repair process for turbine casting

Journal of Metallurgical Engineering, Volume 2 Issue 2, April 2013 www.me‐journal.org

61

Repair Process for Turbine Casting SGK.Manikandan1*, D.Sivakumar2, K.Prasad Rao3, M.Kamaraj4 1,2Indian Space Research Organization, Trivandrum 3,4Indian Institute of Technology Madras, Chennai

*[email protected]; [email protected]; [email protected]; [email protected]

Abstract

Alloy706 is used as exhaust casing material in gas turbines of

cryogenic rocket engines in view of its excellent properties.

The defective zones in super‐alloy castings usually are repair

welded in solution treated condition to obtain good room

temperature and high temperature properties. However a

need came‐up to carry out in‐situ repair welding in one of

the gas turbine castings made of aged Alloy 706 in cryogenic

rocket engine with a limitation that the component cannot be

subjected to any heat treatment cycle subsequent to the

repair welding. Hence it became necessary to develop

process and process parameters with reduced strength

mismatch and HAZ free from micro‐fissures in order to

obtain required mechanical properties and further high

temperature service associated with high pressure hydrogen

environment. This has been approached with process control

using reduced heat input and alloying control.In order to

reduce the heat input, two processes such as GTA welding

and GTA braze welding have been adopted with multi pass

technique.

Keywords

GTA welding, GTA Braze welding, Heat input, Alloy 706 casting

Introduction

The nickel‐iron based super‐alloy, Alloy706 (706Eq) is

used as exhaust casing material in gas turbines of

cryogenic rocket engines because of its good

mechanical properties up to intermediate

temperatures (~700°C) and excellent manufacturability.

This precipitation‐strengthened alloy contains FCC

austenitic matrix (γ‐phase) which is a Ni‐Fe‐Cr solid

solution. The predominant strengthening precipitate is

the coherent, ordered A3B type γ’ phase (Ni3Al‐L12

crystal structure) and γ” phase (Ni3Nb‐DO22

structure). The volume fraction of γ’ phase is more

than that of γ” phase due to higher Ti content. A stable

δ phase (Ni3Nb‐DOa structure) and ordered A3B type

η phase (Ni3Ti‐ordered hcp structure‐DO24 structure)

can also be formed in 706Eq depending on processing

conditions. Detailed micro‐structural analysis of Alloy

706 has been reported by Moll et. al., and the heat

treatment behaviour has been discussed by Heck et.

al.,. Super‐alloy castings are usually repair welded in

solution treated condition to obtain good room

temperature and high temperature properties.

However, in this case there existed in a need to carry

out in‐situ repair welding in one of the gas turbine

castings of aged 706Eq in cryogenic rocket engine with

a limitation that the component cannot be subjected to

any heat treatment cycle subsequent to the repair

welding. Since the casting consists of rotating elements

and soft seals nearby with close tolerances, no in‐situ

heat treatment operations can be done. In the repaired

zone of casting, there will be a mismatch in hardness

value between adjacent base material and fusion zone

(FZ) with HAZ.

Moreover, the cast material is more prone to micro‐

fissuring due to segregation. The solvus temperature

of Laves phase in 706Eq is 1020°C. Optimized

mechanical properties were obtained with a maximum

solution treatment temperature of 1000°C. And hence

Laves phase are not fully dissolved. The presence of

MC type carbides and Laves phase in the inter‐denritic

region of the casting leads to the formation of inter‐

granular eutectic type liquid during welding due to

rapid heating. This is a favorable condition for HAZ

micro‐fissuring. Hence there is a need to develop

process and process parameters with reduced strength

mismatch and HAZ free from micro‐fissures in order

to obtain the required mechanical properties and

further high temperature service associated with high

pressure hydrogen environment. C.L.Ou et. al., has

described the advantages of repair brazing for age

hardenable stainless steel to fix shallow cracks. Since

the defective zone present in the rocket engine consists

of various materials with range of melting

temperature which cannot be furnace brazed.

J.H.G.Mattheij et. al., described the repair process for

various types of defects and/or damages in the gas

turbines in which brazing and allied processes are

considered to be standard repair/regeneration

technique. Hence an alternate method of braze

welding process is selected.

www.me‐journal.org Journal of Metallurgical Engineering, Volume 2 Issue 2, April 2013

62

Repair welding simulation experiments have been

conducted in two modes with multi‐pass Gas

Tungsten Arc (GTA) welding and GTA Braze welding

with 120 μm thick braze foil (by employing special

attachments to avoid oxidation of braze foil) using

rapid heat extraction by external heat sink. Slow but

moderate heating rate and rapid cooling rate have

been employed to avoid micro‐fissuring. Both

weldment and braze weldment were qualified to 100%

X‐ray radiography. Micrographs of GTA weldment

reveals HAZ micro‐fissuring while GTA braze

weldments are free from the same. Comparable

mechanical properties for both the cases have been

achieved at room temperature and at high

temperature (477°C). Ductility of braze weldment is

better by 31% than that of weldment at room

temperature. Strength mismatch has been observed to

be less in braze weldment than that in weldment. To

understand the mechanism, experiments on heat

transfer during process have been conducted and it

was found that the instantaneous cooling rate is

relatively higher while the instantaneous heating rate

is slow in GTA braze welding than that in GTA

welding. Slower heating rate ensures complete dis‐

solution of precipitates and thereby minimizing the

incipient melting of precipitates in HAZ, while rapid

cooling rate refines the micro‐structure

Based on the experimental studies on process and

process variables, process parameters have been

developed for braze welding. The same have been

successfully employed for the in‐situ repair welding of

aged 706Eq gas turbine castings in cryogenic rocket

engine. The braze weldment has been successfully

subjected to high temperature service associated with

high pressure hydrogen environment. This paper

describes about the simulation trials, characterization

of fusion zone, mechanical property evaluation and

implementation of process.

Experimental Setup

Repair welding simulation experiments have been

conducted in two modes with multi‐pass Gas

Tungsten Arc (GTA‐W) welding and GTA Braze

welding (GTA‐BW) in a 4mm thickness cast billet of

Alloy706 in heat treated condition. Details of the heat

treatment are as follows:

Homogenisation treatment at 1125°C±25°C

with 3+0.5Hrs holding time and then air cooled

Solution treatment at 1000°C±10°C with

3+0.5Hrs holding time and then air cooled

First step aging treatment at 750°C±10°C with

8+0.5Hrs holding time and then air cooled

Second step aging treatment at 650°C±10°C

with 8+0.5Hrs holding time and then air cooled

GTA‐BW process was employed with 120 μm thick

braze foil, by developing special attachments to avoid

oxidation of braze foil. Rapid heat extraction was

achieved through external heat sink in order to avoid

heat build‐up. Heating rate is rapid in welding

processes than that in casting processes. This leads to

incipient melting of MC type carbides and promotes

HAZ micro‐fissuring and/or liquation cracking. Hence

slow to moderate heating rate and relatively rapid

cooling rate have been employed to avoid micro‐

fissuring. The complete elimination of the defective

zone is not possible, since the rotating parts have very

minimum radial clearances and there exists in a

possibility of contaminants entering the turbine

housing, which leads to catastrophic failure. Thus a

back‐up of 0.83mm was retained in the 4mm thickness

casting shell as shown in Figure‐(i). A J type groove

was prepared so that the accessibility in root of weld is

better and bead width is also lower. This type of weld

joint configuration reduces shrinkage stresses, thereby

minimizing the susceptibility for HAZ micro‐fissuring.

FIG. (i) DETAILS OF WELD JOINT AND GROOVE

The repair welded samples were qualified with 100%

X‐ray radiography and Dye penetrant test. Transverse

section samples were prepared and polished with

standard procedure. After polishing, samples were

electrolytically etched with 10% oxalic acid at 5‐7VDC

for 2min. Material characterization techniques such as

optical metallography (OM) and scanning electron

microscopy (SEM) were used. Weldment alone was

separated from the fabricated specimen and X‐ray

diffraction analysis (XRD) was carried out in both

transverse direction and longitudinal direction using


63

Cu‐Kα radiation. In order to distinguish the texture

effect during solidification, XRD was carried out in

both longitudinal and transverse directions. To

understand the mechanism and heat build‐up

characteristics, heat transfer experiments have been

conducted during joining process. Scheme for

temperature measurement is shown in Figure‐(ii)a.

Twelve numbers of thermocouples were capacitor

discharge welded on the top surface of the cast plate.

The first 5 nos of thermocouples were positioned at a

distance of 2mm from the bead edge and the

subsequent 4nos and 3 nos of thermocouples were

placed 5mm and 10mm away from the first row of

thermocouples respectively. These temperature

measurements (T1 – T12) were used for studies by

cooling rate and temperature gradient. K type

thermocouples were fixed by capacitor discharge

welded on the surface. To avoid the temperature rise

due to radiative heat transfer from the welding arc,

thermocouples (TA1 to TA 12) were insulted with high

temperature, thermally insulative ceramic cement.

Thermal data were recorded with 1 sec sampling rate

using data logger.

Tensile test samples were prepared according to

ASTM E8M for room temperature and ASTM E21‐09

for high temperature (477°C). Details of tensile test

samples are shown in Figure‐(ii) b and c. Tensile

testing was carried out in the as welded condition as

per ASTM. Fractographic analysis of the tested

samples has been carried out.

FIG. (ii)A SCHEME FOR TEMPERATURE MEASUREMENT

FIG. (ii) B TENSILE TEST SAMPLE FOR 23°C

FIG. (ii) C TENSILE TEST SAMPLE FOR 477°C

Details of the chemical composition for base material,

filler metal and braze foil are as given in Table‐I.

TABLE I DETAILS OF CHEMICAL COMPOSITION

Description Alloy706 Braze foil Filler metal

C 0.04 ‐‐‐ 0.06

Cr 14.2 19 16

Ni 42.23 ‐‐‐ 62

Ti 1.73 ‐‐‐ ‐‐‐

Mo 1.77 ‐‐‐ 16

Al 0.45 ‐‐‐ ‐‐‐

Nb 2.66 ‐‐‐ ‐‐‐

V 0.31 ‐‐‐ ‐‐‐

Co ‐‐‐ 8.5 ‐‐‐

Mn <0.5 35 2

B ‐‐‐ 0.1 ‐‐‐

Si <0.5 0.8 0.5

Fe Balance 1.5 5

Braze foil contains higher manganese content.

J.R.Davis stated that Mn is beneficial in minimising

micro‐fissuring. Additions of upto 9wt% Mn is used in

commercial welding electrodes. As stated in AWS

Brazing manual, Mn improves wetting and bonding.

Since wettability is improved by addition of Mn, the

crack healing will be effective. And at the same time

the bonding is ensured. It also aids for subsequent

heat treatment. Addition of Co promotes volume

fraction of ’ and increases antiphase boundary

energy of / ’. Addition of Mo decreases lattice

mismatch of / ’. Since the objective of this study

includes minimising HAZ micro‐fissuring and filling

of the crack, braze foil is selected with higher Mn

content. Similarly the welding filler metal has lower Fe

and Cr content, which reduces susceptibility for the


64

formation of Laves phase, and higher Mo content

provides better solid solution effect. Since post weld

aging is not possible, more solute elements are added

through filler metal.

Results and Discussion

Microstructure Analysis

The particular casting underwent the qualification test

in which the service temperatures were shooting up

(1000°C) beyond the aging temperature (750°C) for a

short time of the order of less than 100ms. A phase

change occurs in Alloy 706 at the service temperature

of 1000°C with the time hold of 8 hrs. Since the

exposure time at such high temperature (1000°C)

during the qualification testing is very much negligible,

it is assumed that the casting experienced only a

thermal shock. In order to further decide upon the

repair process and assessment on phase change, if any,

the thickness and hardness survey on the defective

zone and adjacent zones were carried out. The details

are given in Table‐II and Figure‐(iii).

TABLE II HARDNESS AND THICKNESS SURVEY

Location

Measured thickness (mm)

As cast condition After emery

polish

Spot I 3.91/4.04 3.68/3.89

Spot II 3.24/3.60 3.19/3.43

Spot III 3.26/3.50 ‐‐NR‐‐

Spot IV 3.19/3.38 ‐‐NR‐‐

NR‐Not required, since these locations were non‐defective

zones and hence emery polshing was not done

FIG . (iii) SCHEMATIC FOR HARDNESS AND THICKNESS

SURVEY

The defective zone was analysed with in‐situ

metallography to work out the repair plan and

branched cracks were found in the casting as shown in

Figures –(iv) (a) and (b). The thickness of the casting

shell is 4mm and the surface of the casting was

prepared for in‐situ metallographic analysis. Since the

defect was found in the assembled condition with

closer to rotating parts, the severity of the defect in the

thickness direction could not be revealed by any of the

NDT methods. Any NDT method can only give results

with overlapped signals/image from the rotating

assembly below the casting inner‐wall.

FIG. iv (a) And (b) OPTICAL MICROGRAPH OF THE CRACK

MORPHOLOGY IN THE DEFECTIVE ZONE OF CASTING (ALLOY

706)

The base material microstructure reveals (Figures‐v

and vi) dendritic structure, which indicates the

limitation in the homogenization temperature as

already mentioned. Also the micrographs show the

Laves phase and MC type carbides. The presence of

these particles has been confirmed with X‐ray

diffraction analysis of base material.

FIG. v MICROGRAPH OF AGE

HARDENED BASE MATERIAL

SHOWS ONLY MC TYPE

CARBIDES

FIG. vi MICROGRAPH OF

AGE HARDENED BASE

MATERIAL SHOWS ONLY

LAVES PHASE

In the case of braze welding, braze metal flows by

capillary action through the existing crack and fills the

cavity, whereas in the case of welding route, crack can

be completely eliminated with the full penetration.

The full penetration in the joint is not allowed, in view

of the restriction of lower clearances in the inner

rotating assemblies. Hence partial penetration was

employed. Optical micrographs of the repair

simulations with braze welding and welding are

shown in Figures‐ vii and viii respectively.

(a) (b)

MC Laves


65

FIG. vii OPTICAL MICROGRAPHS OF BRAZE WELD FUSION ZONE

FIG. viii OPTICAL MICROGRAPHS OF WELD FUSION ZONE

FIG. ix SEM IMAGES OF WELD HAZ

FIG. x: SEM IMAGES OF BRAZE WELD HAZ FREE FROM

LIQUATION

FIG. xi: SEM IMAGE OF WELD

FUSION ZONE

FIG. xii: SEM IMAGE OF BRAZE

WELD FUSION ZONE

FIG. xiii SEM IMAGE OF WELD

FUSION ZONE SHOWING

CRACKS

FIG. xiv SEM IMAGES OF BRAZE

WELD FUSION ZONE FREE

FROM CRACKS

Grain boundary liquation observed in welds is shown

in Figures‐viii and ix. Upon solidification, this leads to

HAZ micro‐fissuring. N.L.Richards et. al, and

O.A.Idowu et. al, stated that the reason for this

liquation phenomenon is the thermal stresses induced

in HAZ by heat input and thermal gradient further

enhances incipient melting of MC type carbides at

grain boundary during welding process. S.Kou et. al,

inferred that HAZ micro‐fissuring does not occur,

when sufficient grain boundary liquid is not available

or in the presence of excessive grain boundary liquid.

Feng et al. developed the correlation for weld cracking

with the stress–strain evolution during weld cooling.

They considered only transverse and longitudinal

stresses as a function of weld cooling. Cracking at a

position will be promoted if a weak microstructure

and/or a sufficiently high tensile stress exists. Dye et al.

proposed a numerical method for the prediction of the

processing conditions to produce defects during

welding such as constitutional liquation, solidification

cracking and a centreline grain formation. Mayor et al.,

investigated the characteristics of Inconel 718 and 706

joints, welded by GTAW (gas tungsten‐arc welding),

finding an excellent weldability and good tensile

strength at both room and elevated temperature. The

thermal effects of welding operations have been

observed to affect the desirable structure of the heat‐

treatable alloys by producing heat‐affected zones with

poor mechanical properties. Li et al., stated that the

microfissuring sensitivity is influenced not only by the

chemical composition of the alloy but also by the

thermal stresses arising during cooling. Increased heat

input prevents HAZ micro‐fissuring by reducing

thermal stresses at HAZ. The thermal stresses decrease

with reduction of welding speed and thickness of the

material. The strain rate in the HAZ is related to the

welding speed. The thermal stresses are developed at

fusion zone and HAZ by the difference in the

temperature distribution during metal joining

processes. When increasing the heat input, the

temperature gradient becomes shallow and hence the

difference in temperature between fusion zone at the

start of solidification and HAZ is reduced. In this case,

the weld cooling rate is reduced, which promotes

more of Laves phase. Heat input rate can be varied

with welding speed for the given heat input. This

reduces the heat input rate even for higher heat input.

Hence a compromise is made between heat input,

thermal gradient and cooling rate by manipulating the

welding speed. Therefore liquation phenomenon fails

to be observed with braze welding process, (Figure‐x),


66

because the heating rate is moderate and cooling rate

is slightly higher though heat input falls on the lower

range in both welding and braze welding modes. Thus

the incipient melting of grain boundary carbides is

minimized with braze welding process. Detailed

discussion on the effect of heating rate and cooling

rate is described using heat transfer studies in

subsequent section. Ferro et al., has studied the effect

of high energy density welding process (EBW) on

HAZ micro‐fissuring of wrought Inconel 706. It is

found that HAZ micro‐fissures are formed due to the

thermal stresses developed during cooling. He also

compared the numerical model of other researchers

with EBW process where heating rate and cooling rate

are higher than that with GTAW process. It is inferred

that slower heating rate reduces the incipient melting

of carbides and moderate cooling rate with extraction

of heat using external heat sink reduces the build‐up

of temperature. And hence the thermal stresses are

balanced.

TABLE III LAVES VOLUME FRACTION USING IMAGE ANALYSIS

Process Laves volume fraction ±2%#

Braze welding 11.5

Welding 27.1

# Values are average of 30 nos of SEM micrographs

SEM micrographs of welded sample (Figure‐xi) show

blocky, thick and continuous Laves network. Braze

welded microstructure shows fine and globular Laves

particles (Figure‐xii), which is the indication of

increased under‐cooling. Further probing of the

microstructures of fusion zone (FZ) in certain samples

of welding route shows cracks in the inter‐dendritic

region (Figure‐xiii). The fusion zone of braze welding

does not exhibit any crack in any of the samples as

shown in Figure‐xiv. It is clearly seen that the cracks

propagate through the inter‐dendritic region in FZ.

The volume fraction of inter‐denritic region laves

phase has been calculated using image analysis.

Details are given in table‐III.

X‐ray Diffraction Analysis

The X‐ray diffraction studies on braze welded sample

have shown minor peaks of laves phase compared to

that in welded sample (Figure‐xv). Thereby it is

confirmed that the volume fraction of laves phase is

less than that in braze welded sample. Since the braze

foil contains Mn, manganese carbides are likely. Traces

of Eta phases are observed.

FIG. xv X‐RAY DIFFRACTION PATTERNS FOR BASE MATERIAL,

WELDED SAMPLE AND BRAZE WELDED SAMPLE

In general, braze welded sample contains high

temperature carbides which inhibit grain growth at

elevated temperatures. Welded samples show

presence of laves phase and also the peaks are

prominent. Details of phases with crystallographic

orientations are listed in Table‐IV.

TABLE IV INDEXING OF XRD PATTERNS

Description Observed

diffraction details

Indexed diffraction details Legend in

Figure‐xv

d‐spacing

(A0)

I/I0 Crys.

Struc.

Crys.

planes

Lattice para.

(A0)

Base Material

2.072 100 Ni3Al

FCC

(111) (200)

(220)

aγ=3.597

aγ’=3.571

1.801 63.35

1.2718 24.72

3.101 13.34 Ortho

Ni3Nb

(011) a=5.116

b=4.26

c=4.565

1.147 1.21

2.53 0.68 FCC

NbC

(111) (220)

(311) (222)

a=4.471

1.648 4.56

1.395 1.63


67

1.272 24.72

2.072 100 FCC

TiN

(200) (111)

(220) (311)

a=4.238

2.53 0.68

5.111 3.36

1.272 24.72 Braze W

elding

2.052 100 �

FCC (111) (200)

(220)

aγ=3.597

aγ’=3.571

1.786 31.34

1.266 17.91

3.101 8.11 Ortho

Ni3Nb

(011) a=5.116

b=4.26

c=4.565

1.147 3.08

2.54 0.68 FCC

NbC

(111) (220)

(311) (222)

a=4.471

1.621 4.56

1.361 1.99

1.293 0.73

2.125 10.35 FCC

TiN

(200) (111)

(220) (311)

a=4.238

2.442 0.17

1.537 5.51

1.266 17.91

2.052 100 BCC

Fe9.64Ti0.3

6

(110) (211)

(200)

a=2.877

1.159 6.92

1.451 7.97

2.317 7.45 FCC

CrC

(111) (200)

(220) (311)

(222)

a=4.03

2.052 100

1.451 7.97

1.293 0.73

1.147 3.08

1.938 2.72 Hex.

Ni3Ti

(202) (201)

(004) (203)

(205)

a=5.096

c=8.304

2.125 10.35

2.052 100

1.786 31.34

1.361 1.99

2.052 100 FCC

Mn23C6

(511) (422)

(531)

a=10.59

2.125 10.35

1.786 31.34

Welding

2.057 100 �

FCC (111) (200)

(220)

aγ=3.597

aγ’=3.571

1.789 39.41

1.267 20.12

3.037 38.24 Orthorh.

Ni3Nb

(011) a=5.116

b=4.26

c=4.565

1.125 1.36

2.491 15.2 FCC

NbC

(111) (220)

(311) (222)

a=4.471

1.602 20.82

1.336 1.02

1.296 4.26

2.098 18.29 FCC

TiN

(200) (111)

(220) (311)

a=4.238

2.491 15.2

1.525 9.85

1.267 20.12


68

2.057 100 BCC

Fe9.64Ti0.3

6

(110) (211)

(200)

a=2.877

1.125 1.36

1.434 15.12

2.283 19.86

FCC

CrC

(111) (200)

(220) (311)

(222)

a=4.03

2.098 18.29

1.473 2.37

1.296 4.26

1.125 3.08

2.057 100 BCT

Ni4Mo

(121) (310)

(002)

a=5.72

c=3.564

1.873 18.75

1.789 39.41

1.92 10.72 Orthorh.

Ni3Mo (211) (020)

a=5.064

b=4.224

c=4.448

2.098 18.29

2.219 0.64

2.057 100 Hexago

nal

Fe2Ti

(112) (103)

(201)

a=4.796

c=7.833

1.92 10.72

2.219 0.64

2.057 100 Hexago

nal

Fe2Nb

(112) (103)

(201)

a=4.813

c=7.849

2.283 19.86

1.92 10.72

Heat Transfer Studies

To understand the mechanism of the segregation and

the microstructure formation during repair welding,

temperature measurement has been employed. The

instantaneous cooling rate is slightly high (30.55°C/s)

in braze welding which minimizes the segregation

during solidification. The instantaneous cooling rate in

welding is found to be 26.45°C/s (Figures‐xvi to xix).

Hence segregation is moderately higher in the welding

route.

As already discussed, compromise between thermal

stresses and cooling rate minimizes HAZ micro‐

fissuring. It also controls the formation of Laves phase.

Braze welding process for the set variables exhibits a

maximum instantaneous cooling rate of 30.550C/s and

a corresponding heating rate of 30.50C/s. The heating

rate ranges from 6.95 to 130.10C/s and cooling rate

ranges from 2.15 to 30.55 0C/s for braze welding

process. An average heating rate of 41.54 0C/s and

average cooling rate of 15.65 0C/s are observed in braze

welding process.Similarly in the welding process for

the set variables, maximum instantaneous cooling rate

is of 26.450C/s and a corresponding heating rate is

30.50C/s. The heating rate ranges from 6.95 to 67.750C/s

and cooling rate ranges from 2.15 to 26.45 0C/s for

welding process. An average heating rate of 31.2 0C/s

and average cooling rate of 8.57 0C/s are observed in

braze welding process.

50 100 150 200 250 300

50

100

150

200

250

300

Te

mp

era

ture

(癈

)

Time (sec)

T1 T2 T3 T4 T5 T6 T7 T8 T9 T10 T11 T12

1150 1200 1250 1300 1350 1400 1450 1500

0

100

200

300

400

500

Tem

pera

ture

(癈

)

Time (sec)

dT/d

t(癈

/se

c)

max

=30.55癈 /sec

T=247.4癈

-35

-28

-21

-14

-7

0

7

14

21

28

35

T2Alloy 706 - Braze welding

FIG. xvi COOLING CURVES FOR

BRAZE WELDING

FIG. xvii COOLING RATE IN

BRAZE WELDING FOR THE SET

PARAMETERS

1200 1300 1400 1500

100

200

300

400

500

Te

mp

era

ture

(癈

)

Time (sec)

T1 T2 T3 T4 T5 T6 T7 T8 T9 T10 T11 T12

0 100 200 3000

100

200

300dT

/dt(癈

/se

c)

TE

MP

ER

AT

UR

E (癈

)

Time (sec)

Welding of Alloy 706

-30

-20

-10

0

10

T1

max

=26.45癈 /sec

T = 199.7癈

FIG. xviii COOLING CURVES FOR

WELDING

FIG. xix COOLING RATE IN

WELDING FOR THE SET

PARAMETERS

From the experiment results, it is evident that equal

amounts of instantaneous heating and cooling rates

are effective in controlling the segregation. The area

under the heating and cooling curves is the

corresponding energies spent for dis‐solution and

segregation respectively. The heating rate energy is

directly proportional to dis‐solution and cooling rate


69

energy is inversely proportional to segregation.

Heating rate energy in braze welding is 50.95 Joules

and 25.65 Joules for welding. Similarly cooling rate

energy in braze welding is 43.34 Joules and 102.35

Joules for welding process. Hence in braze welding,

there is a reduction in segregation in the inter‐

dendritic region. Since energy is equalized, thermal

stresses are also balanced. Thus HAZ micro‐fissuring

is avoided.

Mechanical Testing and Fractographic Analysis

The room temperature tensile test results of braze

welded samples show comparable joint strength with

that of welded samples. The ductility of braze welded

joints is higher than that of welded samples by 31%.

The details of the room temperature and elevated

temperature (477°C) tensile tests are shown in table‐V.

TABLE V MECHANICAL PROPERTIES AT 298K & 750K

Process

Test

temperat

ure

(°C)

Mechanical properties# Failure

locatio

n

UTS

(MPa

)

0.2%Y

S

(MPa)

%

Elongatio

n

Welding 23 622.67 359.33 13.47

HAZ 477 517.67 317 12.73

Braze

Welding

23 600 313 17.6

477# 480 305 12

# Dual necking observed. All the test values are the average of three

samples

Room temperature mechanical properties of repaired

specimens have been compared with those of as cast

samples. 89.46% weld efficiency with welding route on

ultimate tensile strength and 86.21% weld efficiency

with braze welding route on similar conditions was

found.

Room temperature tensile test fracture surfaces of

braze welded samples (Figures‐xx a & b) show more

of dimple structure, which reflects in the ductility of

fusion zone. In case of welded sample (Figures‐ xxi

a&b), mixed structures (Cleavage and dimple) are seen.

The volume fraction of dimple structure in braze

welded sample is higher than that in welded sample.

Similarly, the elevated temperature tensile test fracture

surfaces of braze welded samples (Figures‐xxii a & b)

show trans‐granular fracture surface with wavy slip

mode. This mode of deformation is promoted by the

presence of non‐shearable carbides. In case of welded

sample (Figures‐xxiii a & b), inter‐granular fracture is

promoted by the presence of brittle laves phase.

FIG. xx (a) & (b) SEM FRACTOGRAPHS OF BW SAMPLE

TESTED AT 23°C

FIG. xxi (c) & (d) SEM FRACTOGRAPHS OF BW SAMPLE

TESTED AT 477°C

FIG. xxii (a) & (b) SEM FRACTOGRAPHS OF WELDED

SAMPLE TESTED AT 23°C

FIG. xxiii (a) & (b) SEM FRACTOGRAPHS OF WELDED

SAMPLE TESTED AT 477°C

Conclusion

Based on the experiments conducted on the aged cast

alloy 706 billets with braze welding and welding

processes, the following can be concluded:

Joint strength by braze welding route is

comparable with that obtained in welding

route

Ductility of braze weldment is better than that

in fusion zone and thereby having more of

creep crack growth resistance


70

In‐situ cracks can be repaired by braze welding

process with minimal damage to the base

material and lesser strength mismatch in case

of as welded condition

ACKNOWLEDGEMENTS

The authors would like to thank Director/ LPSC and

AD/LPSC (M) for giving this opportunity to work on

this problem. Meanwhile authors thank

Shri.CH.Kunhikamaran for his valuable technical

guidance during this work. They would also

appreciate, Dr.P.Ramesh Narayanan,

Dr.SVS.N.Murthy, Dr.VMJ.Sharma,Dr.Biju.S.Nair and

Mr.Sudharsan Rao of VSSC, Mr.Thomas Tharian and

Mr.P.Palvannan of LPSC for their support in micro‐

structural characterization.

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