to 150°C (302°F) decreased the cooling rate in the 800 to 500°C (1472 to 932°F) range there was no clear-cut effect on HAZ hardness values. Critical preheating temperatures avoiding failure in a 1000 min holding period during implant testing were much higher than those required during full-scale testing. An obvious problem lies in the comparison of a long term (1000 min) implant test with a short term (15 min) full-scale testing situation. Preheating temperatures to avoid implant specimen failure during a 15 min test period can be estimated by interpolation as noted below, since there is a linear relation between preheat temperature and the logarithm of the time (t) to failure in implant testing (up to 1000 min): • Steel 3A preheat (°C) = 17.8 *i t(min) + 32

temperature

• Steel 57 preheat f Q - 1 6 . 1 Mm,n) + 44

temperature

• Steel 45 preheat (°C) = 12.8 In t(min) + 30

temperature

WELDING RESEARCH SUPPLEMENT j 253-s

CRITICAL RUPTURE STRESS 600

CRITICAL RUPTURE STRESS

N/mm2 600

500

N/mm^

500

•V. 400

400

300

300

•15

•17

•19

•21

•23

25

300

400

500

HAZ HARDNESS HV400 Fig. 12 — Relation between critical rupture strength in implant testing and Fig. 13 - Relation between critical rupture strength in implant testing and HAZ hardness PCM Correlation coefficients are 0.92 (steel 3A), 0.90 (steel 57) and 0.94 (steel 45). From these relations, the estimated preheat temperatures preventing implant specimen failure during a 15 min holding period were 80, 88 and 65°C (176, 190 and 149°F) for steels 3A, 57 and 45. Restrained Root Cracking (Schnadt-Fisco Test). The critical preheat values determined for the three steels tested in the full-scale program were 35, 22 and 35°C (95, 72 and 95°F) for the steels 3A, 45 and 57, respectively.

Discussion Full-Scale Testing As pointed out, the sequence of testing adopted in the present program represents a particularly severe situation. In particular, the sustained loading over a 15 min period is a very exacting test, and should represent a significant factor of conservatism in predictions derived from this type of testing. On the other hand, however, the loading by the hydraulic jacks was a relatively smooth, gradual process, while the manipulation of large masses by relatively insensitive equipment in the field can lead to significant inertial loads. These effects may be expected to offset each other to some extent, and the practical implications of the present program are thus likely to be reasonably representative of the circumstances arising during field construction. Certainly, empirical experience in Canada has led to the widespread adoption of preheat tem-

254-s | AUGUST 1982

peratures around 100°C (212°F) for the manual welding of these materials; when these preheats are conscientiously applied and carefully maintained, the incidence of cracking has been negligible. The cost of this practice can be estimated as lying in the region of $10 to $20 per weld. If this is compared with the cost of even a limited number of cut-outs, not to mention the quality implications of a high incidence of detected cracks, it must be concluded that a generous preheat represents a sound investment. It may be considered that special attention to hot-pass techniques, to ensure that the 6 o'clock region is completed prior to the application of the second lifting cycle, may be sufficient to avoid cracking difficulties. Such practices will certainly be helpful, but may still be insufficient, in the absence of adequate preheat. In particular, it should be considered that very severe conditions (heavywall pipe, high lift) may cause cracking in the last weld deposited, which cannot be protected by hot-pass deposition. In fact, unpublished work by Stelco on the welding of 18.3 mm (0.72 in.) WT, Cr 483 pipe at low heat inputs (