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Proceedings of PVP2005 2005
ASME Pressure Vessels and Piping Division Conference
July 17-21, 2005, Denver, Colorado USA
PVP2005-71108
THERMAL SHOCK CRACKING: DESIGN AND ASSESSMENT GUIDELINES
Professor John W.H. Price
Monash University, Department of Mechanical Engineering
PO Box 197, Caulfield East, Vic 3145, Australia.
john.price@eng.monash.edu.au.
ABSTRACT
Repeated thermal shock cracking is common in the
operation of pressure equipment where water and steam are
present. Surprisingly it is not directly covered in the ASME
Boiler and Pressure Vessel code nor in fitness-for-purpose
recommended practice such as API 579.
An example of thermal shock stresses occurs when hot
surfaces exposed to splashing of cold water. This eventually
may lead to crack nucleation and crack growth. However not
all thermal shock cracks lead to failures (such as rupture, leak
or in more brittle material fragmentation), indeed the most
frequent situation is that the cracking arrests at a depth of a few
millimeters.
This paper presents a unique experimental study and
analysis the information being gained from this study in terms
of design guidelines and crack growth mechanisms. In the
experiments, cracks are initiated and then grown in low carbon
steel specimens exposed to repeated thermal shock. The test-
rigs achieve large thermal shocks through the repeated water
quenching of heated flat plate specimens. The effect of steady
state loads on the growth and environmental effects due to the
aqueous nature of the testing environment are found to be
major contributors to the crack growth kinetics.
The most important findings are that are that the
conditions leading to both the initiation and the arrest of cracks
can be identified and that the depth of a starter notch
contributes little to the crack propagation.
Key words: Thermal shock, cracking, initiation, growth,
arrest, design, fitness for purpose
INTRODUCTION
Down shocks often occur when low temperature fluid
strikes an already hot surface. Another common situation is
where there is rapid depressurisation such as can be caused by
sudden leaks or valve operations. Depressurisation causing
thermal shock also occurs when high pressure water is vented
through orifices.
Down shocks induce a very high skin tension stress on the
component which decreases rapidly through the thickness of
the component, as shown on Figure 1. In normal operation
there is usually also a primary mechanical load due to pressure
and dead weight loading which is applied to the component.
An important recent failure is reported by Alexander et al. [1
In this work, stresses created as a result of thermal shock
are distinguished from stresses which occur during uniform
thermal expansion. Thermal shock stresses decay rapidly both
in time and across the section as shown on Figure 1, whereas
thermal expansion stresses are more constant with time and
across the section. Thermal expansion stresses are in effect
similar to cycling mechanical stresses. Thermal shock stresses
can be regarded as Classification F, Peak stresses in Table 4-
120.1 of ASME VIII Division 2 [2].
Figure 1 Distribution of thermal shock and
mechanical stress across a section.
Thermal shock driven cracking is one of the most
common cracking phenomena observed in many types of
pressure equipment such as those in electricity generating
boilers, nuclear plant, steam turbine auxiliary plant and other
situations [3,4,5]. Thermal shock cracking tends to start at
geometrical discontinuities as shown on Figure 2.
Copyright © 2005 by ASME
2
Figure 2. (a) Cracking from external corners such as at penetrations (b) Cracking at internal corners
Figure 2(c). Examples of geometrical details affecting thermal shock cracking from power stations. Left at the
intersection of two drain lines. Right in an economiser inlet header.
This paper is based on experimental and theoretical work
done at Monash University. The experimental work is
currently limited to carbon steel operating at temperatures
below the creep range in a water/steam environment.
Experimental work
Two thermal fatigue test rigs have been purpose built at
Monash for the investigation of crack initiation and growth due
to repeated thermal shock loading. The experimental work that
is being conducted is unique in several ways and uses full-scale
test specimens which mimic plant conditions [6]. There have
been other experimental tests of the thermal cracking
phenomena but for various reasons they have not simulated the
conditions in operating plant. This has produced a number of
limitations to the results from such testing, in particular they
have not detected the importance of mechanical loading and
environment on the growth rate of cracks [7;8].
The tests at Monash take several months to complete as
the specimens are put through thousands of cycles which take
15 to 20 minutes each. The specimens are alternatively heated
in a convection furnace and then cooled by sprays of cold
water. Approximately one-dimensional conditions exist at any
one crack because of use of attached thermal masses as shown
on Figure 3. The growth of the cracks is recorded every
hundred or so cycles by removing the specimen from the
furnace and examining it under a microscope.
Recent experimental work at Monash has included a new
variety of starting notch sizes and a horizontal test arrangement
as shown on Figure 3. The horizontal arrangement permits a
larger number of higher temperature tests to be conducted than
does the vertical furnace arrangement. The atmosphere in the
vertical furnace has some convection movement of heat which
means that higher parts of the specimen are hotter than lower
parts of the specimen. This means only the top is subjected to
the maximum thermal shock. The horizontal furnace has some
different features; in particular, liquid water is present in the
crack for longer after the shock, which may affect corrosion
rates.
In association with this experimental work there are a
number of thermal shock cracks in power plant that have been
Copyright © 2005 by ASME
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observed on an occasional basis over a number of years. These
have been used to confirm the work.
The stress intensity factor produced by thermal shock
stresses is dependant not only the stresses indicated on figure 1
but also on the depth of the crack (including any notch). This
has been determined using the techniques of ASME Section XI
and are shown on Figure 4. Note that the combined stress
intensities are modified by plastic zone correction factors.
! "
# " $ %!& !
&'( )! * + !
$
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Figure 3. Experiment design. Left, vertical furnace, right horizontal furnace
Figure 4. Maximum stress intensity factor profiles during 7sec shock from 370ºC, with and without 90MP a primary
load.
0
10
20
30
40
50
60
0
5
10
15
20
Distance from quenched face (mm), d
Stress Intensity Factor ("SIF") (MPa.m^.5)
Total SIF
SIF due to mechanical loading
SIF due to thermal shock loading
Copyright © 2005 by ASME
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RESULTS
Many cracks have now been grown on both the
horizontal and vertical rigs and have covered a wide range of
conditions. Some of the factors that have been varied include:
• Peak temperature (T ºC). Temperatures from 240ºC
to 400ºC have been used.
• Constant stress applied (P MPa). This is generally
either zero or 90 MPa stress.
• Quench time (Q seconds). This is the time for which
cold water is sprayed on the specimen. From theoretical work,
peak stress at 3.5 mm was achieved after 7 seconds
• Starter notch, ao (mm). Starter notches from 1 mm to
3.5 mm have been tested and cracks grown from all such
notches.
A selection of the main results in the form of crack
depth, a, versus crack growth rate is given for the two furnaces
in Figures 5 and 6. Figure 7 shows an enlarged view of a
thermal shock crack.
Figure 5. Crack growth rate versus crack length (includes notch depth) for a number of cracks in the vertical furnace
experiments. T = maximum cycle temperature (°C), P = primary mechanical load (MPa), Q = quenching time (time of
water application) in seconds. Notch depth, ao, is 3.5 mm for all cases.
Figure 6. Data generated during horizontal rig experiments. Crack growth rate versus crack length for various notch
depths, ao.
Crack Growth Rate vs. Crack Length
0.000E+00
2.000E-07
4.000E-07
6.000E-07
8.000E-07
1.000E-06
1.200E-06
1.400E-06
0246810
Crack Length 'a' (mm, includes notch depth)
Crack growth rate 'da/dN' (m/cycle)
Q=10, T=400, P=90 Q=10, T=400, P=0
Q=10, T=370, P=90 Q=10, T=370, P=0
Q=10, T=330, P=90 Q=10, T=330, P=0
Q=10, T=280, P=90 Q=10, T=280, P=0
Q=10, T=240, P=90 Q=10, T=240, P=0
Q=7, T=370, P=90 Q=7, T=370, P=0
Q=7, T=330, P=90
Conservative
da/dN=B(a-L)
Best Fit
Copyright © 2005 by ASME
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(a)
(b)
Figure 7. A thermal shock crack at 20 times magnification from (a) the side, (b) fracture surface.
The various growth regimes are indicated. 90MPa loaded specimen with top temperature of 400C.
DISCUSSION
General observations
The main observations from this work are as follows:
• All of the cracks that have been grown have arrested.
• Cracks on specimens with primary loadings have at
first accelerated in growth rate, reached a plateau of growth rate
and then arrested.
• Cracks growing without primary loading have tended
to grow at slower rates and arrest earlier.
• Most cracks have arrested in less than 3 mm growth.
A few cracks growing from the theoretically most severe notch
tested, (ao= 3.5 mm) have grown further to a maximum of 6.2
mm.
Both the physical appearance of the cracks and the growth
rate data suggest there are some simple major features of the
crack growth.
1. Growth in the early stages up to about 1 mm exhibits
the features of plastic tearing – raised and less corroded
surfaces
2. Growth with a large amount of corrosion..
3. Growth region where the corrosion is limited and there
is sometimes branching of the crack (Paris and arresting).
Growth stages of the cracks
To simplify the picture of what is being observed consider
Figure 8. This figure shows two sets of data presented for two
specimens from the vertical furnace which differ only in that
one has a mechanical stress of P = 90 MPa applied.
Paris
Law
Corrosion
dominated
Arresting
Loaded specimen
Notch
~1mm
Plastic
Plastic
Arrest
Paris
Corrosion
Notch
Final inspection break
Copyright © 2005 by ASME
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0.00E+00
2.00E-07
4.00E-07
6.00E-07
8.00E-07
1.00E-06
2 3 4 5 6 7 8 9 10
Crack Length, a mm
T=370, P=90, D.O.=8
T=370, P=0, D .O.=8
HSF
LEFM
da/ dN m/ cy cle
Corr osion dominat ed growt h
Figure 8. Two cases differing only in applied mechanical
loading. The applied 90 MPa loading shows a higher
level of HSF growth and an area of corrosion dominated
growth. A picture of this crack is shown in Figure 7. Data
drawn from Figure 5, 7 second quench.
Environmental effects can be seen by the knee in the
fatigue growth curve first presented by Austen and Walker [9]
and also found in BS7910 [10]. Figure 9 presents some of our
data on this curve. Data with higher primary stresses
apparently have an increased environmental effect.
1.E-07
1.E-06
1.E-05
10 100
∆
∆∆
∆KI (MPa.m1/2)
da/dN (m/cycle)
Experimental (R=0)
Experimental (0.25<R<0.3)
Experimental (0.3<R<0.35)
Experimental (0.35<R<0.45)
Gabetta et al model (1990)
R=0.5
0.4
0.3
0.2
0.1
0.0
Figure 9. Smoothed experimental crack growth data
plotted against a Gabetta et al. [11model prediction,
allowing for the effects of environment and primary load.
Experimental data for DO=8 ppm plotted.
INITIATION OF THERMAL SHOCK CRACKING
A design guideline
The circumstances that can lead to thermal shock cracking
should clearly be avoided during design. In addition, existing
plant should be assessed for the possible presence of thermal
shock cracking. It is thus important to have a guideline for
determining the conditions necessary for the prevention of
thermal shock cracking. There is currently no such guideline in
the standards though it is basically fatigue cracking with a high
stress.
For the initiation of thermal shock cracking, the following
key factors are required:
a. a geometrical feature (stress concentration) and
b. a number of thermal shock cycles of sufficient magnitude.
Crack initiation can be complex, so the proposed
guidelines involve a three level approach as is seen in other
codes such as BS 7910. Each of the levels from 1 to 3 will be
intended to have a decreasing amount of conservatism. The
choice of level to be used in an analysis depends upon the input
data available and the conservatism required. The three levels
are summarised below.
LEVEL 1 is a simplified go/no-go type analysis requiring
the absolute minimum of input data. This approach is the most
conservative in its results as all stress concentrators are
assumed to behave as worst-case notches.
LEVEL 2 is a slightly enhanced procedure allowing for
the effects of some geometries and the number of thermal shock
cycles expected in a component lifetime. Conservatism at this
level is moderate.
LEVEL 3 is an advanced analysis procedure that requires
more detailed input data. Conservatism is at its lowest level in
this analysis. Improved analysis includes
• A detailed temperature profile across a specimen during a
shock to establish surface stress ranges and
• Cyclic stress strain relationships at geometrical
discontinuities to accurately allow for the effects of stress
concentrators.
• Improved fatigue initiation curves (S-N curves) for the
particular material and stress state.
A procedure for each level is suggested in the following
sections.
Level 1 Analysis
At this level, the maximum theoretical thermal stress
amplitude developed during a thermal shock (Sm) is compared
to the allowable stress amplitude based on the design
considerations of the ASME Boiler and Pressure Vessel Code.
The effects of geometric discontinuities (stress concentrators)
are allowed for using a stress concentration factor kf.
A Maximum Theoretical Thermal Stress (Smax)
The maximum thermal stress amplitude generated during
a thermal shock transient is calculated by equation (1):
(
)
( )
f
mk
TE
S⋅
−
∆
=
ν
α
12
max (1)
where: Smax is the maximum thermal stress
amplitude generated during a thermal shock.
E is the elastic modulus
is the coefficient of thermal expansion
Tm is the maximum possible magnitude of thermal
shock
I s Poisson’s ratio
kf is the stress concentration factor due to geometric
discontinuities. This has values in the range 1 to 5.
With E, and being material constants, the only remaining
unknowns are Tm and kf. Methods for determining these
parameters are given below. Note that in this work, the
assumption of temperature independence of E and is
acceptable as long as the product of the two does not vary by
Copyright © 2005 by ASME
7
more than 10% over the expected temperature range. This is
the case for carbon steel between 100C and 370C.
B Magnitude of Thermal Shock (Tm)
The maximum magnitude of the thermal shock (Tm) can
be determined from design data based on the worst possible
thermal shock condition (Tmax – Tmin), where Tmax and
Tmin correspond to the maximum and minimum possible
temperatures of the process fluid passing through the
component during operation.
Level 1 assumes instantaneous change of temperature at
the surface of the metal, that is, infinite heat transfer
coefficient. This is normally approximately correct for water
on clean steel where heat transfer coefficients of around 10
kW/m2K are seen. It can be shown that at this level there is
little change in thermal shock stresses as the heat transfer
coefficient increases up to infinite. In level 3 this conservatism
could be removed.
C Stress Concentration Effects (kf)
To include the effects of geometry and other stress
concentrators such as machining marks, suitable values for kf
need to be selected. In accordance with the fatigue design rules
of the ASME Boiler and Pressure Vessel Code, Section VIII,
Division 2 the maximum value for kf that is generated by a
severe notch (or corner) is 5.0. Similarly, for a flat surface free
of machining marks, kf can be taken as 1.0. It is recommended
that in a Level 1 analysis, a value of 5.0 be selected for all
discontinuities unless clear evidence of the lack of sharp
machining marks on the component can be made. Values of kf
less than 5.0 will need to be justified using methods similar to
that used in a Level 2 analysis.
D Crack Initiation
The generation of thermal shock cracks is indicated if the
half maximum theoretical stress amplitude (0.5 Smax) exceeds
the allowable design stress amplitude Sa. The allowable design
stress amplitude may be taken from the S-N design curves
contained in Section VIII, Division 2 of the ASME Boiler and
Pressure Vessel Code (Figure 5-110.1 for the case of carbon
and low alloy steels). This curve, as is well known, is
generated from entirely different data using mechanical tests.
For the purposes of this work the curve is adopted as a useful
and convenient input which we have found successful under the
conditions described. Generation of a new curve specifically
for thermal shock would be a large additional task which might
be contemplated at Level 3.
Example
Using this method the permitted Tm can be calculated
from transposed equation (1)
f
a
mk
S
E
T
α
ν
)1(2 −
=∆ (1)
With Sa from ASME VIII Div 2 Fig 5.110.1, typical
carbon steel values and a kf of 5, it is found that the permitted
Tm for 1000 cycles is 123 C and for 100,000 cycles is 30 C.
It is to be noted that there are no effects from steady load or
environmental conditions, and this is to be expected for fatigue
initiation.These results can be compared to the EPRI guidelines
for economiser headers [12] which limit Tm to 21C. The
guidelines presented here will normally permit much higher
levels of Tm. However these results can still be restrictive
since thermal shocks exceeding 100 or 200C are often seen in
power plant and this forces us to consider all the conservatisms
in the analysis by introducing a level 2 and 3 analysis.
1.1.1.1 Level 2 Analysis
Possible improvements include
• Magnitude of Thermal Shock (Tm) determined from
measured or calculated data at the component surface
during a thermal transient.
• Stress Concentration Effects (kf) determined theoretically
or experimentally.
Level 2 analysis should allow Tm values 2 to 3 times
higher than level 1 but sharp notches will always require kf =
5.0.
1.1.1.2 Level 3 Analysis
At this level the local strains developed in a component
during a thermal shock transient are determined using advanced
techniques. This level of analysis may require the complete
time-temperature history of the component during a thermal
shock transient as well as accurate representations of stress
staring effects and any geometry.
Appropriate actual S-N fatigue data curves can be used.
These curves remove the conservatism that is built into the S-N
design curves in the ASME code. Better still specific data
curves could be generated for the specific purpose.
This procedure may allow a further factor 2-3 increase in
the permitted Tm when compared to level 2 but again sharp
notches will always require kf = 5.0.
FITNESS FOR PURPOSE CASES
Assessment of Growth and Arrest of cracks in a thermal
shock field
Assessment of more recent experimental data suggests
that the proposals for the growth and arrest guidelines can be
simplified from those previously suggested [13. There are four
mechanisms to be considered in the growth of thermal shock
cracking.
1. Growth due to stress exceeding the tensile strength of the
material – Plastic growth or High strain fatigue (“HSF”)
region
2. Growth due to corrosion dominated growth mechanisms.
3. Paris growth region (affected by environment)
4. Final failure, if the crack does not slow down to the point
of arrest there is plastic collapse of the remaining ligament
in carbon steels.
Plastic growth region
Plastic growth of the crack occurs over about the first 1
mm or so in our experiments. This is largely the region in
which the stress exceeds the tensile strength of the material
(which is specified as 410 MPa in our specimens). In the
Copyright © 2005 by ASME
8
plastic growth zone the growth may be covered by a growth
law proposed, for example, by Skelton [14].
( )
LaB
dN
da −=
(3)
where a is the depth of crack (including starter
notch)
N is the number of thermal shock cycles.
B and L are constants to be determined experimentally.
In the current work a conservative form of this equation
has been determined. L is set at the depth of the starter notch
(3.5 mm) and the slope B has been set by statistical fitting to
the slope of all the data in the HSF zone. Equation 3b is the
solid line on Figure 5. The dashed line can be placed over all
the data so far obtained and in this case means setting L at 2.8
mm.
(
)
34 105.31088.4 −− ×−×= a
dN
da
m/cycle (4)
Corrosion dominated growth.
Once the stress starts to fall then growth may occur by
corrosion. This growth regime was not seen on many of our
specimens especially those with no applied mechanical loads.
Corrosion dominated crack growth is characterized by a growth
regime which is not a function of the stress intensity range, but
is dominated by the factors which control corrosion, in
particular the diffusion of species through the crack. In this
work all the relevant factors have not yet been investigated but
there is data for the following issues.
• Crack opening. If there is no primary stress corrosion
dominated growth does not appear to occur. This is
attributed to the fact that the crack will only be opened for
diffusion for short periods of the cycle.
• Crack depth. The deeper the crack the more time the
diffusion takes and corrosion virtually stops after the crack
grows to certain length (for the shorter cycle lengths studied
here).
• Cycling time. Though only partially examined so far in this
study this is clearly relevant in the corrosion dominated
region.
• Concentration of species. The experiments reported here
control pH and dissolved oxygen (DO) in the spray water.
Only introductory investigations of the effect of species
concentration by varying DO have been attempted. .
The proposed guideline requires two key definitions as
shown in Figure 9.
a. Definition of when corrosion dominated growth occurs.
This takes the form of when the thermal shock stress
intensity exceeds a certain level. This level will take
account of both applied mean stress and crack depth.
b. Defining a value of crack growth rate per cycle in the
corrosion dominated growth region.
a. The corrosion dominated growth occurs in region.
This region can be defined following Gabetta et al. [15]. This
type of growth occurs when the applied K is greater than a
critical value termed Kc. Kc is defined by them as
dt
da
R
R
EK y
ASCR
c+
−
⋅⋅=∆ 1
11
2
τσ
α
(5)
Where ASCR is a constant proposed by Gabetta
et al. which relates an “active surface creation rate” (ACSR) to
the rate of change of crack mouth opening distance during a
load cycle
R is Kmin/Kmax.
τ
is the cycle time, 7 seconds in the current case,
da/dt is a crack growth rate with time.
Currently the Monash experiments cannot distinguish all
these factors but there is enough data to provide the following
interpretation for the conditions.
R
R
Kc+
−
×=∆
1
1
1045.1 3
2
(6)
For the current case the criterion for corrosion dominated
growth could be further simplified to the form:
mean
K
K725
>∆ (7)
In this form it is readily seen that if Kmean is low or zero
then no corrosion dominated growth will occur. Using Figure 4
it can be seen that the condition will only be satisfied for a
range of crack growth in the region of a = 5 to 10 mm in these
experiments.
Crac k dept h, a
da/ dN
Growt h p er c ycle
Corr osion dom inat ed growt h
SIF ind ependant
growt h
End of c orr osion do mi nated
growt h, Growt h lim it ed by
diff usion
HSF, No corrosio n
gro wth
Paris growth
Corr osion dom inat ed growt h region
Fi
gure 9. Corrosion dominated growth region for thermal
shock cracking.
b. Crack growth rates. Once corrosion dominated
growth occurs, the growth rate is dominated by time of opening
and primary stress and its rate is not related to K [16]. The
growth rate in the corrosion dominated region for the current
experiments which have a cycle time of about 15 minutes is
given by
Copyright © 2005 by ASME
9
7
107 −
×=
dN
da
(8)
Paris law type growth
This can also be termed the LEFM growth region. As the
crack grows the stresses fall (See figure 1) and corrosion (if it
occurs) is no longer the dominant growth mechanism since the
diffusion lengths are too long. At this point and the growth
continues in a form covered by LEFM. In this region the
growth rate is found to fit the “Paris Law” equation [15].
( )
m
KC
dN
da ∆= (9)
where C and m are constants which are dependent on the
environment, the mean stress and the stress range.
This equation has been examined and found to fit the data
providing some care is taken in defining the stresses involved
in K. The values of C and m can be found from testing.
For the data collected in this study, the “freely corroding
marine” crack growth equations in BS 7910[10] tend to be
adequately conservative. Alternatively a curve fitting exercise
to the data produced the equation.
( )
89.5
16
1028.5 K
dN
da ∆×= −
(10)
da/ d
N
Crack
Growth per
cycle
Stress dominated gro wth, no cor rosion region
Plast ic zone
HSF
Paris L aw LEFM
growt h
Crossover point, str ess falls belo w
tensi le str ess
Figure 10. Stress dominated growth of thermal
shock cracking (without corrosion growth regime).
Final failure
Because this study is dealing with carbon steels in fairly
warm conditions below the creep level, final failure if a crack
keeps growing will be by normal plastic collapse normally
detected by leaking. This can happen when mechanical stresses
are high and so must be considered in an assessment guideline.
The limits which would seem to be appropriate are covered by
the failure assessment diagram (FAD) of BS 7910. There has
been no through thickness failure case in the Monash tests.
However it is known from industrial case studies that leaks are
observed in plant when high stresses are present.
CONCLUSIONS
This work reported here has demonstrated the factors which
cause and drive thermal shock cracks.
Design and operation: ASME Boiler and pressure vessel
code and EPRI guidelines
A key interest area is design guidelines for vessels subject
to thermal shock. The guidelines we are investigating will
provide the basis for design and operation of pressure vessels
with sharp geometrical features which may be damaged by
thermal shock. There is little information about this case in
ASME.
EPRI guidelines [12] for assessing the initiation of
thermal shock cracking are highly conservative. Using the
simplest level of our guidelines (level 1), typical carbon steels
with a sharp notch are permitted Tm for 1000 cycles is 123 °C
and for 100,000 cycles is 30 °C. If we use level 3 analysis and
remove notches from the component being considered than the
Tm permitted would rise probably by a factor of 5. These
results can be compared to the EPRI guidelines for economiser
headers which limit Tm to 21ºC. The guidelines developed
above will thus in most cases permit higher levels of Tm.
A key result of this work is that the features such the
sharp corners illustrated in Figure 2 should be avoided in
design. Such corners are not stress concentrators in the way
normally considered in design since they may not be highly
stressed in normal operation. Such sharp corners frequently
occur simply because that is the cheapest construction method.
However, it would be appropriate for the ASME Boiler and
Pressure Vessel Code to dissuade the use of the features where
thermal shocks could occur and corners (such as shown in
Figure 2a and 2b) are found on many figures in the code.
Growth guidelines for thermal shock cracking.
The discovery of thermal shock cracking often by opening
up a neglected part of a plant and observing it visually can be
very disconcerting for the operators. Often the immediate
response is to carry out expensive repairs or replacements.
These could in fact be pointless.
It has long been known that many thermal shock driven
cracks arrest at fairly shallow depths and only a few eventually
leak. The experimental work presented here starts to indicate
the growth mechanisms involved in thermal shock crack
growth and present the possibility of fitness for purpose
assessments of thermal shock cracks when they are found in-
service.
This work has identified that there are four mechanisms to
be considered in the growth of thermal shock cracking.
1. Growth due to stress exceeding the tensile strength of
the material – Plastic growth or high strain fatigue (“HSF”)
region
2. Growth due to corrosion dominated growth
mechanisms where the growth rate is constant and independent
of K. Corrosion occurs while the crack is open and not very
deep. The proposal is that this occurs when the applied range
of stress intensity factor, K exceeds some critical value KC.
Copyright © 2005 by ASME
10
The value of Kc depends on a number of factors not fully
investigated in this study such as time of crack opening and
oxygen content of water..
3. Once the stresses start falling and corrosion rate falls,
Paris growth occurs (affected by environment). Thermal shock
cracking is likely to arrest in this region if there are no high
primary applied stresses.
4. Final failure occurs only if primary stresses are high
enough. This occurs by plastic collapse of the remaining
ligament in carbon steels.
ACKNOWLEDGEMENT
This work has been conducted with the assistance of an
Australian Research Council grant with contributions from
HRL Technology Ltd, Optima Energy, Western Power,
Hazelwood Power, Loy Yang Power, Edison Mission, Pacific
Power and EPRI.
REFERENCES
1 C Alexander, J Frey and S. Shin, Evaluation of the failure in
the Texas Genco W.A. Pariish Unit #8 cold reheat line, 4th
Int Conf., Advances in Material Technology for Fossil
Power Plants, Hilton Head, 26-28 Oct 2004, EPRI. Palo
Alto.
2 American Society of Mechanical Engineers, Boiler and
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3 R.B. Dooley, and W.P. McNaughton, Boiler Tube Failures:
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4 H.W. Ng and C.K. Lee. ‘Remaining life of a vessel
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5 G. Yagawa and K.Ishihara, ‘Cleavage and ductile thermal
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6 B.B. Kerezsi, A. Kotousov and J.W.H. Price, ‘Experimental
apparatus for thermal shock fatigue investigations’,
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7 E. Vitale and M. Beghini, ‘Thermal shock fracture
experiments on large size plates of A533-B Steel’,
International Journal of Pressure Vessels and Piping. Vol.
46, pp. 289-338, 1991.
8 D.J. Marsh, ‘A thermal shock fatigue study of type 304 and
316 stainless steels’, Fatigue of Engineering Materials and
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9 I.M. Walker and F.F. Walker, Quantitative understanding of
the effects of mechanical and environmental behaviour on
corrosion fatigue crack growth behaviour, The influence of
environment on fatigue, I Mech Conf, London, 1977, pp 1-
10.
10 British standards, BS 7910:1999 – Guide on Methods for
Assessing the Acceptability of Flaws in Fusion Welded
Structures. London: BSI, 2000.
11 G. Gabetta ., C Rinaldi and D. Pozzi , A model for
environmentally assisted crack growth rate.
Environmentally Assisted Cracking: Science and
Engineering, ASTM STP 1049. Philadelphia: American
Society for Testing and Materials, 1990, p.266
12 G.G. Stevenson, Guidelines for the prevention of
economiser inlet header cracking in fossil boilers, GS-5949,
EPRI, Palo Alto, 1989.
13 J.W.H. Price and B.B. Kerezsi, Potential guidelines for
design and fitness for purpose for carbon steel components
subject to repeated thermal shock, Int J pressure Vessels
and Piping, February 2004.
14 R.P. Skelton, Growth of short cracks during high strain
fatigue and thermal cycling. Low Cycle Fatigue and Life
Prediction, ASTM STP 770. Philadelphia: American
Society for Testing and Materials, 1982, p.337-381.
15 Gabetta et al., ibid.
16 Kerezsi BB, Price JWH and Ibrahim R, A Two-Stage
Model for Predicting Crack Growth, Due to Repeated
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