Effect of Minor Change in Composition on the Toughness of Weldmetal for Repair of Turbine Blades made of Martensitic Stainless Steel

1
Science and Technology of Welding and Joining 2008, Volume 13, No. 2, 159-
166
Effect of Minor Change in Composition on the Toughness of Weldmetal for
Repair of Turbine Blades made of Martensitic Stainless Steel
C.R. Das, S.K. Albert, A.K. Bhaduri, G. Srinivasan and V. Ramasubbu
Materials Joining Section, Materials Technology Division,
Indira Gandhi Centre for Atomic Research, Kalpakkam 603102, India
Abstract
Repair welding procedure for cracked turbine blades, made of 13Cr-2.6Ni-1.1Mo
martensitic stainless steel, has been developed using gas tungsten arc welding (GTAW)
process and a twin-wire filler metal. The twin-wire consists of tack-welds along the length of
the two filler wires, a 1.5 mm diameter ER 16-8-2 and a 2.0 mm diameter ER 410. A two-
stage post-weld heat treatment at 675°C for 2 h and 615°C for 4 h, such that the first heat
treatments is above the Ac1 temperatures of the weld metal and second is just below its Ac1
temperature; has been found to be adequate for good mechanical properties. The weldment
has a good combination of transverse-weldment strength and weldmetal toughness, with its
room temperature yield strength and Charpy V-notch impact toughness being similar to that
of the turbine blade material.
Keywords: Twin-wire filler metal; Repair welding; Turbine blade; Two-stage post-weld
heat treatment; martensite; Impact toughness.

2
1. Introduction
Good mechanical properties, adequate corrosion resistance and relatively low cost
made martensitic steels important materials for various turbine components like blade, shroud
and tenon1-2
. As welding is seldom used for fabrication of turbine components, consumables
of matching composition for these turbine materials are not readily available3
. Therefore, if
welding has to be carried out for these materials, austenitic SS welding consumables, such as
E / ER 308, 308L, 316L, 309L or martensitic SS consumables conforming to AWS
specification E/ER 410 and E/ER 410NiMo are generally used. In the case of former, no
post-weld heat treatment (PWHT) is usually employed4
and strength of the weld joint is
lower than that of the base metal. Further, dilution of the austenitic stainless steel weld
deposit results in austenitic-ferritic-martensitic microstructure in the weldmetal, very close to
the weld interface with consequent poor ductility and toughness4
for this zone. Martensitic
SS consumables are chosen when it is required to match the strength of the weldmetal with
that of the base metal. The 410 SS consumable is used when strength of the weldmetal is of
major consideration, while the 410NiMo consumable is chosen when both strength and
toughness are important. The 410NiMo weldmetal is known to have some retained austenite
[] during cooling from the normalising temperature and hence a two-stage tempering heat
treatment is recommended to ensure the tempering of the martensite formed during cooling
from the first-stage of heat treatment5
.
Various turbine components, such as blades, shrouds and rotor, made of martensitic
SS often develop cracks during service6-7
necessitating either replacement or repair. Weld
repair of cracked components is by far more economical than component replacement, if it is
ensured that the properties of the repair-weld are comparable to those of the wrought
material. The present study was carried out to develop a repair welding procedure for cracked
low-pressure steam turbine blades made of 13Cr-2.2Ni-1.1Mo martensitic SS, welding
consumable with desired chemical composition and PWHT. The welding trials included use
of commercially available ER 410NiMo consumables and specifically designed twin-wire
consumables consisting of ER 16-8-2 and ER 410 filler wires. The paper describes the results
of the present study and discusses how welding consumables and PWHT were selected for
carrying out repair welding of turbine blades.
2. Experimental
The compositions of the turbine blade material and that of the weldmetal obtained
from ER 16-8-2, ER 410 and ER 410NiMo filler wires were analysed using optical emission

3
spectrometer. Two different combinations of twin-wires were employed – one with 2.0 mm
diameter wires of ER 16-8-2 and ER 410, and the another with 1.5 mm diameter ER 16-8-2
wire and 2.0 mm diameter ER 410 wire. In this paper, these filler wires are designated as
TW-1 and TW-2, respectively for simplicity of the discussion. The twin-wire welding
consumable, shown in Fig. 1, was produced by intermittent tack welding of ER 16-8-2 filler
wires with ER 410 filler wire by gas tungsten arc welding (GTAW) process using a current of
35 A and voltage of 8.5 V. The chemical composition of the TW-1 and TW-2 weldmetal
produced by GTAW process, were also analysed using the optical emission spectrometer.
Reproducibility of the TW-1 and TW-2 weldmetal was checked using three weld coupons
made from both the filler wires.
The Ac1 temperature of TW-2, 410NiMo weldmetal and turbine bladed were
estimated indirectly based on hardness measurement, as a function of tempering temperature
of the normalised material. Specimen of dimension 15×15×10 mm3
were extracted from the
respective weld deposits and the blade material for carrying out these heat treatments.
Initially, material were normalised at 1000°C for 1 h, followed by air cooling and then
tempered from 200 to 850°C for 1h and then material is cooled in furnace. Further,
transformation temperatures of these materials were estimated from the results of differential
scanning calorimetry (DSC).
Weld pads of turbine blade that was in service for few years were fabricated by the
GTAW process using TW-1, TW-2 and ER 410NiMo filler wires and welding parameters
used during this fabrication are given in Table 2. Based on filler wire’s designation used in
this paper, the weld pads are referred as TW-1, TW-2 and 410NiMo, respectively. After
welding weld pads were given post-heating treatment at 250°C for 1 h and then furnace
cooled to room temperature. The TW-1 weld pad was subjected to two different PWHT – a
single-stage PWHT of 735°C for 1 h and a double-stage PWHT, initially at 735°C for 1 h
followed by furnace cooling to room temperature and then heating at 600°C for 4 h and again
furnace cooling to room temperature. The TW-2 and 410NiMo weld pads were also subjected
to another PWHT, initially at 675°C for 2 h and cooling to room temperature in furnace.
Subsequently, second-stage PWHT which consists of 675°C for 2 h and cooled to room
temperature followed by 615°C for 4 h and cooled to room temperature. After PWHT, the
weld pads were examined by ultrasonic testing, and no defects were observed in the weld
joints.

4
X-ray diffraction was taken from the as deposited TW-2 and 410NiMo weldmetal.
Microstructural examination and microhardness measurements were carried out on
weldments in as-welded and PWHT conditions. Face- and side-bend tests were carried out on
the weldments after PWHT. Tensile tests of transverse weld round tensile specimen were
carried out after PWHT at a strain rate of 3.2×10–4
s–1
for different temperatures ranging from
room temperature (23°C) to 550°C. Charpy V-notch impact tests were carried out on full-size
specimens with the notch in the weldmetal oriented along the welding direction. It was also
carried out for heat treated base material. Fractography of the impact tested specimens were
carried out using scanning electron microscope (SEM).
3. Results
3.1 Chemical composition:
The chemical composition of the turbine blade, ER 16-8-2, ER 410 ER 410NiMo,
TW-1 and TW-2 weldmetal are given in Table 1. From the table it is clear that chemical
composition of TW-2 weldmetal is closer to that of turbine material than that of the TW-1
and 410NiMo weldmetal. The reproducibility of the TW-1 and TW-2 weldmetal was
confirmed from the chemical analysis of three separately prepared samples from each filler
wires and the values reported (Table 1) are that of these three average. Very close observation
of the table 1, shows TW-1 and 410 NiMo weldmetal contain more silicon than the TW-2 and
turbine blade material. In addition to this, the TW-1, TW-2 and 410NiMo weldmetals contain
higher nickel (5.0–5.3%) than that of the turbine blade material (2.6%). On the other hand
Creq and Nieq values calculated using the schaeffler diagram8
for the TW-2 weldmetal closely
matches to that of the turbine blade material (Table 1). Chemical compositions of TW-1 and
TW-2 filler wire are not significantly different except silicon and few austenite stabilising
elements, which has contributed to the difference in the Creq and Nieq (Table 1).
3.2 Transformation temperature Ac1 and Ac3
To determine the Ac1 transformation temperatures before selecting the heat treatment,
the as-normalised (1000°C for 1h) blade material, TW-2 and 410NiMo weldmetal were
tempered for 1 h between 200–850°C. Variation in hardness with the tempering temperature
for the weldmetals and the turbine blade material is shown in Fig. 12. Figure shows with
increase in tempering temperature the hardness of the materials decreases continuously.
However, hardness decrease is significant only at and around 500°C and it reaches minimum
at 617, 627 and 727°C for the TW-2, 410NiMo and turbine blade material respectively. But it

5
begins to increase again with increase in temperature, at which austenite begins to form with
heating and upon cooling this austenite transforms to martensite resulting in an increase in
hardness of the material. The temperature at which this transition takes place can be
considered as ~ Ac1 transformation temperature of the material.
Ac1 temperatures estimated from the DSC plots are shown in Fig. 13. It shows offset
at which transformation start and it can be considered as Ac1 transformation temperature for
the respective. These estimated temperatures for the respective material are 608, 601 and
651°C respectively.
3.3 Microstructure and hardness
The microstructure of the as deposited TW-1 weldmetal consist of martensite and
inter-dendritic retain austenite. After single-stage of 735°C/1 h PWHT (Fig. 3) the
microstructure of weld interface of this weldment shows weldmetal is fully martensitic while
the heat-affected zone (HAZ) is a mixture of fresh martensite and tempered martensite. The
difference in the sizes of hardness indentation marks made on this post weld heat treated
weldment, at two differently etched locations indicate highly etched region is tempered
martensite and light etch region is fresh martensite. The indentation size in the fresh
martensite location is smaller then the tempered region. The fresh martensite is formed during
cooling cycle of first PWHT. However, during 2nd
stage of the two stage heat treatment
(735°C/1 h + 600°C/4 h) this fresh martensite gets tempered in both the weldmetal and HAZ
(Fig. 4) and it has been reflected in the microhardness distribution taken on the weldment
after single and double stage heat treatment. Microhardness profiles (Fig. 5) taken after
735°C/1 h PWHT across the weld interface of this weldment shows hardness is significantly
high at about 550 VHN in the weldmetal, but it varies widely in the HAZ with the peak
hardness of 560 VHN. This high hardness decreases (350-375 VHN) considerably after two-
stage 735°C/1 h + 600°C/4 h PWHT in both the weldmetal and HAZ indicating martensite
got tempered effectively.
The microstructure of the TW-2 weldmetals in the as-deposited condition also (Fig. 6-
a) consists of martensite and fine inter-dendritic retained austenite, but their volume fraction
is being lower compared to TW-1 weldmetal. In the same way microstructure of 410NiMo
weldmetal also consists of martensite and coarse inter-dendretic retain austenite in addition to
the delta ferrite as shown in Fig. 7-a. Figure 6-b shows microstructure of TW-2 weld
interface after two-stage 675°C/2 h + 615°C/4 h PWHT and Fig. 7-b shows 410 NiMo weld

6
interface microstructure after same heat treatment. Hardness profiles across the weld interface
of TW-2 weldment in the as-deposited and PWHT conditions (Fig. 8) shows high hardness in
the as-welded condition both in the weldmetal and HAZ of about 400 VHN and 450-
500 VHN, respectively. After PWHT, hardness decreases considerably with weldmetal
hardness reducing to about 300 VHN after the two-stage PWHT. The hardness profiles across
the weld interface of 410NiMo weldment (Fig. 9) shows reduction in hardness both in the
weldmetal and HAZ on PWHT it is similar to that observed in TW-2 weldment, except
marginal reduction in hardness from single-stage to two-stage PWHT.
3.4 X-ray Diffraction:
XRD pattern taken from the as deposited TW-2 and 410NiMo weldmetal shows presence of
strong martensite and weak austenite peaks in both the weldmetal. The intensity of the
austenite peak is low due to their lower volume fraction in the deposit as compared to the
martensite volume fraction. Figure 10 shows XRD pattern taken from the as-deposited TW-2
weldmetal and Fig. 11 shows the enlarge view of the same XRD pattern, which clearly shows
presence of all austenite peaks in the weld deposit.
3.5 Bend test
The TW-1 weldment, which was subjected to single-stage PWHT at 735°C for 1 h
and the two-stage PWHT at 735°C for 1 h + 600°C for 4 h failed in guided bend test with
fracture occurring at almost 0° bend angle in the former and 120° bend angle in the latter.
The 410NiMo weldment subjected to two-stage PWHT fractured at a bend angle of 45°. Only
TW-2 weldment subjected to the two-stage PWHT successfully passed 180° guided bend
tests.
3.6 Tensile test
Results of room-temperature transverse-weld tension tests after single- and two-stage
PWHT, along with that of the turbine blade material at different conditions, are shown in
Fig. 14. The tensile strength values reported are average of two tests at each test temperature,
with the strength value obtained for each test conditions being very close to each other. All
the transverse-weld specimens failed in the base metal irrespective of PWHT it was subjected
to. The yield strength (YS) of TW-1 weldment after single-stage PWHT was lower than that
of weldments after the two-stage PWHT. The variations in YS and UTS as a function of
testing temperature for TW-2 and 410NiMo weldments show that both YS and UTS are
consistently higher for the TW-2 weldment.

8
except for Ni, Mo and Si. Presence of high Ni in the weldmetal demand judicial selection of
two stages PWHT for the weldment made of this consumable4
.
Though hardness profile across the weld interface, shown in Fig. 9 indicates reduction
in hardness even after first stage PWHT microstructure observation does not shows
significant tempering. After two stages PWHT microstructure shows presence of tempered
martensite. But the impact toughness of the 410NiMo weldmetal after two-stage PWHT is
considerably lower than that of the blade material. Hence, this consumable could not be
considered for the proposed repair welding using the heat treatment cycle chosen in this
study. However, use of this material needs optimisation of PWHT to obtain suitable
toughness in the weldmetal, which matches with that of the base metal toughness.
The reason for poor toughness of the 410NiMo weldmetal even after two-stage
PWHT and the failure of the weld joint during guided bend test is attributed to presence of
delta-ferrite in the weldmetal as revealed in the microstructure (Fig. 8(b)) and in-appropriate
selections of PWHT temperature and time, which could not temper that martensite adequately
for improving the toughness. In spite of very high Ni content (5.3 %) in the weldmetal
presence of very low carbon content (0.02 %) in the weldment could not counter balance the
ferrite stabilising element effect. These is reflected in the Nieq of this weldmetal, which is
low (5.8) compared to the blade metal (6.8). This low Nieq could not counter balance the
effect of high Creq (13.6) and results in stable delta-ferrite in the weldmetal down to room
temperature. Another reason for low toughness of the weldmetal is high Si content in the
weldmetal. It is well known that toughness of ferritic steel weldmetals decreases with
increase in Si content []. It is also reported that silicon increases tempering resistance of the
material.
The option of making filler wire by tack welding the two different filler wires of
ER 410 and ER 16-8-2 along their length were considered when it was evident that
weldmetal made from ER 410NiMo filler wire cannot match the toughness of the base metal
even after two stage PWHT used in this study. Initially TW-1 filler wire, made using the two
filler wires of same diameter, was chosen. This weldment was subjected to 735°C/1h PWHT
chosen based on PWHT employed for 12Cr martensitic stainless steel9
. However, the fully
martensitic structure of the weldmetal, as revealed from the indentation mark on the
microstructure (Fig. 4) and the hardness profiles (Fig. 6) after this PWHT indicated that the
PWHT temperature was above the Ac3 temperature of the weldmetal. It is also evident from

9
the hardness vs. tempering temperature plot and DSC result (Fig.12 and 13) that this PWHT
temperature is also above the Ac1 of the blade material. The second-stage PWHT of
600°C/4h tempers the weldmetal and HAZ microstructure resulting reduction in their
respective hardness but reduction in hardness in the weldmetal and HAZ is different, due to
difference in their transformation temperature (Fig.). In spite of substantial reduction in
hardness during tempering the weldmetal toughness is far below that of the base metal; as a
result weldments failed to pass the guided bend tests. It may be noted that even after this
PWHT the hardness of the weldmetal is quite high (~350 VHN) and this indicated that the
tempering temperature is too low for this material. Although the Ac1 temperature of this
weldmetal had not been estimated, the high Nieq of 8.3 of this weldmetal indicated that its
Ac1 temperature would be quite low, and hence it would not be possible to increase the
PWHT temperature of second stage much beyond the currently employed temperature of
600°C. This argument is further supported by marginal raise of Ac1 temperature (617°C) for
TW-2 filler metal with lower Nieq of 6.9 than that of TW-1 filler (Nieq = 8.3). It can be
mentioned here that normalising heat treatment at 1000°C for 1h were carried out to
homogenise the chemical composition of the weldmetal so that macro in-homogeneity of the
material chemistry can be avoided. This has been reflected in narrow scatter band in the
hardness plot (Fig.12). In addition to inadequate tempering, high Si content of the weldmetal
also could have contributed to low toughness of this weldmetal. Because of the variation in
Creq and Nieq, selection of PWHT temperature and time is become difficult for obtaining the
adequate mechanical properties in the weldmetal and weldment.
As lower transformation temperature is one of the reasons for the poor toughness of
TW-1 weldmetal, it was decided to decrease Ni content of the weldmetal by using a lower
diameter ER 16-8-2 filler wire for making the twin wire. Hence, TW-2 twin-wire was
prepared using 1.5 mm diameter ER 16-8-2 and 2.0 mm diameter ER 410 filler wires. The
chemical composition (Table 1) reveals that both Ni and Mn content in the TW-2 weldmetal
decreased, with corresponding decrease in Nieq from 8.1 to 6.9 in the TW-2 weldmetal
compared to TW-1 weldmetal. Further, the choice of the two-stage PWHT employed for
410NiMo / TW-2 weldmetal ensured the first-stage PWHT temperature is above the Ac1
temperature of the weld metal and the second-stage PWHT temperature are below its Ac1
temperature. In the first stage tempering martensite which formed during cooling part of
normalised / annealing got tempered and martensite form by the transformation of retained
austenite. In the second stage of tempering this fresh martensite which formed during cooling

10
from the first stage of heat treatment got tempered. As a result, toughness of the TW-2
weldmetal improved considerably and importantly values was similar to those of base metal
obtained after two stage heat treatment. However, adequate toughness could not be obtained
in the 410NiMo weldmetal. Delta-ferrite, which is presence in 410NiMo weldmetal would
have contributed to low toughness of that weldmetal, is not present in the TW-2 weldmetal.
The low Si content in the TW-2 weldmetal could be another reason for the improve
toughness of weldmetal. However it is not clear how Si content in the weld metal reduced to
this level because both ER410 and Er16-8-2 filler wires have higher Si content than present
in the weld metal (Table 1).
The above discussion clearly brings out how commercially available welding
consumables may not be suitable for repair welding of the turbine blades as it would be
difficult to match the weldment properties of the repaired weldmetal to that of the turbine
blade material by suitable heat treatment. This limitation could be overcome by use of twin-
wire consisting of ER 16-8-2 and ER 410 filler wires and by using lower diameter ER 16-8-2
filler wire than that of ER 410 filler wire for making this twin-wire filler wire. The
composition of the weldmetal, similarly microstructure and properties of the weldmetal could
be achieved by appropriately choosing the temperature of the two stages PWHT of weldmetal
could be tailored to provide adequate toughness for the weldmetal. Choice of temperatures
for two-stage PWHT is another factor that was important to ensured desirable weldmetal
properties.
5. Conclusions
For carrying out repair welding of damaged turbine blades made of 13Cr-2.6Ni-
1.1Mo martensitic stainless steel, commercially available filler wires like ER 410NiMo are
found to be unsuitable to achieve toughness similar to that of turbine blade. However,
combining two different commercially available filler wires into a twin-wire, and using this
as a welding consumable, it is possible to produce weldmetal that closely matches the blade
material both in composition and in properties. The studies also show that a two-stage
PWHT, with the temperatures carefully chosen based on the transformation temperatures of
the weldmetal and base metal is critical in achieving good toughness. Further, the high Si
content and presence of delta-ferrite in the microstructure are found to be detrimental to the
toughness of the weld metal.

11
References
1. R.J. Castro and J.J. de Cadenet: “Welding metallurgy of stainless and heat-resisting
steels”, Cambridge University Press, 1968
2. F.B. Pickering: “Physical metallurgy and design of steels”, Applied Science
Publishers, 1978
3. A240/A240M-05: Standard specification for chromium and chromium-nickel
stainless steel plate, sheet, and strip for pressure vessels and for general applications,
ASTM
4. A.W. Marshall and J.C.M. Farrar: Welding in the World., May/June 2001,45, 19
5. Stainless Q&A, Welding Journal, 89, 2000
6. A. Atrens, H. Mayer, G. Faber and K. Schneider: “Steam turbine blades”, in:
Corrosion in Power Generating Equipment, ed. M.O. Speidel and A. Atrens, Plenum
Press, New York, NY, 1983, 299
7. R. Viswanathan: “Life assessment of high temperature components”, ASM
International, Materials Park, OH, 1989, 265
8. L.E. Lancaster, Schaeffler diagram for Fe-Cr-Ni weldmetal showing approximate
regions, Sixth edition, 1999, 318
9. C.R. Das, V. Ramasubbu, A.K. Bhaduri and S.K. Ray, Repair welding of cracked
turbine shroud using matching composition consumables, Science and Technology of
Welding and Joining 10(1) (2005) 110–112
10. A.M. Barnes: Microstructural stability of creep resistant alloys for high temperature
plant application, eds. A. Strang, J. Cawley and G.W. Greenwood, The Institute of
Materials, IOM communications Ltd, 1998, 339

13
Figure Captions
Fig. 1: Tack-welded twin-wire TW-2 filler wire
Fig. 2: Schematic of the two-stage PWHT sequences used
Fig. 3: Microstructure of as-deposited TW-2 weldmetal
Fig. 4: Microstructure of TW-1 weldment after single-stage 735°C/1 h PWHT
Fig. 5: Microstructure of TW-1 weldment after two-stage 735°C/1 h + 600°C/4 h PWHT
Fig. 6: Microhardness distribution across weld interface in TW-1 weldment after single-
stage and two-stage PWHT
Fig. 7: Microstructure of TW-2 weld interface after single-stage 675°C/2 h PWHT
Fig. 8: Microstructure of weld interface in (a) TW-2 and (b) 410NiMo weldments after
two-stage 675°C/2h + 615°C/4h PWHT
Fig. 9: Microhardness distribution across weld interface in TW-2 weldment after different
PWHT
Fig. 10: Microhardness distribution across weld interface in 410NiMo weldment after
different PWHT
Fig. 11: Variation in hardness with tempering temperature for 1040°C/1 h normalised
turbine blade material, and 410NiMo and twin-wire TW-2 weldmetals
Fig. 12: Room temperature tensile strength of different transverse-weld specimens and base
metal
Fig. 13: Variation in transverse-weld yield strength (YS) and ultimate tensile strength (UTS)
with test temperature for 410NiMo and TW-2 weldments after two-stage 675°C/2 h
+ 615°C/4h PWHT
Fig.14: Charpy V-notch impact toughness of turbine blade material and different
weldmetals after single- and two-stage PWHT
Fig. 15: Fracture surface of impact tested specimen for twin-wire TW-2 weldmetal after
two-stage 675°C/2 h + 615°C/4 h PWHT

14
Fig. 1: Tack-welded twin-wire TW-2 filler wire
-2 0 2 4 6 8 10 12 14 16 18
0
100
200
300
400
500
600
700
800
Second Stage
First Stage
Post Heating
Temperature,C
Time, h
735
0
C /1h + 600
0
C/4h (for TW-1)
675
0
C/2h + 615
0
C/4h (for TW-2 and 410NiMo)
Fig. 2: Schematic of the two-stage PWHT sequences used
Fig. 3: Microstructure of TW-1 weldment after single-stage 735°C/1 h PWHT

15
Fig. 4: Microstructure of TW-1 weldment after two-stage 735°C/1 h + 600°C/4 h PWHT
-2 0 2 4 6 8
300
350
400
450
500
550
600
Base metal
HAZWeldmetal
Hardness,VHN200g
Distance across the weld interface, mm
TW-1, 735
0
C/1h
TW-1, 735
0
C/1h + 600
0
C/4h
Fig. 5: Microhardness distribution across weld interface in TW-1 weldment after single-
stage and two-stage PWHT
Weldmetal
Weld interface

16
Fig. 6: Microstructure of as-deposited TW-2 weldmetal and weldment showing interface
(after double stage PWHT) microstructure
Fig. 7: Microstructure of as-deposited 410NiMo weldmetal and weldment showing weld
interface (after dowle stage PWHT)
-4 -2 0 2 4 6 8
250
300
350
400
450
500
550
Hardness,VHN200gm
Distance across the weld interface
As Welded
TW-2, 675
0
C/2h
TW-2, 675
0
C/2h +615
0
C/4h
Weldmetal Base metalHAZ
Fig. 8: Microhardness distribution across weld interface in TW-2 weldment after different
PWHT

17
-4 -3 -2 -1 0 1 2 3 4 5
300
350
400
450
500
550
Hardness,VHN200gm
Distance across the weld interface, mm
As Weld
410 NiMo 675
0
C/2h
410 NiMo 673
0
C/2h + 615
0
C/4h
Weldmetal
Base metal
HAZ
Fig. 9: Microhardness distribution across weld interface in 410NiMo weldment after
different PWHT
0 20 40 60 80 100
0
500
1000
1500
2000
2500
3000
TW-2
(220)α
(211)α
(200)α
(110)α
Intensity,Arbitary
2θ
Fig. 10 XRD taken from the TW-2 weldmetal

18
40 60 80 100
0
10
20
30
40
50
60
70
80
90
2900
2905
2910
2915
2920
2925
(311)fcc
(220)fcc
(200)fcc
(111)fcc
Intensity,Arbitary
2θ
TW-2
Fig. 11XRD taken from the TW-2 weldmetal
0 200 400 600 800
20
25
30
35
40
45
Turbineblade
410NiMo
TW-2
Hardness,RC
Tempering temperature
0
C
Twin wire
410NiMo
Turbine blade

20
0
100
200
300
400
500
600
700
800
900
1000
1100
1200
1300
1400
1500
1600
Double Stage
heat treatment
Single stage
heat treatment
As Received
base metal
410NiMo
Double StageTW-2
Double Stage
TW-1
Double StageTW-1
Single Stage
Strength,MPa
Samples
YS UTS
Fig. 14: Room temperature tensile strength of different transverse-weld specimens and base
metal
0 100 200 300 400 500 600
500
550
600
650
700
750
800
850
900
950
1000
TensileStrength,MPa
Testing Temperature,
0
C
YS, TW-2 weld joint
UTS, TW-2 weld joint
YS, 410NiMo weld joint
UTS, 410NiMo weld joint
Fig. 15: Variation in transverse-weld yield strength (YS) and ultimate tensile strength (UTS)
with test temperature for 410NiMo and TW-2 weldments after two-stage 675°C/2 h
+ 615°C/4h PWHT

21
1 2 3 4 5
0
10
20
30
40
50
60
70
80
90
410NiMo
675
0
C/2h +
615
0
C/4h
TW-2
675
0
C/2h +
615
0
C/4h
TW-1
735
0
C/1h+
600
0
C/4h
TW-1
735
0
C/1h
Turbine
blade
ImpactToughness,J
Fig.16: Charpy V-notch impact toughness of turbine blade material and different
weldmetals after single- and two-stage PWHT
Fig. 17: Fracture surface of impact tested specimen for twin-wire TW-2 weldmetal after
two-stage 675°C/2 h + 615°C/4 h PWHT

Effect of Minor Change in Composition on the Toughness of Weldmetal for Repair of Turbine Blades made of Martensitic Stainless Steel

Recommended

Recommended

More Related Content

What's hot

What's hot (20)

Viewers also liked

Viewers also liked (12)

Similar to Effect of Minor Change in Composition on the Toughness of Weldmetal for Repair of Turbine Blades made of Martensitic Stainless Steel

Similar to Effect of Minor Change in Composition on the Toughness of Weldmetal for Repair of Turbine Blades made of Martensitic Stainless Steel (20)

Effect of Minor Change in Composition on the Toughness of Weldmetal for Repair of Turbine Blades made of Martensitic Stainless Steel