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Development of a Dissimilar Temper Bead Welding Procedure for an Amine
Tower Repair
Conference Paper · January 2008
DOI: 10.1115/PVP2008-61779
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Proceedings of PVP2008
2008 ASME Pressure Vessels and Piping Division Conference
July 27-31, 2008, Chicago, Illinois, USA
PVP2008-61779
DEVELOPMENT OF A DISSIMILAR TEMPER BEAD WELDING PROCEDURE FOR AN
AMINE TOWER REPAIR
Matt A. Boring, P.E.
Edison Welding Institute
1250 Arthur E. Adams Dr
Columbus, Ohio 43221
mboring@ewi.org
ABSTRACT
An amine tower was inspected and was shown to have wall
loss over the majority of the circumference. A repair plan was
developed which included welding onto the tower while the
tower remained in operation. The repair plan consisted of two
steps the first of which was welding two Inconel 625 rings to
the SA516-70N steel tower. Secondly, a stainless steel sleeve
was welded to the Inconel 625 rings to cover the corroded area
to contain the tower contents if the tower wall was breached
and to add structural support until the next shut down. The
typical requirements for welds that could possibly be exposed
to amine service are outlined in the API Recommended Practice
945 document (API RP 945). API RP 945 recommends a postweld heat treatment (PWHT) to reduce hardness and relieve
stress but since the planned repair was to be made in-service,
PWHT was not preferred. The objective was to develop a
welding procedure to attach the Inconel 625 straps to the
SA516-70N steel amine tower without using PWHT.
A temper bead welding technique was used to successfully
qualify a welding procedure with hardness values below 200
Brinell without PWHT, in accordance with 2006 NBIC and
2004 ASME Boiler and Pressure Vessel Code Section IX. The
temper bead welding procedure required strict welding heat
input control and weld bead placement. The heat input can be
monitored by controlling the welding parameters or by using
the run-out ratio diagram which was developed during this
project. The temper bead passes need to be deposited in such a
manner that the weld toe of the temper bead is within 3/32 in.
(2.4 mm) of the weld toe of the initial layer.
When manually welding Inconel 625 to SA516-70N with
SMAW electrodes it is recommended that Inconel 182
electrodes be used rather than Inconel 112 electrodes. This
Dr. Yu-Ping Yang
Edison Welding Institute
1250 Arthur E. Adams Dr
Columbus, Ohio 43221
yyang@ewi.org
recommendation is based on the increase ability of the Inconel
182 electrodes to form manganese oxides, which remove the
oxygen from the weld metal. This in turn helps to avoid local
brittle zones which caused the Inconel 112 weld to fail
qualification.
INTRODUCTION
An ultrasonic scan of an amine tower discovered wall loss over
the majority of the tower’s 90-in. circumference. The corrosion
was found in 14 feet of the total 65-foot high tower. The repair
plan included welding two Inconel 625 rings, one above the
corrosion and one below the corrosion, to the SA516-70N steel
tower. A stainless steel repair sleeve would then be welded to
the Inconel 625 rings to cover the corroded area, contain the
tower contents if the tower wall was breached, and add
structural support until the next shut down, which was
scheduled in one year.
The typical requirements for welds that could possibly be
exposed to amine service are outlined in the API
Recommended Practice 945 document (API RP 945). API RP
945 suggests that the hardness of any carbon steel weldment in
Amine service should not exceed 200 Brinell. One method
commonly used to reduce the hardness of completed carbon
steel weldments is to perform a post-weld heat treatment
(PWHT). Since the planned repair was to be made while the
tower remained in service it was preferred that the welding
procedure did not include PWHT. It is important to note that
PWHT relieves residual stress as well as reducing the hardness
of the weld zone and that API RP 945 states that limiting
hardness alone will not prevent stress corrosion cracking and
residual stresses need to be addressed.
The objective of the project was to develop a welding
procedure, using standard industry practice, which could be
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used to attach the two Inconel 625 straps to the SA516-70N
carbon steel amine tower with a resulting hardness of less than
200 Brinell. If possible, the welding procedure should not
require a PWHT to achieve the recommended hardness values.
In addition to the procedure qualification, the welding
application was simulated using thermo-mechanical models to
predict the resulting residual stresses after the welds were
completed.
TECHNICAL APPROACH
First it was necessary to determine if a weld deposited using the
shielded metal arc welding process (SMAW) could be made
onto SA516-70N steel with an acceptable hardness level
without PWHT. Once the technique was defined, a weld was
fabricated and tested to qualify the welding procedure in
compliance with 2006 National Board Inspection Code (NBIC)
and the 2004 ASME Boiler and Pressure Vessel Code Section
IX. Once the welding procedure qualified, the parameters were
used in the residual stress modeling to predict the residual
stress on the inside surface of the amine tower underneath the
weld.
Hardness Trials
A commonly applied welding technique that is used to reduce
the hardness in a carbon steel weldment is a temper bead
technique. A temper bead technique relies on the heat from
subsequent passes to temper the heat-affected zone (HAZ) of
the previously deposited passes.
The hardness trials were carried out on SA516-70N steel in the
horizontal welding position to simulate the application. The
chemistry of the SA516-70N steel used in the hardness trials
had a higher carbon equivalent (CE) value (CE = 0.422) than
the SA516-70N steel used in the construction of the amine
tower (CE = 0.396). The difference in chemistry added to the
conservatism of the results.
The hardness trial welds were deposited using (SMAW) with
Inconel 112 electrodes. The initial layer of the temper bead
technique consisted of several beads deposited side by side
using 3/32-in. (2.4-mm)-diameter electrodes. The initial layer
was lightly ground, prior to depositing the temper bead layer,
until approximately 3/32-in. (2.4-mm) of the un-ground weld
bead remained at the weld toe. The temper bead layer was
made using 1/8-in. (3.2-mm)-diameter electrodes with the
beads deposited approximately 3/32-in. (2.4-mm) from the
weld toe of the initial layer.
Four different variations of the temper bead welding technique
were used to determine which technique would achieve the
required hardness. The four variations included welds with and
without preheat and two different heat input ranges for the
temper bead passes (Table 1). The hardness of the HAZ for
each weld was measured using a Vickers 10-kg hardness
indent. There were 14 indents made in the HAZ of each
hardness trial weld and an additional two indents in two of the
weld sections.
TABLE 1. HARDNESS TRIAL WELDING VARIABLES
Weld
ID
II-Lo
IV-Hi
PH-Lo
PH-Hi
Heat Input Range, kJ/in.
Initial Layer
Temper Layer
10-15
17.5-22.5
10-15
Min. 25
10-15
17.5-22.5
10-15
Min. 25
Preheat
(°F)
(°C)
70
21
70
21
350
177
350
177
Procedure Qualification
The procedure qualification was performed in accordance with
the 2006 National Board Inspection Code (NBIC) and the 2004
ASME Boiler and Pressure Vessel Code Section IX. The
testing requirements included two tensile tests, four bend tests,
and a hardness test.
A 350°F (177°C) preheat temperature and a 450°F (232°C)
interpass temperature were used during the qualification. The
weld joint was a complete penetration V-groove with a 30degree bevel preparation (60-degree included angle). One side
of the joint consisted of 0.5-in. (12.7-mm) SA516-70N steel
(P1) and the other side consisted of 0.5-in. (12.7-mm) Inconel
625 (P43). The SA516-70N steel was also used for the backing
bar. The CE value of the SA516-70N steel was 0.431.
The SA516-70N portion of the weld joint was initially welded
using the same temper bead welding techniques (i.e., electrode
diameter, bead placement and grinding) and parameters that
were shown to produce hardness values below the 200 Brinell
limit from the hardness trials. Care was taken to not allow any
portion of the temper bead pass to contact the SA516-70N base
material. The temper bead layer was then lightly ground to
improve the fit up between the SA516-70N steel and the
backing bar prior to completing the groove weld.
Two welding procedures were evaluated during the course of
the project. The difference between the two procedures was the
welding consumable that was used. The first attempt at
qualifying the welding procedure was done using an Inconel
112 SMAW electrode. The second attempt at qualifying the
welding procedure was done using an Inconel 182 SMAW
electrode.
Residual Stress Modeling
The SMAW procedure that was qualified for the amine tower
repair was thermo-mechanical modeled to determine the
residual stress distributions on the inside surface of the amine
tower. The model variations included the welding direction
(i.e., circumferential or axial), bead size and cooling conditions
as a result of the welding preheat and tower contents.
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The amine tower model had an internal diameter (ID) of 89-in.
(2.26 m), a wall thickness of 0.5-in. (12.7 mm) and a height of
64-in. (1.6 m). The actual amine tower was larger than 64-in.
(1.6 m) in height but the height was limited to reduce the model
size and simulation time while still providing required the
structure stiffness. Inconel 625 ring was 6-in. long. To make
the results conservative the non-welded end of the ring was
fixed in the model. To properly model the residual stress
distribution both a thermal analysis and stress analysis need to
be performed.
The heat flux distribution, as shown in Figure 3, is expressed in
Eq. 1.
Figure 1 shows the two dimensional (2D) axisymmetric model
used to simulate the welding repair procedure for the thermal
analysis of welds deposited circumferentially.
WP1
FIGURE 3. HEAT FLUX DISTRIBUTION
WP2
Inconel 182
Inconel 625
Inconel 625
Inconel 182
q ( x, y , z , t ) = f
SA516-70
0.5’’ thick
ID=89’’
SA516-70
1
Weld procedure 1
−3 y 2
a2
b2
e
−3[ z + v (τ −t ) ]2
c2
e
(1)
The variables a, b, and c are the semi-axes of the ellipsoid, η is
heat efficiency, Q is the power of arc welding process, and t is
the time. Heat loss to the air and to the fluid in the tower was
considered by simulating heat convection.
64’’
The two weld cross-sections in Figure 1 show two possible
circumferential repair procedures. The difference in the
procedures was the bead size of the temper bead layers. Weld
procedure 1 (WP1) used six passes in the Layer 1 and 4 passes
in Layer 2. Weld procedure 2 (WP2) used five passes in the
Layer 1 and 3 passes in Layer 2. The difference in the bead
size between WP1 and WP2 is approximately 16%. The fillet
welds for both welding procedures had identical weld sizes (six
passes).
The weld deposition sequences for the two
circumferential welding procedures are shown in Figure 2.
13 15
11 12 14
7 8 9 10
2 3 4 5 6
abcπ π
e
−3 x 2
6’’
FIGURE 1. FINITE ELEMENT MODEL
16
6 3Qη
14
Fillet
Layer 2
Layer 1
1
11 13
9 10 12
7
8
6
2
3
4
5
To evaluate the effect of welding direction on the residual
stress distribution two more variations of the weld model were
developed where the heating was applied in the axial direction.
Figure 4 shows the welding sequences for a two-layer temper
bead procedure and a one-layer temper bead procedures.
6
6
3
1
3
1
5
2 4
Two layers
5
2 4
One layer
Fillet
(a) Two-layer temper beads
Layer 2
(b) One-layer temper beads
Layer 1
Weld procedure 2
FIGURE 2. WELDING SEQUENCE FOR WP1 AND WP2
DEPOSITED CIRCUMFERENTIALY
A moving arc analysis was used to apply the welding heat to
the circumferential weld models. The moving arc analysis
procedures used a simplified Goldak’s double ellipsoid model.
FIGURE 4. WELDING SEQUENCE FOR WP2 DEPOSITED
AXIALLY
As indicated by Figure 4, the axial weld passes were lumped
into two layers or a single layer. The weld layers were then
heated simultaneously with an equivalent heat input
(approximately 75%) of the heat input used in the
circumferential model. This type of thermal analysis (i.e.,
lump-pass analysis) was been used successfully in the past (ref.
1). The heat input for the lump-pass analysis procedure is less
than the heat input in the moving arc analysis because less heat
is lost in the lump-pass analysis. The fillet welds for the axial
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model trials were deposited with the same heat input and
moving arc analysis that was used in the circumferential model.
It is important to note that the one-layer model is not
representative of a temper bead procedure. It was analyzed to
gather more data on the effect of bead size on the residual
stress distribution.
140
True Stress (ksi)
100
Table 2 lists the welding parameters and heat inputs used for
the thermal analysis. The heat input used for Layer 1 of the
circumferential model was smaller than Layer 2 to achieve the
tempering effect. Also, a higher heat input was used for the
WP2 case because of the increased bead size. The axial weld
models used temper bead layer heat inputs comparable to the
circumferential temper bead layers using the lump-pass
analysis. The fillet weld heat inputs were identical for both the
circumferential and axial fillet welds.
TABLE 2. WELDING PARAMETERS FOR THE THERMAL
ANALYSIS
Weld ID
Layer 1
WP1
Layer 2
Fillet
Layer 1
WP2
Layer 2
Fillet
Layer 1
Layer 2
Fillet
Layer 1
Fillet
Current,
amps
Voltage
, volts
1000F
80
1400F
60
40
1800F
20
0
0
5
10
15
20
25
30
35
True Strain (%)
Fig. 4. Temperature dependent true stress-strain curves of INCO182
FIGURE 5. TEMPERATURE DEPENDENT TRUE STRESSSTRAIN CURVES OF INCO 182
Heat
input,
kJ/in.
RESULTS AND DISCUSSION
The results of each portion of the project is described in the
following sections.
9.5
5.6
8.8
6.3
4.4
8.8
10.0
23.2
15.1
15.1
29.8
15.1
Hardness Trials
The individual hardness values of the HAZ for the four
variations of the temper bead welding technique evaluated are
listed in Table 3. For reference, the hardness values for the
base material were within the range of 151-161 Brinell.
7.5
17.4
15.1
12.5
15.1
Travel
Speed, ipm
Circumferential Model
65
24.4
88
24.6
88
25.1
66
24.0
88
24.8
88
25.1
Axial Model
72F
600F
120
TABLE 3. WELDING VARIABLES FOR HARDNESS TRIALS
88
25.1
8.8
88
25.1
8.8
Weld ID
II-Lo
The thermal-mechanical analysis was performed by applying
the predicted temperature history from the thermal analysis and
the proper boundary conditions from the specific application.
Melting and solidification in the weld zones were considered
with the *ANNEAL command in commercial finite element
software ABAQUS.
For proper stress analysis the model requires the temperature
dependant properties of the materials. Figure 5 shows the
temperature dependent tensile stress-strain curves of the
welding consumable (i.e., Inconel 182). The temperature
dependant mechanical properties for SA516-70 and Inconel
625 can be found in ref. 2 and ref. 3, respectively.
IV-Hi
PH-Lo
PH-Hi
Indent
Loc.
Left Toe
Mid
Right Toe
Left Toe
Mid
Right Toe
Left Toe
Mid
Right Toe
Left Toe
Mid
Right Toe
Brinell Hardness
240
192
196
268
177
205
188
161
177
210
166
166
234
189
193
235
179
201
183
165
185
202
159
164
217
202
204
232
192
192
188
171
203
194
157
165
222
188
216
182
206
197
182
180
228
184
161
167
206
239
183
195
177
232
169
170
A majority, 90%, of the high hardness values highlighted in
Table 3 are located near the weld toe. The mounted weld
sections were inspected and it was discovered that the
inconsistency was mainly related to the placement of the
temper bead weld relative to the weld toe of the initial layer.
The importance of the proper placement of the temper bead
pass can be illustrated by the hardness values taken from Weld
PH-Hi (Figure 6).
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procedure qualification. Inconel 182 was selected because of
the increased amount of manganese, relative to Inconel 112,
which is commonly added to consumables as a deoxidizing
agent.
FIGURE 6. MACRO OF PH-HI HARDNESS TRIAL
The hardness values of the left toe of Weld PH-Hi are higher
then the hardness values of the right toe even though both weld
passes were made at similar heat inputs [22.6 and 21.9 kJ/in.
(0.89 and 0.86 kJ/mm), respectively]. The temper bead pass
that was used to temper the left toe was deposited more than
3/32 in. (2.4 mm) from the weld toe of the previous pass
resulting in hardness values exceeding the 200 Brinell limit.
On the other hand, the temper bead pass that was used to
temper the right toe was deposited less than 3/32 in. (2.4 mm)
from the weld toe of the previous pass resulting in hardness
values less than the 200 Brinell limit.
The results from the hardness trials indicated that welds made
with a 350°F (177°C) preheat temperature along with a temper
bead welding technique, with a minimum heat input value of 25
kJ/in. (0.98 kJ/mm) for the second layer of the temper bead
passes, would result in a weld which would have hardness
values below the 200 Brinell hardness limit. Even though two
of the hardness values located at the left toe of the PH-Hi
hardness trial weld were above 200 Brinell this was related to
the welder technique which can be corrected. The confidence
in these temper bead procedure is further increased due to the
higher CE SA516-70N steel used during the hardness trials
compared to the SA516-70N steel from which the amine tower
was constructed, 0.422 and 0.396 respectively.
Procedure Qualification
Two attempts were made to qualify a SA516-70N steel to
Inconel 625 temper bead welding procedure. The difference
between the two procedures was the welding consumable used.
The first attempt was made with an Inconel 112 welding
consumable. The procedure qualification weld passed the two
tensile tests but half of the bend samples failed due to cracking
(Figure 7).
A brief analysis of the material chemistries indicated that the
failure was probably caused by brittle zones in the weld metal
which could be caused by weld metal oxides that are retained
upon solidification. To address this issue a second Inconel
welding consumable, Inconel 182, was selected for a second
FIGURE 7. ROOT CRACKS IN THE INCONEL 112 PQR
QUALIFICATION BEND SAMPLES
The Inconel 182 qualification weld was made using the same
heat input requirements that were used for the Inconel 112
qualification weld. The bend tests showed no signs of local
brittle failure and the tensile samples also exceeded the
minimum base metal values. Hardness transverses were
performed on the Inconel 182 qualification weld and none of
the hardness indents exceeded the 200 Brinell hardness limit.
It is important to reiterate that the SA516-70N steel which was
used during the procedure qualification had a higher CE value
(0.431) than the SA516-70N steel used during the hardness
trials and the SA516-70N steel used for the amine tower
construction.
Heat Input Control
The temper bead technique relies heavily on the heat input
from subsequent passes to temper the HAZ of previous passes.
One of the main difficulties with temper bead approaches is
assuring that the heat input used during welding is sufficient
and within the procedure requirements.
There are two common methods used to control welding heat
input. The first method requires the actual welding parameters
[current (amps), voltage (volts) and travel speed (inches per
minute)] to be monitored and recorded.
The recorded
parameters are then used to calculate the heat input of each
welding pass by using Eq. (2).
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Heat _ Input =
4.
60 * Voltage * Current
1000 * Travel _ Speed
(2)
The second method is the run-out ratio. In addition to
producing the welding arc, the welding current also heats the
electrode prior to the electrode melting. As a result, the higher
the welding current the more easily an electrode can be melted.
This relationship between the welding current and the rate at
which a SMAW electrode can melt is the basis for the run-out
ratio (ref. 4). The run-out ratio can be calculated by dividing
the weld length by the length of electrode consumed [Eq. (3)].
Run − Out _ Ratio =
Weld _ Length
Electrode _ Consumed
(3)
Run-Out Ratio
It is important to note that the run-out ratio is dependent on the
size and type of the electrode. In other words, the ratio for one
size/type of electrode cannot be used for a different size/type of
electrode. The run-out ratio which was developed during this
project for the two different diameter Inconel 182 electrodes is
shown in Figure 8. The run-out ratio is typically used to
bracket a heat input range or to specify a maximum or
minimum heat input range. When the run-out ratio is used to
bracket in a specific heat input range the welder should target
the midpoint of the specified run-out ratio.
1.4
1.2
1.0
0.8
0.6
0.4
0.2
0.0
Where the horizontal line intersects the run-out ratio
axis is the ratio that is needed to achieve the specified
heat input
Residual Stress Modeling
The two models shown in Figure 2 were analyzed to study the
effect of welding procedure on residual stress. To simplify the
first approach, the fluid flow inside the tower was not
considered. The effect of the inside fluid-flow cooling is
introduced below. The preheating was simulated by uniformly
heating a 6-in. (152.4 mm) height of the tower around the weld
area to 350ºF (177ºC).
During welding, the interpass
temperature was limited to 450ºF (232ºC). During welding, the
air cooling was modeling with heat convection.
Figure 9 shows an example of the predicted temperature
distributions for a bead in the second temper bead layer. The
temperature evolutions are at time intervals when the arc was
approaching the bead, when the arc was near the bead, and
when the arc passed the bead.
3/32 Diameter
1/8 Diameter
8
10
12
14
16
18
20
22
24
26
28
30
Heat Input, kJ/in.
FIGURE 8. RUN-OUT RATIO FOR 182 SMAW ELECTRODES
Using a run-out ratio monitor welding heat input can be
described in a few simple steps:
1. Specify which diameter electrode to use
2. Draw a vertical line through the specified welding
heat input value
3. Draw a horizontal line through the intersection point
of the vertical welding heat input line and the run-out
ratio curve
FIGURE 9. TEMPERATURE EVOLUTION DURING WELDING
USING A MOVING ARC ANALYSIS
Figure 10 compares the axial and hoop stress distributions
between WP1 and WP2. The axial stresses (i.e., top two
images of Figure 10) are bending-type stresses which were
induced by the weld shrinkage in the axial direction. The outer
surface appears to be in compression and the inner surface
appears to be in tension.
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Axial stress
Stress (ksi)
The preheating was simulated by heating a 6-in. (152.4 mm)
length of the outer surface of the tower to 350ºF (177ºC). Due
to the inside surface cooling, the inside surface temperature of
the tower was only able to achieve 315ºF (157ºC) as shown in
Figure 12. During welding, the interpass temperature was
limited to 450ºF (232ºC).
Hoop stress
0 Distance
0 Distance
WP1
WP2
7’’
FIGURE 10. COMPARISON OF STRESS DISTRIBUTIONS
BETWEEN WP1 AND WP2
Stress (ksi)
Figure 11 compares the axial and hoop stress magnitudes on
the inner surface. The predicted hoop stresses caused as a
result of welds deposited using WP1 had a slightly wider
distribution than that of welds deposited using WP2. The axial
stresses were very similar between the two procedures. In
summary, the two procedures produce very similar stress
distributions when cooling from the in-service amine tower was
not applied to the model. The peak values of both axial and
hoop stress are very close with the hoop stress distribution of
WP1 being slightly wider than WP2.
50
45
40
35
30
25
20
15
10
5
0
350ºF
7’’
315ºF
FIGURE 12. PREHEAT SIMULATION FOR INSIDE SURFACE
COOLING COMPARISON
For Case 1 a heat convection coefficient for the air was applied
on both the inside and outside surfaces. For Case 2 a heat
convection coefficient for the air was applied on the outside
surface and a heat convection coefficient 100 times higher than
air (100hair) was applied on the inside surface (Figure 13). The
heat convection coefficient, 100hair was based on experimental
data in which a vessel was welded on the outer surface with
water flow inside the vessel, ref. 5.
Hoop
Air cooling (hair)
Axial
WP1
WP2
Liquid forced cooling (100hair)
No heat convection during welding
0
0.5
1
1.5
2
2.5
Distance (inch)
6’’
6’’
FIGURE 11. INSIDE SURFACE STRESS COMPARISON
BETWEEN DIFFERENT WELDING PROCEDURES
The next portion of the circumferential modeling effort was to
simulate the effect the in-service amine tower had on the
residual stress distribution of the weldment. Since the previous
model results indicated that both WP1 and WP2 produced
similar results, only WP2 was selected for cooling study. For
the comparison, the WP2 model was rerun without and with
inside surface cooling (Case 1 and Case 2 respectively), which
simulated the actual cooling conditions of the amine tower
repair.
FIGURE 13. SET-UP FOR INSIDE SURFACE COOLING
COMPARISON
Figure 14 shows the predicted maximum temperature
distribution for Case 1 and Case 2. It was predicted that higher
temperatures can not transfer to the inside surface due to the
cooling potential of the amine tower products. As a result a
higher temperature gradient between the outer and inner
surfaces was observed for Case 2.
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Copyright © 20xx by ASME
the axial and hoop stress for the Case 1 and the solid lines show
the axial and hoop stress for Case 2. The results indicated that
the hoop stress was reduced significantly with inside surface
cooling, but the peak axial stress increased under the same
conditions.
FIGURE 14. TEMPERATURE DISTRIBUTION FOR INSIDE
SURFACE COOLING COMPARISON
Figure 15 shows the predicted axial stress distributions for
Case 1 and Case 2. The stress values are higher for Case 2
which could be a result of the higher temperature gradients
between the outer and inner surfaces. The higher temperature
gradients tend to produce more bending deformation resulting
in more inner surface tensile stresses. Figure 16 shows the
predicted hoop stress distributions for Case 1 and Case 2. The
inside surface hoop stresses for Case 2 are significantly
reduced. The reduction of the inside surface hoop stress could
be due to the lower inside surface temperature caused by the
fluid flow cooling the inside surface.
FIGURE 17. INSIDE SURFACE STRESS DISTRIBUTION FOR
COOLING CONDITION COMPARISON
In an attempt to reduce both the axial and hoop stress a third
analysis was performed. The third analysis included the inside
surface cooling but the simulated welding was done in the axial
direction which was different from the previous two analyses.
Changing the welding direction was proposed because results
of previous research showed that overlay welding in the axial
direction reduced the residual stress distribution on the inside
surface of pipes (ref 6 and 7). The application of overlay
welding is similar to the first passes of a temper bead welding
procedure. The two cases described in Figure 4 (two-layer and
one-layer) were analyzed to determine if the welding direction
effected the residual stress distribution on the inside surface.
Stress (ksi)
FIGURE 15. AXIAL STRESS DISTRIBUTION FOR INSIDE
SURFACE COOLING COMPARISON
Hoop Stress
Axial Stress
25
14.5
Two-layer temper beads
-2
7.6
One-layer temper beads
FIGURE 18. STRESS DISTRIBUTION FOR THE TWO-LAYER
AND ONE-LAYER APPROACHES
FIGURE 16. HOOP STRESS DISTRIBUTION FOR INSIDE
SURFACE COOLING COMPARISON
Figure 17 compares the inside surface stress resulting from
Case 1 and Case 2 model predictions. The dashed line shows
Figure 18 shows the axial and hoop stress distributions for the
two-layer and one-layer approaches after only the temper beads
were deposited. Both the axial and hoop stresses were
significantly lower for the one-layer approach compared to the
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two-layer approach. The peak axial tensile stresses on the
inside surface were reduced from 25 ksi (172 MPa) (two-layer)
to -2 ksi (-14 MPa) (one-layer). The peak hoop stress on the
inside surface was reduced from 14.5ksi (100 MPa) (twolayer) to 7.6 ksi (52 MPa) (welding in the one layer). It is
believed that the one-layer approach is more effective in
reducing the stress on the inside surface because of the larger
axial shrinkage induced by the larger bead size. The axial
shrinkage induces bending, which causes tension, but also
induces compression on the inside surface. The tension and
compression on the inside surface counteract and if the
compression is strong enough, low stress will be observed as
shown in Figure 18.
50
Two-layer
One-layer
40
Axial
Stress (Ksi)
30
20
10
Hoop
0
-10
0.0
0.5
1.0
1.5
2.0
2.
-20
-30
-40
Distance (inch)
After simulating the fillet weld it was found that the inside
surface axial stresses increased. The increase in the axial
stresses was believed to be due to the additional bending
induced by the completed fillet welds. Figure 19 shows the
stress distributions for the fillet weld made onto the one-layer
model. The peak axial stress increased from -2 ksi (-14 MPa)
without the fillet weld to 26.6ksi (183 MPa) after the fillet
welds.
Stress (ksi)
FIGURE 20. STRESS COMPARISON BETWEEN ONE-LAYER
AND TWO-LAYER APPROACHES
CONCLUSIONS
Temper bead welding procedures can be used to produce welds
with hardness values below 200 Brinell hardness. However, it
is imperative that the temper bead passes are deposited in such
a manner that the weld toes of the temper bead deposit is within
3/32 in. (2.4 mm) of the weld toe of the initial layer.
When manually welding Inconel 625 to SA516-70N with
SMAW electrodes, it is recommended that Inconel 182
electrodes be used rather than Inconel 112 electrodes.
One-layer temper bead
Axial Stress
26.6
Hoop Stress
7
FIGURE 19. FINAL STRESS DISTRIBUTION FOR THE ONELAYER APPROACH
Figure 20 compares the axial and hoop stresses of the two-layer
and one-layer approaches after the fillet weld were simulated.
The figure shows that both the axial and hoop stresses in the
one-layer approach are lower than the two-layer approach.
Comparing Figure 20 to Figure 17 shows that depositing the
temper layers in the axial direction produces less inside surface
axial stresses than welds deposited in the circumferential
direction.
As a result of the temper bead welding technique, the heat input
of the welding process for the welding procedure developed
during this project requires more control than other typical
welding procedures. The heat input can be monitored by
controlling the welding parameters or by using the run-out ratio
diagram which was developed during this project.
The tensile axial stress on the tower inner surface is mainly
induced by welding the INCONEL 625 rings to the temper
beads. The axial stress is a bending-type stress with the
compression near the outer surface and tension near the inner
surface.
The welding of INCONEL 625 rings to the temper beads has
less influence on the hoop stress of the tower inner surface.
The hoop stress on the inner surface of the tower is mainly
induced by welding the temper beads.
It is critical to have strong cooling inside the tower to achieve
low tensile stress or compressive stress on the inner surface for
the temper-bead welding and Tee fillet welding.
Welding the temper beads along the tower axial directions is
better than welding the temper beads along the circumferential
direction from the standpoint of reducing the axial stress on the
tower inner surface.
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For the temper-bead welding along the tower axial direction,
large bead sizes will help reduce the axial and hoop stress on
the inner surface of the tower.
REFERENCES
1. Yang, Y. P., Brust, F. W., and Kennedy, J. C., 2002,
“Lump-Pass
Welding
Simulation
Technology
Development
for
Shipbuilding Applications,”
Proceeding of ASME Pressure Vessels and Piping
Conference, 4-8 August, 2002, Vancouver, British
Columbia, Canada.
2.
Brust F. W., Yang Y. P., and Scott P. M. , “Evaluation
Of Reactor Pressure Vessel (RPV) Nozzle to Hot-leg
Piping Bimetallic Weld Joint Integrity for the V. C.
Summer Nuclear Power Plant”, Contract Number –
NRC-04-97-052, Job Code W6775.
3.
http://www.specialmetals.com/documents/Inconel%20
alloy%20625.pdf
4.
Jackson, C. E., and Shrubsall, A. E., Welding Journal,
Vol. 29, pp. 231s-242s (1950).
5.
Yang Y. P., and Dong P., “Pressure Vessel Residual
Stress Mitigation”, Battelle Research Report, 2002.
6.
Brust, F. W., and Rybicki, E.F., 1981, “Computational
Model of Backley Welding for Controlling Residual
Stresses in Welded Pipes,” Journal of Pressure Vessel
Technology, 103, August, pp. 294-299.
7.
Brust, F. W., and Kanninen, M. F., 1981, “Analysis of
Residual Stresses In Girth Welded Type 304 Stainless
Steel Pipes,” Journal of Materials for Energy Systems,
3(3), pp. 56-62.
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