The Formation of Chevron Cracks in Submerged Arc Weld Metal
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The Formation of Chevron Cracks in
Submerged Arc Weld Metal
Chevron crack formation is a multi-stage process, and its
morphology is explained in terms of a model of crack
nucleation within shear bands intersecting columnar grain
boundaries
BY D. J. ALLEN, B. CHEW AND P. HARRIS
ABSTRACT. The cause of chevron crack- longitudinal-vertical sections —Fig. 1. In measures. Similar observations were
ing in multipass submerged arc weld plain view, the cracks are approximately made by Chitty and Brown (Ref. 4), who
metal has been investigated by welding transverse to the welding direction. suggested that the tendency to cracking
trials and fractographic studies. The Chevron cracks are partly intercolum- increased with section thickness. The
quenching procedure of Tuliani was used nar and partly transcolumnar. Near-verti- consensus view at the time was that
in a series of test welds in which the same cal "intercolumnar" components, run- chevron cracking was probably caused
flux and wire were used throughout, but ning mainly through proeutectoid ferrite by hydrogen, associated with the use of
the weld metal hydrogen content was and thus approximately following prior hygroscopic agglomerated fluxes. As a
varied by baking the flux. austenite grain boundaries, are linked by result, the flux manufacturing process
Chevron cracking was observed at short horizontal transcolumnar compo- was modified in 1971 by raising the flux
high hydrogen leveis, but cracking was nents. The intercolumnar and transco- baking temperature to the region of
entirely eliminated at low hydrogen lev- lumnar segments tend to alternate and 800°C (1472°F), which reduced the initial
els. Electron fractography demonstrated produce a stepped morphology on a moisture content and also reduced the
that both the intercolumnar and transco- microscopic scale, while maintaining the tendency to moisture reabsorption on
lumnar crack components were formed overall average 45 deg inclination to the exposure. As a result, chevron cracking
at low temperature. Thermal etching and vertical on a macroscopic scale. ceased to be a problem for some
surface smoothing were observed on Chevron cracking was first recognized years.
some crack surfaces, but these effects in submerged arc welded structural steel Unfortunately, in 1974-1975 a new
were shown to be caused by reheating fabrications in 1969 (Ref. 1, 2) and outbreak of chevron cracking occurred in
on deposition of subsequent weld beads occurred in several submerged arc weld- submerged arc welded oil production
and thus did not imply a high crack ments within the next t w o years (Ref. 3, platform components and heavy section
formation temperature. 4). This initial outbreak of chevron crack- engineering fabrications. According to
Ultrasonic and magnetic particle ing was associated with the introduction the flux manufacturer, the cracking coin-
inspection were used to detect chevron of a particular commercial agglomerated cided with a temporary increase in flux
cracks and the sensitivity of these tech- flux of high basicity. This type of flux moisture content associated with the
niques was determined by comparison produced cleaner weld metal and introduction of new drying equipment.
with optical microscopy. The morpholo- improved notch toughness compared Considerable confusion arose, because
gy and preferred locations of chevron with the previously used fused fluxes, but there were several instances when
cracking were investigated by optical was susceptible to moisture pick-up. hydrogen control measures such as flux
metallography. A new model of crack Hamilton (Ref. 3) found two types of drying and raised preheat and interpass
nucleation in shear bands is used to crack: short, isolated intercolumnar temperatures did not overcome the
explain the observations. cracks aligned with the proeutectoid fer- problem (Ref. 5). It is still not clear wheth-
rite, and larger "staircase" cracks com- er these measures were inadequately
posed of a series of intercolumnar cracks controlled or were simply insufficient.
Introduction
linked by transcolumnar cracks. He noted These difficulties and uncertainties gave
The occurrence of chevron cracking that the fractographic features did not rise to a burst of research activity, which
has only been recognized as a problem correspond with the traditionally ac- is reviewed below as a background to
for the UK welding industry within the cepted picture of hydrogen cracking, but the present work.
last decade. The cracks appear most nevertheless found that the problem
commonly in multipass submerged arc could be overcome by hydrogen control Review of Previous Research
mild and low alloy steel weld metals.
They are typically 0.25-3 mm (0.01-0.12 Tuliani (Ref. 6), in collaboration with
in.) long and may be recognized by their D. /. ALLEN. B. CHEW, and P. HARRIS are Taylor and Farrar of Southampton Uni-
characteristic stepped or staircase mor- Research Officers, Central Electricity Generat- versity, was able to reproduce chevron
phology and by their overall 45 deg ing Board, Marchwood Engineering laborato- cracking using a relatively simple labora-
inclination to the plate surface as seen in ries, Southampton, England. tory test. Cracking was obtained by intro-
212-s | JULY 1982
Chevron
i — crack
surface
Subsequent
room temp,
(ductile)-
fracture
Fig. I - Example of chevron cracking: A — cracking in a vertical longitudinal section (weld 7, section A3), X5.6; B - surface observed in SEM after
crack
breaking open at room temperature, X26.5 (Reductions as follows on reproduction: A - 12%; B— 14",,)
ducing a nickel-bearing welding wire (i.e., the Tuliani test procedure with a molyb- als, but long vertical cracks with few
electrode) and quenching the weld after denum-bearing instead of a nickel-bear- transcolumnar components were found
each run. This form of test has been used ing welding wire. This produced stepped, in higher strength weld metals. The verti-
in much of the subsequent work, includ- 45 deg angled cracks. However, unlike cal crack components of 45 deg cracks
ing the present research. Tuliani studied the cracks studied by Tuliani, these fre- were observed to be not exclusively
laboratory and industrial cracks by optical quently showed extensive fine branching. intercolumnar. Their fracture surface
metallography, scanning electron micros- Tests in which slices were cut from the morphologies were commonly quasi-
copy (SEM) and transmission electron weld after each bead was laid indicated cleavage; some possible indications of
microscopy (TEM). that cracks appeared in the lower runs thermal faceting were found but were
Optical metallography showed that only after several upper runs had been not considered significant as they only
intercolumnar crack components were deposited. This was rationalized in terms appeared in small areas. These regions
often of appreciable width but that trans- of the build-up of hydrogen and residual were apparently "smoothed" by reheat-
columnar components were always thin. stress during the course of welding, ing (Ref. 13). In associated work, Crouch
TEM examination of carbon replicas indi- assuming purely cold cracking. (Ref. 14) found that a high heat input
cated that the intercolumnar fracture sur- Farrar and Taylor (Ref. 8, 9) attempted promoted chevron cracking.
faces were generally smooth and feature- to avoid the possible complications of Wright and Davison (Ref. 5) made a
less, but showed evidence of thermal reheating effects by examining crack sur- comprehensive study using the Tuliani
faceting and grain boundary grooving. faces taken from the final bead. They test with SD3 1Ni wire and 28 commer-
This was taken to be definitive evidence found ductile regions and also smooth cially available submerged arc fluxes.
of crack formation at high temperature. areas on intercolumnar crack surfaces, These included fused and agglomerated
However, the cracks did not coincide with some evidence of solidification types and covered a wide range of basic-
with the dendritic solidification structure structure. Thus, their observations sup- ity. Twelve of these fluxes produced
as revealed by a selective etchant; also, ported the hot component hypothesis. cracked welds. Correlations were sought
they did not exhibit surface films or Hart (Ref. 10) studied examples of between cracking and flux particle size,
precipitates. chevron cracking from production and weld bead shape, flux and weld metal
Solidification and liquation cracking laboratory sources, using SEM. He compositions, flux moisture content, and
were, therefore, eliminated, and a "duc- observed features such as quasicleavage weld metal hydrogen content.
tility dip" mechanism involving decohe- and microvoid coalescence fractures. Flux moisture content and weld metal
sion of the austenite grain boundaries These were taken to indicate a hydrogen hydrogen content were found to be
was suggested. The transcolumnar frac- cracking mechanism. significant; a particularly clear correlation
ture surfaces, however, showed ductile Mota, Jubb and Apps (Ref. 11-13) stud- was observed between cracking and
shear dimples with no evidence of ther- ied chevron cracking in multi-run shielded hydrogen content. Of the fluxes which
mal faceting. Accordingly, they were metal-arc welds, using a water-cooled jig. initially gave < 10 ml NTP H 2 /100 g weld
considered to be a secondary interlinking Basic electrodes produced chevron metal, all but one produced crack-free
phenomenon occurring later at low tem- cracking at intermediate weld metal welds and the exceptional flux yielded a
peratures. The stepped crack morpholo- hydrogen levels. Severe microfissuring higher carbon equivalent weld deposit.
gy, and the considerable differences in occurred at high hydrogen levels, while Cracking was most prevalent with
width of intercolumnar and transcolum- cracking was generally reduced or elimi- agglomerated fluxes of high basicity.
nar segments, were cited as supporting nated when electrodes were baked to However, it was also observed with
evidence in favor of this two-stage yield low hydrogen levels. It was sug- fused and/or low basicity fluxes. It was
model of crack formation. gested that chevron cracks form at inter- also shown that chevron cracking could
Subsequently, Keville (Ref. 7) argued mediate hydrogen levels because higher be induced by wetting a flux which had
that the thermal etching effects observed hydrogen levels promote extensive fis- previously yielded sound weld metal.
by Tuliani could have been produced by sure nucleation, and thus reduce the However, a trial using the reverse proce-
the reheating (on deposition of subse- tendency for subsequent crack growth. dure of baking a crack-prone flux down
quent weld beads) of cracks which The cracks maintained an overall 45 deg to 8.6 ml H 2 /100 g did not eliminate
formed at low temperatures. Keville used orientation in medium strength weld met- cracking.
WELDING RESEARCH SUPPLEMENT I 213-s
Workers outside the UK have also would be expected to eliminate the cold metal-arc welding where the heat input is
observed cracks which, in the present cracking component, but would leave a lower. The expected trend of decreasing
terminology, may be recognized as chev- residue of vertical intercolumnar hot hydrogen level with increasing flux bak-
ron cracks. Killing and Orlikowski (Ref. cracking. However, if chevron cracking ing temperature is clearly observed in
15) found such cracks in submerged arc were purely cold cracking, it would be Fig. 2.
welds, observed a variety of cleavage- possible to eliminate both the vertical and Two tests to check for moisture regain
type fracture surfaces, and attributed the horizontal components of cracking and in the holding oven were included, since
cracking to hydrogen. achieve sound welds by means of hydro- it was realized that because of the time
gen control alone. required to make the cracking test welds,
Present W o r k the flux would have to be baked on the
day before welding. The results show
Summarizing previous work, welding Experimental W o r k that holding the baked flux at 150°C
trials have generally indicated that chev- (302°F) for up to 24 h does not cause any
Flux Baking Trials
ron cracking is a form of hydrogen- significant increase in weld metal hydro-
induced cold cracking in weld metal; It was necessary to carry out a prelim- gen content.
metallographic studies, however, have inary flux baking exercise in order to
shown several features which do not anticipate the likely weld hydrogen levels
Production of Test Welds
correspond with the traditional picture of associated with the cracking tests. A 200
hydrogen cracking. The choice of correct kg (441 Ib) stock of the agglomerated Eight welds were produced according
remedial action against chevron cracking basic flux OP41TT was obtained from to the procedure given by Tuliani (Ref. 6).
during fabrication is made difficult by this Oerlikon Electrodes Limited, which Two 380 X 150 X 25 mm (15 X 5.9 X 1
conflict of evidence. The present work agreed to withdraw the experimental in.) mild steel test plates were butt-
was aimed at resolving this difficulty by material from their standard production welded together using a 16 mm (0.63 in.)
testing the hypothesis that the interco- sequence before the normal final high- root gap and 20 deg included angle — Fig.
lumnar components of the cracks form at temperature bake. It should be noted, 3. The plates were pre-set so that, after
high temperature. therefore, that all observations of hydro- movement during welding, they became
The plan was to make a series of welds gen levels and cracking reported in this flat as required for ultrasonic testing and
with the same flux-wire combination, but work apply to this experimental batch of sectioning. The welding conditions,
to use a flux which had been withdrawn flux, and not to the regular production which were the same as those used in the
from manufacture before high tempera- grade of OP41TT. hydrogen determinations and in Tuliani's
ture baking, and to bake this flux at Samples of flux were each baked for work, are given in Fig. 2.
several selected temperatures to provide 1h at a series of temperatures ranging After each run, the assembly was left
a range of moisture contents and hence from 150 to 850°C (302 to 1562°F). to stand for specified times (following
hydrogen levels for the different welds. Weld metal hydrogen determinations Tuliani) of 5, 10 and 15 minutes (min) for
In this way, a series of test welds could be were carried out on single-pass sub- the bottom, center, and top third of the
produced in which weld metal hydrogen merged arc deposits using these samples joint respectively, to assist diffusion of
content would be the only variable and with Oerlikon SD3 1Ni wire according to hydrogen out of the deposit. Then,
all other composition and process vari- the method given in BS 639, but with a before the next run was laid, the joint
ables would be held constant. non-standard copper jig to clamp the was deslagged, quenched in water to
If the high temperature crack compo- base strips on which the deposit was laid. 20°C (68°F), and then carefully dried.
nent hypothesis was correct, a progres- This change was necessary because the The exact procedure was as set out by
sive reduction in weld hydrogen content BS jig is intended for use in shielded Wright and Davison (Ref. 5), except that
OP41TT "Special batch" 3 8 0 x 1 5 0 x 2 5 mm
mild steel base plates
* Baked as indicated
plus24h at 150°C SMAW
sealing runs
J. Spread of results (3)
O Mean value
Welding conditions -
500 A
30 V
6.3 mm/s
Thermocouple
hole 3mm deep
1 6mm root gap
250 350 450 550 650 750 850 450 x 50 x 12 mm
Temperature of lh bake, ° C mild steel backing
Fig. 2 —Hydrogen potential as a function ol flux baking tempera- bar
ture Fig. 3 - Details of weld test assembly, after Tuliani (Ref. 6)
214-s | JULY 1982
length of the w e l d p r e p a r a t i o n . The
Table 1—Weld Hydrogen Content and Cracking remaining runs r e q u i r e d t o fill the g r o o v e
(13-14 passes in all) w e r e laid only along
Weld half the w e l d length. The o b j e c t w a s t o
flux hydrogen,
CRACKS determine whether overlaying affected
baking
MPI crack f o r m a t i o n in the l o w e r runs.
Weld temp., 100 g
number °C weld metal Ultrasonics All Long(b)
Examination of Test Welds
2 440 11.2 49 174 7
3 530 7.6 (c) 124 2 O n c o m p l e t i o n , w e l d s w e r e left for
4 720 3.4 0 0 0 b e t w e e n one and t w o weeks before
5 370 16.6 34« 385 20 ultrasonic inspection w a s carried o u t . This
6 550 4.3 0 0 0 i n v o l v e d machining o f f the backing strip
7 350 17.7 60 2160 82 f o l l o w e d b y examination using Wells-
8 465 7.6 27(a) 155 10 Krautkramer U.S.M.B.2 e q u i p m e n t w i t h
9 350 18.5 39(a) n.d. n.d.< d '
C D 1 0 / 5 c o m p r e s s i o n w a v e p r o b e s (5
10 350 n.d.'d> 90< » a
n.d. (d > n.d.< d '
M H z ) and M A P 6 0 and 7 0 d e g , T M A P 45,
(a) Original ultrasonics data referred to non-standard weld sizes: these numbers have been scaled up accordingly for comparison 6 0 , a n d 70 d e g shear w a v e p r o b e s (4
with the other welds. M H z ) . T h e n , f o l l o w i n g Tuliani, w e l d s
(b) Appearing as > 2mm length in MPI records, and thus including cracks of true length > 1mm, see text.
|c) Ultrasonic inspection incomplete: see "Results" section. w e r e sectioned vertically along the cen-
(d) n.d. — not determined ters of each o f t h e t w o t o p beads. This
e x p o s e d f o u r longitudinal-vertical sec-
tions, labelled A 1 , A 2 , A 3 , A 4 in succes-
sion across the w e l d , w h i c h w e r e g r o u n d
Table 2—Chemical Composition of Filler Metal Wire and Weld Metal, %
flat and s u b j e c t e d t o magnetic particle
inspection (MPI).
Mn Ni Cr Mo Cu
T w o MPI techniques w e r e used — o n e
Wire
in w h i c h A C c u r r e n t w a s passed t h r o u g h
SD3 INi 0.11 1.95 0.39 1.56 0.18 0.09 0.024 0.027 0.38
Weld 2 0.10 1.51 0.35 1.37 0.12 0.06 0.014 0.027 0.23 the s p e c i m e n , the s e c o n d involving a
Weld 4 0.10 1.50 0.33 1.33 0.10 0.07 0.014 0.026 0.22 magnetic y o k e — t o d e t e c t transverse and
Weld 5 0.10 1.52 0.35 1.23 0.10 0.06 0.014 0.027 0.21 longitudinal defects, respectively. Perma-
Weld 6 0.10 1.50 0.38 1.37 0.12 0.06 0.016 0.027 0.23 nent records w e r e o b t a i n e d o n adhesive
Weld 7 0.10 1.56 0.37 1.34 0.10 0.07 0.015 0.026 0.22 labels placed o n the s p e c i m e n surfaces
b e f o r e energizing.
O p t i c a l m e t a l l o g r a p h y w a s used t o
those authors used slightly shorter h o l d exceeding 50 m m (2 in.). Eight t h e r m o - check the N D T findings a n d d e t e r m i n e
times b e f o r e q u e n c h i n g . couples w e r e placed at various depths the sensitivity of MPI, t o study the crack
Flux baking t e m p e r a t u r e s f o r each a n d locations t o c o n f i r m that an e v e n m o r p h o l o g y , a n d t o p r o v i d e quantitive
cracking test w e l d w e r e selected (using distribution of baking t e m p e r a t u r e w a s data o n the o c c u r r e n c e and size distribu-
t h e results of the baking trials) t o give achieved. T h e flux was sieved b e f o r e use tions of cracks in different beads a n d
target h y d r o g e n levels. D u r i n g t h e p r o - t o r e m o v e particles greater t h e n 2 m m microstructural zones. Electron f r a c t o g r a -
d u c t i o n of each w e l d , f u r t h e r h y d r o g e n (0.08 in.) in diameter. A f t e r baking, the p h y , using TEM examination o f c a r b o n
determinations w e r e carried o u t at the flux was s t o r e d o v e r n i g h t at 1 5 0 ° C replicas and SEM studies, p r o v i d e d fur-
beginning, m i d d l e , and e n d o f w e l d i n g . (302 °F) a n d used directly f r o m the tray t o ther i n f o r m a t i o n o n t h e fracture m e c h a -
These h y d r o g e n values w e r e averaged t o m a k e the test w e l d : only sufficient flux nisms. Sample chemical analyses (Table 2)
give a measure of the h y d r o g e n potential f o r o n e r u n w a s r e m o v e d f r o m the tray s h o w e d n o significant c o m p o s i t i o n a l dif-
associated w i t h each particular cracking at any given t i m e . The plate a n d cooling ferences (other than h y d r o g e n level)
test w e l d —Table 1. The actual h y d r o g e n w a t e r t e m p e r a t u r e s prior t o each w e l d b e t w e e n different w e l d s .
c o n c e n t r a t i o n s w i t h i n the multipass w e l d run w e r e r e c o r d e d , and the plate t e m -
can b e assumed t o b e different f r o m , but peratures immediately b e f o r e q u e n c h i n g Results
p r o p o r t i o n a l t o , the h y d r o g e n levels w e r e also n o t e d .
Incidence of Cracking
measured in single-pass w e l d s . W e l d s 9 a n d 10 d i f f e r e d f r o m the rest
T h e flux was b a k e d in 4 5 0 X 4 5 0 m m in that only six runs ( w e l d 9) or f o u r runs Cracks w e r e d e t e c t e d b y N D T in all
(17.7 X 17.7 in.) trays filled t o a d e p t h not ( w e l d 10) w e r e d e p o s i t e d along the full w e l d s e x c e p t n u m b e r s 4 a n d 6. W e l d 3
MPI Ultrasonics cf
2000 60
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o 40
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o
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1000 /
/ o /
o
/ E 20 - /
£ / 3
2
/
Z /
200 - 8-- S
0 -0-0- _L J
0 -o-a^- J
5 10 ~T5 55
5 10 15 20
[ H ] In w e l d m e t a l , m l N T P / l O O g d e p o s i t e d metal [ H ] in w e l d m e t a l , ml N T P / l O O g d e p o s i t e d metal
Fig. 4 — Cracking as a function of test weld hydrogen content: A - detected by MPI; B — detected by ultrasonics
W E L D I N G RESEARCH S U P P L E M E N T I 2 1 5 - s
1) and morphologies did not appear to
Table 3—Correlations Between MPI and Optical Crack Detection Showing That MPI change significantly. However, there
Overestimates Crack Size were considerable random variations in
the incidence of cracking. Detailed
Measured Number of cracks per studies (Ref. 16) showed a general
Observation crack size "apparent weld bead '(a)
absence of cracking in the top beads and
technique range 1 2 3 4 5
a marked fall-off in crack density towards
MPI > 2 mm 0 0 9 2 3 the sides of the weld; however, no other
Optical > 1 mm 0 0 6 2 4 systematic correlations were revealed
between cracking and position within the
MPI 1-2 mm 0 1 36 11 7
weldment. These studies also showed
Optical 0.5-1 mm 1 2 31 10 14
only minor variations in cracking density
MPI All cracks 0 2 56 15 30 between the reheated and as-welded
Optical > 0.5 mm 1 2 37 12 18 regions.
Optical < 0.5 mm, wide 0 1 29 9 11
Welds 9 and 10 provided information
Optical < 0.5 mm, narrow 1 26 82 55 290
on the effects of overlaying on crack
(a) "Apparent w e l d beads are those shown bv MPI — see Results" section formation. In weld 9, the uppermost
beads in the part-filled region were beads
5 and 6, while in weld 10 the uppermost
bead was bead 4. In weld 9, beads 5 and
contained a region of gross porosity Fig. 5B were the intercritical reheated 6 appeared uncracked in the half-filled
which was not examined by ultrasonics, regions, and not the fusion boundaries. region, and contained a few small cracks
but MPI revealed cracking in and around Hence, the apparent "beads" identifiable in the fully-filled region; beads 3 and 4,
the porous region. The results of MPI and in Fig. 5B did not coincide with the true however, were highly cracked, with no
ultrasonics were in agreement on all oth- weld beads. apparent difference between the half-
er welds —Table 1; they showed that the To assess the sensitivity of MPI, the filled and fully-filled regions. In weld 10,
severity of cracking increased with cracks on weld 5 section A3 were bead 4 contained several 2-3 mm (0.08 —
increasing hydrogen potential (Fig. 4) and counted and classified in terms of size 0.12 in.) cracks within the fully-filled
decreasing flux baking temperature. using the MPI magnetic yoke method, region, but was crack-free in the part-
Specimens from the A3 sections of welds and assigned to the relevant apparent filled region.
4 and 6 were thoroughly scanned for "bead" using the MPI AC current
cracks at a magnification of X250 in the records. These results were compared to
optical microscope after etch-polishing crack counts on the same section Crack Morphology
and etching in 2% nital. Two side bend obtained by optical metallography, The cracks showed all the characteris-
tests were also carried out on each of showing apparently poor agreement. tics generally attributed to chevron crack-
these welds and no cracks of any kind However, reasonable correlations be- ing. The stepped morphology was com-
were found. tween the t w o sets of results can be mon (Fig. 6) but not universal; some
The t w o MPI techniques gave comple- obtained if it is assumed that MPI overes- cracks were mainly intercolumnar, and
mentary results. While the magnetic yoke timates the sizes of cracks —Table 3. This there were also numerous microcracks
method (Fig. 5A) gave better chevron is plausible because the magnetic field (20-500 /xm) within the grain boundary
crack detection, the AC current method perturbation will be larger than the crack ferrite. The cracks were often clustered.
(Fig. 5B) also indicated the weld bead causing it. Thus, it appears that a 1 mm Microcracks often appeared in horizontal
sequence. Comparison with optical met- (0.04 in.) length crack is shown as ~ 2 bands —Fig. 1A. They ran along weld
allography showed that the dark zones in mm (0.08 in.) on the MPI records, and beads, apparently uninfluenced by the
that MPI detects all cracks over microstructure formed after reheating.
~ 0.5 mm (0.02 in.) in length plus many Intercolumnar microcracks were also
of the wider microcracks. Subject to observed in "staircase" arrays —Fig. 7;
these limits of sensitivity, therefore, MPI these were similar to full-length chevron
gives a reasonable measure of the sever- cracks but without transcolumnar inter-
ity of cracking. linking segments. Larger chevron cracks
Despite order of magnitude variations were often associated with microcracks
between welds in the total number of near their tips, again positioned in the
cracks, the crack size distributions (Table "staircase" orientation but not linked to
~i:..^:^':'-.y
. /
/
1
.
I Fig. 6 — Chevron crack morphology, showing
Fig. 5 — Typical cracking records obtained by intercolumnar and transcolumnar components
MPI (weld 2. Section A2): A — magnetic yoke with an unlinked microcrack at the foot of the II•
method; B — AC current method. X2.5 (re- "staircase." X77 (reduced 30% on reproduc- Fig. 7 —Unlinked microcracks in a "staircase'
duced 52",, on reproduction) tion) array. X75 (reduced 29% on reproduction)
216-s JULY 1982
In the as-welded regions, cracks were
stepped (Fig. 1B) and intercolumnar and
transcolumnar surfaces were readily dis-
tinguished. The intercolumnar segments
showed cleavage-type fracture surfaces.
These, however, differed from the frac-
tures produced at liquid nitrogen temper-
atures in that they were more broken up
into flakes, exhibited holes and protru-
sions, were more curved and had less
clearly developed river patterns— Fig. 10.
These features may be evidence of a
limited amount of ductility. Flat grain
boundary facets were not observed, indi-
Fig. 8 —A crack X arrested by the fine-grained Fig. 9 —Crack "melted off" at fusion boundary
reheated region, in line with a microcrack Y in above. X29 (Saspa-Nansa etch) cating that intergranular fracture did not
the coarse-grained reheated region, X22 (re- occur. The transcolumnar segments
duced 30% on reproduction) exhibited rough ductile shear fracture
surfaces —Fig. 10. Some cracks showed
temper colors when examined by eye,
Transcolumnar Transcolumnar indicating slight oxidation. These did not
appear very different from cracks with
non-oxidized surfaces when examined in
the SEM.
In the reheated regions (including the
intercritical regions), cracks were also
stepped and were often slightly oxidized.
The oxidized intercolumnar segments
were fairly similar to those observed in
the as-welded regions, except for some
fine-scale roughening on the cleavage
facets, which may have been caused by
transformation on reheating. By contrast,
non-oxidized intercolumnar segments
I*" were quite different to those observed in
the as-welded regions. Smoothed sur-
Intercolumnar faces appeared in patches of varying size
Fig. 10 — Fracture morphologies in the as-welded regions (SEM): A — X490;
-X2720 (A and . within most of the unoxidized interco-
reduced 34 % on reproduction)
the main crack —Fig. 6. the arrested main crack X, in the coarse-
In the as-welded regions, vertical crack grained reheated region.
segments generally followed the colum- Cracks tended to be confined within a
nar grain boundaries, but were not single bead. Many cracks ended at the
entirely confined to the proeutectoid fer- fusion boundary of the bead above and
rite. These crack segments were often had blunted tips —Fig. 9; this indicated
quite wide —Fig. 6. The overall directions that parts of these cracks had been
of the cracks appeared to depend on the melted off during deposition of subse-
columnar grain orientation — that is, when quent weld beads. However, cracks also
the columnar grains pointed forward often arrested close to the fusion bound-
along the weld, the cracks usually also ary as they approached it from above. A
pointed forward. However, the cracks few of the longest cracks did cross the
maintained a constant overall inclination fusion boundary and continue into the
of approximately 45 deg, despite marked coarse-grained reheated region of the
variations in columnar grain inclination. bead below, but these cracks were
The crack step spacing was of the same sharply reduced in width as they entered
order as the columnar grain width, the lower bead —Fig. 8. Thus, the longest
~ 5 0 - 1 0 0 )j.m. Thus, the columnar cracks observed in welds 2-8 were only
microstructure clearly exerted a strong slightly longer than the bead depth. A
influence on crack formation. few longer cracks, 10-15 mm (0.39-0.59
In the reheated regions, many cracks in.), did appear in the full-height regions
exhibited the stepped structure, and thus of welds 9 and 10.
appeared to have formed prior to
reheating. However, they were often
Electron Fractography
wider than cracks in the as-welded
regions, probably as a result of widening Several chevron cracks were examined
during the reheating cycle. There were in the SEM after breaking open at room
indications that delayed cracks could temperature or at liquid nitrogen temper- Fig. 11 — Fracture morphologies in the
have been arrested by the ductile fine- reheated regions (SEM): A - smoothed area
ature. In describing these observations, it
showing grain boundary grooving and indica-
grained reheated region — Fig. 8. A com- is necessary to distinguish between tions ot thermal faceting. X2640; b-
mon feature which is also shown in this cracks in the as-welded and reheated smoothed (prompt) and unsmoothed (de-
figure is the occurrence of a microcrack, regions, and between oxidized and non- layed) fracture areas in juxtaposition, X2620 (A
marked Y, directly ahead of the line of oxidized fracture surfaces. and B reduced 38",, on reproduction)
WELDING RESEARCH SUPPLEMENT I 217-s
low temperature —Figs. 10, and 12. O n
the other hand, the smooth surfaces of
Figs. 11 and 13 were found only on
non-oxidized cracks from the reheated
region. These appear to have been
formed by diffusional surface smoothing
of a cleavage-type fracture, and their
appearance can be attributed to the
effects of reheating. Slight oxidation
appears to pin the surface and prevent
surface diffusion, thus preserving the
original crack features during heating.
It can be shown theoretically that sur-
Fig. 12 — Fracture morphologies in the as-welded regions (TEM): A — intercolumnar cleavage. face smoothing, thermal faceting and
X2440; B - transcolumnar ductile dimples, X3200 (A and B reduced 29% on reproduction) grain boundary grooving can be pro-
duced in multipass welds during quite
moderate temperature reheat cycles, and
that the observed grain boundary groove
widths are consistent with peak reheat
lumnar crack segments (Fig. 11), typically and Davison (Ref. 5). In their tests, crack- temperatures of ~ 500-900°C (932-
covering half the area. However, the ing was usually eliminated at hydrogen 1652°F) (Ref. 17). The patchy nature of
remaining intercolumnar areas showed levels below 8.5-10.5 ml NTP/100 g. In surface smoothing on non-oxidized
sharp cleavage-type fractures. Mota's work (Ref. 13) crack-free welds cracks from the reheated region cannot
A view of a junction between were only obtained at < 3.7 and < 7.3 ml be due to variations in peak reheat tem-
smoothed and unsmoothed regions (Fig. NTP/100 g for SA and SMA deposits perature. This is because smoothed and
11B), demonstrates the real difference in respectively. unsmoothed surfaces appear only a few
appearance. Grain boundary grooving on These differences between workers microns apart (Fig. 11B), and differences
the smoothed surfaces was fairly com- could readily be accounted for by differ- in the reheat thermal cycle must have
mon—Fig. 11 A. There were also indica- ences in experimental procedure. Wright been insignificant. Therefore, it is neces-
tions of thermal faceting in many of the and Davison used much reduced heat sary to assume a combination of prompt
smoothed areas, but only a few rather inputs for their hydrogen determinations, cracking (providing cracks which are
coarse facets were clearly resolved by while the continuous cooling technique open during reheating) and subsequent
SEM. A comparison between the used by Mota could have resulted in delayed crack growth which produces
smoothed and unsmoothed intercolum- higher retained hydrogen contents. It unsmoothed fracture surfaces by crack-
nar surfaces suggests that their underlying should be remembered that hydrogen ing after reheating has occurred.
features are the same, but that all the determinations in single-pass welding In summary, the metallographic evi-
sharp surface features on the un- tests do not give more than a guide to the dence does not imply crack formation at
smoothed cracks have been rounded off actual hydrogen concentrations which high temperature and is entirely consis-
on the smoothed cracks. build up during multipass welding. tent with a low temperature hydrogen-
TEM examination of carbon replicas, The "safe" hydrogen level for avoid-
taken from the cracks discussed above, ance of chevron cracking must be
again showed unsmoothed intercolum- expected to vary in practice with welding
nar cleavage-type fractures and transco- procedure, restraint, and consumable *lt is emphasized that at present there are no
lumnar ductile shear dimples on and base metal composition.* The British or International Standards for the deter-
as-welded region cracks —Fig. 12. effects of these parameters on cracking mination of weld hydrogen in submerged arc
Smoothed surfaces were found only on welding. The data for hydrogen in weld metal
in the weld metal are much less well
given here should not be confused with values
replicas from reheated region cracks. understood than is the case in HAZ to be found in the technical literature pub-
Thermal faceting and grain boundary hydrogen cracking. There is a need for lished by consumable manufacturers, for
grooving were detected at high magnifi- more work in this area. which different methods of determination may
cation on some of these smoothed areas have been used. Where information on the
(Fig. 13) but were not found on uns- Fractographic Evidence drying process, hydrogen potential of consum-
moothed areas. Details of these features ables or welding procedure is required, users
are described elsewhere (Ref. 17). The cracks in the as-welded regions should seek advice directly from the manufac-
showed clear evidence of formation at turer.
Discussion
The Cause of Chevron Cracking
The original aim of this work was to
isolate any "hot cracking component" of
chevron cracking by progressively reduc-
ing the weld metal hydrogen content. All
cracking was eliminated by this means (as
shown in Fig. 4 and Table 1) and con-
firmed by careful optical examination and
bend-testing of the crack-free welds. At
hydrogen contents corresponding to test
values less than 5-7 ml NTP/100 g depos-
ited weld metal (attained at flux baking
temperatures of about 550°C and
above), cracking was eliminated in this
work. Fig. 13 —Fracture morphologies in the reheated regions (TEM), showing grain boundary grooving
Similar behavior was found by Wright and thermal faceting: A-X16600; B-X13800 (A and B reduced 28% on reproduction)
218-s | JULY 1982
assisted cracking mechanism. Maximum shear
Verfica stress directions
Crack Detection
Preferred
The comparison between optical and shear band
MPI examination techniques shows that Weld
direction
MPI is a reasonable tool for detecting direct
Columnar
chevron cracks over 0.5 mm (0.02 in.) grain
long on ground surfaces. However, since
boundaries
the surface beads contain few or none of
the cracks, MPI would be of little value in astic zone
the absence of sectioning. Ultrasonic rack
inspection is capable of detecting chev- Longitudinal
ron cracks in all weld locations, but tensile stress
appears to have an identification limit of
about 2 mm (0.08 in.) —Table 1. Because
of this, it might be feared that ultrasonic A ) Microcrack nucleation ( g ) Blunting and arrest
inspection could allow a heavily micro-
cracked weld to be passed as sound. This Dislocation
did not occur in the present work, since pile-up Shear band
microcracks were generally associated
with larger, ultrasonically detectable
cracks.
It should be emphasized that chevron
•&K-. Nucleation
cracks will not generally be detected
Direction W site of
unless an appropriate ultrasonic inspec- Crack 3
of shear
tion technique is selected, using 45 deg
angled probes to inspect the weld met-
al.
Crack Formation Sequence
Crack 1
The observations described earlier
may be used to relate the times at which /
crack nucleation and growth occurred to \Cj Crack formation within D) Macrocrack formation
the sequence of welding operations. Lit-
the shear band
tle cracking is found in the last bead to be
deposited. This is shown most clearly by Fig. 14 — Proposed mechanism of chevron crack formation associated with dislocation shear bands:
the observations on welds 9 and 10. It A—microcrack nucleation; B — blunting and arrest; C—crack formation within the shear band;
D — macrocrack formation
suggests that a critical requirement for
crack formation in a given bead is that
there should be one or more overlaying
beads. growth probably also occurs. The obser-
A Mechanism for Chevron Cracking
Few cracks are likely to be formed vations of smoothed and unsmoothed
during the initial quench. However, many fracture surfaces in juxtaposition The above observations give some
cracks definitely form prior to the reheat- (Fig. 11 B) directly demonstrate crack indication of the reasons for the stepped
ing cycle which takes place on deposition growth both before and after reheating. crack structure. However, a full explana-
of the overlaying bead. This is shown Thus, cracking appears to begin early in tion of the crack morphology requires a
by the observations of cracks with the welding sequence and continue dur- more detailed model.
smoothed surfaces within the reheat ing and after completion of the weld, The restraint provided by the cracking
region and by the common observation with individual cracks experiencing more test assembly, and the transverse nature
of cracks "melted off" at the fusion than one stage of growth. of the cracks, indicate that the highest
boundary above the bead containing the The way in which this discontinuous stress is in the longitudinal direction. O n
crack —Fig. 9. form of crack growth takes place may this assumption, Keville (Ref. 7) pointed
It is not entirely clear why cracks perhaps be inferred from the observa- out that the 45 deg crack orientation
should be largely absent from a bead tions of clustered microcracks —Figs. 6 coincides with the direction of maximum
which has just been deposited and and 7. These clusters appear to represent shear stress. However, she did not
quenched, yet appear in considerable intermediate stages in the formation of a advance a specific crack formation mech-
numbers before the bead experiences a macroscopic chevron crack. Thus, a anism. It will be shown here that a model
reheating cycle. It may be noted, howev- series of intercolumnar microcracks may (adapted from that of M o t a - R e f . 13) of
er, that there is generally one weld run first form in a staircase array as in Fig. 7, crack nucleation in shear bands intersect-
interposed between a given bead (e.g., and subsequently link up by transcolum- ing the columnar grain boundaries can
bead 7) and its overlaying bead (in this nar shear. Alternatively, a crack may account for the main morphological fea-
case, bead 9). One speculation is that grow by nucleation of a microcrack tures of chevron cracking.
cracks may form in bead 7 ahead of the ahead of the main crack tip as suggested The suggested sequence of events is:
arc while bead 9 is being deposited, in by Fig. 6, followed by transcolumnar 1. Longitudinal tensile stresses build up
regions which are then subjected to ten- linking back to the main crack. In either as the weld bead cools, and approach or
sile stresses but are not close enough to case, the appearance of the cluster sug- exceed yield point magnitude (Ref. 18).
the arc to be heated above hydrogen gests a discontinuous, multi-stage crack Plastic strain is initially concentrated in the
cracking temperatures. growth mechanism in which microcrack soft intercolumnar proeutectoid ferrite.
While many cracks form before nucleation seems to be aided by the When enough hydrogen is present, the
reheating, considerable subsequent crack presence of nearby cracks. ductility of the proeutectoid ferrite is
WELDING RESEARCH SUPPLEMENT | 219-s
much reduced, and so an intercolumnar readily explained in terms of the shear metal hydrogen cracking. The present
microcrack may nucleate (Ref. 19, 20) — nucleation model. Shear bands in one work provided no evidence to indicate
Fig. 14A. bead are unlikely to coincide with shear that hot cracking is even partially respon-
2. As the crack grows, the size of the bands formed in neighboring beads, sible. The thermal etching and surface
plastic zone at the crack tip will tend to owing to differences in columnar grain smoothing effects observed on chevron
increase. Hence the crack may be orientation and in the times at which crack surfaces are attributed to the
blunted and arrested. The plastic zone deformation takes place. Consequently, reheating of cracked regions on deposi-
may then form the nucleus of a long- cracks are often confined within one tion of subsequent weld beads.
range transcolumnar shear band at 45 bead. Larger weld beads would be 2. Chevron crack formation is a multi-
deg to the longitudinal direction (Ref. expected to contain more well-devel- stage process. The morphology of crack-
21) —Fig. 14B. (The reason why one oped shear bands; this may account for ing can be satisfactorily explained in
particular 45 deg direction is preferred is the correlation between chevron crack- terms of a model of crack nucleation
discussed later.) ing and high heat input observed by within shear bands intersecting the
3. Plastic flow within the shear band Crouch (Ref. 14). columnar grain boundaries.
then leads to dislocation pile-ups at the In Fig. 8, crack X in the as-welded 3. Cracks over 2 mm (0.08 in.) long
neighboring columnar grain boundaries. region is in line with a microcrack Y in the were detected by ultrasonic inspection.
The stress concentration at the head of coarse-grained reheated region. This can Since optical metallography established
the pile-up nucleates a second, separate be associated with variations in ductility that microcracks occurred only when
intercolumnar crack, either directly by within the weld bead. The fine-grained longer cracks were also present, the
causing decohesion at a weak zone such reheated region is too ductile to crack, ultrasonic technique successfully identi-
as an inclusion, or indirectly by promoting but the 45 deg shear band passes fied crack-free welds in this case.
further plastic flow in the proeutectoid through it to cause microcracking in the
ferrite. Hydrogen transport to the colum- coarse-grained reheated region. Acknowledgments
nar grain boundary by moving disloca- Mota and Apps (Ref. 12, 13) found a The authors are indebted to Mr. J.
tions assists crack formation (Ref. 13, transition from chevron cracking to "ver- Eariss, who developed the MPI tech-
22). tical" intercolumnar cracking with in- niques used in this work and carried out
4. The second intercolumnar crack in creasing weld metal strength. They sug- the welding program. Thanks are also
turn grows (upwards and downwards) gested that, because plasticity is limited in due to Oerlikon Electrodes Limited,
and blunts. In this way, an array of the harder materials, an intercolumnar which supplied the flux and wire elec-
intercolumnar crack segments forms by crack does not blunt and arrest and can, trodes for this investigation. The authors
repeated nucleation on columnar grain therefore, propagate along the columnar also acknowledge useful discussions with
boundaries within the shear band. As grain boundary without interruption. Dr. J. M. F. Mota and Professor R. L. Apps
each segment relaxes the longitudinal Mota noted that a decrease in the of the Cranfield Institute of Technology.
tensile stress on either side of it, the base amount of proeutectoid ferrite occurred
of crack 2 tends to arrest at about the together with the increase in strength and References
same height as the top of crack 1 — Fig. transition to vertical cracking. It seems
1. Thomas, S. N. C. 1969. Contribution to
14C. likely that proeutectoid ferrite plays a discussion: Conference on cracking in welds.
5. The crack array forms a zone of significant part in suppressing vertical Met. Const. I(2s): 142.
weakness. This promotes further plastic cracking by the blunting mechanism, thus 2. Cotton, H, C. 1969. Contribution to dis-
flow between the microcrack tips. Hori- promoting chevron cracking. cussion: Conference on cracking in welds.
zontal transcolumnar cracks therefore The model thus suggests that chevron Met. Const. 1(2s): I44.
form by ductile shear, assisted by hydro- cracking is a special form of weld metal 3. Hamilton, I. G. 1972. Trends in user
requirements for welding consumables. Inter-
gen. These cracks link the intercolumnar hydrogen cracking. Chevron cracking national conference on welding research re-
microcracks together to produce a mac- involves a significant amount of deforma- lated to power plant: 285-292. England: Univ.
rocrack —Fig. 14D. tion, while vertical weld metal cracking of Southampton.
This model successfully accounts for and heat-affected zone hydrogen crack- 4. Chitty, A., and Brown, I. M. 1972. Weld-
the 45 deg crack orientation and pro- ing do not. This suggests that there may ing consumables for the turbine-generator
vides a plausible explanation for the dis- be important differences in hydrogen industry. International Conference on Welding
connected, multi-stage growth behavior cracking behavior between weld metals Research Related to Power Plant: 293-309.
observed. and heat-affected zones. In a multipass England: Univ. of Southampton.
weld, the weld metal undergoes much 5. Wright, V. S„ and Davison, I. T. 1978.
Implications of the Mechanism more deformation during welding than Chevron cracking in submerged arc welds.
Conference on Trends in Steels and Consum-
Several features of the crack morphol- does the HAZ. Thus, HAZ hydrogen
ables for Welding: paper 38, London. Also
ogy can now be rationalized. The crack cracks form by a brittle mechanism in (abridged) in Met. Const. 11(3): 129-133.
orientation is clearly affected by micro- microstructural zones of high hardness, 6. Tuliani, S. S. 1976 (April). A metallograph-
structure as well as stress. There are t w o but weld metal hydrogen cracks may be ic study of chevron cracks in submerged arc
directions of maximum shear stress at 45 able to form in softer material via the weld metals. CEGB Report R/M/R234. Also in
deg from the vertical (Fig. 14B) and a relatively ductile chevron cracking mech- Welding Res. Int. 6(6): 19-45.
marginal difference in shear resistance anism. Thus, carbon equivalent formulas 7. Keville, B. R. 1976. An investigation to
derived from studies of HAZ cracking are determine the mechanisms involved in the
between these two different directions
unlikely to be directly applicable to weld formation and propagation of chevron cracks
could result in almost all the deformation in submerged arc weldments. Welding Res.
being concentrated in a single shear metal cracking.
Int. 6(6): 47-66.
band. The proeutectoid ferrite in the Chevron crack length is often re- 8. Farrar. R. A. 1977. The nature of chevron
present material forms easy shear zones stricted to the size of the weld bead. cracking in submerged arc weld metals. Weld-
along the columnar grain direction. Shear While it would clearly be unsafe to ignore ing Res. Int. 7(2): 85-89.
along these soft zones is, therefore, max- the possibility of longer cracks, this may 9. Farrar, R. A., and Taylor, L. G. 1977
imized by formation of shear bands in the be the reason why chevron cracking (Nov.). A metallographic study of chevron
45 deg orientation which is closest to the does not appear to have caused any cracking in submerged arc weld metals. Weld-
columnar grain direction. Thus, the cracks major failures of components in service. ing and Metal Fab., 45(9): 575-578.
also have this orientation. 10. Hart, P. H. M. 1978. Weld metal hydro-
Conclusions gen cracking, Weld. Inst. Res. Bull. 19(11):
The tendency of chevron cracks to 320-324.
arrest at the fusion boundary is also 1. Chevron cracking is a form of weld 11. Mota, |. M. F., |ubb, |. E. M.; and Apps,
220-s | JULY 1982
R. L. 1978 (Nov.). Chevron cracking: initiation Investigation of damage caused by the pres- Steels for Fast Reactor Steam Generators,
and propagation. Welding and Metal Fab. ence of hydrogen in submerged arc weld paper 78. London: BNES.
46(9): 625-627. metal. Schweissen u. Schneiden 29(8): 286- 19. Watkinson, F. 1969. Hydrogen cracking
12. Mota, I. M. F.; Apps, R. L; and lubb, |. E. 288. in high strength weld metals., Welding journal
M. 1978. Chevron cracking in manual metal- 16. Allen, D. I., Chew, B„ and Harris, P. 48(9): 417-s to 424-s.
arc welding. Conference on trends in steels 1980. The formation of chevron cracks in 20. Konkol, P. )., and Domis, W. F. 1979.
and consumables for welding: paper 18. Lon- submerged arc weld metal. CEGB Report RD/ Causes of grain-boundary separations in elec-
don. M/R295. troslag weld metals. Welding lournal, 58(6):
13. Mota, |. M. F. 1979. Chevron cracking 17. Allen, D. |. 1981. Thermal etching and 161-s to 167-s.
in steel weld metals. Ph.D. thesis. England: smoothing of weld fracture surfaces at high 2 I. lino, M. 1978. The extension of hydro-
Cranfield Institute of Technology. temperatures. Metals Technol., 8 (10), 395- gen blister-crack array in linepipe steels. Metall.
14. Crouch. S. |. 1978. The influence of 404. Trans.. 9A(II): 1581-1590.
heat input on weld metal transverse cracking. 18. )ones, W. K. C , and Alberry, P. |. 1977 22. Savage, W . F.; Nippes, E. F.; and Toku-
M.Sc. Thesis. England: Cranfield Institute of (lune). The role of phase transformations in the naga, Y. 1978. Hydrogen induced cracking in
Technology. development of residual stresses during the HY-130 steel weldments. Welding journal,
15. Killing, R., and Orlikowski, P. 1977. welding of some fast reactor steels. Ferritic 57(4): 188-s to 126-s.
WRC Bulletin 273
December, 1981
Design Implications of Recent Advances in Elevated Temperature Bounding Techniques
by J . S. Porowski, W. J . O ' D o n n e l l a n d M . Badlani
Recent advances in bounding {i.e., limiting) techniques and simplified methods of analysis for
components operated in the creep regime are used herein to obtain some very useful design guides.
Damage mechanisms are determined for a wide range of dimensionless design parameters, operating
pressure and cyclic thermal conditions, and material properties.
Publication of this report was sponsored by the Subcommittee on Elevated Temperature Design of the
Pressure Vessel Research Committee of the Welding Research Council.
The price of WRC Bulletin 273 is $10.00 per copy, plus $3.00 for postage and handling. Orders should
be sent with payment to the Welding Research Council, 345 E. 47th St., New York, NY 10017.
WRC Bulletin 274
January, 1982
International Benchmark Project on Simplified Methods for Elevated Temperature Design and Analysis: Problem II—The Saclay
Fluctuating Sodium Level Experiment; Comparison of Analytical and Experimental Results; Problem III—The Oak Ridge Nozzle to
Sphere Attachment
by H. Kraus
Problem II. Recently, experimental results became available on the second benchmark problem on
simplified methods for elevated temperature design and analysis: the Saclay fluctuating sodium level
experiment. These are compared to previously published numerical and analytical results in WRC Bulletin
258, May 1980.
Problem III. The Oak Ridge Nozzle to Sphere Attachment is analyzed by finite element computer
programs and by approximate analytical techniques. The methods are described and the results
obtained by each are compared. No experimental data are available.
Publication of these reports was sponsored by the Subcommittee on Elevated Temperature Design of
the Pressure Vessel Research Committee of the Welding Research Council.
The price of WRC Bulletin 274 is $10 per copy, plus $3.00 for postage and handling. Orders should be
sent with payment to the Welding Research Council, 345 East 47th St., New York, NY 10017.
WELDING RESEARCH SUPPLEMENT I 221-sYou can also read
