ISIJ International, Vol. 38 (1998), No. 12, pp. 1379-1386

Fracture Properties of Multipass SubmergedArcSteel Produced by Using Flux Cored Filler Wire

Weld of HSLA

P. K, GHOSH.P. K. SINGHand N. B. POTLURl

Welding247667,

Research Laboratory, Department of MechanicaiIndia.

(Received on June

and Industrial Engineering. University of Roorkee. Roorkee-

77. l998, accepted in final form on August 20. 1998)

Multi pass submergedarc welding of 25mmthick structural steel plates has been carried out at different

energy input using flux cored filler wire and basic agglomerated flux resulting the weld deposit havingchemical composition confirming that of a high strength low alloy (HSLA)steel. Microstructure and hardnessof different microstructural regions of the multipass weld has been studied. Tensile properties of the welddeposit has been found out by carrying out tensile test of the weld joint being fractured from the weld,Fracture toughness (Ka) and fatigue crack growth properties of the weld are studied and correlated withthe welding energy input, microstructure and ultimate tensile strength of the weld. The increase in energyinput has been found to reduce the hardness, ultimate tensile strength and yield strength, but to enhancethe ductility andKaof the weld primarily due to its influence on microstructure of the weld deposit. However,the increase in energy input has been found to reduce the daldN at higher AKbut, it has been found to

enhancethe sameat lower AK. TheKaand da/dNof the weld are found to be we[1 correlated with its tensile

strength, where an increase in tensile strength reduces the Ka but enhancesthe da/dN at AKhigher than

30 MPaJF~.

KEYWORDS:submergedarc welding; flux cored filler wire; HSLAsteel weld deposit; multipass weld;

energy input; weid characteristics; microstrvcture; hardness; tensile properties; fatigue crack growth rate;

fracture toughness (Ka)-

l. Introduction

The high strength low alloy (HSLA) steels have be-

comewideiy popular in fabrication of various structuresin manycrltical and non critical applications due to high

economyand more allowable design stresses permitted

by their signlficantly enhancedyield strength. The fab-

rication of these structures largely involves different

weiding processes, where the multipass submergedarcwelding (SAW)is commonlyused during joining of thick

steel sections. Asuccessful and economical use of HSLAsteel weld very muchdepend on proper selection ofwelding consumableand possibility of using hlgh heatinput welding process allowing high deposition rate.1 ~ 3)

TheSAWprocess has been successfully applied in manyoccasions of multipass welding of thick steel plates, wherethe weld properties are significantly governed by the

microstructure of weld deposit.2-7) Thus, to enhancethe range of heat input of the SAWprocess in producingthe multipass weld of HSLAsteel a characterlsation of

weld properties with respect to its microstructure and

energy input of the process at different welding param*eters is very muchnecessary. This Is because the failure

of welded structures is often found to be initiated

from the we]d metal and fracture propertles of a weld

1379

slgnificantly depend on morphology of the weld de-poslt,7,8) which in multipass weld is primarily char-

acterised by the proportion of the dendritic and re-

heat refined regions9) present in it. The morphology of

a multipass weld is largely governed by the weld thermalcycle and characteristics of deposition dictated by the

energy input and welding parameters of the process.9)

Further the characteristics of deposition also varies

slgnificantly with the type of welding electrode, such as

a solld or flux cored electrode, used in a welding process.In recent years the use of flux cored electrode as filler

wire is very muchfavoured by manyindustries primarily

due to its ability to produce goodquality weld of adesired

chemical composition wlth high deposition rate. But,

hardly any work has been reported so far on the varlous

aspects, as discussed above, of a multipass submerged

arc weld produced by using flux cored filler wire.

By keeplng a pace wlth the modernconcepts of struc-

tural design and economy the growing interest and

awareness about the safety and reliability of a welded

structure, especially under dynamic loading, has madeit imperative to understand the fracture mechanicsbehaviour of a weldlo) along with its conventional

mechanical properties. But a very limited number ofliterature is available so far on fracture mechanics

O1998 ISIJ

ISIJ International, Vol. 38 (1998), No. 12

Table l. Chemical composition ofthe filler wire. (wt"/*) Table 3. Welding parameters.

C Mn Si Cl' Ni Mo

0.08 1.5 O.35 2.508 0,4

Welding current Arc voltage Welding speed Energy input(A) (V) (cm/min) (kJlmm)

Tabie 2. Chemical composition of the baslc agglomeratedflux. (wt'yo)

400450500450450

2828282828

4040403035

l .68

1.89

2. l2.522. 16

Si02+Ti02 Ca0+MgOA1=03+MnOCaF, Basicity index

15 40 20 25 3. 1

2ef

)~8

150

Fig. 1.

l~lr=,^,=

L~

mmSchematic diagram of the weld groove.

behaviour of a multipass weld and its correlation withthe microstructure, weld thermal cycle and weidingparameters. In this investigation an effort has beenmadeto study the effect of energy input of welding,varled with a change in welding parameters, on micro-structure, fatigue crack growth rate and fracturetoughness of multipass weld, deposited by submergedarc welding process using flux cored steel wire electrodeof a high strength low alloy steel.

2. Experimental

2.1. WeldingSubmergedarc welding of structural steel plates of

dimension 400 x 150 > 25mmwas carried out using3.2mmdiameter flux cored filler wire and basic ag-glomerated flux. Chemical compositions of the filler

wlre and agglomerated flux are shownin Tables I and2respectively. Thechemical composltion of the filler wiredeposit weld metal confirming the composition of aHSLAsteel. To minimise final distortion of the weld,the base plate and the backing plate was tack weldedwith a pre-bend angle of 8' using manual metal arcwelding process. The welding was carrled out by mul-tipass deposition in a groove having a backing plateof dimension 430x 40 x lOmmas schematically shownin Fig. 1. Prior to welding the groove surface and its

adjacent region of the base plate wasthoroughly cieanedmechanically and by applying acetone to remove the

presence of rust or any oil and grease on the faying sur-face. During welding the energy input was varied bychanging the welding current and speed as shown in

Table 3, whenthe electrode polarity was malntained asdirect current electrode positlve (DCEP).

After welding the run on and run off portions of the

weld wasdiscarded and the specimensfor metallography,

C 1998 ISIJ 1380

tensile testing, fatigue crack growth rate testing andfracture toughness testing, were collected by saw cut-ting from the stable welding region of the joint.

2.2. Chemical Analysis and MetallographyThe chemical composition of weld deposit was

analysed under atomic absorption spectrometer andinfrared carbon-sulphur analyser using the chips drilled

out from it. Transverse section of the weld wasprepar-ed and polished by standard metallographic procedureand etched in 5"/* alcoholic nitric acid (nital) so]utionto reveal the microstructure of weld deposit for study-ing under optical microscope.

2.3. HardnessTest

Hardness measurementof the weld was carried outby Vicker's diamondindentation along the weld centre-line, at a load of lOkg, as revealed on the transversesection of weld joint prepared for the metallographicstudies. Thehardness of various microstructural regionssuch as the coarse and fine dendritic regions and the

coarse and fine graln regions of the weld wasalso studiedby Vicker's mlcrohardness indentation at a load of 200g.

2.4. Tensile Test

Tensile properties of the weld joint, such as its ultimatetensile strength ((T~), yield strength (cry~) and elongation,

were studied using standard (ASTME8M-89b) roundtensile specimen havlng weld at its centre as schemati-cally shown in Fig. 2. The test was carried out in auniversa] testing machineoperated at a cross headspeedof I mm/min. Elongation of the specimen was studiedat a gauge length of 40mmwith the we]d at its centreand the yield strength was determined at 0.2"/, off-set

strain on the stress-strain diagram.

2.5. Fatigue Crack Growth Rate Test

Fatigue crack growth characteristics of the weld metal

wasstudied using 15mmthick C(T) specimen, which wasprecision machined from the weld in accordance withthe ASTME647-88aby keeping the orientation of notchalong the direction of welding. A schematic diagram ofthe C(T) specimenhas been shownin Fig. 3. The test wascarried out on pre-cracked specimen. The initial cracklength to width ratio (ao/ W)of the specimenwaskept as0.25

.During pre-cracking under dynamicloading the ratio

of crack length (a) to width of the specimen (W) waskept in the range of 0.3-0.32 using compliance methodof crack length measurement,where the stress intensity

factor (K) was maintained constant by reducing the

maximumload under load control modeof the machine.The fatigue crack growth (FCG) tests were performed

•~

[SIJ International, Vol.

-co

40

90

SCALE: mmFig. 2. Schemrtrc dl l~l un of the tenslle speclmen

+

(~dI\

*~ roro

+

so

7s

[~l

38 (1998), No, 12

~lSCALE: mm

Fig. 3. Schem'atic diagram of the C(T) specimen.

under cyclic loading in accordance wlth the procedureoutlined in ASTME647-88ausing fully automatedclose

loop servo-hydraulic universal testing machine. The test

was carried out in open air laboratory environment at

room temperature and uslng a sinusoidal wave formhavlng cyclic frequency of 15 Hz. The test was carried

out by constant amplitude methodat a given minimumto maximumstress ratio (R) of O. l, where the minimumandmaximumstress intensity enhancedwlth the increase

of crack length. Here also the compliance methodwasadopted for measurementof crack length, where the Kvalue was determlned from the corre]ation of load andcrack length at a given stress intenslty. A11 the tests

were conducted under load control modeof operationof the machine wlth a load control card range beingselected based on speclmen thickness, thereby mlnimis-ing any error arising from load control. A MTSbased

computersoftware wasused to conduct the test and storethe test results of (i) cycle number, N, (ii) maxlmum10ad, (iii) minimumload and (iv) crack length. Thesoftware wascapable to calculate the crack length based

on CMODgaugedlsplacement and to adjust automati-cally the minimumand maximumload. A fifth ordercompliance polynomial was used to correlate the ratio ofchange in crack openlng displacement (COD)to changein load with the crack aspect ratio (a/W) to determinethe crack length a. A visual crack length measurementwas also periodically carried out to ensure that the

curvature of the crack length is within acceptable limits.

Thecrack length along wlth the corresponding minimumand m'aximumload and cyclic counts (N) were saved in

data file at certain Interval. The tabulated data wasusedfor evaluation of crack growth rate (da/dN) and the

stress intensity factor range (AK). The fatigue crackgrowth characteristics of the weld meta] deposited at

different energy input was plotted as log(cialc!N) vs.

1381

logAKand analysed in the light of the Paris law.

(da/clN) = C(AK)*"....

..........(1)

whereCandmare the material constants. Theconstants

Cand ill for a crack growth rate were determined byfitting the data points to the power law curve (Eq. (1))

using standard software of linear regression analysis.

2.6. Fracture ToughnessTest

The fracture toughness of the weld metal was also

studied using C(T) specimen (Fig. 3), having notch orl-

entation a]ong the dlrection of weldlng. The specimens

weremachinedas per the specification of ASTME399-83.The test wascarried out also by confirming the standardtest method in accordance to ASTME399-83 usingfatigue pre-cracked specimen, having crack length a/ W=0.5, under a fully automatedclosed loop servo-hydraulictesting machine. The maximumload at the end of the

pre-cracking was 12.5kN. A MTSbased computersoftware wasused to conduct the test and store the test

results of load vs, displacement across the notch. The10ad PQcorresponding to the 5o/o apparent incrernent ofthe crack extension wasestablished by drawing a secantline at a specified deviatlon from the linear portion ofthe record. Thecrack length wasmeasuredafter fracture

of the specimens to the nearest of 0.50/0 at the threepoints such as at the centre of the crack front, midwaybetween the centre of the crack front and end of thecrack front on fracture surface of the specimen. Anaverage of this three measurementsof crack length andthe load PQwere used for evaluation of the plain-strain

fracture toughness KQas follows.

KQ=(PQ/BW1/2)..f(a/W) ..................(2)

where

./'(a/ W)= [(2 +a/ W){O.866+4.64a/ W-13.32(a/ W)2

+ 14.72(a/ W)3- 5.6(a/ W)4}]/[ I - (a/ W)]3/2.(3)

PQ=Load (kN)

B=Specimenthickness (cm)

W=Specimenwidth (cm)

a=Crack length (cm)

Theevaluated KQwasverified for valid Klc by estimating

B or a;~2.5 (KQ/aYs)2, where (TYS is 0.20/0 off-set yield

strength of the weld deposlt. It was observed that the

KQrs not equai to Klc as It does not fulfil the conditionof thickness requirement of the specimen, which is used

as less than the minimumthlckness required for obtaining

a valid Klc' Thus, the fracture toughness properties ofthe weld has been studied in terms of the KQ.

3. Results and Discussion

3.1. Chemical Composition of the WeldChemical composition of the weld metal deposited at

different energy input, varied with a change in weldlngcurrent and speed, has been shownin Tab]e 4. The table

depicts that the weld metal confirms the chemica] com-position of a HSLAsteel and the change in welding

energy input is havlng insignificant infiuence on chemi-

(c,) 1998 ISIJ

ISIJ International, Vol. 38 (1998), No. 12

Table 4. Chemical composition of the weld deposit.

Welding condition

Energy Input (kJ/mm)

Chemical composition (wt"/.)

C MIl Si Cl' Ni Mo P S

l .68

l .89

2, l2, 16

2,52

O13

0.13

O.13

0.13

0.12

l 75

l .80

l .70

1.80

l .86

0.34

0.35

0.33

0.35

0.36

o.97

0.95

0.900.95

0.96

2.03

l .98

1.93

l .98

2.06

O48 0.01 6 0.0050.48 0.019 O0060.45 0,0 15 0.0060.48 0.019 0.0050.49 0.0 17 0.005

Fig. 5.

Fig. 4. Microstructure of the multipass weld deposited atdifferent energy input of (a) 168 kJ/mm, (b)

l ,89 kJ/mmand (c) 2. IkJ/mm.causedby the variationIn welding current at a given welding speed and arcvoltage as mentioned in Table 3,

cal composition of the weld. This is because the small

amotmt of dilution of base materia] through the weldbeads deposited at the groove surface becomespracti-cally negligible to create an appreclable variation in

chemical composition of a comparatively large massofthe multipass weld.

3.2. Microstructure of the WeldMicrostructure of the welds prepared at different en-

ergy input of 1.68, 1.89 and 2,1 kJ/mm, resulting froma change in welding current to 400, 450 and 500A re-spectively at a given welding speed of 40cm/min, hasbeen shown in Figs. 4(a), 4(b) and 4(c) respectlvely.

Similarly at different energy input of 2,16 and 2.52

Cc', 1998 ISIJ 1382

Microstructure of the multipass weld deposited atdifferent energy input of (a) 2,16kJ/mm and (b)

2.52 kJ/mm, caused by the variation in welding speedat agiven welding current and arc voltage as mentionedin Table 3.

kJ/mm, varied with a change in welding speed to 30'and 35cmlmin at a glven welding current of 450A, themicrostructure of the weld has been shownin Figs. 5(a)

and 5(b) respectively. The micrographs presented inFigs. 4 and 5 clearly reveal the presence of differentmicrostructural regions consist of coarse dendrite (CD)region, fine dendrite (FD) region, coarse grain reheat(CGR) reglon and fine grain reheat (FGR) region in

the matrix, which is commonlyknownmorphology ofa multipass weld.9) The typicai morphology of theseregions are moreclearly revealed in the micrographs ofhigher magnification as presented in Fig. 6. The mi-crographs shown in Figs. 4and 5also depict that thevariation in energy input changes the morphology andamountof different microstructural regions in the ma-trix. In general it is qualitatively marked that the in-

crease in energy input enhancesthe amountof dendritealong with comparative coarsening of its morphology.This has happened because the variation in weldingcurrent and speed changes the geometry of weld beadand weld therma] cycle which significantly affects the

morphologyandextent of the dendritic and reheat refinedreglons in the matrix.9)

3.3. Hardnessof the WeldHardness (HV IO) of the weld has been found to be

reduced almost lineally with the increase of energy in-

put as shown in Fig. 7. However, the hardness of a

ISIJ International. Vol. 38 (1998), No. 12

~~i:r.~ ~~"~

:~~;.'~~h.~

~i~t~*

: ':~t.~~.~

~i

S~.~

f

' *~J

IC,

~,

,, '(::*

~=*_~i*

,CD .-,~

~'

~.f.~

t

leO._uEv, (b) { '10O urra

*~(a)

z:r

tntnUJzoa:

IoocuZ

460

420

380

340

300

lAO

FINE DENDRITE :VHN= 448•48- 6S•83. In EGFINE CRAINREHEAT :VHN=430•1S-S8,98.InEGCOARSEDENDRITE :VHN=420•33-88•18 [nEGCOARSEGRAINREHEAT:VHN=375•39-58•23.inEG

+A

~\ A---_LL_____J

~\_ \.~\\_\~\,~\ ~~~\~~--~L

1,6

Fig. 6. Morphologyof dlfibrent microstructural reglons of themultipass weld revealed at a comparatively highermagnification; Coarse dendrite (CD); Fine dendrite(FD); Coarsegrain reheat (CGR)and Fine grain reheat(FGR).

Fig. 8.

1•8 2•O 2•2 2•4 2•6

EG,ENERGYINPUT (kJ/mm)

Effect of energy input on microhardness of differentmicrostructural regions of the weld.

2.8

3801200

zIcnulUJzaoc:

I

340

VHN= 410 -35 EG

3001•61•8 2•o 2•4 2•62•2

EG,ENERGYINPUT (kJlmm)Fig. 7. Effcct orenergy input on hardness ofthe wdd.

multipass weld is primarily governed by the microhard-ness of different microstructural regions, as stated earlier,

and their proportionate amountin the matrix. Thepro-portion of different mlcrostructural regions In a multi-

pass weld v'aries with the welding parameters govern-Ing the thermal cycle and geometry of weld deposit.9)

The microhardness of different microstructural regionshas also been foLmd to be reduced logarlthmically withthe increase of energy input (Fig. 8) causing a gross re-duction In hardness of the weld as depicted in Fig. 7.

It Is also observed in Fig. 8, that at a given energy inputthe fine dendrite (FD) region of the matrix possesseshighest hardness followed by a reduction in hardness ofother microstructural regions in order of fine grain reheat(FGR) region, coarse dendrite (CD) region and coarsegrain reheat (CGR)region. The difference in nature ofvarlation in hardness (HV 10) of the weld with the

energy input in reference to that observed in case ofmicrohardness of various microstructural regions ofthe weld possibly primarily attributed to the presenceof all kinds of microstructural regions in the matrix atdifferent proportion at different energy input.

3.4. Tensile Properties of the WeldJoint

The effect of energy input on tensile properties of theweld joint has been shownin Fig. 9. During tensile testthe specimens are found to fracture from the weld andthus the properties measuredby the tensile test primarilyrepresent the tenslle properties of weld deposit. In

~2IH,~)

zUJO~hU)

1000

800

600

400

A UTS(aTu) :l Y.S.(crlys):

e E[.

criu = 1254•18 - 90•909 EGoTys = 841•773 ~ 53•030 EcE[. = S•114 EG* 2•31S

ee

4s

35 •:~

ZOH

25 OZOJUJ

15

1•6

Fig. 9.

1•8 2• O 2•2 2•4 2-~EG

,ENERGYINPUT (kJlmm)

Effect of energy input on ultimate tensile strength, yieldstrength and elongation of the weld.

agreement to someearlier observatlons,1 1~ 13) the figureshows that the increase in energy input comparativelyreduces the ultimate tensile strength and yield strengthbut enhances the elongation of the weld. This maybeprimarily attributed to enhancementin amountof den-drite as well as coarsening of its morphology9'14~16)with the increase in energy input, as discussed above.The increase in coarse dendrite content of the matrixwith the increase of energy input makesit comparative-ly softer (Figs. 7 and 8) and consequently weaker in

agreement to the observations revealed in Fig. 9.

3.5. Fatigue Crack Growth Properties of the WeldThe fatigue crack growth rate characteristics of the

welds deposited at different energy input has been com-pared as shown in Fig. 10. The figure depicts thatthe rate of increase of the fatigue crack growth rate(da/dN) with the enhancementof stress intensity factor

range (AK) significantly reduces with the increase of

energy input. The figure also showsthat at a given stressintensity factor range the increase in energy input mark-edly enhances the fatigue crack growth rate especially

at the lower range of AK of the order of about

1383 C 1998 ISIJ

ISIJ International, Vol. 38 (1 998), No. 12

~~~:

EEz1:'

\e~

ld2

ld3

10~4

ld5

ld6

1d7

10

ENERGYINPUTe 1•68kJ/mmA 1•89kJ/mmI 2•10kJ/mm• 2•16kJ/mmo 2•52kJ/mm

.~il'- /71' '

-fl:/ '

l'

~i~

,

10~2

10~3

10~4

1J5

10~6

10

10~7

DELTA(K),(NPq/?Fr )

The relationship of fatigue crack growth rate wlth

stress intensity factor range for the weld prepared at

different energy input.

ov:~v\EEZ\c'

~

Fig. lO.

6-e--- m

c

10~7

l0~3

e\ \

10~4

5

4

\

\ \ \ \ \cN\ \ \ \

l

I

I 10~8

10~9,,,

u>u\E10~10 E

L)

10~5

EI e \\

\e \\

10~6

1

l

JL

3

2

\ e\ \\ \m=8•431 -6•547. In EG

C= -44•542 x ell•28. EG

10~11

Al

A

11•6 I2•42•21•8 2•O

EG,ENERGYINPUT(kJ/mm)

2

Fig. Il.

A

~-- __J___ __- A

610~12

Effect of energy input on growth exponent ,n andconstant Cof the Paris law of fatigue crack growthrate of the weld.

l0-30 MPa~/~T.Thesebehaviours of the (da/dN) vs. AKrelationships with a change in energy input are further

corroborated by the logarithmic decrease of m andexponential increase of C, of their Paris law relationships,

with the increase of energy input as revealed in Fig. Il.

At a given AKthe influence of energy input on fatigue

crack growth rate characteristics of the weld, has beenshownin Fig. 12. The figure shows that at a given AKof 15 and 30MPa~/rmthe crack growth rate enhanceswith the increase of energy input but, the rate of increase

in crack growth rate reduces with the enhancementof

~lAK. However, at a further higher AKof 45 MPa mthe crack growth rate has been found to be loweredsignificantly with the increase of energy input.

Fatigue crack growth rate characteristic of a metallargely dependsupon its tensile properties. 14) Theincrease

ee

ee AK=15MPaVFf,A AK=30h,tPclVl~T

l AK=45MPavliT

5do/dN = 2•524x 10~~EG~3•989xI~ 5da/dN= 629Sxl0~3' EG~1•316 x

140~6

do/dN =1 4S5xlCr -4•223xlO~ .EG

Fig. 12.

6 2O 2•62•41•8 2•2

EG,ENERGY[NPUT (kJlmm)

At a given stress intensity factor range the effect of

energy input on fatigue crack growth rate of the weld,

in energy input has been found to reduce tensile strengthbut to enhanceductility of the weld and thus, it reducesthe Paris law exponent mof the crack growth rate withrespect to AKas shownin Fig. 11.

With the decrease of

energy input the increase in tensile strength of the weldprimarily resulted the significant increase of crack growthrate in it, especially at the hlgh AKof the order of45MPaJ~,, as shown in Fig. 13. The noted influence

of tensile strength on the crack growth rate also confirmthe earlier observation.14) However, the effect of tensile

strength on enhancementof crack growth rate in the

weld has been found to be more strongly of oppositenature, of decrease in crack growth rate with the increase

of tensile strength, as one goes on lowering the AKtoJ~T30MPa mand AKof 15MPa~/rm,as depicted also in

Fig. 13. These behaviours of crack growth rate withthe change in AKmaybe primarily attributed to somemicroscopic features of the matrix affecting the fracture

mechanics. Thesemicroscopic features maybe primarilyidentified as the extent of formation of cleavage facets

and/or initiation, growth and coalescence of micro-voidin the matrix under various test conditions.14'17) At a

J~10werAKof the order of 15MPa mthe appearanceofcleavage like facets in the matrix largely dominates the

fracture mechanism, the extent of which reduces withstrengthening of matrix by the increase of tensile strength.

Thus, at a lower AKthe increase in tensile strength

reduces the fatigue crack growth rate significantly. But,

at higher AKof the order of 45MPa~/rmthe micro-void coalescence causing the plastic blunting of the cracktip possibly primarily controls the fatigue crack growthrate, which occurs comparatively moreat higher tensile

strength of the weld causing an increase in fatigue crack

@1998 ISIJ 1384

ISIJ International, Vol. 38 (1 998), No. 12

10~3

;//~/'~~=r

_A A~ ~lr--

10~4

AA

:~u\EEz~\a e

10~S

e

e AK=15~lPaV~T

A AK=30h,lPuvifrI AK=4SMPaViTe da/dN=2,808xl0~4-2•S18xl0~7alu

L do/dN=7•107xl0~4-5•457xl0~7ol e

10~61 da/dN=4•890xlO~SoTu~4•S30xl0~3

1020 1040 1060 1080 1100

a~u (MPa)

100

~I~ gO=cs

x801

Ka 74667 14496 i~ EG

ee

Fig. 14.

•6 1•8 2•4 2•62•O 2•2

EG, ENERCYINPUT (kJlmm)

Effect of energy input on fracture toughness of theweld.

95

~~ 85zoy

Ka=171•930 - 0•081 aTU

Fig, 13. At a given stress intensity factor range the effect ofultimate tcnsile strength on fatigue crack growth rateof the weld,

growth rate.17) The fracture mechanismat AKof theV~~order of 30MPam possibly lies in the transition re-

gion of both the mechanisrns reversibly affecting thefatigue crack growth rate and thus becamepractically

independent of tensile strength of the weld.In this context it maybe noted that, although the rate

of fatigue crack propagation is largely a function of microlevel features of the matrix, but In the present case ofcharacterisation of fatigue crack growth properties of themultipass we]d it was found appropriate to correlate it

with the broad basedqualitative microstructural featuresof the weld, affecting its tensile strength, because ofhighly heterogeneous nature of distribution of FD, CD,FGRandCGRregions in respect to the crack path in thematrix. However, the comparative influence of the mi-crostructure and tensi]e strength of the weld on theoverall rate of propagation of the fatigue crack in it

maybe studied by quantitative analysis of the micro-structural features of the weld, which was beyond the

scope of the present investigation.

Besides a]1 the aspects dlscussed above the presenceof residual stresses in weld deposit also affects the behav-iour of fatigue crack propagation by modification ofstress field at the crack tip. But, in the present studies

on fatigue crack growth characteristics of the weld therole of residual stresses has not been considered becausethe residual stresses in the interior layers of multipasswelds are generally released significantly by thermo-mechanical processes. 18)

3.6. Fracture Toughnessof the WeldThe effect of energy input on plain-strain fracture

1385

7SI0201040 1060 1080 1100

oTiu (MPa)Fig. 15. Effect of ultimate tensile strength on fracture

toughness of the weld.

toughness (KQ) of the weld has been shownin Fig. 14.

The figure showsthat the increase in energy input of theorder of 0.84kJ/mm relatively enhances the fracturetoughness of the we]d by about 6MPavrm

.The en-

hancementof fracture toughness with energy input is

primari]y caused by the decrease in tensile strength ofthe weld as revealed in Fig. 15. The figure depicts thatthe increase of fracture toughness of the weld by this

~ramountofabout 6MPam is caused by a decrease inits (T** of the order of 70MPa. This behaviour is in

agreement to an earlier observationl6) and also in theline of basic relationships of fracture toughness with themorphological nature affecting the flow characteristicsof a matrix. An increase in susceptibility to plastic flowwlth a decrease in tensile strength enhances the fracturetoughness of a matrix,16,19) which maybe the primaryreason that uitimately governed the fracture toughnessof the weld in present investigation.

4. Conclusions

A variation in welding energy input affects thehardness, tensile properties, fatigue crack growth rate(da/dN) characteristics and fracture toughness (KQ) ofmultipass flux cored submergedarc weld primarily dueto its influence on morphology of the weld consists ofdifferent amountof coarse and fine dendrites and coarseand fine reheat refined grains. The increase in energyinput reduces the hardness, ultimate tensile strength andyield strength but enhances the ductility and KQof theweld. However, the increase in energy input has beenfound to reduce the da/dN at higher AKof the order of

J~~45 MPa mbut to enhancethe samemoresignificantly

with the lowering of AKto 30 and 15 MPa~/rm.TheKQ

and da/dN of the weld are found to be well correlated

with its tensile strength, where an increase in tensile

strength reduces the KQbut enhances the da/dN at AK

@1998 ISIJ

ISIJ International, Vol.

higher than 30 MPaJHT.

REFERENCESl) P. T. Oldland. C. WRamsay.D. K. Matlock and D. L. Olson:

We!d. J., 68 (1989), No. 3, 158-s.

'_) J A. Gianetto, N. J. Smith. J. T. McGratt and J. T. Bowker:Wc'lc!. J., 71 (1992). No. I l, 407-s.

3) F. Matsuda, Z. Li. P. Bernasovsky, K. Ishihara and H. Okada:We!cl. Wo,'!d. 29 (1991), No. 9/10, 307.

4) P Youngyuth. P. K. Ghosh, P. C. Gupta. A. K. PatwardhanandSatya Prakash: Inl. J. Joi,1. Mater., 5(1993), No. 5, 31.

5) P. Youngyuth. P. K. Ghosh. P. C. Gupta. A. K. Patwardhanand Satya P]'akash: Mater. T,'ans. JIM, 34 (1993), No. 6. 533.

6) P. Yongyuth. P. K. Ghosh, P. C Gupta, A. K. PatwardhanandSatya Prakash: Int. J. Join Maier., 7(1995), No. 2/3, 87.

7) P. K. Ghosh, P. NageshBabuand P. C. Gupta: ISIJ 1,7!., 34(1994). No. 3, 280.

8) T. Kunio and K Yamada: Microstructural Aspect of the

Threshold Condition for Non Propagating Fatigue Cracks in

Martensitic-ferritlc Structure, ed, by Fong. Fatigue Mechanism.ASTMSTP-675, (1979).

9) P. K. Ghosh, P. Youngyuth. P. C. Gupta. A. K. Patwardhan

38 (1 998),

lO)

l 1)

12)

13)

l4)

i5)

16)

l7)

18)

l9)

No. 12

and Satya Prakash: ISIJ In/., 35 (1995), No. l, 63.

S. T. Rolfe and J. M. Barsom: Fracture and Fatigue Control in

Structures, Prentice-Hall Inc., Englewood Cliffs, NewJersey,

(1997).

G. M. Evans: We!d. J., 61 (198•_), No. 4, 125-s.

G. M. Evans: Wc'!d. J., 62 (1983), No, Il. 313-s.

K K. Das, G. L. Datta and B. Sasmal: Proc, ofSymp. on recentdevelopment in welding metallurgy and process, BARC,Bombay,(1985), 36.

H M. Hussain, P. K. Ghosh, P. C. Gupta and N. B. Potluri:T,'ans. Inc!ia,1 Ins!. Met., 50 (1997), No. 4, 275.

H. M. Hussain, P. K Ghosh, P. C. Gupta and N. B. Potluri:

WeJd. J., 75 (1996), No. 7, 209-s.

H. M. Hussain, P. K Ghosh, P. C Gupta and N. B. Potluri:hl!. J. Joi,1. Ma!er., Communicated,(1998).

R. W. Hertzberg: Deformation and fracture mechanics ofengineering materials. 3rd ed., John Wiley & Sons, NewYork,(1989), 530.

D. Radaj: He'at Effects of Welding, Springer-Verlag, Berlin.

(1992), 9.

G. E. Dieter: Mechanical Metallurgy, 2nd ed., McGrawHill

BookCo., NewYork, (1976), 273.

@1998 ISIJ 1386