o I
r
welding
e ( J Pergamon
m
the Worldflx
Pnntcd
rg
m
Soudagc dan, Ie
Great Bntam All
’
Monde Vol 15 No 6 pp 17B.-1lj() 199S Copyright o It IIW/IIS
nghl’ 1’(-er.edl1mpnme en Grdnde Bretagne
Elxwer El%c ier S, Smemc ienle Lld Lid (1(141-==lihWS S9 51) .0(I(1
W43-22U(95)OOM-8
TMCP steels and their welding
by
N
/
S A
B)
Yunoka(Japan)>
1. Introduction
contmuous hot
features have ansen in the weldmg of TMCP steels They include HAZ hardness. HAZ softemng, cold crackmg. HAZ toughness, post weld heat treatment, weld metal toughness, and solidification cracktng All of these items are more or less related to HAZ hardenabtltty Therefore, this review discusses the items focussmg on HAZ hardenability The objective of this review is to md welding a
new
2. ’ TMCP steels
? I ’°&dquo; u of
T41CP °f’
temperature
temperature
(Ar,) The than that for the rolhng of
temperature conventional steels, refinement ofof the rolling of the refinement steels, which is the facilitates
the
microstruc-
Such mtcro-alloymg elements as molybdenum, mobium, vanadmm, and titanium expand the temperature range for austenue non-recrystalhzation Therefore, these elements are indispensable in the CR process and among them, mobium is mostly used tures of hot rolled
plates
It has been well known that the coohng rate dunng transformation governs the metallurgical microstructures of the transformed steels This concept was first realtzed m a RQ (reheat quenching) process m 1935 [I]J Then, RQT (reheat quenchmg and tempenng) was tntroduced, enabhng us to manufacture highly tough highstrength steels In 1962, accelerated coohng on run-on tables m a
successfully employed
was
m
the UK
the
TMCP (the rmo- mechanical control process) is a combined technology of CR and AcC This technology was first mtroduced on an mdustna) scale to manufacture steel plates m 1980 m Japan TMCP makes it possible to concurrently control not only the coohng rate but also the start and halt temperatures of water cool109 It thus can control the compostnonal ratto of the transfonnation products such as femte, pearhte, baintte and martensite at a desired level and produce steel plates with the preferable balance of strength and toughness Because of its capabihty of producmg steels with a flexible selecuon of steel mechamcal properties, TMCP has been able to manufacture steel sheets, shape steels and statnless steels other than the plates 2 2 Plate
produced by Plates of 490 MPa tensile strength (TS) grades §. MPa of AcC white(§§ of TS590 §3( A <- type, §§§ £p§£ plates # pro-
are . ,. _, ..y .., are some an TMCP duced by this process TS490 MPa grade plates metallurgically consist of femte-bamrte or femte-pearlite which are strengthened by the lower temperature transformation and higher contents matnx due to of carbon in a cooling can steels be reduced in TMCP elements °’ Therefore. alloy by an °Y amount correspondmg to the strength mcrement due to TMCP As shown 10 Fig 2, the carbon equivalent can be reduced in TMCP . that of of nonnahzed steels 10inthe from that steels by 0 04 to 00° 08% from normalized steels thesame same
femte
the accelerated
£’t§)&dquo;§ 1/§jf£fl§£i§fl)$) £?
£° Z? ° th bygrade 004 rade t[3]o [3] strength
metallurgtcal structures of TMCP steels are hIghly metallurgical highly refined as e of convent.onat m F.g wnh those conventional steets. steels as shown .n Fig 3 P &dquo;’&dquo;’ compared This results in a significant improvement in the base metal toughness For example,a Charpy fracture appearance transition temThe
in their carbon confor the stress avoidance of corrosion mainly cracking at the HAZ dunng operation In this respect, the use of TMCP is essenual for the production of hne-pipe steels. Moreover, HIC (hydrogen induced crackmg) other than the HAZ stress corrosion
tent
Weldmg and Jointing Research Centre m Japan 375
i
Perature (VT,,) goes down to -100°C 10 TS490 MPa TMCP teels for use in offshore structures This is an improvement of 60 C in vT&dquo; from normalized steels of the same strength level Steels for low temperature use such as LPG storage tanks need an arrest capability to prevent bnttle fracture from propaganng into a base steel As shown m the ESSO tests results of JIS SLA360 (RQT) and niobium beanng TMCP (CR-AcC) steels for LPG storage tanks in Fig4 [4], TMCP steels dtsplay considerable improvements in fracture arrest capability This presumabiy results from the refined grams transfonned from elongated ’ ° in m hot rolhng the austemte non-recrystathzatton temperatures
Line-pIpe steels definitely demand reduction
IlSf[IW- 1251-94 (ex doc IX-1739-94) recommended for publication by f[W Commission IX &dquo;Behaviour of metals subjected to weldmg&dquo; as the author’s own work N Yunoka is with the
This
solute
of hot I sows h t the h e cdasstficauon I a1 Ii IcatIon of hot rolling Figure shows Figure) rolhng processes for steelI in of the thermal term-, of the terms CR (controlledrolhng) plates in thermal history hi
was
ongma) technique of on-fine accelerated coohng (AcC) immediately followmg hot rolhng
[2)
TMCP (Thermo-Mechamcal Control Process), whtch is a process of controlled hot rolling followed by on-hne accelerated cooltng, was first employed to produce steel plates m 1980 Smce then. many types of steels, mcludmg a sheet steel, a shape steel and a stainless steel, have been manufactured by TMCP The accelerated cooling strengthens steels and can reduce the use of hardenable alloying elements. resultmg in hardenability reduction and thereby the improvement of weldablhty However. TMCP steels somewhat differ from conventional steels 10 chemical compostuon and they are sensitive to some types of heat treatments
stnp mill
’
,
RO
N
[A
i
O m
:;
Ar3
-
l
§
¡j E <.J 1)
8
Ac,
------- Ac,
Arl
RC)-T
1 Reheat Quenched Temper /ed
(Normalmed)
CR AcC
A9./ 3 ::·
Ar3 -
1 3 r,’ Ar AI Ar-;s:
Arl Ar
’k
1
-
N
(As Rolled)
1B1
T
A&dquo;B&dquo; A&dquo;B’3(’
AR
8.E (D§ p
--- -- AC3 AC3
-Ac3 AC3
Ar3 _
1B1
2u,
-.-
Arl
CR
CL
,
Ar r 3
A Arl
-
Al Art T T Ar53
T
Aci Aci
----
Ar,·
:.
, CR CR-AcC
CR
CR-DLO
CR-DO
DO-T
QuenchedI Controled Cooled Rolled (CR-Accelerated/Direct , - Tempered
I Thermal histonesof plate hot
Fig Plate Thick 50
to
f
550 -
500 -
N( Norm )
°’ 450 -
400 -400
.. a Z
i
TMCP
YS
35° 350
A j
T
____Ac Aci ----
:
.-: Arl-’LlfB’
DLO-T
(Lameller ,
Quenched
rolling processes
N( Norm ) N(Norm) N(Norm
possible to produce steels with low YR (yeld yeld strength [YSI/tensile strength [TS]) The low YR steel
TMCP makes it
a 300 -
rauo
! >
r3,; ,’
case is direct the with reheat conventional quenchmg (DQ) Compared quenchIng (RQ), the mechanical properties of CR-DQ steels are improved even though the cooling rate m quenchmg is the same m both processes This improvement anses from the refinement of martensite lath structures after transformation by DQ [6] Furthermore. precipitation hardening by copper, molybdenum niobium and vanadmm can be effectively undertaken m DQ-T In DQ, steels are retained at higher temperatures than In RQ, leading to the enhancement of solid solution of these alloy elements Then. the solute elements pr:’clpltate to strengthen steels dunng tempenng (T) after DQ
TS
TMCP
§
elements are on-line accelerated cooled This
100mm
600 -
£
,
AND THEIR WELDING TMCP STEELS
376
250 -
1
.32
I
I
1
36
I
I
I 40
1
I,
I. 44
CE=C+Mn/6+(Cr+Mo+V)/5+(NI +Cu)/15
Fig 2 Relation and °° &dquo; conventional
c n equivalent an between carbon equ
line-pipe
o stee strength steel strengt of
TMCP
metallurgically consists of a ferrite phase and a bainite phase formed by DQ from the femte-and-austenite dual phase zone (between Ar, and Ar, transformation temperature) This type of DQ is called DLQ In DLQ steels, soft femte phases first yield in
tensile testing while lowenng the yeld strength, and then hard bainite phases resist ductile fracture, elevating the tensile strength Thus, low yeld ratio (YR) results The low YR steelss are used in structural members of high-rise bmldmgs in earthquake prone countnes In order that earthquake energies may be absorbed by plastic deformation of steel members with low yield
steels carrying wet H¡S nch gas or oil The wet H2S environments enhance entry of hydrogen into strength [6] The HIC around steel tends inclusions to HIC steels, initiating propagate along segregation bands such as pearhte bands As shown m the metallurgical structures of Fig 3, TMCP steels are A normalizing (N) process is being replaced by TMCP (DQT) in metallurgically charactenzed by a small populauon of pearle producmg heat resistant pressure vessel steels of 2 25CR-1Mo and 3CR- I Mo types The satisfactory sohd solution of niobium is bands This is because carbon itself is reduced and because the accelerated cooling does not give carbon sufficient time to diffuse achieved by high temperature heating Then, the solute niobium to form pearlite Therefore, the HIC resistance property is slgOlfiprecipitates and finely disperses dttnng the temper treatment folin TMCP steels cantly Improved lowmg DQ The niobium precipitated DQT steels possess improved properties of the creep fracture strength [7]. Also, an of consist or martensite a Microstructures single phase a bainite- attempt has been made to employ TMCP to manufacture 9CrB IMo steels, one of the important femuc heat resistant steels [8]’ martensite mixed phase when steels with abundant hardenable
cracking
occurs In
t ; .
=
, -
378
t
TMCP STEELS AND THEIR WELDING
2 l Sramless steel AUlemtlc stainless steel,, are subjected to the carbides solid olution heat treatment for improvement of their corroson resistance
However, soltd soluuon
reduces steel strength and austenitic stainless steels are, thus, not appropnate for use as structural members TMCP techniques have been employed to strengthen stainless steels without reducing the corrosion teststance propentes Since austenitic stainless steels do not transform dunng hot rolhng, strengthemng by low temperature transforma-
products
uon
is not
500 -
However, controlled
rolhng
at
-
-
o 400
;
not
hot
0
-
E300300-B -
stainless steels 131
0o
including wide flange beams and ruls are formed 10 simple section shapes, and thus a uniform thermal control m rolltng and accelerated coohng is difficult However. TMCP been introduced to manufacture
&dquo;
B
steels [ l4] For employment of TMCP is desired to improve the toughness of a wide flange beam core where a flange and a web intersect In convenaona( hot rolhng, crystalhne grains coarsen because of the high hot rolhng fimshmg temperature due to the heavy secuon at the core. resulung in significant deterioration of toughness at the core portton
shape
t
’―― .*――――-** CE=0.37 64
*8
o
,
’C)
C CEE== 13’ 0 33
0
CE=036
cW
sb *
w
-
i
0
CE = 0 38
-
:
f¡
.
steels
as mstance,
- - - 43--e‘_ -’S―o**’’- 8 - - _- CE=0 -B
the
grain refinement causMoreover, accelerated coohng retards CI that a soltd soluuon treatment can be tic
;
o
Shape steel
Shape
+ 15
5
_ -
non-recrystalhzlOg temperatures facilitates
25
Ni+Cu
Cr+Mo+V
+ +----
treatment
expected
ing steel strengthening carbide precipitation so omitted for TMCP au stem
Mn
=C+=C+ 6
CE CE
,
-
CE
=
CE
=
_
0
30
__ ’***%*
200 -
0.28
,
_ -
the
I,
I
I
10 I 30 I 50 I I
I
,
100
130
I
i
I
I
I!
Bead Length Length (mm) (mm)
Fig6
Relation between bead
length and EtAZ maximum
hardness
: 1
171
: ,
i
Rails,
8
of pearhte microstructures with a eutecgeneral, toid carbon content of 0 8% Pearlite is charactenzed by the htgh abrasion resistance needed on the heads of rails Controlled cooltng is required after hot rolhng 10 order to obtain the full pearhttc structures with fine lamellar spaces In fact. rapid coohng results in martensite formation and. conversety. slow coohng forms coarsened pearhnc structures On-hne controlled coohng which utlhzes the latent heat of hot rolling has been developed for rail production [15] High temperatures from the latent heat can be obtained before coohng unltke reheat-and-coohng m conventional hot rolltng The employment of on-lme cooling results 10 rail heads with deeply hardened portions So far, the improvement of the resistance agamst abrasion and fatigue has been a pnmary concern for rail manufacturers But toughness improvement has been attempted, especially for rails for cold countnes. through controlled rolltng in the lower limit temperatures of an austenne consist
10
region [16] 3. HAZ
.
building However, those shorter than 50 mm but longer than 10 mm became perrrussible provided fabncators use TMCP steels of TS490 MPa grades wtth CE&dquo;w, (IIW carbon equivalent) not higher than 0 36% The hardness
was
Hv400 (Vickers hardness) and when ustng TMCP steels
still hmned
arc
stnkes
to
be
not more
than
allowed
even
were not
,
’
For the avoidance of cold cracking, the HAZ hardness is often limited to Hv300 or Hv350 in the welding fabncanon of offshore structures and ice breakers The maximum permissible HAZ hardness is 22 In the Rockwell C scale (Rc22), which is equtvalent to Hv248. for steel Plpehne weld HAZs to be subjected to moist H,S envtronments Nevertheless. HAZ is Itkely to harden when weldtng offshore structures and pipelmes, because low weld heat inputs have sometimes to be inevitably employed Normahzed steels have difficulty sausfymg the HAZ hardness limttatton, and consequently TMCP steels are exclusively used for offshore structures and line-pipes une-pipes
I
, ’
,I
j’ ,
I
hardenability and hydrogen cracking of TMCP steels 3 I 2 HAZ
1. tZ HAZ hardenublrr_v 3 ). hardenabiliv
hardenabtliiy and hardness estimation
I
A HAZ hardens 3.1 1Limitation
of HAZ hardness
For two years after 1983, the Japan Shipbuilding Research Committee (SR193) investigated the performance of TMCP steel welds, focusing on whether or not TMCP steel plates of TS490 MPa grades could be used for shipbuilding [17] As to HAZ hardness, the SR193 committee examined the relationship between the maximum HAZ hardnesses and the bead length. Figure 6 shows the expenmental results indicating that the cooling rate increases and thereby HAZ hardnesses increase with a decrease in the weld bead length The HAZ hardness tends to rise more rapidly m the lower carbon equivalent steels but their hardnesses remain at a lower level than those of the higher carbon equivalent
steels.
Weld beads
not
longer than 50
mm were
not
permitted
in
ship-
most at a region close to the fusion hne where the coarsening of austemte grams occurs and hardenability resultantly elevates Figure 7 shows the dependence of HAZ maximum hardness on the weldmg cooling rate or the weldmg coohng time between 800 and 500*C, r,s HAZ hardnesses decrease fairly smoothly with increasing t, in all the femtic steels including TMCP steels as shown m Fig 7 A HAZ becomes full martenstnc and its hardness levels off at the highest value in the cooling times of tL,5 shorter than the tm (point M) in Fig 7
.
The hardenability does not
necessanly descnbe the absolute level of hardness of quenched steels but it represents their Lkelthood to be martensittc or the facility of martensite acquisition. Therefore, the HAZ hardenabthty may be denoted by tM, the longest coohng time by which full martensite is acquired In other words, the longer tM is, the tugher HAZ hardenabtlity results tit! is given for hypoeutectoid steels (18]
-
TMCP STEELS AND THEIR WELDING
379
&dquo; -
600 .no
- M( tM HM) 7 x>
W.=--.... Hv=HB- 2 HM 2 20 HBerctanxl
―――-. 0
2
I = Q 4 , c-
log 2tM HV=---8rctBnIX)
0
A
400 4UU --
8 ° 0
300 300 -
_
Measured Measure A516 Gr 70 olate te
m
-
9 -
06 ,..-O6-
.=
7 7
i
-09-
m
HB) We)
B.
10 -
12 -
0
-t0- 0 .
0
,1
I
’
i
5’ 10o
’
,
change with
respect
to
i
o
i
0
0
-
°
-
Ot1
022
£-
6 -
I
oa4
033
U u
05
(%)
,
50
Coolnp Time Between 800 and 500t
7 HAZ hardness
*
-6 8
I
Area Occuped by Inclusiors
1
10 o
-6
’
’_
100 · &dquo;, i
II
T
!BB(ts 8(tB
-
Fig
L. BB - 06- !
tAA
D-
-
:: 10
04 -
2
X-4 oo tB - oe tM
,
_
t. , -
-
Lab -melt Steel
B9 -02-{B - -i
2-
11 §
Nb f V ree
0 -
HM)
500 -
1
1 S0 Mn i 014C-045m150Mn
too
Fig
50o
8 Effect of steel cleanliness
on
HAZ hardenability [24]
to(s)
weldmg coohng ume [) 8)1
depleuon occurs around the inclusion, resulung in the enhancement of carbon diffusion and femte transformauon around the inclusion
log t, =4
60 CEI - 2 08
(1)
presumed reason why steels with more sulphur harden their HAZ Whereas, niobium tends to co-segregate with manganese and retains m a manganese depleted zone, cancelhng the effect of manganese depletton It follows that femte nucleauon is not facihtated by inclusions In niobium beanng steels [25] The effect of tnclustons of HAZ hardenabthty is complex This less
Si
CEI = CP +-
Cu
Mn +
24
-
+
-
+
5 15
6
Ni
Cr(1 -0 16·ICr)
-+
+
122
+
8
is a
in
Mo
-
(2)
+AH
4
cp
= C (C
0 3),
Cp
=
C6
+
0 25 (C > 0 3%)
Equation (2)
is the carbon equivalent descnbmg the HAZ harden0I! is the hardenablhty increment due to boron and where abihty. steel cleanliness Boron significantly influences AH at a very low content It is reported that the boron effect is more significant m lower nitrogen steels [ 19] TMCP steels include, in general, reduced nitrogen for their HAZ toughness improvements There are some high strength steels which effectively uttlize boron to increase their strength together with the HAZ toughness improvement Hardenable elements are reduced in these steels, and therefore the HAZ hardenability mcrement by boron is designed to be cancelled
i
TMCP steels are improved in their base metal properties as well as in their weldability This improvement is attnbuted to the development of not only the process technology to control hot rolling and cooltng but also the steel making technology controlhng micro-alloy elements precisely Therefore, TMCP steels are generally very pure and clean It was reported that clean steels and especially low sulphur steels harden in their HAZs more than conventional steels [20-23] This is based not on expenments but on expenence Recently, the effect of steel cleanlmess on HAZ hardnesses was tnvesttgated using laboratory-melt steels with varying sulphur and oxygen contents [24]. Figure 8 shows the experimental results indicating that the effect of steel cleanlmess is obviously recogmzed in steels without niobium and vanadium but not in niobium beanng steels at a steel cleanlmess of 0 lo. It should be noted that 0.19’o steel cleanhness means intolerably duty for structural use The effect of steel cleanliness on the HAZ hardenabihty (Afr7 is considered neglIgIbly small for the normal structural steels, as long as oxygen is less than 50 ppm, and sulphur is less than 0 02%
Some inclusions act as femte nucleation sites dunng coohng after welding and reduce HAZ hardenability. Specifically, a MnS inclusion plays an important role in the femte nucleatton, MnS tends to precipitate on an existing inclusion and manganese
A prease formula to estimate HAZ maximum hardnesses is needed when determining the weldmg conditions to sausfy hardness
hrmtauon requirements or when designing the cherrncal compostnons of a steel to be welded under the hardness hmitation. Many formulas have been proposed [26-31 ] The author considers that a formula [ 18] denved under the detailed investigation of HAZ hardenability is the most reltable for a wide vanety of femuc steels Furthermore, this formula can be extended to estimate HAZ hardness after PWHT [32, 33] and weld metal HAZ hardness
[34]
3 13 Hardness
of resistance spot rselds
Resistance spot weldmg
usually employed
automobile sheet welding As a quality assurance test for spot welds. a tensile shear test and/or a cruciform tensile test are performed TSS (tensile shear strength) and CTS (cross tension strength) are given as a function of i (sheet thickness). d (nugget size) and a (base metal tensile strength) as follows TSS
= A d&dquo;
CTS where A and B
is
=
B
t 6
[35]
t d. [36]
m
(3) (4)
are constants
nugget of the appropnate size with respect to sheet thickness is formed, rupture m tensile shear tests always occurs outside the nugget and TSS predicted from Eq (3) is constantly When
a
achieved. This rupture mode
is called &dquo;plug failure&dquo; In the cruofonn tensile testing. CTS predicted from Eq. (4) is also obtained when plug failure takes place However, rupture often occurs inside a nugget. CTS is, m this case, much lower than that predicted from Eq (4) and concurrently, considerably scatters in its value It ts known that many factors including the welding condilion and the steel propentes (strength and compostuon) affect the scattenng of CTS values [35, 36, 37]
following empmcat equation of a carbon equivalent type gives the transition from the plug rupture to the in-nugget rupture The
[37-40].
-
.
,
380
AD THEIR %%’ELDING TNCP STEELS
Si
Mn
C+ - + -+?P+.lS?0_’r 30
A copper electrode function% weldmg Thus. spot welds cool
01
the
(5)1
20
500 as a
weldmg
heat sink
rapidly,
m
resistance
spot
developed by
tams
very low carbon and
ness
dtstnbuuon of
i
CTS
is
weld of this steel
as
shown
m
Fig
9
strength sometimes strength [46]
I
I
Low C-Cu
(873Kxx 60s)
1
A-.-A- *1’ -―r-i-)N*D’<.. ,
XL
)
200 -
Low C-Cu
(873Kx 60s)
Or-t 100 -
Nugget + I
Fig
I
I
0*0―0
!
Base Metal Base
HAZ
Nugget
044
..
I
II
I
i
066
08 8
Distance from Fusion Line
(mm)
2 02
4 04
02
0
9 Hardness astnbunon
In resistance
t
1 0
2 1
, spot welds [42..w1
TMCP (z5mm thick)
[44],
encountered
m
The SR 193 research group of the
Shipbuilding Research Insntute softening on tensile strength, resistance fangue strength, bending strength, and buckhng strength of the welded joints of TS490 MPa grades TMCP steels for shipbuilding [45]. It was clanfied that the weld joint tensile strength m small test pieces was 90% of the base metal strength the effect of HAZ
and the softened HAZ width was not more than the plate thickness under the weld heat inputs from 14 kJ/mm (submerged arc welding with three electrodes in tandem) to 61 kJ/tnm (electro-
o
n
<
<
:I: ,
flush-butt
becomes much lower than the base metal
Irput t39kj/mm
Heat
CEnw= 033 200 -
submerged arc welding. electro-slag weldmg weldmg may soften HAZs of TMCP steels which are strengthened by accelerated coolmg, as shown in Fig 10 [45] Flush-butt weldmg is employed to shop weld rails In a recently developed high strength rail whose head is strengthened by controlled cooling. HAZ softening takes place by flush-butt welding This softening can be recovered to some extent by air forced cooltng immediately after welding Flushbutt welding and DC (direct current) contact weldmg are also employed to produce wheel nms for automobtles The butt-weldedd wheel nms are cold-formed, the cold fonnablhty is affected by the weld hardness distnbution, especially by the width of softened HAZ [46] High strength hot rolled steel sheets of a dual phase type are also manufactured by one type of TMCP for use m automobile strength members This steel softens m its HAZ even m spot weldmg featured by rapid cooltng and the weld joint
slag welding).
B
180 >
and
investigated
c-rQ
I
:!
) (
is
substantially improved
high heat mput
I
/
:
type of
3 2 HAZ softening The
Medium C-Mn
1
300 -
superb m spot weldability This copper beanng steel is strengthened after press forming and weldmg by means of a copper precipitation hardenmg heat treatment of 600°C x 60 s or thereabouts Mamly because of the umfonn harda
1r _Medmm C-Mn
Mold steel!
TMCP alsoprecipitated steel has strengthmembers Thissteelconone
<
&dquo;*
T
_’ developed ply
steel has been
B_ r
B1i
III
precIpitated
/
400 -
An austenrte highly retaining steel is produced by TMCP This steel has been developed for the high strength members of automobiles Normally, this type of steel must contam high carbon for austenite stabilization However, m-nuggèt rupture with low CTS is more Itkely m spot welds of higher carbon steels The reduction in carbon is imperative for spot weldabihty Recently, an austemte retaining steel with reduced carbon as low as 0 20% has been [ 12] Post spot heating, that is additional current supafter spot welding, is an effective method to temper a hardened weld so that CTS propentes can be improved
A copper
0
3&
reponed as short as sheets [41 ] Thts rapid
r&dquo; is
ofI mm thick steel makes a coohng nugget and HAZ around the nugget fully martenstuc. whose hardness is determmed by the carbon content alone However. Eq (5) includes elements other than carbon This implies that the transition of nugget rupture mode is mfluenced m a complex manner not only by the weld hardness but also by the HAZ width (determined by HAZ hardenability) and hardness distnbution [42, 43] Figure 9 shows the hardness distnbution of spot welds of vanous types of steels [42. 43] 10
13,
t6o -
°
140 -
P
o
?
B
CD
P y
P
OD <
.
120 -Metal :7al -+
HAZ
Weld Metal
I
i
I
I
20
10°
0
10
,I
10
Distance from Fusion
Fig 1451
10 Hardness distribution of
high
-!t HAZ
Base
Metal -¡-:t:1
I
I
I
0
10
20
LIne (mm)
heat mput weld of TMCP steels
The SR research group conducted tensile tests on mde
plates
with softened HAZs as wide as the plate thickness The results revealed that the reduction of the tensile strength due to HAZ softemng is not recognized because of a mde plate constraint effect The mde plate tests more successively represent fracture in the actual steel construction than small size tensile tests It was thus concluded that HAZ softening is not detnmental m the welded structures of TMCP steels [45]
Concern has arisen over the occurrence of fracture of welded . structures at a low level of overall strains because of the strain concentration at softened HAZs Some researchers, from thts concern, advocated an overmatching joint in which the weld metal
strength is higher than that of a base metal [47].
In order to mvesthe effect weld of ugate overmatching, a fimte element analysis was conducted to clanfy the behavtour of CTOD (crack np open-
---
-
·
.
381J
TMCP STEELS AND THEIR WELDING
the surface notches on T-joints and butt joints subjected to high strain of 0 40rlr .i8 CTOD In this analysis is not a crmcj! CTOD value representing the metallurgical
disptjcement)
mg
at
’
toughness but
a
mechdnlcal value such
as
-- Test to:>
C
H)I i)7kJ/’m
elongation
’ M-3Smmth!et
of butt joints CTOD
case
was
03
mm
under the under-
match joints with a degree of overmatching (yeld strength of weld metal/that of base metal ) of 0 90, white CTOD was 0 15
under the overmatch joints with a degree of overmatching of means that overmatching is preferable However. the analysis of T-omts with higher inherent restraint than butt Jomts revealed that CTOD is 0 3 mm irrespective of the degree of overmatchmg Furthermore it is reponed that CTOD is not affectedd by the degree of overmatching at all but governed by HAZ toughness m the case of bnttle fracture without any plastic deformation 1491 Although the effect of the degree of overmatchmg seems to be not significant some degree of overmatchmg is desmed especially In butt joints mm
I 2 This
high high input shows thatresutts ofsoftening Ftgure tmplyng )nput weids affects TMCP hardly fangue .-.... ° .-, because so strength presumabty mostiy because fatigue governed presumably strength fatigue is governed mostly by shape a Jomt, Ie, reinforcement It was further revealed that HAZsoftening softening has remforcement It was further reveated that HAZ substanttal effect the bendtng strengths and the )) shows the results of I1 HAZ This Th
welds of TMCP
heat
Ftgure
stee steels,
_
a
y
ar
a
ts
y of a weld has no
the
resistance
on
e
buckhng strengths [45]
In order to take measures for HAZ soft.. , steel composition. ntobium addmon is effecuve the aemng from
.......r because of oe prectpUauon hardentng dunng coolmg after htgh because of its precipitation hardening dunng cooling after high heat input weiding Hv mcreases of fact, As aa matter matter of increases by fact. Hv by)IS5 at the welding As softened HAZ by the addition of 0 017r/c niobium under welding wtth a 23 kJ/mm with heat input kJ/mm heat 50] input [50]
3 33 Cold 3 Cold crocA:Mg crackrng
3 3.1 Cold cracking suscepnbrlrn of TMCP steels The
susceptibility to cold crackmg, one type of hydrogen crackwas investigated using B-groove self restraint cracking tests
mg, for the TMCP steels and the normalized steels of TS490 MPa grades This testing provides the critical preheat temperature, that is the minimum preheat temperature necessary to prevent cold cracking, as a parameter of cold cracking susceptibility of a steel The critical preheat temperatures found m the tests are plotted agamst CE,j, (an IIW type of carbon equivalent) in Fig 12 [51] The result reveals that the necessary
TMCP steels
is
preheat temperature
lower by approximately 100°C than that for
Symbol SYmbol TMCP Cornent 500 -
2 400 --
.
B BXQ
: 400 -
for nor-
Butt
25
0
.
Butt
20
Fillet
z5
200 -
=
<)
I
10’ 10 o
I
. I
I...Ii
o
- 100
――**―r-’*** ’.
30
.... ’
’,
’
25
,
’
’
-o
’I
,’ .
35
’,
1
o
’,
’ ..
’,
’,
.. ’
45
cE ac+,i6+(cr+eno+v)i5+(N+Cu)/15
Fig 12 Relation between critical preheat temperature against ing and carbon equnalent [51]
cold crack-
on of the reduction of the carbon TheSR)93 steelsThe TMCP steets the TMCP SR193 researchgroup group examequivalent In the equ)va)entm of weld ,ned ...... ,, metal..rf. dtffustble hydrogen and plate thickthe effect,. e ness on cold cracking in one-passhonzontal fillet welds The results are shown in Fig 13 [17] Cold cracks m one-pass fillet e the condition w,eldmg were not initiated
malmed steels enttrely because .....
rr
f....
results
are
hydrogen hydrogen ment ’ent
FIg 13 [17] Cold cracks
shown
under
m
one-pass fillet
of diffusible
mt/tOOg as high htgh as 2525mU100 g (bygtycennedtsptaceglycenne dtsplacethickness are not method),provided provtded CE&dquo;W CE.w and and the the p)ate plate thickness contents as
as
higher 0 36% and this and 25 mm, mm, respectively higher than 036% respecttvely OnOnthethebasis ofbasis ofthis of a ,. use j j was electrode r...... the result, high hydrogen approved provide ed thee CEil&dquo; of a stee osteel IS not higher than 0 36°k ThIs high hydrorate and features a electrode gen high deposition provides preferable concave fillet welds without undercut imperfections However, this type of electrode is not permitted to be employed for overhead fillet welding and mufti-pass weldmg, because the former is manoeuwed always with low heat inputs and the latter is liable to weld metal crackmg
.
,
...
,
·
30 30
S* 0 ng :Ao
0
0 0
g
o
A
0
*
’6 r .
25 N
0
0
,-.
O 0
0
0 Cf
o 0
0 A
g 15 - 36
&dquo;
)
or 8
/A
C#° A
O
e O · ·
O
·
6
6 Crack CE
43
No
m
W :f
&dquo;
ø
I
I
.
I..,.I
o° IDs
t... I
10’ to,
i....ii
10’ to’
NF (Hz)
I1 Results of tugh heat mput welds of TMCP steels
CES 0 36 0 36
o
F8t.- LIfe.
Fig
150
0
260 -
10 io
00 oo
-
0
0
0
° 0
5 -
CE S 38 36
36
, I... I’
0
Creck Crxk l/l/7T/tj’/77p 0 . cA A
No
―
-
100
co
cR
ozo c 50
/’
c r:
gJ,, ’. &dquo;’O &dquo; 0 200 I! 01 < ’ .&dquo;, B,c 0v - - :i i --r to._. -
...-
°
300-
-
* *O
TMep(cx-M:)-
S
Jomt Thick(mm)
1
r*’
ja 350 -
g
Norm f-Norm
5_10 TMCP Strb
20 ~CD 20
Plste Plate
0
450... i 400- -
. 0
0 .
BM<<
,i3000° --
Weid Weld
- 20°
T5 !IOOMPo Gr50
In the
(0
H I a 17kJ/mm 25 - 38mm thl&dquo;&dquo;
500 -
*
’
I
I
10
I
I
to
0
A A
43
-
I
I
20
Ftg13 Dependence of fillet weld cracking metal hydrogen and plate thickness [ 17]
30 on
carbon
I
I
-
40
equivalent.
weld
’
182
TMCP STEELS AND THEIR WELDING
125
’_ CoIJ c rucl.rng vusc eptibilir%
3 3
rnder
1521 published
100-
E
1§
m
ongmated
vi
=f0 a
preheat for TS490 MPa steels and TS785 methods based on the Pw-t, [58], Ctthe given t,m [61], and AWS D1-90 Appendix IX [64] cntena if their relevant carbon equivalents are the same The preheat given from the above cntena is too conservative for TS490 MPa grades and may be, contranly, insufficient or nsky for TS785 MPa grade steels BSS13S [65] is very convenient because necessary preheat temperatures are given in figures and tables However, this method is based on CEllw, and thus it is considered inappLcable for low-carbon. low-alloy steels As a matter of fact. BSS 13S is based on the premise of carbon steels and carbon-manganese steels but not low-alloy steels Recently. a chart method [66] has been proposed which is based on the CEN carbon equivalent This method considers the effects of steel composition, weld metal hydrogen, welding heat input plate thickness, weld metal strength and joint restramts Figure 14 is one example of the necessary preheat predicted by this method 3 33
8 -
same
is
in
Mull/-pass
weld metal
for conventtonal steels because carbon content is generally lower in weld metal than m steel and weld metal transforms mto femte before the steel HAZ does, resulung in the enhancement of
case
11 bl It.&dquo; Table Various types off carbon equivalent ,.....ity .. , of,steet , weidability 0
Formula
A
15
5e §g + hg..- + ;-u + L. + .r + hLo +-v59 &dquo;―c
17
60 7 23 16 9 8 CE(Gr8YIII.)-c+Mn_+k+M2+Nb+Y 15 C 1°* &dquo; C +k 25+V TO&dquo;IT+#+§+§b+ic W 20 ’ 40° ’ 6
18j
+ C pcw - c PC
B
15
BB 75C 50D
50 _
B
IOO,G 1oo*c
B
B 150ot 0 ’
B
B
B
125C
B
B
<
.
B
B
5’C 25 t
B. -
1
1
0
i
!
4 0
033
Carbon
1
066
0 5
¡
Equivalent. (CEN)
Frg 14 An example of predicted
necessary
preheat temperature [661
weld metal with lower hydrogen solubtltty to an austenite HAZ with higher solubtlrty On the other hand, TMCP steels are produced generally with low HAZ hardenabtltty, i e , their HAZ may transform from austenite to femte earlter than weld metal This situation is opposite to the case shown m Fig 1 S, and diffusion from a weld metal to HAZ thus dtrrumshes [68] Therefore, weld metal cracking night be more ltkely m TMCP steel weldmg than expected
hydrogen diffusion from a femte
,
;
methods reported to detennme the necessary preheat temperature for avoiding multi-pass weld metal cracks [69-71] The following is one of these [71] There have been
Tp (°C)
some
=
0 524 oB
+
277
(6)
log Hoc - 482
,
where 6B is the weld metal tensile strength (MPa) and Hoc is weld metal diffusible hydrogen by gas chromatography (ml/100 g)> This procedure does not consider the effect of weld metal height (the total weld metal thickness) and this effect should be taken into account as mentioned in the report [69]] TMCP greatly improves steel weldability, while wetdabihty of a weld metal remains unimproved because its strength and toughness have to be provided not through a thermo-mechamcal control but under an as-welded (as-solidtfied) condition The necessary preheat, therefore, should be determined based on the carbon equivalent of a weld metal rather than that of a steel, especially for high strength steel weldmg As estimated from Eq (6), weld steets metal cracking we)d)ng of TS490 MPa steels crackjng hardly occurs in weldmg
I
I
CE (Stout 1) 1000-C-(-Ma+Cr 6 10 +Mo +A.-+gu) 20 40
a
16
H H H
I I’
a
TB/ B.
i
H’.%’% H+ H+ H+ yB Metsl MeIel
H+ H+
’
6
076+0.25 ,eml20(C-0I2)1
Weld
BB
B
16 6
CE (Scam 1 ) 1000·C· I d + ip + + ) 1g CEHC+A,(C)·I+e^+5+p+Cr+tdo+ND+V+581 C
//
14
19
,
i
-
- **-***
5
040 20 to 10
Where A(C)
B B
75 B
B
B
Ref
CE<))w)-e-t-’-.C&dquo;K. c<-v CE(WES)C+++N-’+5r+M-°+ cE(ston o)=c+L&dquo;e^-++i’ ·io ‘ o’ Q
B
B B
yg-B B
B
cracking
Figure 15 demonstrates schemallcally how hydrogen diffuses from a weld metal to the HAZ dunng welding [67] This is the
Group
B
―**- ,
-
MPa steels
B
B
25 -
cold cracking susceptlblhty of both carbeen proposed bon steels and low-carbon, low-alloy steels It should be noted that a CE,tw type of carbon equivalent also descnbes HAZ hardenabtltty because it is similar to Eq (2)
For tnstance, the
I
ys(WM)=400MPa
-
1B1
to assess
Many methods have been proposed to determine the necessary preheat temperature in steel weldmg [58-63] However, many a problems remain 10 their appltcanon to actual welding practice
,i
,
H&dquo;w = 5ml/100gDM
HI) =17kJ/’ H -) 7kJ/mm
the concept of carbon in 1940, mdices many equmalent evaluating the cold cracking ot steels have been reported av shown m Table 1 susceptibility They are roughly divided into three groups, the first is of a CEI, from Deardens carbon equivalent the sectype which ond is of a Pcm type which regards carbon as more important than the first group and the third is CEN carbon eqmvalent in which the significance of alloy elements vanes depending on a carbon content CEnw has been long used m steel specifications as a weldability yardstick CE&dquo;w well assesses cold cracking susceptibility of carbon steels and carbon-manganese steels, while Pcm better evaluates that of low-carbon, low-alloy steels CEN has Smce Dearden and O’Neil
,
.
,
,I
H A
’Z
y
i
H+ H+
.‘
).
Base Metal
,
Il
Fig tEsumatJon of hydrogen diffusion m welds dunng weldmg [67]
I
;
i
¡
,
,
TMCP STEELS AND THEIR WELDING
383
welding TMCP steel,, with low hydrogen welding matrnal, preheat is thu,, unnerese.trv w in TS.190 MPa Lride weld109 eXlept in the case of uhra-heavy thick plate welding When
,
w.-
-.
1:;
toughness of TNICP steels
4. HAZ
’,
46
1
-,Ok
,
UB _
..
v’
· y
,
. - B..
’
GBF 4
lOne-pan
s
1
’
;, ’, ’ -Oar
4
eld HAZ
The microstructure of a HAZ vane-, depending on the steel chemical composition and the weld heat input and HAZ fracture toughness chinges corresponding to the microstructural change Figure 16 shows the dependence of Charpy fracture appearance
temperature (I T,,) on the weld cooling time for the coarsened gram HAZs of one-pass welds in HAZ thermal simulation tests [72) In the small heat input welding and thereby the short cooling time weldmg, the microstructures are of lower bainitic (BL) and its toughness is sausfactonly high. Ie, I’T&dquo; Is very low HAZ toughness most degrades in the weld cooling times (r,) between 10 and 30 s in the case of TS490 MPa grade steels The microstructures consist of upper-bainite (B.) and femte side plates (FSP) in these weldmg cooling times
y
V.
transition
17 shows typical examples of microstructures at coarse in high heat input weldmg [73] Gram boundary ferHAZs gram rite (GBF) precipitates mainly along the pnor-austemte grain boundanes and FSP and BL grow towards inner grams from GBF Sometimes. femte is nucleated inside grams and it is called intragranular femte There are two types of intragranular fernte, one is polygonal ferrite (IPF) and the other is acicular fernte (IAF) When welding of high strength steel of a TS785 MPa grade, BL is nucleated because of their high hardenability However, when the welding heat input employed exceeds a cer-
,
SP
-
: ° !
:.
GN
,
!.
r
-! .
Figure
8
taIn
level. then HAZ
and
severe
microstructures
upper-balmuc (BL,)
become
results The limitation of a weld heat input (4 IJ/mm or thereabouts) must be stnctly maintained in welding of TS785 MPa steels
8
degradauon
of
toughness
In one-pasn welding, i T&dquo; of a HAZ increases t e HAZ toughness decrease,, as weldmg heat input increases as shown in Fig 16 Howeker the absolute leBel of toughness is always higher in the lower-carbon lower-carbon equivalent steels Figure 18 shows the relauowhtp between the HAZ toughness In terms of the cntical CTOD and CE,,, for TS490 MPa steels This relation also implies that low carbon steels provide higher HAZ toughness although the excessive reduction of carbon results in the severe degradauon of HAZ toughness [74] Figures 17 and 18 show that the reduction of carbon and/or carbon equivalent is beneficial in the improvement of HAZ toughness and it is, thus, essential to employ TMCP which can reduce carbon equivalent or hardenability without the reduction of the steel strength
12C- 3OS,-I.38Mn- ozrro- 36CEiiw &dquo;
350 -
0 08C- 25s,-I.39Mn- 32CEllw
>
(t:>
A 06C- 28S,-125Mn- 15Cu- 2ON,- 29CEIIW
s°
so
o
E
-
300-
I-’ vco
r - ----ð
.
:
smwa.a HAZ -
co
1623K(I350t:>
Temp
-60 -50
-
Z
! 200 -
-o
-
I
2M -
;
-_.--o
Â’,
.
-iI
1.. 1.1
5
10
. I’&dquo;
’,
.. ’
I. IIII
so
’.
I
,
soo
100
Weld Cooling Time bet- 800 .e
’I
soo>r
te -5 (4)
Fig 16 Dependence of Charpy ftacture appearance transition temperature welding cooling time (74)
on
I
=,
. ‘·,11e
a
F
..
&dquo;
_ IPf ’ I IPF
-
’
T
’ FIg 17 Designation of HAZ
microstructures
[73]
Figure 19 shows a dependence of the HAZ toughness (VT&dquo;) on the effective grain diameter of a HAZ mIcrostructure 10 HAZ thermal history simulation tests of TS490 MPa steels [75] The effective gram sme, which is shown as d in Fig 19, is called the fracture facet unit, Ie, a unit step of brittle fracture propagation The toughness degradation caused by the generation of Bu and FSP microstructures is partly due to the coarsened size of the fracture facet unit Since the carbon solubility is lower in femte than in austenite, carbon is expelled from transformed fernte dunng transformation The thus expelled carbon tends to segregate at the boundanes between laths of Bu or FSP, and the austemte-to-fernte transfotirtatton retards at the boundanes This results in the formation of the nuxed microstructure composed of non-transformed austemte and transformed martensitc This nucrostructure is called MA (martens ite-austenite constituent) MA is much_ harder than the surrounding matnx and, therefore, it facilitates the initiation of bnttle fracture and lowers HAZ toughness Reduction In carbon is the most effective way to dinunish MA In
,
I
TMCP STEELS AND THEIR WELDING
3h4
J
,
350 ’
(C)
Q - 50
325 -
B,
300 300 U E
0
’0-
0 Q
’B.../
°
:
-
Q
275
j>
0
0
-0 0o
0
0 Q
a
E
250 -
---
- - 25
8
QQ
_
VI
0
0
- -50 ’
200 -
-
I
25 CE
=
Fig
Q
225 -
..
I
I
I
30
35
40
-75
C + M n/6 + ( Cu + N ) / 15 +( Cr + Mo +
18 Relation between critical CTOD and welding cooling time [74JJ
Fig
transition
/5
temperature of HAZ
200
500 400 300 I
Y 320 -
2 r-
c
,
Q -O t !:::. . T T I I - - N 8 ( 10 ppm
Ct) 40 40
Ti -B
-B
reponed
is
Q
prevent
.
:
femte nucleation also
i ,
; ’
of HAZ toughness
HAZ MlI.ltl-pass Muln-pass weld HAl
d
_
d
280 28°
2 .
d
W 260 -
+
-
zso - _
1350
1400
of
I
I
180 -
I 160 =140 I
160 -
’
a
multi-pass 14 50(*C)
-/’ ...&dquo;
,,;>=--
.I’
....,-- -
140 -
‘20 FSP FSP - - -20
<
I
It
I
(*C) ()
,...
0 0 0
. > 300
(;8B’ 1 t e 1 -60 00411141-60 -
130 5 130S tg -
1! E
4
0 06
1
1
0 08
010
>
-
12 12 0012
-
...
jib-
0 Ti -N
I-
,&dquo;&dquo;
,,’
280 -
260. -
240 m240 -
20 20
,
0260 -
60
Fig 19 Relation between Charpy transition temperature and fracture at fracture Imllallon facet umt unit al initiation site [75] (75) other words, it is also essential to employ TMCP from the point of HAZ toughness improvement.
,
300 - A !:::. Ti-B TI-8
-40 - 40
CN’
- 40
0 Ti-O
-
220 -
’
I
-
240 240
weld
180 -
-
-00
GBF
o 04
I
austerute grams at the HAZ
- 20
-° -
LL
facilitating
that inclusions
The microstructures of a coarsened HAZ
-
0 Ti -N ( 40 ppm S )
_
(Dtom
’
(75)]
42
-
IAF -
plate nucleated at Ti-oxide [75]
S )
300 -
E CL 9
I
TI
-
I
’
100
femte
from coarsening [80. 82] This is the so-called pinning effect The refinement of austenite grains leads to the refinement of FSP and B, growing from austenite grain boundanes and thereby to the improvement
<
150
Intragranular
[81] Oxides among the IAF are inclusions thermally stable and maintain their nucleating capability of fernte nucleation at HAZ close to the fusion hne over 1450*C dunng weldmg On the other hand, fernte nucleation sues other than oxides melt near the fusion line Therefore, a newly developed Ti-oxide dispersed steel is appropnate for high heat input weld 109 as shown In the expenmental results of Fig 211
effectively
’ d &dquo; ’ Gram Effective &dquo; Size.
20
growmg IAF nucleated at Ti-oxides
It
.= >
I
i
‘
2525 25
B0
EE
~
v
·..
-
0
’
,,’
,
&dquo;
.’
--20
.
’
/
- -40
’
vlew-
Intragranular acicular fernte (IAF) is featured by its refined grams and its fracture facet unit is always small Therefore, IAF retains improved fracture toughness Some types of inclusions are considered to act as IAF nucleation sites, they may be REM oxysulphide + BN [76], MnS + TIN [77-79], calcium oxysulphide [80J. and Ti-oxide [75, 81Figure 20 is a microphotograph of
m0 g 220 -
ti LL
200
A-
-
g
--60 I
I
I
1673
1623
1623
1723
1723
1673 Peak
ture
_
-
Temperature (K)
21. Effect of peak temperature at HAZ on and hardness [75]
Charpy transition tempera,
,
TMCP STEELS AND THEIR WELDING
385
change from location to location depending on the extent of the mu)t)-therma!htston. Figure 22 schemaocatty shows the change HAZ microstruLtures in a muttt-pass weld of a TS490 MPa steel I 1831 The HAZ whose microstructure I, coarsened by the first thermal history (the first weld) change, into different structures corresponding to the thermal histories of the subsequent weld passes When the peak temperature of the second pass is immediately above the Ac, transformation temperature. the microstructure is refined because of the so-called normalizing effect (Fig 22B) However when the second heat peak temperature is less than Ac,. the coarsened HAZ is only tempered In the case of a reheating peak temperature between Ac, and Ac, (austenite-femte dual phase region). the diffusible carbon generated by dissolved carbides concentrates in the austemte at around the peak temperatures and hardenability nses in the austenite Some high carbon austenite is transformed into martensite dunng cooling and MA is formed dispersively or in an island-like man-
(*C at dual phase reheated zone id (mm)1-&dquo;-’
C
=
4 12 [ MA (%)j - 14
+
(7)
in
ner
elements such as carbon boron and molybdenum raise and facilitate MA formation Therefore. it is to reduce these elements for HAZ toughness 1871 MA decomposes to some extent dunng reheating under Ac, by subsequent weld passes because of their tempering effect Silicon The
allo,.
hardenabt!ny
HAZ desired
800
600
50 _*&dquo;
HT490
Peak Temp
23 shows the relation of cnnca) CTODs to the reheating temperature in thermal history simulation tests [841 As descnbed above, the cntical CTOD is the worst due to the generation of MA (island-like martensite) when reheated to the dual phase region around 800°Cm the case of TS490 MPa steels The region reheated to the dual phase also exists approximately 3 mm apart from the first pass fusion hne (Fig 22E) This region is another locally embnttled zone due to MA other than the embnttled zones along a coarsened HAZ The HAZ microstnictures of a TS785 MPa steel with high hardenability are generally of lower bainite (BL) at rather low heat input welding However, once HAZ is reheated immediately above Ac,, the microstructure changes from B, to B, because of low resulting from fine gram size of the onginal BL structure [85]
(1400t:)
-
g
governed by
not
0
01
t -
that the effect of MA
is more
dominant
U ’
0 0
-
0
f
1
I
I
I
I
1000
1200
1400
1600
Cycle (K)
Peak Temp at 2nd
23 Relation between
peak temperature of
second weld pass
on cnti-
[841
coarse-grain HAZ I (CGHAZ) ’ ,’J g ytJ j
,’,.l’
,,’,
B―
’
Fig
B
B
A
8
cal CTOD of HAZ
3 bead
t B A B*9&dquo;’B B f fA
t
1 st Cycle
A A
B A
-
0 01
[86]
BA B
fB
n
_
8
: 0 o0 05 -
only the effective
gram diameter (d) but also the total volume of island martensite (MA) The following relation represents the HAZ toughness (%,T,) as a function of the MA volume and the fracture facet unit,
Imply 109
-
C 7ï;
Q B
-
SJ
re
8*T A
0 5 -
h1 ardenbiltyD is
1 st Cycle 1673K
at
r:1
Figure
&m A
HT785 A
-
10
1400 (*C)
1200
&dquo;Ic 0
-
along the pnor austemte gram boundanes (Fig 22C)>
The HAZ embnttlement
1000
’’!’’’!’!I
B
I
i
I Fir* -grain HAZ
(FGHAZ)
2 bead
BBBy BB iS? B
bead
c
ed
coarse grain zone
IRCG)
’&dquo;’-&dquo;’ D subcritically ’’t
BBBBBMS<
intercritically reheat-
’’.v-.r’.BB’
Fig 22 Microstructures of mulu-pass weld HAZ [83]
reheated ’&dquo; zone
-
·
.
TMCP STEELS ANL) I HtIR H TLDIN(,
396
retards MA decomposition because it stabihzes MA [87 981 Therefore the reduction ot ,ilicon result- m the improvement of toughness of mulll’pJ&dquo; H .1&dquo;Z, 1801 Due to their fine size acicular femte (AF) can disperse carbon concentrated zones dunng reheating to the dual phase region and then the MA formation is diminished 187) The steels effectively utilizing mclusions such as Ti-oxides provide satisfactorily toughness with their mulu-pass welds because their HAZs mostly
,.,
consist
of fine IAF
.
-
;
If)
- ·
CR-AcC
>
A
340 -
_
g
œ
{! .
CR
(t)
0
AR
.
high
,:
20 (*C) II
0 -120 -100-80-60 -40 -20 I I I I I I
// - 60
320 -
’
300 -
K +40K
JI 220 -
/
.II’
/
·
-
- 20 20 -
/
-0
/ O · O
?
t
/
*I
40
/
--20
i -1 1: 1
0/ · A ./
-
-40
- -
-60 60
-
-80
./’
523K x 3600s
A
markedly
/O
(40C) , /’1 00 0 (40*C
Ea
MA forms locally embnttled zones and they are the main cause 280 of the occasional occurrence of very low values dunng a number III of CTOD tests It was once indicated that the TMCP steels may 260 be susceptible to local embnttlement an their HAZs This controversy emerged presumably because TMCP was developed just 240 when CTOD testing, whtch facilitates the detection of local embnttlement, was prevailing and CTOD tests were thus performed predommantly on TMCP steels In fact, TMCP can prectsely control the hardenability enhancing elements and the MA200 decomposmon retardtng elements at a desired level, concurrently ti 180 considering the steel strength and toughness It is certam that the is whose HAZ TMCP steels manufactures of employment A 1B improved with respect to not only Charpy charactenscnuca) CTODs tics but also critical
//
(250’C x 1hr) - -100 -
))))))I
t t t t
260
280
I
300
I
Fracture Apperance Trans Temp . vTrs (K)
5.
Reheating characteristics of TMCP steels
Ftg
5.1Stram aging g
healing characteristics
or
spot flame heating Microstructures
locally heated heaung they are subjected to, at
the
change dependtng on the toughness and strength also vary The Japan Shipbuilding Qualtty Standard stipulates 600’C as the maximum allowable heaung temperature for flame heating followed by water coohng However, heaung up to 900*C may be permitted if regions
and thus their
the heated portton The mechanical
m
base metal toughness under stram agmg
(3]1
is air
cooled
properties
to
were
i
Post weld heat treatment (PWHT) including normaltztng (N), normalizing and tempenng (NT) and quenchmg-and-tempenng tQT) are not permitted for TMCP steels. which are strengthened by accelerated cooling Hot formtng is also not allowed but warm forming may be performed with the necessary cauuons being taken Stress rehevmg PWHT is appitcabie to TMCP steels JIS (japan Industrial Standard) pressure vessel steels whose thickness is 38 mm and over have to be subjected to stress rehef PWHT m order to reduce welding residual stresses and to improve matenal propemes This is ruled on the premtse of the use of normalized steels and as-rolled steels For TMCP steels wth improved base metal toughness. PWHT may be performed only for residual stress
In steel fabncatton, fabrication, steels are formed to a desired shape by the use of local plastic deformation or thermal contraction caused by
hne
Change
5 3 PWHT characteristics
Steels are often subjected to reheattng mcludtng flame heattng after bemg cold formed to several per cent m fabncanon At this ttme, base steels degrade m their toughness due to strain agmg Figure 24 shows changes in Charpy fracture appearance transition temperatures when subjected to 250°C x 3600 s agmg treatments after 5% cold formmg [3] The degree of degradation due to stram agmg is 20*C or thereabouts and it does not differ between convenuonal steels and TMCP steels Therefore, the absolute level of toughness after stram agmg is higher in TMCP steels than m convennonal steels because of the improved base metal toughness in TMCP steels 5 2 Flame
24
500*C and then
water
cooled
examined for normalized steels,
controlled rolled steels and TMCP steels when they are subjected to local heaung with varymg maximum heaung temperature and water coohng start temperature [89] All the steels were strengthened to some extent and their toughness was reduced while the degree of toughness degradation did not differ from steel to steel All the steels were most degraded when heated to the dual phase temperature (between 600 and 700*C) and then air cooled, because of the resultant coarsened nucrostructures mixed with femte, bairate and island martensite Water quenching from these temperatures must be avoided and the rule of air coohng to 500*C should be kept- TMCP steels are also supenor to conventional steels 10 flame heaung charactensucs because of their improved base metal toughness
rehevmg relieving
26 shows the expenmental results of the effect of PWHT on TMCP steels [90. 91] is seen that the welding residual stresses are satisfactorily removed by PWHT of 550*C without significant reducuon of their strengths, mstead of PWHT of 600*C This result suggests that the PWHT condition may be relaxed from 600’C to 550’C for TMCP steels The addition of small contents of ntobtum is very effective through precipitation hardenmg agamst the strength reduction dunng PWHT But care must be taken to prevent precipitanon embntthng of the steels caused by excessive addition of niobium [92]
Ftgure
6. Weld metal
properties in TMCP steel welding
6.1.
Toughness of weld metal
The
welding
Ti-B system is extensively used for Hardenabihty of this welding matenal is
matenal of
a
high heat input weldmg controlled W1thm the appropnate range depending on welding heat mputs to be employed, so that its rrucrostructure mostly consists of fine acicular femte (IAF) nucleated from Ti-oxtdes while impeding the precipitation of grain boundary femte (GBF) and upper baitute (Bu) to a minimum The appropnate amount of free
-
197
TMCP STEELS AND THEIR WELDING
- J As Welded tl 5501rx I .
2h
2h
-0 600’t1Ix 2h (’C)
.t.}
it―――K
-1
250
·___Sll._____________·1 _ -60 lr----<>---------FL -80
-
:
>200 200 -
O---lL
-
--40
+2mm
_
--100 -100
66000 ---
-
0)
165
B ’B
’.
MM -!! M
v B
-
-
.1C)
As We!ded
873 K
x
m
properties
-
* A***
,,,,,,, YS
p
200 -
S-Mn-Mo-T-A8 S,-Mn-T,-B E 0 Q 0
·
·
J
m
!
0
r
100 -
100 -
-S i
p
0 0
0 0 <:xö
0
00 ’bOO
__
o
-
0
I
25
10
I
I
I
I
30
35
.40
45
toughness and strength of Ti-B
I
I
I
i
i
04
06
08
1
I
I
003
007
In Steel Plate (%) I
I
017
012
I
I
027
032
I
021
poor weld metal
thus.
be matntatned at results
cannot
toughness
a
destred level and
Many elements in a base metal affect, through dtlutton the strength and toughness of the Ti-B weld metals m high heat input welding In parttcular, aluminium in a base metal has a substantial effect, as shown in Fig 27 [94] Al-oxides are not capable of nucleating destrable IAF.
weld metal
a
unhke Tt-oxtdes The de-oxtdanon molten weld metal proceeds m order of affintty to
followed by tttamum, stltcon and then manganese Therefore, the remamng oxygen after the oxtdauon of aluminium should be retained to a suitable level depending on base metal aluminium content in order to form effective Ti-oxtdes In the excessive aluminium region, all the oxides are AI-oxtdes and no Ti-oxides are formed As a result, the weld metal toughness significantly degrades Conversely, an excessively low alununium content results in the fonnatlon of not only Ti-oxtdes but also undesirable Si-oxides and Mn-oxides, and the weld metal toughness also worsens oxygen,Ie, aluminium is first de-oxidized
Cb 0 C 5
CE = C+St/24+Mn/6+Mo/4
Fig 26 Dependence of carbon equivalent [93]
--too
02
AI Content
reaction m
* &dquo;
50 -
::
00
0
(598)
26 shows the dependence of the strength and toughness of the Ti-B weld metals on thetr carbon equivalent [93] At a low level of carbon equivalent, the microstructure consists mostly of GBF due to lack of hardenabtltty, while it consists of upper baintte (BL) m the high carbon equivalent Toughness m both regions is low In the intermediate region, desired hardenability is acquired, resulting in ACF microstructures with high toughness Lack of hardenability In the Ti-B weld metals may be caused by base metal dilution when welding of TMCP steels whose carbon equivalent (CE&dquo;&dquo;,) is not higher than 0309é In this case. welding materials whose hardenability is prepared to a somewhat high level must be used
A= ABo.-rPS-SD―
&dquo;
’&dquo;°
Figure
800 -
Ii) 400 -
0
B weld metals,
7200s
25 Companson of TMCP steel and convenuonaf steel after PWHT (90, 91 )]
-
/
--60
Al Content m Weld Metal (96)
Fig
TS
/
0
/
Fig 27 Relation between Charpy transition temperature of Ti-B weld metal and alunumum content (94]
Y’――’=<
823K x 7200s
_-40
‘ (588) 0 120mm
t80 -
N( Norm )
A
B.
(566) (5054)
’°°’200
LL
’
,S 600 -
B
r-―――――――― P 0 TMCP
/
() T S (M Pa)
0 i550) O
ma
_ -o――_
-
g
f
220
TS(MPa) /m
(545)
¡:
·
-
500 -
æ
# co <11
!
-
-
-
m
E 9
/
4 kJ 4 66kJ/mm
5)
240-240
S
550 ..
a>
(t)
637)637)--20 -20
300 ppm Oxygen
c
_ -
f
l4Mni6Mo-Ti-B 14Mn-16Mo-T!-8
20
O---
-
f
260 -
-
on its
boron is necessary to prevent the nucleation of GBF The amount 6 2. Solidification cracking of weld metal of necessary free boron vanes depending on the amounts of oxygen, mtrogen, aluffunium and tttantum Also, the amount of Ti- A root weld m pipeline gtrth welding with V-grooves is consideroxides necessary for the nucleation of IAF is determined by a ably diluted by base metal because of the strong arc penetration of root welding even though weld heat mput is low Since the sensitive balance of utantum, oxygen, alumimum, silicon, and metal into weld metal dilution the HAZ hardness lirrutauon is stnngent m the pipe girth weldmg, as base Furthermore, manganese becomes tttunense m high heat mput wetdmg. 1D which the dllualready descnbed, a carbon content is usually reduced m hne-pipe tion rate may be as high as 50% The cherrucal composition of Tisteels through TMCP It is known that high carbon weld metals
-
,
TMCP STEELS AND THEIR WELDING
388
generally ,u,,ceptible to ’ohdlflcallon cracking However weld metals woh e1(ce&dquo;lvely reduced carbon are also vulnerable 10 ,ohdlficallon cracking 1951[
2 A J De Xrdo Proc Int Ssmp on Awelerored Cnnlme of Rolled Steel (Edued b G E Ruddle). CIM. winnipeg Canada
Solidification cracking is more likely as the weldmg velocity increase, Therefore, a root weld in pipe girthh weldmg is very susceptible to solidificauon cracking because root welds are made in a vertical-down manner at very high speed Figure 28 is one example of a solidification crack in the low carbon region In weldmg of the ultra-low carbon TMCP steels for line-pipe use This type of solidification crackmg occurs in V-groove cellulose electrode weldmg when the weld metal carbon is less than 005% [96] The reduction of carbon content in a steel to a level less than 0 02% leads to unexpected coarsening of austenite grains in HAZs Therefore, high grade line-pipe steels contam carbon at 0 05% or thereabouts for concurrent consideration of the avoidance of solidification cracking and the HAZ hardness hmltallon
3 Monkawa.
are
’ ’ ’ ’1. ;
t
e
1988
B10nyama
pp 83-90. 1986 (in &
7 Tsuchida. Yamaba, 1989 fin Japanese)
(in Japanese)
.
9 Takechi
No 7
CAMP. Vol 2. pp 1724-1727,
Yamaguchi
Shiga
CAMP, Vol 2, pp 1728-1731. 1989
Tetsu-to-Hagane, Vol 68, No 9, pp 1244-1255 1982 (in Japanese) 10
s
Vol 55
6 Ohashi, Mochituki. Yamaguchi Settetsu Kenkw, Vol 334, pp 17-28, 1989 (in Japanese)
j/
‘
of TMCP Steelf for
5 Terashima. Furuklml J Jpn Weld Soc pp 411-418. 1986 (in Japanese)>
8 Matuszaki. Saito,
..
Japanese)
Steel Institute Properties Preuure Vessels, 1986 (in Japanese) 4
Japan Iron
Jpn Weld Soc. Vol 55 No 2
ltoh J
Ikenaga. Taklta, Mizui CAMP, Vol 2.
p 759, 1989 (in
Japanese) E R Parker, D Fahr, R Bush Trans Am Soc Met , Vol 60. pp 252-259, 1967 11V F
12
Zackay.
CAMP, Vol 1, No 3. p 945, 1988
Tsukaya. Kamel, Sakai
.
2mm Fig
28 Solidification crack
In
pipe
girth
root
weld of very low carbon
steel 1961
13 Yamamoto. Kobayashi, Hondda CAMP. Vol 2. No 5. p 1732, 1989 (in Japanese) 14 Ida, Takeshima 1988 (in Japanese)
15
a
O
Fujimoto CAMP,
Kageyama, Sugmo,
Vol
I
No 5, p
1509
Fukuda Seitetsu Kenfsu, Vol
7. Conclusion
pp 2-14. 1988 Un
It has been possible to manufacture high quality of steels by a precise control of chemical composiuons together with employ-
Tetsu-to-Hagane, Vol 73. No pp 1162-1169, 1989 On Japanese>
of TMCP TMCP steels are not permitted to be subjected to heat treatments in which the temperature employed exceeds the steel transformation temperature TMCP steels are supenor to conventional steels with respect to strain aging embnttlement, flame heating embnulement, toughness after PWHT, HAZ hard-
16
329
Japanese)
Wada, Fukuda
9.
ment
ening, cold cracking susceptibility, and HAZ toughness However, there are some problems concerning HAZ softening, weld metal toughness and solidification cracking when welding of TMCP steels
17 Japan Shipbuilding Research Association Report No 100. Research on HT50 High Strength Steels by New Process, 1985
(in Japanese)
18 N Yunoka, M Okumura, T No 4, pp 217R-223R, 1987 19 N Yunoka Advances
In
Kasuya
Metal Const, Vol 19,
Welding MetallurgB, pp 51-64,
AWS-JWS-JWES, 1990
Acknowledgements 20 N Smith, B I
for Dr S Aihara, Dr Y Hagtwara, Dr H T Saito, Mr Y Hom and Dr M Okumura of the Steel Research Laboratones of Nippon Steel for their expert
The author
is
Tamehiro, Mr
grateful
Bagnall
Metal Const , Vol
I, No 2,
pp 17-23, 1969 21 E J Ridal Metal Const , Vol
3,
assistance He is also thankful to Prof De Meester of Universitd Cathohque de Louvain and Dr C Shiga of Kawasala Steel for their advice on the revision of this manuscnpt.
22 D McKeown, P Judson, R No II, pp 667-673, 1983
L Apps
References
23 P Hart Metal
No 11, pp 413-417, 1972
advice and
Melloy, C W Vol. 5, p. 896, 1967 1 GF
Roe and R D Romenl: Industnal
Heating.
:
Metal Const , Vol 15, ¡
Const , Vol 18, No 10, pp 610-616,1986
24 Okumura, Kasuya, Yunoka: Quarterly J 6, No. 1, pp 144-150, 1988 (in Japanese)
Jpn.
Weld Soc , Vol
-
-
TMCP STEELS AND THEIR WELDING
’_5 Yamamoto Ha7e Mjtsudj
p 776
1998
110
389
ChIJI1BB.J
CAMP Vol I No 3
Hagmara 8th OM4E,
49 Kubo Bakano
CAMP, Vol .1 p 919 1991 (in
JJp.¡nt:,e)t
Beckert.R Holz pp 344-346 1973 26 M
S( l/IIelHtedlnl1..
27 N Yunoka S Ohshtta. Tamehiro Symp the 80’, AWRA Melbourne Australia. 19811
Vol
23 No 8
Ppelme Weldmg
Japane,,e)
50 Amano Shiga, TJnJkJ Symp on Weldmg Metallurgy of TMCP Steel, JWS.pp 101-I 10 1985 (in Japanese)> m
Koba_vashi Symp
51
Weldmg Metallurgy Un Japanese)
of TMCP Steels.
52 J Dearden. H 0 Neil Trans Inst Weld
Vol 3 No 10,
on
JWS pp. 177-184. 1985 28 C Duren IIW Doc IX-1356-85 29 PJBoothby Metal Const , Vol
vol III 1989
.l8 S Machida Y
17, No 6, pp 363-366,
pp 203-:!14, 1940
1985 53 H Suzuki, H Tamura IIW Doc IX-285-61
30 Suzuki Quarterlv J Jpn Weld Soc . Vol 4, No I, pp 90-95. 1986 (in Japanese)> 31 Terasaki Nomura, Kitada Quarterh J 6, No I, pp )39-)43. )988()n Japanese)
8
Jpn
Weld Soc , Vol
55 Ito, Bessho J Jpn Weld Soc . Vol 37, No 9, pp 983-991. 1968 (tn Japanese)
32 M Okumura. N Yunoka, T Kasuya. H.J U Cotton Proc on Stress Rehetrng Heat Treatment of Welded Steel Construction, 56 B A Graville Conf Welding of HSLA Structural Steels, 1987. Sofia, pp 61-68 Pergamon Press, Oxford ASM/AIM. Rome. Italy, 1976 33 Yunoka. Okumura J 1990 (tn Japanese)
Jpn Weld Soc , Vol 59,
No 2, pp 6-9,
57 K Lorenz, C Duren Conf on Steels for Linepipe and Ptpelne Rttings. TMS. London, pp 322-332, 1983
34 K Shmada Y Hom, N Yunoka Weld J. Vol 71, No 7, pp 253s-262s, 1992
58
Nakata Saji. Fukuda, Saito JWS Resistance Weldmg 35 Research. RWS-54C-82, 1982 (tn Japanese)
&dquo; ° IX-970-76 59 S Matsui. M InagakJ IIW Doc
36 JWS Reststance Nakata. SSaji, Fukuda, Welding Resistance 36 Nakata. r Fukuda, Satto Weldtng
Research, RWS-54D-82. 1982 (in
8
54 R D Stout. R Vasudevan, A W Pence Weld J , Vol 55, No 4, pp 89s-94s. 1976
Japanese)
110. Bessho J Jpn Weld Soc, Vol 1 l34-I 144 1969 (m Japanese) pp ’
60 V PaBaskar J S S
1982J 60 -56--6-. V P’ yskar
Kirkaldy Kirkaldy
38, No
10.
’
’
Scand J J MfM//. Scand Metall , Vol
I 11, 1,
PP
37 Nishi, Saito. Yamada, Takahashi Settetsu Ketzku, No 307 p 56. 1982 (in Japanese)
61 N Yunoka. H Suzuki S Ohshlta Weld J . Vol 62. No 6, pp l.l7s-153s 1983
38 Nose Tanaka, Sato JWS Resistance 149-78, 1978 (in Japanese)
62 R A J Karppt. J Runsila, M Toyoda, K Vannamen Scand J Wetall. Vol 13, pp 66-74. 1984
39 Yamauchi. Koh JWS Resistance 79, 1979 (m Japanese) 40 J W Michell
41
Svmp
on
Welding Research
Welding Research,
RW-
RW-166-
Bessho J Sumttomo Metal, Vol 26, No 2, p 182, 1974 (in
Tanaka, Nomura, Kokubo Tetsu-to-Hagane, Vol 68, No 9, 1437, 1982 (in Japanese) p
42
Nakata, Nishikawa, Ohsawa JWS Resistance Weldmg Research, RW-153-78. 1979 (in Japanese) 43
Vol 2, p 759, Taktta, M.zu,. 759, Mtztn, iCsh.da Ktshtda C CAMP, Vol Ikenaga. Takita, Ikenaga, (in Japanese)
44
Yapma Symp
on
Japanese)
46 Shtnozakr 128th
Nishiya
1989 (tn
Japanese)
1989
Welding Metallurgy of TMCP Steels, JWS,
pp 140-150, 1985 (in
Memonal Lecture, pp 111-153,
47. R Denys Proc Intl Conf on Evaluation of Materials Performance in Severe Environments, ISIJ, Kobe,
pp 1013-1027, 1989
on
Weldmg and
AWS DI 1-90 Structural Steel Code, Appendix XI, 1990
Micro-Allottig, 75, 94, 1975
Japanese)
45
63 C Duren. K Niederhoff Proc 3rd lnt Conf Performance of Pipelines TWI, London, 1986
65 British Standards lnsmute Specification for process of arc welding of carbon and carbon manganese steels. BS5135, 1987
Kasuya
66 N Yunoka, T 67 H J &dquo; 68 J’
Granjon
IIW Doc
IIW Doc IX-1740-94
IX-830-73,
IX- 1649-91 IIW Doc -’-&dquo; &dquo;
Defoumy
69 Yatake, Yunoka, Kataoka J _’96, 1981 (in,, ,_ Japanese) pp 291
No ’ Vol ’ 50, ° Soc ’, ° 3, Jpn Weld °
70 PH M. Hart Weld J, Vol 65, No 1, pp 14s-22s, 1986 71 N
Okuda, Y Ogata, Y Nishiyama- Weld. J , Vol 66, No 5,
pp 141ÿ-147s, 1987
72. Tanaka pp
Nishiyama 77-105, 1989 (in Japanese) 128th
73 Honi: 128th
Nishiyama
1989 (in Japanese)
Memorial
Lecture,
JISI,
Memonal Lecture, JISI, pp 39-76,
_
TMCP STEELS AND THEIR WELDING
390
J Njkantsh. Komiru Fukada (3«<-r Vol 4. No 2 pp 447F52 1986 Un Japaneu1
74
Jprr
Hc/J Snc
75 K Yamamoro S ’1J,uda. T Hazc. R Chyma Pnx Symp on Residual and Unspecified Elements in Steel AST11 STP1042 pp 266-284 1987
76 Funakosht Tetsu-ro-Hagane Vol 63 No 2 pp 105-&dquo;4 1977 1 in Japanese)> 77 Ohno CAMP. ol 73, No
8.pp 9-1-101 1987 (in Japanese)>
79 Furusawa J Sumnomo Metal, Vol 40. No 1, pp 39-.f8, 1988 (m Japanese) C
Chiga
Pre-Assembly Symp
Proc
Welding/Joining/Coating and Surface Vol 1, pp 207-212, 1974
IIW
Y Terada. R
Chtjnwa.
82 C Shtga,
Y Saoo 1 st Jpn-USSymp on Advances We)d)ngMeta))urgBAWS-JWS-JWES.pp 295-324. 1990 83 Haze Sertetsu
on
in
, No 326, pp 36-44, 1987 (in
Japanese) 84 T Haze S Aihara IIW Doc IX- 1432-86 85 Nakao, Oh%ato No. Nishi QuarterhJ 5. No 3 pp 410-422, 1987 (in Japanese)
OMAE. Vol 3
87 S Aihara K OkJmoio Prcx Intl Conf on the MetaXurgy Welding and Quaithcation of Bhcroai)oyed Steel Weldment
AWS, Houston, pp 401-425. 1990
gg SJ Barnard Adiancev rn P/!tuca/M<-fM//Mt TMS pp 33-37 19811 .4/t
and
89 Watanabe NKK Tech J . No I 1 2. pp 63-68. 1986 (m
90 Tanaka, Shikauchi CAMP, Vol 3. No 6. p 1971, 1990 (in
Japanese) 91 Tanaka. Ohmsh) CAMP Vol 3. No 6, p 1972, 1990 (in
on
Modification, IIW Darlian.
H Tamehiro 2nd Intl Conf HSLA Steels, j 6-ff 519-524, 1990
81
Inil Cont 7th
Japanese)
78 Chou Krm JWS Annual Meeting. Vol 43, p 311 1988
80
86 K Uihino Y Ohno Proe pp 159-165 1987
Jpn Weld Soc. Vol
Japanese) 92 Sato. Takeda. Kanaya Quarrerh J No 4, pp 691 99, 1984 (in Japanese)
Jpn Weld Soc , Vol 2.
93 Okuda, Wada. Tanaka Symp of Welding Metallurgy TMCP Steels, JWS, pp 200-209. 1985 (m Japanese) 94 Y Honi. S Ohkita. M
.. Research. ASM Tennessee. USA
of
..„.. pp
413-417, 1986
95 Masumoto, Ozaki J Jpn Weld Soc , Vol 42, No 7. pp 674-684, 1973 (in Japanese) 96 S Ohshtta. N Yunoka T Kimura Weld 7. Vol 62, No 5. pp )29s-)36s. )983
!
: t
,
