Recent Topics of Welding Metallurgy Relating to Hot Cracking and Embrittlement in Iron and...

Recent Topics of Welding MetallurgRecent Topics of Welding Metallurgy Relating to Hot Cracking y Relating to Hot Cracking

and Embrittlement in Iron and Nickand Embrittlement in Iron and Nickel-base Alloysel-base Alloys

Lab. Material Joining Process

Kazutoshi Nishimoto

Department of Manufacturing ScienceGraduate School of Engineering

Osaka University

Osaka University

Contents


1. Background

2. Prediction of Degree of Embrittlement　　■ 475 ℃ Embrittlement　　■ Sigma Phase Embrittlement3. Mechanism of Weld Cracking　　■ Solidification Cracking in Laser Welding　　■ Ductility-dip Cracking4. Summary

Osaka University

Prevalent problems in welds of iron-base and nickel-base alloys


Hot cracking

Cold cracking

σ phase embrittlement

Embrittlement by grain coarsened ・ 475℃ embrittlement

Nie

q＝

%N

i＋30

×%C

＋0.

5×%

Mn

Creq ＝ %Cr ＋ %Mo ＋ 0.5×%Si ＋ 0.5×%Nb

δferrite content

Osaka University

The use of new alloys or new welding processes

Need for researches to understand their response to these problems.

■■ New welding processes such as laser welding may cause changes in a susceptibility to weld cracking that requires further investigation. ■■ Invar alloy which has recently become widely used in cryogenic plants, is found sensitive to hot cracking, but its mechanism is not clarified yet.⇒

Prevalent problems in welds of iron-base and nickel-base alloys

Hot cracking

Cold cracking

σ phase embrittlement

Embrittlement by grain coarsened ・ 475℃ embrittlementN

ieq

＝%

Ni＋

30×%

C＋

0.5×

%M

n

Creq ＝ %Cr ＋ %Mo ＋ 0.5×%Si ＋ 0.5×%Nb

δferrite content

Osaka University

The use of new alloys or new welding processes

Need for researches to understand their response to these problems.

■■ Embrittlement is also a serious problem in weldments of especially ferritic or duplex stainless steels. ■■ Although many investigations have been conducted into the material behavior producing embrittlement, rather few of these are useful for predicting the degree of embrittlement of the alloys during welding and/or in post-heat treatment. ⇒

Contents


1. Background


Osaka University


Osaka University

Spinodal Phase Decomposition in Chromium Containing Iron base Alloys

5nm

Cα’’

Cα’

C 0

CCα’’

Cα’

C 0

Nucleation and growth

Spinodal decomposition

Initial Middle Final

■ ■ When ferritic or duplex stainless steels containing more than about 20 % Cr are exposed to temperatures of 673-823K, they may suffer from "475 ℃ embrittlement", which somewhat limits the operating temperatures of their applications.

(a)

(b)

G

G

0 1CαCα’ Cα’’

α

α’+α’’

Spinodal decomposition

2G

C2 0

(a) Free energy curve(b) Phase diagram and Spinodal curve


Osaka University

Theoretical Analysis of Spinodal Decomposition during iso-thermal Process

cCr

t = D cCr

xx – 2

K 3cCr

x3

x

Cahn-Hilliard's non-linear diffusion equation

Fourier expression of diffusion equation

Q(h)t = – h 2

{ D0 + 2h22 K0 Q(h) + 12

D1 R(h)

+ 13

D2S(h) + 14

D3T(h) + }

– 2h4{K1U(h) + K2V(h) + K3W(h) + }

■■ The Cahn-Hilliard non-linear diffusion equation is one of the most useful approaches to spinodal phase decomposition.

■■ Recently, Miyazaki proposed a general formula with a Fourier expression of this non-linear diffusion equation. However, these approaches are meant to be used, for isothermal heat-treatment, and cannot be directly applied these to a phenomenon during the welding process.


Osaka University

The Method of Analysis for Spinodal Decomposition in thermal cycle process

cCr

t = D cCr

xx – 2

K 3cCr

x3

x

Cahn-Hilliard's non-linear diffusion equation (extended)

Interdiffusion coefficient

Gradient energy coefficient

D= M(Cr) 2G

cCr2 = M0c FecCr

2GcCr

2

K = 0 M(cCr) [(cCr,T) + {(cCr,T) /cCr}cCr]

Fig.1 Flow chart of computer simulation of phase decomposition

Input of parameters , cCr, thermal cycle

Input of initial composition-wave Q0(h)

Replacement of Fourier waves for convolution

Fourier transformation FFT

Rf = QfQf, Sf = RfQf, Tf = SfQf Uf = k3QfQf, Vf = UfQf, Wf = VfQf

Inverse Fourier transformation IFFT

Calculation of ŽQ(h)/Žt

Q(h)t+t = Q(h)t + {ŽQ(h)/Žt}tt

Output of Q(h) & graphing

Display?

YES

NO

Calculation of temp. & material constants

Completion of thermal cycle?

NO

End

YES

t = t + t

■■ Developed the method of analysis for the decomposition in thermal cycle process by extending the Cahn-Hilliard non-linear diffusion equation to this processes and applied it to a computer simulation of phase decomposition for 30Cr-2Mo steel.⇒


Osaka University

Two dimensional Evolution the Cr-rich phase induced by Spinodal Decomposition in 30Cr-2Mo steel

cCr

t = D cCr

xx – 2

K 3cCr

x3

x

■■ In the early stage of decomposition ,until the 2nd cycle, composition variations develop monotonically with time; however, they periodically fluctuate until the spinodal decomposition has further progressed.

■■ On the basis of thus calculated results, we tried to predict the degree of embrittlement due to the spinodal decomposition.


Osaka University

Theoretical approach for prediction of 475°C embrittlement in 30Cr-2Mo steel

Cut-through model (Mott-Nabarro's equation)

Hv : Hardness increment, R : Radius of precipitates, K : Constant, V : Volume fraction of precipitates, : Stiffness, N : Numbers of dislocation, : Misfit between matrix and precipitates

HV V 4 / 3N1/ 63/ 2

10ln

1

V

3/ 2

KR1/ 2V 4 / 3 ln1

V

3/ 2

vTE KR1/ 2V 4 / 3 ln1

V

3 / 2

Relationship between ΔHv and ΔvTE

■■ The change in hardness ΔHv due to the phase decomposition well agree with the value of R1/2V4/3{ln(1/V)}3/2 which is a hardenability parameter derived from Mott-Nabarro precipitation hardening theory. ■■ This fact suggests that hardening in this case follows the theory proposed by Mott-Nabarro.


Osaka University

Theoretical approach for prediction of 475°C embrittlement in 30Cr-2Mo steel

Cut-through model (Mott-Nabarro's equation)

Hv : Hardness increment, R : Radius of precipitates, K : Constant, V : Volume fraction of precipitates, : Stiffness, N : Numbers of dislocation, : Misfit between matrix and precipitates

HV V 4 / 3N1/ 63/ 2

10ln

1

V

3/ 2

KR1/ 2V 4 / 3 ln1

V

3/ 2

vTE KR1/ 2V 4 / 3 ln1

V

3 / 2

Relationship between ΔHv and ΔvTE

Relationship between R1/2V 4/3{ln(1/V)}3/2 and ΔvTE

■■ On the other hand, experimentally determined the functional relationship between the change in the transition temperature of the Charpy impact energy ΔvTE, and that in the Vickers hardness ΔHv.⇒


Osaka University

Example of the Calculated value of ΔvTE in the triple pass GTA weldment of 30Cr-2Mo steel

■■ The high value ofΔvTE due to the 475℃ embrittlement can be clearly recognized in the HAZ near the bottom of the plate on the 2nd/3rd pass welding, and it becomes dominant as the weld pass progresses.

■■ It can be also seen that the severely embrittled zone corresponds to the a position that has undergone triple heatings to about 800K.⇒

1st

pa

ss w

eld

ing

2n

d p

ass

we

ldin

g3

rd p

ass

we

ldin

g

Contents


1. Background


Osaka University

Embrittlement due to sigma phase Precipitation


Osaka University

■■ Sigma phase precipitation, which degrades not only mechanical properties but also corrosion resistance in alloys, is well known, but still a serious problem in stainless steel weldments. ⇒

Tem

pera

ture

(℃

)

Tem

pera

ture

(℉

)

WT.%Cr

NCr

800

700

600

500

400

0 0.2 0.4 0.6 0.8 1.0

700

800

900

1000

1100

1200

1300

1400

1500

100806040200

α α’

σ

α+σ σ+α’

α+α’

X X XX

X X XX

X X XX

Phase diagram of Iron-Chromium Alloy

Microstructures of super duplex stainless steels (heated at 1073 K for 1.8 ks)


Osaka University

NAS64 SAF2507 DP3W

50µm50µm 50µm

α

γ

α

γα

γ

Material

SAF2507

DP3W

NAS64

C

0.022

0.02

Si

0.36

0.32

Mn

0.90

0.51

P

0.021

0.023

S

0.0009

0.020

Ni

6.86

6.85

Cr

25.55

25.12

Mo

3.82

3.17

N

0.24

2.23

Cu

0.095

0.46

W

0.281

0.013 0.900.710.026 0.001 6.4524.60 3.28 0.16

PRE/ PREW*

41.9

42.7

37.9

■■ The microstructures of the super duplex stainless steels heated at 1073 K for 1.8ksec, which demonstrate sigma phase precipitation. ■■ Sigma phase precipitated mainly at delta/gamma boundaries in these steels.

Sigma phase precipitation Curves in super duplex stainless steels


Osaka University

■■ Sigma phase precipitation phenomenon follows the Johnson-Mehl type of kinetic equation in the case of weld metals of austenitic stainless steels.⇒

NAS64 SAF2507 DP3W

Aging time (s)

50454035302520151050101 102 103 104 102 103 104 103 104 105

Aging time (s) Aging time (s)

Are

a f

rac

tio

n o

f σ

ph

as

e (%

)

Aging temperature1073K1123K1173K1223K

Kinetics of Sigma Phase Precipitation ---Johnson-Mehl equation---


Osaka University

y = 1 – exp (– k tn) y = 1 – exp (– k tn) log ln

11 – y

= n log t + log k

log ln1

1 – y= n log t + log k

■■ A good linear relationship is found between the aging time and the fraction precipitated, which indicates that the sigma phase precipitation in duplex stainless steels also follows the Johnson-Mehl type kinetic equation. ⇒

NAS64 SAF2507 DP3W

Aging time (s)

1.0

101 102 103 104

Lo

gln

1/(1

-y) 0.5

0

-0.5

-1.0

-1.5

1.0

0.5

0

-0.5

-1.0

-1.5

-2.0

-2.5

1.0

0.5

0

-0.5

-1.0

-1.5

-2.0101 102 103 104 102 103 104 105

Aging time (s) Aging time (s)

NAS64 base metal

n=1.32

SAF2507 base metal DP3W base metal

n=1.62 n=0.879

Aging temperature1073K1123K1173K1223K

Prediction of the Amount of Sigma Phase during thermal cycle process by additivity rule


Osaka University

T1

T2

T3

Δt1 Δt2 Δt3time

Δt1 Δt2 Δt3

T1

T2

T3

Tem

per

atu

re

time

F = fmax(1){1-exp(k(1)Δt n)}　 + fmax (2){1-exp(k(2)Δt n)}　 + fmax (3){1-exp(k(3)Δt n)}　 + ・・・

F： Saturated volume of precipitation

f(t) = fmax(t){1-exp(k(t)Δt n)}

Additivity rule

■■ Applying the additivity rule and assuming that the saturated volume fraction of the sigma phase and the rate constant k vary with temperature, we can calculate the amount of sigma phase precipitated during an arbitrary thermal cycle with this equation.

F fdt fsat 1 exp ktn dt

k k0 exp Q RT

Based on the isothermal kinetics of the sigma phase precipitation

The amount of sigma phase precipitated in SAF2507 during two types of synthetic thermal cycles


Osaka University

Tem

per

atur

e (K

)

1200

Time (s)

373

900

C.R. : 5K/s

H.R. : 80K/s

(G.C.)(G.C.)

: Pattern A : Pattern B

600 120 180 240R.T.

Cycle-1 Cycle-2 Cycle-3

Cycle-1 Cycle-2

6

5

4

3

2

1

00 1 2 3 4 5 6 7 8 9 10

Area

frac

tion

of σ

phas

e (%

)

Number of thermal cycles

SAF2507 base metal

●,▲: Measured

Pattern A

Pattern B

Pattern A Pattern B

■■ The calculated curves agree fairly well with the measured results in both of the thermal cycles.

■■ This correspondence suggests that sigma phase precipitation in duplex stainless steels during the thermal cycle process can be predicted by this computation.

Relationship between the amount of sigma phase and the Charpy impact energy of duplex stainless steels aged at 1173K


Osaka University

■■ In each steel, the Charpy impact energy decreases drastically with increases in the amount of sigma phase.⇒

0 2 4 6 8 10 12 14 16

10

0

20

30

40

50

60

70

80

Area fraction of σ phase(%)

Imp

act

ab

sorb

ed

en

erg

y (J

)

Aging temperature : 1173K

NAS64

SAF2507

DP3W

Calculated amounts of the sigma phase and degree of embrittlement due to sigma phase precipitation


Osaka University

(a) Area fraction of sigma phase in multipass weldment

(b) Decrement in impact absorbed energy in multipass weldment

■■ The most embrittled zone locates in HAZ parallel to the weld interface and the level of the Charpy impact energy in this region is reduced by at most 17J from that of the unaged base metal.⇒

(Under the assumption)

(Under the assumption)

Contents


1. Background


Osaka University

Reasons for this enhancement of hot cracking susceptibility in laser welds

Osaka University

However, hot cracking susceptibility may be enhanced in some cases of the laser welding of stainless steels and nickel base alloys.

Generally speaking, decreasing the welding heat input is one of the most effective countermeasures for preventing hot cracking.

From this reason, laser welding is a preferable safeguard against this problem, because it can provide a lower welding heat input.

■■ Due to a characteristic shape of penetration in laser welds ; 'the key hole type of penetration.■■ Due to the rapid solidification and cooling that takes place during welding with an extremely low heat input.

There are two reasons for this enhancement:

Types and positions of hot cracking in laser welds


Osaka University

Center-line crack

Solidification crack at neck

Inter-granular crack at well

Hole by shrinking during solidification at bead center

Liquation crack at neck in HAZ

■■ In the case of the key hole type of penetration of laser welds, various types of cracking may be experienced.■■ These types of cracking is caused by the strain concentration at the specific part in the welds or in HAZ ⇒

due to a characteristic shape of penetration in laser welds.

Ｈ ot cracking susceptibility may be enhanced in laser welding

Mechanism of solidification cracking

Osaka University

Temperature

Str

ain

TL TS

(No cracking)

(Cracking)

Solidification brittleness temperature range (BTR) Local strain

溶接金属

Crack

Weld metal

■■ In general, solidification cracking will develop under the condition that the thermal strain subjected to the welds exceeds more than the critical value that it can bear. That is, solidification cracking will occur when the strain curve during cooling intersect with the solidification brittleness temperature range ; BTR

■■ In laser welding with a low heat input, the strain rate during cooling will increase, and consequently may enhance the hot cracking susceptibility.

due to the rapid solidification and cooling during welding .

Ｈ ot cracking susceptibility may be enhanced in laser welding

(the second reason)

Relationship between laser traveling velocity and total crack length in laser welds (SUS316L (P+S:0.04%))

Osaka University

60μm60μm

■■ Evidently, an increase in the laser traveling velocity produces a greater susceptibility to hot cracking in laser welds.■■ In addition, note that as the laser traveling velocity rises, the location where hot cracks occur changes from the dendrite boundaries to the center line of the welds.

Two major factors to influence hot cracking susceptibility in laser welds

Osaka University

Temperature

Str

ain

TLTS

(No cracking)

(Cracking )

Theoretical analyses of the liquidus and solidus temperatures during laser welding

Theoretical analyses of the liquidus and solidus temperatures during laser welding

Thermalelastic-plasticanalysis

Thermalelastic-plasticanalysis

BTR

Local strain

■■ The BTR in laser welds will vary because of changes in the liquidus and solidus temperatures due to the rapid solidification.

■■ The strain rate in laser welds will also enhance due to rapid cooling.

Estimation of BTR in laser surface melted region

Osaka University

Ts’

BTR

Temperature

Str

ain

(%)

TL Ts

BTR

TL

Str

ain

(%)

TsTemperature

TL’

BTR

Str

ain

(%)

TL’ Ts’Temperature

Arc welding Laser welding

Effect of rapid solidification

TL decreases due to supercooling Ts varied by micro-segregation of impurity elements

■■ Determined the BTR in laser welds by theoretical analyses of the liquidus and solidus temperatures based on the BTR for GTA welding obtained by the Varestraint test.


Osaka University

Ts’

BTR

Temperature

Str

ain

(%)

TL Ts

BTR

TL

Str

ain

(%)

TsTemperature

TL’

BTR

Str

ain

(%)




TL decreases due to supercooling Ts varied by micro-segregation of impurity elements 2

P2 D2V2 +

mC0 1 – K C

D 1 – 1 – K Iv PV + G = 0

modified KGT model

■■ In order to estimate the liqudus temperature, we calculated the dendrite tip temperature (T*), which corresponds to the liquidus temperature through calculation by the modified KGT model.


Osaka University

Ts’

BTR

Temperature

Str

ain

(%)

TL Ts

BTR

TL

Str

ain

(%)

TsTemperature

TL’

BTR

Str

ain

(%)




TL decreases due to supercoolingTL decreases due to supercooling Ts varied by micro-segregation of impurity elements

Ts varied by micro-segregation of impurity elements

J i = DCi + 1 – Ci

x

C jS knesC j1

L

Thermo-Calc@

■■ To determine the solidus temperature, we have conducted a theoretical analysis on the effect of the micro-segregation of impurity elements during welding on the solidus temperature by using the data-base of Thermo-calc.

Theoretical model for calculation for impurity elements segregation in solidification process

Osaka University

LS

CjS knesCj1

L

knes kes 1

Rv

x

2DS

Ci 2Dst

x2 2i1 i Ci1

B CiB i 1 Ci

B Ci 1B

Distribution of S at liquid/solid boundary

Non-equilibrium coefficient Knes

Diffusion in solid

kes ： Equilibrium coefficient ， Rv ： Solidification speed ， Ds ： Diffusion coefficient

12

i-1

N-1N

i i+1

S L

LS

Solidification

S c

once

ntra

tion

J i = DCi + 1 – Ci

x

Cs

■■ In this analysis,assumed the morphology of a dendrite to be a hexagonal column and evaluated distribution of the solute concentration with a one-dimensional diffusion model in which the solute diffused in the direction perpendicular to the grain boundary⇒

BTR calculated in laser welds of SUS316LOsaka University

P : 0.02% P : 0.03%

1680 1640 1600 1560 1520Temperature (K)

1.0

0.8

0.6

0.4

0.2

0S

trai

n (%

)

LTV (mm/s)204060

1680 1640 1600 1560 1520Temperature (K)

1.0

0.8

0.6

0.4

0.2

0

Str

ain

(%)

LTV (mm/s)204060

■■ The solidus temperature in the laser welds is found enhanced with a rise in the laser traveling velocity due to the increase in the solidification rate. ■■ On the other hand, the liqudus temperature in the laser welds decrease due to supercooling in laser welds⇒

Direction of the the strain analyzed at the surface of the welds


Osaka University

Analysis point and direction

Calculated by Quick Therm

θ=60°θ=35°

θ =50°

Laser traveling velocity : Increase

■■ The thermal strain is another important factor to consider the occurrence of cracking in welds

■■ Used the 3-dimensional thermal elastic-plastic software package "Quick Therm" to calculate the strain formed during welding in laser welds and analysed the strain which is perpendicular not only to the center line of the weld but also dendrite boundaries.

20mm/s 40mm/s 60mm/s

Local strain at the center perpendicular to laser scanning direction

Osaka UniversityLTV : Laser traveling velocity (mm/s)

Center-line1.2

1.0

0.8

0.6

0.4

0.2

01650 1600 1550

Temperature (K)

Str

ain

(%)

LTV=60

LTV=40

LTV=20

1.2

Dendrite boundary1.0

0.8

0.6

0.4

0.2

01650 1600 1550

Temperature (K)

Str

ain

(%)

LTV=60

LTV=40

LTV=20

■■ Examples of calculation which show that the change in the thermal strain occurring during solidification increases with increasing laser traveling velocities. ■■ In contrast, in the case of dendrite boundaries, the strain taking place with a laser traveling velocity of 40mm/s is larger than that under other conditions.⇒


(θ=35°)

(θ=50°)

(θ=60°)θ

Comparison between BTR and strain at the bead center perpendicular to laser traveling direction


Osaka University

0.02 ％ P+S0.04 ％

P+S

0

Str

ain

(%

)

1.2

1

0.8

0.6

0.4

0.2

Temperature (K)

LTV = 20 mm/s LTV = 40 mm/s LTV = 60 mm/s

LTV : Laser traveling velocity

1650 15501600 1650 15501600 1650 15501600

■■ Examine the possibility of solidification cracking in laser welds by superimposing plots of the BTR and the strain produced during cooling in laser welds. ■■When the laser traveling velocity is 40 or 60mm/s, the strain perpendicular to the center line of the welds crosses to the BTR, which means that solidification cracks will occur at the center of the welds in this laser traveling velocity range.

CrackingCracking

SUS316L

Comparison between BTR and strain perpendicular to dendrite growth direction at dendrite boundaries

Osaka University

0.02 ％ P+S0.04 ％

P+S

0

Str

ain

(%

)

1.2

1

0.8

0.6

0.4

0.2

Temperature (K)

LTV = 20 mm/s LTV = 40 mm/s LTV = 60 mm/s

LTV : Laser traveling velocity

1650 15501600 1650 15501600

Cracking

1650 15501600

■■ The strain curve estimated for dendrite boundaries crosses the BTR when the laser traveling velocity equals 40mm/s.■■ This result suggests that solidification cracks will occur at the dendrite boundaries in the welds in this laser traveling velocity range.⇒

SUS316L

Comparison between measured and theoretically calculated conditions to occur cracking

Osaka University

Laser traveling velocity (mm/s)

P+

S c

ont

en

t (m

ass

％)

10 40 50 603020

0.02

0.03

0.04

Center-line cracking (Calculated)Cracking at dendrite boundary (Calculated)

Crack（ Center-line ）Crack(Dendrite boundary)

No crack

■■ Good agreement between these two conditions determined by calculation and experimentals. ■■ These results suggest that the cause of solidification cracking in laser welds is actually the increase in the strain rate during solidification, in spite of the fact that the BTR becomes narrower due to rapid solidification.

SUS316L

Enhanced susceptibility due to solidification mode shift


Osaka University

J.C.Lippold, Weld. J., 73-6 (1994) 129s-139s

■■ There is another factor to be considered which may influence hot cracking susceptibility in austenitic stainless steels in laser welding. ■■ It is known that the solidification mode in austenitic stainless steel weld metals shifts from primarily ferrite to primarily austenite when the solidification rate becomes sufficiently high.

1.8

1000

100

10

1

0.11.5 2.0

Creq/Nieq

Sol

idifi

catio

n ra

te (

mm

/s)

Austenite(A)

Ferrite(F)

F/MA

FAAF

Arc welding YAG welding0.1

0.05

01.4 1.6

Cr/Ni-equivalentS

+P

+B

(m

ass%

)

Crack No crack

Enhanced susceptibility due to solidification mode shift


Osaka University

J.C.Lippold, Weld. J., 73-6 (1994) 129s-139s

■■ Laser welding with a low heat input can provide in some cases such solidification condition to cause solidification mode shift. ■■ Alloys solidified in primarily austenite mode is more sensitive than ones in primarily ferrite mode. ■■ This is another reason for the increased hot cracking susceptibility of stainless steels in laser welding. ⇒

1.8

1000

100

10

1

0.11.5 2.0

Creq/Nieq

Sol

idifi

catio

n ra

te (

mm

/s)

Austenite(A)

Ferrite(F)

F/MA

FAAF

Arc welding YAG welding0.1

0.05

01.4 1.6

Cr/Ni-equivalentS

+P

+B

(m

ass%

)

Crack No crack

Condition for transition of solidification mode

Osaka University

In the case of T*γ>T*

δ In the case of T*δ>T*

γ

T *= T L + (mv ,iC i*– m 0.iC0 ,i)

– 2 / R – V/ – GD / V

FA mode AF mode

■■ In general, the phase which has the higher dendrite tip temperature is more likely to be the primary phase on solidification. Therefore the solidification mode shift can be predicted if the dendrite tip temperature of each phase is known.

Theoretical model for dendrite growth


Osaka University

T = m C0 – C1 + 2R

2P2 D2V

2 +mC0 1 – K C

D 1 – 1 – K Iv PV + G = 0

R = 2

–mC0 1 – K cV

D 1 – 1 – K Iv P– G

T *= T L + (mv ,iC i*– m 0.iC0 ,i)

– 2 / R – V/ – GD / V

K ： Partition coefficientR ： Dendrite tip radiusV ： Dendrite growth velocityΔT ： Undercooling related to the tip radiusG ： Temperature gradientD ： Liquid interdiffusion coefficientP ： Peclet numberIv(P) ： Ivantsov's solutionξc ： Absolute stability coefficientmv,i ： Velocity dependent liquidus slopeΓ ： Gibbs-Thomson parameter

S.Fukumoto, W.Kurz, ISIJ Inter., 37-7 (1997) 677-684

S.Fukumoto, W.Kurz, ISIJ Inter., 38-1 (1998) 71-77

(modified KGT model)

Dendrite Tip Temperature ： T*

■■ Used the modified Kurz-Giovanola-Trivedi (KGT) model, which was extended to multicomponent alloys by Kurz in order to calculate the dendrite tip temperature.■■ According to the model, the dendrite tip radius, R, is expressed as a function of dendrite growth velocity, V, as shown in this equation .■■ For multicomponent alloys, the dendrite tip temperature, T*, is given by this equation. ⇒

Effect of dendrite growth velocity on dendrite tip temperature of ferrite and austenite


Osaka University

1750

1730

1710

1690

1670

16501×10-2 1×10-1 1×100 1×101 1×102

Dendrite growth velocity (mm/s)

Den

drite

tip

tem

pera

ture

(K

)

■■ The dendrite tip temperature in austenite rises above that in ferrite at dendrite growth velocities exceeding 0.9mm/s. ⇒

23Cr-9Ni-0.34N steel

Comparison of calculated solidification mode change with experimental results in laser welds of stainless steel


Osaka University

Bead center

AF mode

FA mode

FAAF

15

10

5

01.3 1.4 1.5

Creq/Nieq

Lase

r tr

avel

ing

velo

city

(m

m/s

)

Predicted condition to yield crack

■■ By using the above mentioned results, you can also predict the risk of hot cracking by calculation assuming that the solidification mode change from FA to AF will enhance cracking susceptibility.

■■ For instance, this figures show the theoretically predicted transition line from FA to AF at the center part of the weld metals for nitrogen containing austenitic stainless steels.⇒


Comparison of calculated solidification mode change with hot cracking susceptibility in laser welds of stainless steel


Osaka University

Bead center

Crack

No crack

FAAF

15

10

5

01.3 1.4 1.5

Creq/Nieq

Lase

r tr

avel

ing

velo

city

(m

m/s

)

Predicted condition to yield crack

■■ The condition to yield AF mode coincide with the condition to occur hot cracking.

■■ It means you can predict the risk of hot cracking through calculation of the mode shift from FA to AF even in laser weld. ⇒


Contents


1. Background


Osaka University

Mechanism of ductility-dip crack


Osaka University

Str

ain

Temperature

BTR DTR

ε1

ε2

ε3

Ductility-dip cracking

■■ Ductility-dip cracking can occur in various alloys which exhibit a loss of ductility below the solidus temperature, when they are subjected to a strain sufficient to produce cracking during cooling in welding.

■■ A ductility-dip crack formed in weld metals is normally very small, and is sometimes called a 'micro fissuring'. ⇒

Ductility curve

Mechanism of ductility-dip crack


Osaka University

Str

ain

Temperature

BTR DTR

ε1

ε2

ε3


■■ Recently Invar alloy has attracted special interest as a suitable material for cryogenic applications, such as fuel transport pipes due to its low thermal expansion coefficient and good toughness at low temperatures.

■■ Invar alloy is however, found to be very susceptible to micro fissuring in multi-pass welds of heavy sectioned pipes.

But, the mechanism of micro fissuring in the weld metals of Invar alloys is still uncertain. ⇒

Ductility curve

Surface of weld metal of Invar alloy after triple bead longitudinal Varestraint

Osaka University

■■ Many cracks in original weld pass which was reheated by a subsequent pass. Cracks preferentially occurred along the columnar grains and/or around the center line of the original weld bead.

Fe-36Ni alloy

Effect of weld thermal cycles on cracking susceptibility


Osaka University

Double-beadTriple-bead

Tot

al c

rack

leng

th (

mm

)

0

1

2

3

1100〜1200

1000〜1100

900〜1000

900〜800Peak temperature range in HAZ (K)

■■ The total lengths of the cracks in the triple-bead test were much greater than those in the double-bead test. We can see that this tendency predominated in the peak temperature range between 1000K and 1100K.

Fe-36Ni alloy (0.011%S)

Effect of S on susceptibility to ductility-dip cracking


Osaka University

Double-bead Varestraint testDuctility-dip crack in first bead

Augmented strain 1.6%2.4%

80

60

40

20

00 0.005 0.010 0.015

S content (%)

Tot

al c

rack

leng

th (

mm

)

■■ The total length of the cracks grew as the sulfur content in the samples increased.

■■ This experimental result demonstrate that sulfur is evidently detrimental to cracking susceptibility in the weld metal.

Fe-36Ni alloy

The result of Auger analysis conducted on the fractured surface in the multi-pass weld metal


Osaka University

Effect of thermal cycles at peak temperature of 1000K on S concentrationTwice ： Average 7% 　⇨　 Three times ： Average 9%

AES spectrum

S

C O

Fe

Ni

100 200 300 400 500 600 700 800 900 1000Kinetic energy (eV)

c/s

-1200

-1000

-800

-600

-400

-200

0

200

400

600

800

■■Sulfur is segregated on its surface. Moreover, the amount of sulfur on the grain boundary increase with increase of welding pass. ⇒

Fe-36Ni alloy

Method of calculation for S segregation at grain boundary during multi-pass weld thermal cycles


Osaka University

Stage I : Solidification process

Stage II : Cooling process (after solidification)

Melting point

Melting point■■ The concentration of sulfur at the grain boundary was analyzed for two stages:

■■ That is,

Stage I: the solidification segregation during welding, Stage II: the grain boundary segregation during the cooling and reheating by the subsequent weld passes.



Osaka University

CjS knesCj1

L

knes kes 1

Rv

x

2DS

Distribution of S at liquid/solid boundary

kes ： Equilibrium coefficient ， Rv ： Solidification speed ， Ds ： Diffusion coefficient

Stage I : Solidification process

Melting point

Melting point

■■ The same model as the one described in the previous section was adopted for the solidification segregation during welding. thus, the concentration of sulfur after completion of solidification was calculated by this equation.



Osaka University

Csgb(t) Cs

gb(0)

Csgb() Cs

gb(0)1 exp

4Dst

12d2

erfc

2 Dst

1 d

CX 2DXt

x2 2i 1 i CX i1

B CX i

B i 1 CX i

B CX i 1B

Solute concentration change

Boundary condition

Stage II : Cooling process (after solidification)

Melting point

Melting point

C ： Concentration of solute, vacancy, complex ，k ： constant ， Ev ： Vacancy forming energy ，Ec ： Binding energy between vacancy and solute

■■ As for the grain boundary segregation during the cooling and reheating by subsequent weld passes, we calculated the change in the sulfur concentration at the grain boundary after solidification by the these equations based on the equilibrium segregation theory.

[ Peak temp. of 3rd pass =1000K ]

Peak temp. of 2nd pass = 1200K Peak temp. of 2nd pass = 1000K Peak temp. of 2nd pass = 800K

Temperature

0 20 40 60 80 100 120 140Time(s)

Temperature

0 20 40 60 80 100 120 140Time(s)

400600800

10001200140016001800

Te

mp

era

ture

(K)

Temperature

0.0

1.0

2.0

3.0

4.0

S c

on

ce

ntr

ati

on

(%)

0 20 40 60 80 100 120 140Time(s)

1200K

1000K 1000K 1000K800K

1000K

IN4 (0.011%S)

IN3 (0.004%S)

IN0 (0.001%S)

Cgbi (IN4)Cgbi (IN3)Cgbi (IN0)



IN4 (0.011%S)IN4 (0.011%S)

IN3 (0.004%S) IN3 (0.004%S)

IN0 (0.001%S) IN0 (0.001%S)

Osaka University

The concentration of S at grain boundaries calculated for multi-pass weld thermal cycles

■■ In the solidification process, the sulfur concentration in the liquid phase rose as the solidification proceeded.■■ During cooling, the sulfur concentration at the grain boundaries first rapidly fell, and then increased again with correspondent to the increase of its equilibrium concentration at the grain boundaries.

Osaka University

[ Peak temp. of 3rd pass =1000K ]

Peak temp. of 2nd pass = 1200K Peak temp. of 2nd pass = 1000K Peak temp. of 2nd pass = 800K

Temperature

0 20 40 60 80 100 120 140Time(s)

Temperature

0 20 40 60 80 100 120 140Time(s)

400600800

10001200140016001800

Te

mp

era

ture

(K)

Temperature

0.0

1.0

2.0

3.0

4.0

S c

on

ce

ntr

ati

on

(%)

0 20 40 60 80 100 120 140Time(s)

1200K

1000K 1000K 1000K800K

1000K

IN4 (0.011%S)

IN3 (0.004%S)

IN0 (0.001%S)




IN4 (0.011%S)IN4 (0.011%S)

IN3 (0.004%S) IN3 (0.004%S)

IN0 (0.001%S) IN0 (0.001%S)

The concentration of S at grain boundaries calculated for multi-pass weld thermal cycles

■■ In the reheating process, the grain boundary sulfur concentrations decreased at temperatures above about 1100 K, due to the reduced equilibrium concentration. ■■ However, in the reheating process in which the peak temperature was less than 1000 K, the grain boundary concentration of sulfur increased again with elevations in its equilibrium concentrations.


Osaka University

Mechanism of ductility-dip cracking in multi-pass weld thermal cycles

Grain boundary

1st pass 2nd pass 3rd pass

■■ The theoretical analysis makes it clear;

1) the sulfur segregation at the grain boundary in multi-pass welds was enhanced by multi-pass weld thermal cycles

2) became dominant when a weld metal was reheated twice at a temperature range between 900 and 1100 K. Varestraint test results have shown that this region was the most susceptible to cracking.

Grain boundary segregation of sulfur

Mechanism of ductility-dip cracking in multi-pass weld thermal cycles


Osaka University

■■ These results suggest that the cause of cracking in the multi-pass welds of Inver alloy can be attributed to decrease in the critical strain of DTR caused by grain boundary weakening due to sulfur segregation

■■ which has been accelerated by the multi-pass weld thermal cycles.

Grain boundary

1st 2nd 3rd

Grain boundary segregation of sulfur

Str

ain

Temperature

BTR DTRDuctility curve

Strain curve


1st

2nd

3rd

Contents


1. Background


Osaka University

SummaryOsaka University

DevelopmentDevelopment

■■Quantitative prediction of microstructure and properties

■■Precise understanding of mechanism

Better Better performance performance of jointof joint

Better Better performance performance of jointof joint

Calculation codes

■■ In the last few decade,mathematical approaches using computer technologies in the welding metallurgy have significantly contributed to its recent developments. Such approaches have enabled a more precise understanding of welding metallurgical phenomena and a more accurate comprehension of the mechanism for weld defects, including weld cracking through a visualization of the results. ■■ Moreover, mathematical approaches have provided important information to control weld defects, which ensures better weldment performance and the reliability of welded joints. Further development of mathematical modeling should be more encouraged.


Osaka University

Thank you for kind your attention!Thank you for kind your attention!

The method for examination of reheat cracking susceptibility

Osaka University


　　　　＜ Test conditions ＞Welding current ： 150AWelding voltage ： 13VWelding speed ： 1.67mm/sAugmented strain ： 1.6, 2.4, 3.6%

1st bead

2nd bead

2nd bead

1st bead

3rd bead

3-bead Varestraint test2-bead Varestraint test

Surface appearance of laser surface melted regions

Osaka University


1mm 1mm

60μm 60μm

Local strain at the center perpendicular to laser scanning directionOsaka University

Lab. of Material Joining Process

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Ç™Ç±ÇÃÉsÉNÉ`ÉÉÇ¾å©ÇÈÇ…ÇÕïKóvÇ ÇÅBLs

er s

cann

ing

dire

ctio

n

Melted region

Center element

The increment of local strain at the center part in laser surface melted region during solidification increases with increasing laser traveling velocities.

Temperature (K)

0

0.2

0.4

0.6

0.8

1

1.2

1690 1650 1610 1570 1530

LTV=20

LTV=40

LTV=60

Str

ain

(％

)Temperature (K)

LTV : Laser traveling velocity (mm/s)

Local strain at dendrite boundaries perpendicular to dendrite growth directionOsaka University



Ç™Ç±ÇÃÉsÉNÉ`ÉÉÇ¾å©ÇÈÇ…ÇÕïKóvÇ ÇÅB






θ θ θ

Lase

r sc

anni

ng d

irect

ion

Element located at 0.3mm from center

Melted region

Perpendicular to dendrite growth direction

Local strain at dendrite boundaries perpendicular to dendrite growth direction (Calculated)Osaka University


θθ

θ


The increment of local strain in the condition of laser traveling velocity of 40mm/s is larger than the strain in other conditions. That is, in this case, not only the laser traveling velocity but also the dendrite growth direction affects the local strain at dendrite boundaries.

LTV=20

LTV=40

LTV=60

0

0.2

0.4

0.6

0.8

1

1.2

1690 1650 1610 1570 1530Temperature (K)

Str

ain

(％

)

Temperature (K)

LTV : Laser traveling velocity (mm/s)

Recent Topics of Welding Metallurgy Relating to Hot Cracking and Embrittlement in Iron and...

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Transcript of Recent Topics of Welding Metallurgy Relating to Hot Cracking and Embrittlement in Iron and...