A Composite Modeling Analysis of
Impression Creep Testing on
Heterogeneous Materials
Model helps predict properties in localized
microstructural zones of weld structures
BY M . A. LISIN, I. D. C H O I , D. K. M A T L O C K A N D D. L. O L S O N
ABSTRACT. The high-temperature mechanical properties of weldments containing discrete microstructural zones were
investigated with the impression creep
test, an indentation creep test that offers
the ability to determine the creep properties of minute microstructural zones by
monitoring the displacement rate of a cylindrical punch as a function of stress and
temperature. The position-dependent
creep behavior across the interface of a
roll-bonded copper-brass laminate, a material which represents ideal two-component weldments, was evaluated. Impression tests at 450° to 525°C (842° to
947 °F) were performed at several locations in the interface region such that the
punch deformation field contained varying amounts of each component. A theoretical model for a multicomponent system was developed and was shown to
correlate with experimental observations
if the deformation field that controlled
punch velocity was assumed equal to the
punch contact area. The implications of
this study with respect to evaluating the
localized creep properties within weldments are presented, and the practical
usefulness of the impression creep test is
discussed.
Introduction
Weldments represent complex heterogeneous materials that contain gradients
in microstructure, composition and properties. Predictions of weldment properties
frequently depend on the ability to char-
M. A. LISIN is with San Diego Gas and Electric,
San Diego, Calif. I. D. CHOI, D. K. Matlock and
D. L. Olson are with the Center for Welding and
joining Research, Colorado School of Mines,
Golden, Colo.
Paper presented at the 70th Annual AWS
Meeting, held April 2-7, 1989, in Washington,
DC.
acterize the properties of each individual
microstructural zone within the heterogeneous structure of the weld. In addition,
prediction of macroscopic properties requires the use of an appropriate composite theory that can accurately describe the
interactions between the various microstructural zones.
The heterogeneous structure of welds
consists of either fine or coarse gradients.
Fine scale gradients result from coring induced compositional changes within individual weld beads (Refs. 1-4). In contrast,
coarser microstructural gradients result
from bead-to-bead variations in thermal
history or due to gradients in both composition and heat from the weld metal to
the base metal (Ref. 5).
The effects of microstructural and compositional gradients on mechanical properties of weldments have been considered in several recent publications (Refs.
6-11). All have shown that a complete
understanding of the overall properties of
welds requires an analysis of position-dependent mechanical properties across the
welds.
Several tests have been developed to
evaluate localized properties within a
weld. These include the use of subsized
samples removed from specific local zones
(Ref. 12), or the use of samples with
KEY W O R D S
Composite Modeling
Impression Creep
Impression Tests
Creep Testing
Heterogeneous Materials
Cu-Brass Laminates
Laminate Composites
Microstructural Zone
Punch Velocities
Activation Energies
microstructures designed to simulate specific zones within the weld. With respect
to high-temperature creep properties, the
impression creep test (Refs. 13-18) offers
the opportunity to evaluate position-dependent properties within a weldment
without the need for specialized sample
geometries. In the impression creep test
the displacement of a cylindrical punch
under a constant applied stress is monitored as a function of time. The resistance
to deformation is controlled by the material immediately below the punch, and
thus, the ability to resolve the properties
of local microstructural zones is limited by
the punch diameter (Ref. 17). The effects
of stress and temperature on measured
displacement rates are analyzed and are
shown to correlate directly to conventional creep data (Ref. 13).
In t w o recent studies, impression creep
results were correlated directly with microstructural gradients across weldments.
Gibbs, ef al. (Ref. 19), in his study of the
creep behavior across austenitic to ferritic
steel dissimilar metal weldments, showed
that the minimum in creep resistance occurs in a zone adjacent to the fusion line
where creep cracking has been observed
in service. Wang, etal. (Ref. 20) evaluated
bead-to-bead variations in the creep properties across a multipass ASTM-A36 steel
weld. Specifically, the creep properties on
a traverse from the cast microstructure of
the last bead, through the reheat-affected
zone of the next-to-last bead, and into the
aged microstructure of the next-to-last
bead were measured. The microstructures associated with this traverse are
illustrated in Fig. 1, and the corresponding
impression creep velocity data are shown
in Fig. 2. The reheat-affected zone of the
next-to-last bead consists of fine equiaxed
ferrite, while the solidified zones of both
beads consist of columnar grains of fine
acicular ferrite surrounded by either elongated or equiaxed proeutectoid ferrite.
The acicular ferrite zone in the next-to-last
WELDING RESEARCH SUPPLEMENT 1159-s
fact that the indenter plastic zone contains
multiple components. In addition, continuous gradients from one zone to the next
make the delineation of individual components difficult. In order to extend the
impression creep test to complete analyses of weld metal with continuous microstructural gradients, an appropriate composite theory is required. In this paper, the
basics of such a theory are developed and
predictions are compared to impression
creep data.
GRAIN REFINED HAZ
AS-SOLIDIFIE
ZONE
An ideal material for study should consist of microstructural gradients with welldefined interfaces between microstructural zones. Furthermore, each zone
REHEATED
^'^ :
MiWy-^-^4
should be larger than the punch diameter
f\S-SOLIDIFIED Jf ~"M'^*'
SSaSfc
v
so that the impression creep properties of
ZONE
' ,~..ft<;t-'•'•"'-•- . ;: X-f
the individual zones can be determined.
To simplify the analysis, a two-component
roll-bonded laminate composite was chosen. In the composite, impression creep
data were obtained in each base metal
/%. 7 — Microstructural zones existing within a multipass SMA W A36 low-carbon steel weldment. and as a function of position across the
Light photomicrograph, nital etch. From Ref. 20.
diffusion-bonded interface.
bead appears to have coarsened due to
aging. The minimum in creep resistance, as
evidenced by a maximum in impression
creep punch velocity, occurs in the finegrained, reheat-affected zone. The observation of a decrease in creep resistance
with grain refinement is consistent with
general creep theory (Ref. 21). These re-
sults on ASTM A36 indicate that the variations in creep behavior across the multipass weldment mirror the variations in microstructure.
To date, impression creep tests on
weldments have been unable to determine properties within a single microstructural zone. This inability stems from the
Table 1—EDS Analysis of the Test Material Compositions (wt-%)
C11000 (12.7mm)
C26000 (12.7mm)
10
Cu
Ni
Zn
Co
99.9
69.2
0.01
0.05
30.6
0.01
0.01
I
' 1
1
Fe
0.02
0.05
1
MMt
0.01
0.14
~\
Grain Refined
HAZ
Reheated A s - S o l i d i f i e d
Zone
Mn
A s - S o l i d i f ied
Zone
»Mm
J» 8
E
*«
6
252 MPa
567 °C
1 . 0 mm I n d e n t e r
x
•
v .
J
fi.
i
-
'A
(
3 4
•
>
3
-
J
2
-
1
0
-
•
i
1
1
2
3
Experimental Procedure
A roll-bonded laminate composite was
fabricated from copper Alloys C11000
(commercially pure copper) and C26000
(cartridge brass, nominally 70% Cu and
30% Zn). The composition and initial plate
thickness of each alloy are shown in Table
1 T w o 102- X 150-mm (4- X 6-in.) plates,
one of each alloy, were encased in a
stainless steel canister attached to an
active vacuum system and preheated to
820°C (1508°F). The canister was hot
rolled on a laboratory rolling mill and the
thickness was reduced 20% in multiple
passes. Metallographic analysis and roomtemperature mechanical testing verified
that a complete bond was obtained (Ref.
22). In addition, prior to creep testing, the
composite was annealed at 625°C
(1157°F)for 4 h.
Impression creep tests were performed
with the test system developed by Gibbs
(Refs. 17,19). A 1-mm (0.040-in.) diameter
molybdenum punch was used. Creep
data for the copper-brass system were
obtained at temperatures between 450°
and 525°C and stresses between 90 and
117 MPa (13 and 17 ksi).
The microstructure of the composite
prior to testing and the deformed zones
below the punch were analyzed with
standard light metallographic techniques.
The copper-brass composite was etched
in a solution of 5g K2O2O7, 20 mL H 2 S0 4 ,
5 mL HCI and 250mL H2O. Chemical analysis across the laminate composite was
obtained with a JEOL JXA-840 scanning
microanalyzer.
Results
D i s t a n c e f r o m H A Z / R e h e a t e d Weld I n t e r f a c e (mm)
Fig. 2 — Steady-state impression creep punch velocity as a function of position within the multipass Two sets of experimental impression
creep data were obtained to provide inSMAW A36 low-carbon steel weldment shown in Fig. 1. From Ref. 20.
160-s | APRIL 1990
. ^
Copper
0.25 mm
Fig. 3 — Light micrographs.
A — The copper component;
B - the interface; C— the brass
component in the copper-brass
lamjnate_
Copper y
VV
,/
•
.. .
-• ' / '
•
-: \
p >
— •-.•.«*•>.; . . » » • -f^W
••»%'.»
=^x
<J^
I
Brass
put to the composite model analysis,
which follows in the Discussion section.
First, impression creep properties of each
base material in the composite were determined. Then, the effect of position with
respect to the interface was evaluated.
Light micrographs of the base materials
and interface are shown in Fig. 3. Both
base metals consist of a duplex grain
structure in the annealed condition. The
interface shows a complete bond and the
base metal grain structures are constant
up to the interface. The zinc and copper
compositions adjacent to the interface
were measured with energy dispersive
spectroscopy (EDS) techniques, and the
results are shown in composition profiles
in Fig. 4. Figure 4 shows that the gradient
in composition occurs entirely within 0.1
mm of the interface. The dimension represents 10% of the punch diameter. Consequently, tests in the interface region
represented primarily the properties of
the t w o components, with only a small
contribution to punch velocity from the
continuous compositional gradient region.
0.25 mm
1
Brass
displacement. Steady-state punch velocities were taken as the average punch velocity between 12 and 36 h. Also note that
the steady-state punch velocity of the
brass is approximately 3.9 times that of the
copper. The creep rates of the t w o constituents are different enough that the influence of one component on the creep
rate of a two-component composite
should be easily detected, and yet the t w o
creep rates are similar enough that they
can both be measured accurately under
40
1
T
identical stress and temperature conditions. It should be noted the microstructures of the as-tested samples revealed
microstructural changes below the punch:
recrystallization in brass and grain growth
in copper (Ref. 22). However, a systematic
microstructural examination showed that
all changes in microstructure developed
within the first 12 h, and microstructures
were stable (Ref. 22) for times greater than
12 h. Therefore, the measured steadystate creep rates reflected a constant mi-
1
1
1
1
1
I
\
30 --
^
•
•
_
_
-
I
I
0.25 mm
•£.- .
n
-
"
n
"
1
1
n
n
100
90 t!
20
1/
A!
D Zn
• Cu
10
- 80
Impression Creep of Base Metal
Typical impression creep curves for the
copper and brass base metals at 475°C
and a punch stress of 117MPa (17 ksi) are
shown in Fig. 5. Both curves show a region
of transient creep, in which the punch velocity (i.e., the slope of the displacementtime curve) decreases with displacement,
and a region of steady-state creep where
the velocity is constant, independent of
a
n
n
"
a
n
D
,/
•
»
T\ •
_
•
.
•
- 70
-
Copperi Brass
i
>
i
i
i
60
-0.2 -0.1 0.0
0.1
0.2
0.3
0.4
Distance from Interface (mm)
Fig. 4— The effect of position on the zinc and copper compositions (EDS analysis) in the interfacial
region of the copper-brass laminate.
-0.4
>
i
-0.3
WELDING RESEARCH SUPPLEMENT 1161-s
each system. An average stress exponent
of 4.0 was calculated for both materials.
The maximum deviation in n was obtained
in the brass system with measured values
between 3.9 and 4.05.
The plots of punch velocity versus
inverse absolute temperature in Figs. 8
and 9 yield values of the apparent activation energies for impression creep. The
slopes are approximately equal at each
stress level, suggesting that the activation
energies remain constant within the stress
range investigated. Furthermore, the fact
that data plot linearly and parallel at each
stress level indicates that the same creep
mechanism or mechanisms dominate over
the entire stress and temperature ranges
investigated. Average activation energy
for copper is 17.4 kcal/mole, and for brass
is 25.2 kcal/mole. Maximum deviation in
the individual activation energies for both
materials at each stress level was less than
7% of the average value.
The significance of the measured n and
Q values for impression creep in each
system can be evaluated by comparing
these results to published Q and n values,
summarized in Table 2, obtained on conventional creep samples. The measured
stress exponent of 4.0, although slightly
lower, is in reasonable agreement with the
published values which range between
4.5 and 5.0.
The published creep activation energies
for brass and copper vary significantly. A
Table 2—Creep Activation Energies and Stress Exponents for Copper and Brass Alloy C26000
Obtained with Conventional Creep Samples
Material
Copper
Temperature
°C
above
405 to
450 to
413
above
500
Brass
Alloy 260
610
610
530
510
-
—
31
30
12.5
—
-
550 to 700
500
48.0
28.0
32.8
27.5
42
49.6
24.8
20.2
Interpreted
Mechanism
41.9
Reference
—
(bulk diffusion)
(boundary diffusion)
(boundary diffusion)
-
-
4.5
28
—
-
5
-
4.5
29
30
31
32
33
(boundary diffusion)
23
23
24
25
27
26
26
31
-
(bulk diffusion)
crostructural condition.
The effects of stress and temperature
on punch velocities were analyzed according to the following equation:
Bo- n exp(-Q/RT)
Stress
Exponent
4.8
4.8
5
4.8
5.0
—
413
500
Activation Energy
kcal/mole
(I)
where v p is the punch velocity, rr is the
true stress, n is the stress exponent, Q is
apparent activation energy for impression
creep, B is a constant, and R and T have
their usual meanings. Previous studies (Ref.
13) have shown that v p , n and Q obtained
from impression creep data relate directly
to conventional creep data.
The effects of stress at constant temperatures on steady-state impression
creep velocities are shown in Figs. 6 and
7 for copper and brass, respectively. Correspondingly, the effects of temperature
at constant stresses are shown in Figs. 8
and 9. The linear relationships shown in
Figs. 6 to 9 indicate that Equation 1 is applicable to both copper and brass for the
test conditions of this study.
From the slopes of Figs. 6 and 7, the
stress exponents, n, were obtained for
10
0.8
Punch Stress
Temperature
Copper
Stress Exponent=4
0.6
u 10
0.4
a
0.2
V-1.08xlCr« mm/s
•
•
10
0.0
10
20
30
Time (Hrs)
40
50
SO
1
i
i
90
100
110
Stress
(MPa)
Punch
4 7 5 °C
5 0 0 °C
5 2 5 °C
120
130
Fig. 5 —Impression creep curves for the unlaminated copper and brass Fig. 6 — The effect of stress on the steady-state impression punch velocity for copper at three test temperatures. The slope is equal to the stress
plate materials tested at 475°C and 117 MPa.
exponent in Equation 1.
162-s|APRIL 1990
10
10
•
•
A
90 MPa
104 MPa
117 MPa
ui
E
E
u 10
u 10
aj
>
Q.
0.
Brass
•
•
•
Stress Exponent=4
Copper
450 °C
475 °C
500°C
Activation Energy
=17.4 kcal/mole
10
100
90
BO
110
120
Punch Stress (MPa)
Fig. 7— The effect of stress on the steady-state impression punch velocity for brass at three test temperatures. The slope is equal to the stress
exponent in Equation 1.
comparison with the data in Table 2,
shows that for both systems the apparent
activation energies for impression creep
compare with the lower values for conventional creep.
With the measured n and Q values for
impression creep of copper and brass, the
preexponential constants in Equation 1
were determined, and the following unified velocity equations were developed
(with v in mm/s and a in MPa).
Vp(copper) = 7.1 X 1 0 _ l 0 a 4 e x p
/ - 1 7 . 4 kcal/mole\
{
RT
J
2
<>
vp(brass) = 5.8 X l O ^ V e x p
- 2 5 . 2 kcal/mole\
RT
j
<3>
These equations were shown to adequately describe all of the impression
creep data of this study.
Impression Creep of Laminate Composites
Punch velocity data at positions that
traverse the interface region are plotted
as a function of location with respect to
the interface in Fig. 10. Each data point in
this figure indicates the position of the
punch centerline with respect to the interface. All tests plotted in this figure were
performed at a punch stress of 104 MPa
(15 ksi) and a test temperature of 475°C
_i_
10
130
1.24
1.26
1.2B
1.30
1.32
1.34
Inverse Temperature x 1000 CK)"1
..36
Fig. 8 — The effect of temperature on the steady-state impression punch
velocity for copper at three stresses. The slope is equal to Q/R in Equation 1.
(855°F). Also indicated in this figure is the
diameter of the indenter. Note that the
indenter diameter is much larger than the
symbols used to plot the data. Thus, a data
point located 0.5 mm (0.020 in.) from the
interface indicates that the punch is entirely within one component, and that the
punch wall is tangent or immediately adjacent to the interface. Two predictions,
based on a theory developed below for
impression creep within a multicomponent system, are also shown.
Figure 10 shows that in a test in which
the punch was located entirely within the
copper but immediately adjacent to the
interface, the punch velocity is approximately equal to the average punch velocity within the copper away from the
interface. This indicates that the adjacent
brass material had no effect on the punch
velocity at this location. This same observation holds true for tests conducted in
the brass. Consider the data point located
at - 0 . 6 2 mm. In this case, the punch wall
is only 0.12 mm (0.005 in.) from the interface, yet the punch velocity is within the
scatter for the average punch velocity
within the brass. These observations show
that only the material immediately below
the punch controls the punch velocity.
creep velocities can be evaluated following the composite modeling approach introduced by Gibbs (Ref. 17). This analysis,
based on an isostrain composite, is summarized below and is a uniform strain approximation of the deformation behavior
within a nonuniform deformation zone.
The applied load in impression creep is
supported by the effective volume of
material that is deformed below the
punch. If this volume includes t w o components, then the applied load, Fp, is:
FP = FT + F2
Fp
aA n
The effect of punch position with respect to the interface on the impression
(5)
where A p is the actual punch area and a
is a parameter (a > 1) that defines the size
of the deformation zone. Correspondingly, the stresses in each zone are:
(6)
°2
Discussion
(4)
where F-| and F2 are the loads supported
by the t w o components.
The average applied stress within the
deformation zone, an, is:
=
F2
A7
(7)
where:
A i + A2 = aAp
(8)
WELDING RESEARCH SUPPLEMENT 1163-s
v p = VT = v 2
10
(12)
At a constant temperature (with K = Bexp
(-Qi/RT)):
Vn-KiOVn
(13)
n
v 2 = K2a2 2
(14)
with ni. = n>, as is the case for the copper
and brass of this study, combining Equations 12, 13, and 14 with Equation 11
yields:
(15)
u 10
•
Solving Equation 15 for v p results in the
following expression, which can be compared to the experimental data:
SL
>
JZ
u
C
=1
Q-
Fig. 9 - The effect of
temperature on the
steady-state
impression punch
velocity for brass at
three stresses. The
slope is equal to
Q/R in Equation 1.
X1
Brass
Activation Energy
=25.2 kcal/mole
J_
1.30
1.32
1.34
1.36
1.38
Inverse Temperature x 1000 (°K)"'
10
1.28
and A, and A2 are the deformation areas
in each component. The area fraction of
each component are:
Aj
X,=
90 MPa
104 MPa
117 MPa
1 .40
(11)
a p = X-|0"i + X2CI2
For this system, the average punch velocity, v p , is equivalent to the punch velocities within each zone, such that:
aA p
(10)
x? =
K2
Combining Equations 6 to 10 with Equation 4 yields the rule of mixtures expression for stress:
(9)
aA n
T
14
Ul
T'
-1—
1
^ i
•
12
•
\E
-^10 -
•
•
^N
\
N
0
X
1
1
B
•
\
\
\
\«\
„
Theory, a=9
Theory. a = l
Exp. Data
"\
\
-
4-1
u
6
0
"
Indenter Dia.
OJ
1^
V P -•.__
> 4
•
•
•
•
C
D
Brass
n
-4
1
1
1
Copp er
1
1
1
- 2 - 1 0
1
2
Distance from Interface (mm)
Fig. 10 — Punch velocity in the interface region of the copper-brass laminate. The horizontal position
of the data indicates the location of the punch centerline with respect to the interface. The punch
stress was 104 MPa, and the test temperature was 475 "C. Predictions of the composite model for
two assumed deformation zone sizes are shown.
164-s I APRIL 1990
x2
(16)
1/r
To utilize Equation 16, two conditions
must be specified: the position of the
punch with respect to the interface and
the effective size of the deformation zone
(i.e. a in Equation 5). Previous analyses,
which have included metallographic observations of surface deformation adjacent to the punch (Refs. 14,17) and slip line
field calculations, have indicated that the
diameter of the deformation zone may be
as much as three times the punch diameter. A deformation zone diameter three
times the punch diameter is described by
an a (in Equation 5) of 9.
To evaluate the data shown in Fig. 10,
t w o predictions of the position-dependent impression punch velocity were
made; one with an a of 1, a calculation
that assumes the zone immediately below
the punch controls the punch velocity,
and one with an a of 9. The significance
of the choice of a on the calculated component area fractions that correspond to
a specific punch position is illustrated in
the following example calculation. With
the interface 0.4 mm (0.016 in.) from the
centerline of a 1-mm (0.040-in.) diameter
punch, the area fraction of the minority
constituent is only 0.052 for a = 1, but is
0.33 for a = 9.
The predicted punch velocities for a
values of 1 and 9 are plotted in Fig. 10
along with the experimental data. Both
calculations show a continuous transition
in the impression creep behavior across
the interface. The continuous transition
extends to further distances from the
interface with an increase in a. A comparison of the experimental data with the
theoretical predictions in Fig. 10 shows
that there is excellent agreement for the
case of a = 1. This correlation implies that
the material immediately below the punch
primarily controls the observed impression punch velocity.
In addition to the copper and brass data
in which the creep rates of the base materials differed by less than one order of
m a g n i t u d e , t h e material i m m e d i a t e l y b e l o w t h e p u n c h c o n t r o l l e d p u n c h rates in a
n i c k e l - c o p p e r laminate c o m p o s i t e , w h e r e
the base metal impression p u n c h v e l o c i ties d i f f e r e d b y greater than t w o orders of
m a g n i t u d e (Ref. 22). The observations in
b o t h t h e copper-brass and c o p p e r - n i c k e l
systems are significant w i t h respect t o i m pression c r e e p testing w i t h i n w e l d m e n t s
containing fine microstructural gradients.
The data p r o v i d e an u p p e r limit t o p u n c h
diameter w h e n a t t e m p t i n g t o characterize
the c r e e p p r o p e r t i e s of m i n u t e m i c r o structural zones. As long as the p u n c h fits
entirely w i t h i n the microstructural z o n e of
interest, the m e a s u r e d c r e e p p r o p e r t i e s
will n o t significantly be a f f e c t e d b y the
adjacent microstructural zones.
A c o m p o s i t e m o d e l i n g a p p r o a c h has
b e e n p r e s e n t e d f o r the analysis of impression c r e e p data in a simple t w o - c o m p o nent system. The theoretical predictions
w e r e s h o w n t o correlate w i t h e x p e r i m e n tal observations w h e r e the d e f o r m a t i o n
z o n e that c o n t r o l s p u n c h w a s assumed
equal t o t h e p u n c h diameter. In m o r e
c o m p l e x systems, such as w e l d m e n t s in
w h i c h the d e f o r m a t i o n z o n e m a y contain
m o r e than t w o discrete microstructural
zones or a c o n t i n u o u s microstructural
gradient, it should b e possible t o predict
localized c r e e p properties f r o m impression c r e e p measurements by using a direct
analog of t h e analysis p r e s e n t e d in this
paper. A n e x a m p l e analysis of this app r o a c h is p r e s e n t e d e l s e w h e r e (Ref. 22).
T h e impression c r e e p test p r o v i d e s a
simple m e t h o d t o evaluate the local c r e e p
p r o p e r t i e s of materials. Several potential
industrial applications exist:
1) T h e existence a n d location of zones
w i t h l o w c r e e p strengths w i t h i n a w e l d m e n t can be d e t e r m i n e d b y a direct c o m parison of p u n c h velocities. This i n f o r m a t i o n c a n g u i d e w e l d i n g p r o c e d u r e and
material selection.
2) Changes in p u n c h v e l o c i t y a n d c o n sequently c r e e p resistance can be determ i n e d as a f u n c t i o n o f time. This allows a
direct c o m p a r i s o n o f the service behaviors of candidate materials, w e l d i n g
p r o c e d u r e s , heat t r e a t m e n t s , etc. T h e effects o f t i m e - d e p e n d e n t microstructural
changes (e.g., carbide f o r m a t i o n , grain
g r o w t h and decarburization) can b e evaluated.
3) Punch v e l o c i t y data can b e used in
an activation e n e r g y analysis. T h e resulting
activation e n e r g y can be used in a LarsonMiller t y p e of analysis for the p r e d i c t i o n of
residual life.
Conclusions
1) Impression c r e e p data have b e e n
o b t a i n e d f o r c o p p e r Alloys C 1 1 0 0 0 and
C 2 6 0 0 0 (brass). Both materials display a
stress e x p o n e n t equal t o 4.0. A c t i v a t i o n
energies of 17.4 k c a l / m o l f o r c o p p e r and
25.2 k c a l / m o l f o r brass h a v e b e e n deter-
m i n e d . These values c o m p a r e f a v o r a b l y
w i t h published values o b t a i n e d f r o m c o n ventional c r e e p tests.
2) C o p p e r - b r a s s laminates h a v e b e e n
successfully f a b r i c a t e d b y roll b o n d i n g .
T h e materials h a v e p r o v i d e d ideal systems
f o r m o d e l i n g of microstructural gradients.
3) Based o n t h e predictions o f a t w o c o m p o n e n t c o m p o s i t e m o d e l , the i m pression c r e e p p u n c h v e l o c i t y appears t o
b e influenced primarily b y material i m m e diately u n d e r the p u n c h . This p r o v i d e s an
u p p e r limit t o p u n c h diameter w h e n perf o r m i n g impression c r e e p tests w i t h i n m i crostructural gradients such as w e l d m e n t s .
Acknowledgments
T h e authors appreciate and a c k n o w l e d g e the research s u p p o r t of the O f f i c e of
Basic Science of the U.S. D e p t . of Energy.
Also, the assistance of M r . Shing-Hoa
W a n g w i t h the w o r k o n A 3 6 steel cited in
the I n t r o d u c t i o n is a p p r e c i a t e d .
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W E L D I N G RESEARCH S U P P L E M E N T 1165-s
WRC Bulletin 344
June 1989
This Bulletin contains two reports covering three-dimensional finite element analysis of 45-deg lateral
branch pipe models.
( 1 ) Three-Dimensional Finite Element Analysis of PVRC 45-Degree Lateral Model 4 (d/D = 0.5,
D/T = 40) under Out-of-Plane Moment Loading on Branch Pipes
By P. P. Raju
( 2 ) Three-Dimensional Finite Element Analysis of 45-Degree Lateral Model 2 (d/D = 0.5, D/T = 10)
under Out-of-Plane Moment Loading on the Branch Pipe
By P. P. Raju
Publication of these reports was sponsored by the Joint Task Group on Laterals of the Subcommittee
on Piping, Pumps and Valves, and the Subcommittee on Reinforced Openings of the Pressure Vessel
Research Committee of the Welding Research Council. The price of WRC Bulletin 344 is $16.00 per copy,
plus $5.00 for U.S., or $8.00 for overseas, postage and handling. Orders should be sent with payment to
the Welding Research Council, 345 E. 47th St., Room 1301, New York, NY 10017.
WRC Bulletin 347
September 1989
This bulletin contains two reports:
( 1 ) Welded Tees Connections of Pipes Exposed to Slowly Increasing Internal Pressure
By J. Schroeder
( 2 ) Flawed Pipes and Branch Connections Exposed to Pressure Pulses and Shock
Waves
By J. Schroeder
Publication of these reports was sponsored by the Subcommittee on Reinforced Openings and External Loadings of the Pressure Vessel Research Committee of the Welding Research Council. The price of
WRC Bulletin 347 is $25.00 per copy, plus $5.00 for U.S. and $10.00 for overseas, postage and handling.
Orders should be sent with payment to the Welding Research Council, Room 1301, 345 E. 47th St., New
York, NY 10017.
WRC Bulletin 349
December 1989
This bulletin contains two reports that evaluate the PWHT cracking susceptibility of several Cr-Mo steels
and several HSLA pressure vessel and structural steels.
Postweld Heat Treatment Cracking in Chromium-Molybdenum Steels
By C. D. Lundin, J. A. Henning, R. Menon and J. A. Todd
Postweld Heat Treatment Cracking in High-Strength Low-Alloy Steels
By R. Menon, C. D. Lundin and Z. Chen
Publication of this report was sponsored by the University Research Committee of the Welding Research
Council. The price of WRC Bulletin 349 is $35.00 per copy, plus $5.00 for U.S., or $10.00 for overseas,
postage and handling. Orders should be sent with payment to the Welding Research Council, 345 E. 47th
St., Room 1301, New York, NY 10017.
166-s | APRIL 1990