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Transcript of Colector Header
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Life Assessment of High Temperature Headers
Greg J. Nakoneczny
Carl C. Schultz
Babcock & Wilcox
Barberton, Ohio, U.S.A.
Presented to:
American Power Conference
April N-20,1995
Chicago, Illinois, U.S.A.
BR-1586
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LIFE ASSESSMENT OF HIGH TEMPERATURE HEADERS
GREG J. NAKONECZNY
Babcock & Wilcox
Energy Services Division
20 S. Van Buren Avenue
Barberton, OH 44203
CARL C. SCHULTZ
Babcock & Wilcox
Research and Development Division
1662 Beeson Street
Alliance, OH 44601
ABSTRACT
High temperature superheater and reheater headers have
been a necessary focus of any boiler life extension project
done by the electric utilities. These headers operate at high
temperatures in excess of 900°F and are subject to thermal
stresses and pressure stresses that can lead to cracking and
failure. Babcock & Wilcox Company’s investigation of
these problems began in 1982 focusing on Pl 1 materials
(1 1/&r-1/2Mo). Early assessment was limited to dimen-
sional analysis methods which were aimed at quantifying
swell due to creep. Condition assessment and remaining
useful life analysis methods have evolved since these
initial studies. Experience coupled with improved inspec-
tion methods and analytical techniques has advanced the
life assessment of these high temperature headers. In the
discussion that follows we will provide an overview of
B&W’s approach to header life assessment including the
location and causes for header failures, inspection tech-
niques and analysis methods which are all directed at
determining the remaining useful life of these high tem-
perature headers.
INTRODUCTION
HISTORICAL PERSPECTIVE
Hlgh Temperature Headers
In 1982 Babcock & Wilcox (B&W) first began its investi-
gation of superheater outlet headers because of cracking
that was found in the headers of several of our utility
customers. The damaged headers were in both once-through
and drum type boilers. Initially the cracked headers were
comprised of only l’/,Cr-*/*MO alloy material (SA335
Pl 1) and had been in operation from 17 to 22 years. Creep
related failure in SA335 Pl 1 material could be explained in
part by changes in the ASME code. In 1968 the code
allowable stress for l’/,Cr-‘lzMo was reduced for high
temperature applications. The allowable stresses at 1000°F
and 1050°F were reduced 16% and 26%, respectively. As
a result headers, as well as piping, designed during the
1950s and early 1960s had the potential to be under de-
signed on the basis of the updated code. The likelihood of
creep degradation increased for older boilers that had been
in operation for an extended period. As a result of this
potential problem B&W initiated a review of all its boiler
contracts which were affected by the code change. Those
units which would no longer meet code for the revised
allowable stresses were identified. B&W established the
Plant Service Bulletin program in which all affected boiler
owners were notified of this potential for header creep
damage. The high temperature header program launched
the condition assessment and life extension programs which
have since become a standard part of a plant’s preventive/
predictive maintenance. As the focus was placed on high
temperature headers it became apparent that 1 1/4Cr-1/zMo
alloys were not the only materials subject to creep rela-
tively early in the materials life. Cracks in headers made of
2l/&r-lMo alloy material (SA335 P22) were also found. It
was clear that the mechanisms leading to the cracking of
these headers could not be explained by simple creep.
Investigations were begun to determine the root cause of
these header problems. Several programs were sponsored
by the Electric Power Research Institute (EPRI) to ascer-
tain causes of header damage, inspection methods and
analysis techniques which would help the electric utilities
in assessing and maintaining their boilers.
Steam Pipe Fallures
On June 9.1985 a major catastrophic failure of a hot reheat
pipe at an electric generating station in Nevada resulted in
the death of 6 workers and serious injuries to numerous
others. The failure occurred in the longitudinal seam weld
of the pipe and resulted in an 18 foot long tear along the
weld line. The pipe material was 1 /$r-l zMo alloy. The
pipe had been in service forjust 14 years pnor to the failure.
Creep was identified as a contributing cause of the weld
failure. Six months later, on January 30, 1986 a second
catastrophic pipe weld failure occurred at an electric utility
generating station in the midwest. Fortunately there were
no deaths, however, numerous injuries of personnel re-
sulted. The failure was a 30 foot long tear of the long seam
weld in a hot reheat steam pipe. The failed pipe was 2*/&r-
1Mo alloy material and had been in operation only 15
years. As with the previous pipe the operating steam
temperature was 1000°F and creep was identified as a
contributing cause of the failure. The occurrence of two
such serious failures in the span of six months coupled with
the fact that they had similar operating conditions but were
of different alloys further focused the attention of the utility
industry on the problems of creep related failures. This
gave further impetus to the growth of life assessment of
heavy wall components such as the headers and steam
piping systems.
HEADER DAMAGE
High temperature headers that most often experience sig-
nificant damage are the superheater outlet headers that
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operate at temperatures near 1000°F. High temperature
headers are generally constructed of 11/4Cr-1/2Mo (SA335
Pll) or 2’/,,Cr-1Mo (SA335 P22) steels. The typical oper-
ating temperatures are well within the creep regime for
both the Pl 1 and P22 materials. Creep is the phenomenon
in which the alloy experiences inelastic strain that is depen-
dent upon sustained stress at relatively high temperature.
Given sufficient time in operation, creep damage will
accumulate from exposure to the normal operating tem-
peratures and stresses seen during sustained (base load)
boiler operation; the high temperature headers have a finite
life due to creep. Cyclic operation, both on/off and load
cycling, can accelerate the accumulation of creep damage.
Boiler cycling introduces the additional damage mecha-
nisms of oxide notching and fatigue. These damage mecha-
nisms, operating together, can significantly reduce the
service life of a header.
Figure 1 illustrates locations where cracking is most likely
to occur in high temperature headers. Cracking has been
found to occur at virtually every weld as well as at the
ligament area between tube stub bore holes. The economic
impact of header damage is a function of both the damage
location and damage mechanism. From the boiler owner’s
perspective, failures which are a precursor to the header’s
end of life are of greatest importance. Early identification
and assessment of this damage is most critical to decisions
regarding the long term reliability and cost to maintain
boiler steam generation. Header damage can generally be
classified as repairable or non-repairable. The majority of
header damage has been found to be repairable such that
header replacement is not required.
Reinforced
Section
We-Ids
Y
Drain
Figure 1 Header locations susceptible to cracking.
Repairable Header Damage
Repairable damage consists of cracks or other damage that
can be weld-repaired. This can include cracking of welds at
support lugs, support and torque plates, branch connections
such as drain line and vent line welds, the outlet nozzle
welds and header girth welds, radiograph plugs, master
handhole cap welds and, depending upon root cause of the
damage, some tube stub-to-header welds. The most fre-
quent incidence of cracking which leads to steam leaks is
in tube stub-to-header welds. Although tube stub-to-header
weld cracks are readily detected and repaired, they nor-
mally result in costly forced outages. Weld cracking at
thermowells, RT plugs, handhole fittings, etc., is often
quite similar to the cracking at tube stub-to-header welds.
Damage at all o f these locations can be caused by creep of
the header along with the differences in the creep strain
rates between the header and connection or fitting. For
example in the case of radiograph plugs which are openings
provided in the header to allow insertion of a radiographic
source for testing of adjacent welds, one type of plug uses
a threaded cap which is seal welded on the OD of the
header. The radiograph plug threads are intended to form
the pressure boundary of the plug. On older superheater
headers subject to creep, the header can swell due to creep
strain, i.e. plastically deforms. The radiograph plug de-
forms much less, or not at all, resulting in stresses and
cracking in the seal welds as well as disengaging of the
radiograph plug threads.
Local differences in yield strength and creep strength
within the different constituents of the various weldments
can produce metallurgical notch effects quite similar to
those of geometric notches. When acting together, global
differential creep rates along with the notch effects of strain
concentration can be detrimental at areas of low ductility
that may exist within the weldment. The cracking or failure
of welds at the various branch connections caused by
header creep is important from the standpoint that it indi-
cates creep strain in the material which might lead to more
serious problems in areas not yet seen. It emphasizes the
need that these high temperature headers be given a com-
prehensive inspection and remaining life evaluation.
Header cracking at outlet nozzle-to-header welds, outlet
nozzle-to-pipe welds and support plate welds can indicate
that additional driving forces or stresses beyond the pres-
sure stress are occurring. In the case of the outlet nozzle, it
is common in most power plants to find problems with the
piping system. Piping loads shift and redistribute during
the plant’s operating life. Failure of piping supports is not
uncommon. All of these factors lead to excessive loads
being imposed on the outlet nozzle and support system of
the superheater and reheater outlet headers. These exces-
sive forces from the piping system produce stresses that
lead to crack initiation on the OD of the header; normally
these cracks initiate at major strength welds. The outlet
nozzle is most susceptible. The higher stresses can also
produce creep in the welds before creep is found at other
locations in the header. For units that are frequently on/off
cycled, the high stress amplitudes can lead to cracking as a
result of fatigue. Damage associated with these higher
imposed stresses is normally on the OD surfaces such that
the damage can be removed and repaired. In such instances,
assessment and correction of piping system support prob-
lems is important if the damage is to be prevented from
returning.
In general, cycling of a boiler, particularly on/off cycling,
introduces cyclical stress and strain that can cause damage
as a result of fatigue. In the special case of the header drain
lines cycling can also lead to thermal shock in the header
material. Most boilers designed in the 1960s and 1970s
were expected to be operated as non-cycling base loaded
units. Although allowances were made for expansion
stresses the designers allowed for relatively low numbers
of cycles. As the electric utilities were forced to begin
cycling many of their plants and boilers due to the changing
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nature of power demand, problems in the piping systems
and boilers have resulted. In high temperature headers
cycling leads to fatigue crack initiation. In addition to the
outlet nozzle damage noted above, fatigue can cause crack-
ing at the support welds, branch connections, girth welds
and tube leg welds. During cold start up of the boiler the
superheater headers are subject to humping as a result of
topto-bottom temperature differences. This humping im-
poses stresses on the various attachments and supports.
Generally the larger boilers have the largest and longest
headers. Thermal expansion is greater and humping is
more likely, and of greater amplitude for these larger
headers. Additionally, for large boilers, the thermal expan-
sion of the superheater outlet headers will place bending
stresses on the outlet tube legs. For frequent on/off cycling
the cyclical bending stresses have caused cracking in the
outlet leg tube stub-to-header welds. Cracks associated
with cycling will occur nearest the header ends where
expansion and bending stresses are greatest. For drain line
connections, on/off cycling can lead to severe localized
damage to the header as a result of thermal shock. In plants
where more than one boiler or header are tied to a common
blowdown tank it has been found that condensate can
sometimes back up through drain lines and enter a hot
header during start up. The resulting thermal shock can
cause fatigue damage to the header immediately adjacent to
the drain connection.
Many of the indications or cracks associated with creep or
fatigue (including thermal shock) as described above can
be repaired. In some cases simply blend grinding will
remove an indication without the need of weld repair. In the
case of drain line thermal shock damage, a header end
section may have to be replaced, however, this repair is
relatively small when compared to the logistics and cost of
complete header replacement. It is important to note that
the damage mechanisms described above have been classed
as repairable in the context of whether repair of the damage
is a possible option. In all cases inspection and life assess-
ment of the header must consider all damage together.
Although local repairs are possible, the presence of damage
in many areas coupled with the presence of creep and the
owner’s experience with forced
outages
may dictate that
header replacement is the best course of action. Retirement
of the header can be driven by economic as well as material
considerations.
From a material standpoint, the problem that most otten
results in the replacement of the high temperature headers
is cracking of the header in the bore hole and bore hole
ligament area. One exception is the possibility of a header
made of seam welded material. For seam welded pipe used
in headers the concern is for creep and catastrophic failure
of the long seam as wa s experienced on hot reheat piping
systems described above. Although at least one header was
replaced as a result of a long seam failure, the majority of
boilers use seamless pipe for the headers.
Non-Repairable Header Damage
In recent years, the utility industry has recognized ligament
(or bore hole) cracking as a significant, life-limiting prob-
Babcock & Wilcox
lem in headers subjected to elevated temperature service.
Ligament cracking is most frequently found in secondary
(or finishing) superheater outlet headers. Severe ligament
cracking, requiring header re
I:
lacement, has occurred in
both 1 /,Cr-l/zMo (Pl 1) and 2 /,Cr-1Mo (P22) headers.
Ligament cracking generally initiates as numerous longitu-
dinal cracks in tube bore holes. Figure 2 illustrates these
longitudinal cracks in the interior of a bore hole. The
ligament cracking of Figure 2 is in a very advanced stage.
These cracks extend (either initially or eventually) to the
inside surface of the header, appearing as a “starburst”
pattern when viewed from the inside of the header; see
Figure 3. Some of these cracks continue to grow along the
inside surface of the header, eventually linking up with
similar cracks emanating from adjacent tube bore holes, as
seen in Figure 4. These cracks continue to propagate,
growing simultaneously from the header ID toward the OD
Figure 2 Advanced ligament cracking.
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Figure 3 Large ligament cracks on header ID.
and between adjacent bore holes, as shown schematically
in Figure 5. Review of Figure 2 reveals that at least one of
the cracks has advanced through almost the entire ligament.
The thermal cycling that results from on/off operation
accelerates both the initiation and propagation of ligament
cracks. Two competing mechanisms are believed to be
responsible for the initiation of the cracks. One of those
mechanisms is referred to as “oxide notching.” High tem-
perature steam in contact with Pl 1 and P22 material pro-
duces oxidation in the low alloy header materials which
forms a brittle oxide scale layer which is mainly magnetite
(Fe,O,). This oxidation occurs during periods of sustained
operation at elevated temperature. The oxide layer grows in
thickness over time. Since the oxide layer is relatively
brittleitisnormalfortheoxidetobegintocrackandorspall
of f in flakes. Normally the major concern associated with
Figure 4 Linking of cracks between adjacent bore holes.
4
Figure 5 Progression of ligament cracking.
the exfoliation of oxide is the solid particle erosion it can
cause on valves and turbine components. However, crack-
ing of the oxide layer due to the temperature and strain
cycles that occur during a shut down and subsequent start
up, exposes the header base metal to oxidizing steam, re-
establishing the initial high rate of oxidation. As this
process continues over time it preferentially oxidizes the
header along the crack in the oxide, eventually forming a
notch for crack initiation.
The other mechanism that contributes in the initiation of
ligament cracking is a combination of localized creep
damage and thermal fatigue damage. These damages are
the result of the significant thermal stresses that are typi-
cally incurred during on/off operation and or during load
cycling. The intended elevated temperature service for
superheater headers results in a relatively low allowable
design stress as dictated by the ASME code in order to
avoid excessive creep deformation. For superheater outlet
headers intended for high temperature service at high
pressure the allowable stresses result in relatively thick
walls. The temperature gradients, and thus thermal stresses,
that result from the thermal cycling during on/off and load
cycling operation, become more severe as the design wall
thickness increases. The area of the header bore hole
penetrations, which act as geometric discontinuities, is also
where the highest local stresses occur from the internal
pressure. Through finite element analyses conducted by
B&W it was determined that bore hole penetrations have a
significant effect on the thermal stresses that occur during
rapid changes in the steam temperature. The effect of
thermal stress at the bore hole locations is two-fold. First,
as with the pressure stresses, the bore hole acts as a
geometric discontinuity which increases the adverse ef-
fects of the thermal stresses. Second, the bore hole open-
ings provide additional heat transfer surface through the
header wall at the outlet legs which can increase the effect
of outlet leg temperature differential. This second effect is
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particularly important because of thermal upsets, or tem-
perature variations that can occur across the width of the
boiler and superheater. Tube temperatures may vary result-
ing in a mismatch between the temperature of the steam
within the bore holes and that within the main cavity of the
header at the same position. Since tube temperatures re-
spond more quickly than the main header to load changes
and firing fluctuations, the tube steam temperature mis-
match is more likely in transient operating conditions, such
as oad changes. As proven through B&W’s finite element
modeling, the localized heating/cooling that results from
this temperature mismatch can be a source of significant
thermal stress. Lastly, the ligament metal temperatures
may locally exceed the design outlet steam temperature for
extended periods of operation. The higher ligament tem-
perature can accelerate creep damage, oxide growth and
crack growth rates.
In general, quantifying the remaining life of high tempera-
ture headers focuses on analysis and prediction of header
crack growth which has been developed using time depen-
dent fracture mechanics and considers the effects of creep.
Programs exist today, such as the PC computer code
BLESS developed through an EPRI sponsored project and
discussed later in this paper, which allow for the prediction
of crack initiation as well as crack growth. However,
detailed operating data for older boilers, which is critical to
the prediction of crack initiation, is normally not available
in sufficient detail. As a consequence most quantified
header life assessments are based upon the predictions of
growth for a pre-existing crack. With the awareness of life
assessment and predictive maintenance. boilers built today
are more likely to incorporate systems that allow for
monitoring of operating conditions so that prediction of
crack initiation and on line assessment of operational
upsets is possible.
FACTORS AFFECTING LIGAMENT DAMAGE
Design Parameters
Several years ago, as part of an EPRI program, B&W
reviewed inspection reports of 376 headers that had been
inspected, by B&W. for ligament cracking. The incidence
of cracking, for different types of high temperature head-
ers, is reported in Table 1. The incidence is seen to be far
greater in secondary superheater outlet headers: 28% ver-
Tabk 1
Header Inspection Results - October 1988 Header Types
Numbsr wl
Numbsr Tube Bors
Inspectsd Cracks K
secondely SH Outlet Headers 157
44
20%
1’4 Cr Material 73 26 36%
2’4 Cr Material 76
17 22%
Operating Temperature r 105OF 14 6 43%
ReheatedSH Outlet Headers 116 2
2%
All Other Headers 101 4 4%
Tabb 2
secondety Supetheater Outlet Header hspection Resuits
CktoberW66-AgeandMaterials
I”, cr-‘I, MO 2’1, Cr-1 MO
Material (Pll) Materlal (P22)
Headsr
Numbsr K With Numbsr 96 With
gervbx Ysara Inapsctsd Cracking Inspsctsd Cracking
2OYearsorLess 13 46% 41 17%
21 lc 25 Years 29 26% 15 40%
26tO3OYMNS 23 52% 10 20%
Morethen3oYears z
8
72 36% 75 22%
Averags Age of hspected Pl 1 Headers
W~lh Damage = 24 Years
Wlthout Damage = 24 Years
Avmgs Age of lnspecbd P22 Headers
with Damage = 22 Years
without Damage = 20 Years
sus only 3% in all other high temperature headers in-
spected. Secondary superheater outlet headers operate at
much higher pressure than reheat outlet headers. As a result
of the higher operating pressure, the secondary outlet
headers are considerably thicker than reheat outlet headers
operating at the same temperature. The greater wall thick-
ness results in more damaging thermal stresses being gen-
erated in the secondary outlet headers. The incidence rate
is reported relative to header age and material type in Table
2. Although the incidence rate is greater in the Pl 1 material,
the rate is still significant in the P22 material. The age of the
header, alone, does not appear to be a determining factor.
For example, the average age of Pl 1 headers found to have
ligament cracking, as well as those in which damage was
not found, wa s 24 years. Similarly, the average age of the
P22 headers found to have ligament cracking was 22 years
while the average age of those in which damage was not
found was 20 years. The incidence of ligament cracking did
show a strong dependence on the bore hole penetration
pattern. Six headers with mixed radial/nonradial bore holes
were inspected and all were found to have ligament cracks.
Only 28% of the 72 headers with radial bore holes that were
inspected were found to have ligament cracks. Similarly,
only 3 1% of the 45 headers with nonradial bore holes were
found to have ligament cracks. Figure 6 illustrates radial,
nonradial and mixed bore hole penetration patterns. It is
noteworthy that 6 of 14 (42%) headers operating at tem-
peratures over 1050°F were found to have experienced
ligament cracking, illustrating the significance of tempera-
ture and its effect on creep.
Radial Nonradial RadiaVNonradial
Figure 6 Header bore hole penetration patterns.
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Bohr Oporatlon
There are three factors, relative to boiler operation, that
influence ligament damage in high temperature headers:
combustion, steam flow, and boiler load. Most boiler
manufacturers design the boiler with burners arranged in
the front and/or rear walls depending upon the size and
capacity of the unit. Heat distribution within the boiler is
not uniform: burner inputs can vary, air distribution is not
uniform; and slagging and fouling can occur. Even if
burners are optimized for equal firing, the temperatures of
the combustion gases exiting the furnace are lower near
the side walls than at the middle of the boiler. This occurs
since the perimeter of the furnace is constructed of water-
cooled tubes and there is greater heat transfer from the
combustion gases near those cooler wall tubes. Air distri-
bution can also vary from side to side, across the unit,
causing unbalanced flow of combustion gases exiting the
furnace. On coal-fired and some oil-fired boilers, slagging
and fouling occur causing biasing of combustion gas flow
and uneven heat absorption in the furnace and convection
passes. The net effect from these combustion parameters is
to cause variations in heat input to the superheater and
reheater.
Typical Header
Tu& Leg
Tube Leg
I
Temperature
107OF
I
(577C)
Left End
Tube Leg Location
Rght End
Combined with the combustion parameters, the super-
heater and reheater experience differences in the steam
flow in individual tubes within the bank. A tube carrying
greater steam flow wil l experience less of a steam tempera-
ture increase than a tube with reduced f low, assuming equal
heat is absorbed by both tubes. Spatial variations in gas
temperature and tube-to-tube variations in steam flow can
combine to result in significant variations in tube outlet leg
temperatures entering the outlet headers. Since the overall
bulk header temperature is close to the controlled outlet
steam temperature, large temperature differences can oc-
cur at tube bore locations. As shown in Figure 7, a 70°F
temperature difference between an individual outlet leg
and the bulk steam temperature is not uncommon, even
under normal base load conditions. It should be noted that
on tangentially comer-fired boiler designs the combustion
gases flow in a cyclonic path within the furnace. As a result
more heat absorption is expected to occur toward the
outside of the superheater such that the temperature distri-
bution will vary from that shown in Figure 7.
Figure 7 Steam temperature variation in a header.
As a consequence of the through-wall temperature differ-
ences and the temperature differences between individual
outlet legs and the bulk header steam temperature, the
header experiences localized stresses much greater than the
stress associated with steam pressure. Further, during in-
creasing and decreasing load changes, the reversal of the
through-wall temperature differences and the reversal of
individual tube leg steam temperatures relative to the
header causes reversal of corresponding stresses at the bore
holpnetrations. These increased and reversing stresses
Boiler start-ups and shut-downs result in significant tran-
sient thermal stresses as a result of the steam temperature
changes in the thick-walled headers. Changes in boiler load
have the effect of further increasing the temperature differ-
ence between the individual tube legs and the bulk header
temperature. As boiler load increases, the firing rate must
increase to maintain pressure. During this transient, the
boiler is temporarily over-fired to compensate for the
combined effect of increasing steam flow aud decreasing
pressure. As a result there is a temporary upset in steam
temperature from individual tube outlet legs relative to the
bulk header temperature. During load decreases the oppo-
site occurs; firing rate decreases slightly faster than steam
flow in the superheater with a resulting decrease in tube
outlet temperatures relative to the header bulk temperature
(Figure 8).
Figure 8 Superheater tube leg temperatures vary with load.
6
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further contribute to the initiation of cracks in the header
along the bore hole penetrations which eventually lead to
premature header end of life. The cracks are oriented along
the axis of the bore hole and propagate along the bore and
across ligaments between adjacent holes, as was shown in
Figures 2 - 4. If not detected in its early stages, these cracks
will eventually propagate through the tube stub-to-header
welds resulting in steam leaks. Bore hole cracking com-
bined with general creep of the header can lead to more
catastrophic stub weld failure as seen in Figure 9.
Figure 9 Superheater header stub failures.
HEADER ASSESSMENT
Assessment of the high temperature headers most often
focuses on nondestructive examination @IDE) followed by
evaluation of the NDE results. As with most condition
assessment programs the project follows several phases
that are geared to the plant outage when examinations and
testing can be performed. B&W fol lows a three phase
program.
Phase I - Pre-Outage Planning
l
Review operation and maintenance history
l Review design characteristics
l
Perform preliminary analysis if required
l
Establish outage inspection/test plan
Phase II - Outage
l
Implement inspection/test plan
l
Perform root cause analysis as needed to ensure all
necessary data is obtained during the outage. Install
instrumentation to support on-line testing if required by
the phase I plan or for root cause analysis.
Phase III - Post Outage Testing and Engineering Analysis
l
Perform final remaining life analysis
l
Conduct operational testing and analysis as required
l
Develop recommendations for follow up - repair, replace,
or reinspect based upon the analysis
For the high temperature headers key information to con-
sider in the phase I review includes the material and design
type. Is it 1 /&r or 21/qCr alloy? Is the header made of seam
welded pipe? Does it have radial, nonradial or a combina-
tion stub geometry? Phase I considerations for operating
characteristics include: temperature, is it designed to oper-
ate at lOOO”F, 1025”F, 1050°F etc. and how well is it
controlled? Are tube outlet leg thermocouples installed and
operable and is data available to be reviewed? Is the boiler
cycled? Ifit is cycled, then how and how often, i.e. is it load
cycled, on/off cycled, and how many times annually and
during its life? In phase I, consideration is given to the
maintenance history. Has the header experienced any sup-
port failures or cracks? Have steam leaks been experi-
enced? If so, where and how often? For example, if leaks
have been a recurring problem at tube stub-to-header welds
then it would be important to know where the leaks oc-
curred and whether the unit was cycled often. In general the
phase I review allows the planners to determine how
problematic the header has been historically, as well ashow
likely it is to be at risk for creep, creep-fatigue and fatigue
related header problems in the future.
For most life assessment projects phase II is limited to
performing the nondestructive testing as well as visual
inspections. In some instances an owner is changing opera-
tion. They may be changing from base load operation to
cycling operation, or, they are planning a major upgrade
such that a more comprehensive engineering study is needed.
In such instances it may be necessary to instrument the unit
for operational testing following the outage. Occasionally,
in addition to NDE, it is necessary to remove samples from
the header to perform material testing and laboratory analysis.
Nondestructive Examinations
Planning for the nondestructive testing is directed to select-
ing the best locations to perform the various types of NDE.
It is important to ensure the locations selected will include
the welds most likely to have experienced damage. The
most common NDE methods used include: magnetic par-
ticle testing (MT), liquid dye penetrant testing (PT), di-
mensional measurement and analysis, oxide measurement
(B&W uses the company’s NOTISe test), metallographic
replication, bore hole ligament exam, internal video probe
or fiber optic probe exam, in-situ alloy analyzer testing,
ultrasonic testing and radiography. In special applications
eddy current testing may also be used. In the majority of
header inspections B&W recommends, in addition to a
thorough visual examination, MT and/or PT for surface
examination of welds, bore hole ligament examination
following oxide removal, metallographic replication, in-
ternal inspections (normally with video probe), dimen-
sional analysis and ultrasonic testing for volumetric exami-
nation of the major welds. Use of the remaining methods is
normally dictated by special considerations determined
during phase I review of the unit or in follow up to problems
identified during the phase II inspections. Guidelines for
determining where to perform the NDE are presented
below. A comprehensive guideline for NDE of headers wa s
preparedbyB&WforEPRIproject 2253-10,AnIntegrated
Approach to Life Assessment of Boiler Pressure Parts.
Refer to volume 6, Guidelines for NDE of Heavy Section
Components, for more information.
Visual Examinations - internal and external should be
performed on all high temperature headers. The goal of the
external visual examination is to identify obvious damage
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and to help target other NDE to areas of suspected prob-
lems. In particular, the visual inspection should include the
support system and welds of the header to identify cracks,
distortion, or in the case of support rods, loose rods which
no longer carry load. Weld inspection is intended to iden-
tify macroscopic cracking associated with creep or fatigue.
Overheating of the header or of the outlet legs can some-
times be seen by discoloration of the metal or by the
presence of excessive scale. Internal inspection of the
header focuses on finding unusual oxide exfoliation. If
ligament cracking is advanced and the cracks are large then
internal inspection aids in determining the extent of cracking.
Nondestructive examination methods are a cost effective
means of identifying cracks and degradation on the sur-
faces of the headers. Critical to the success of NDE is
proper preparation of weld surfaces where the NDE is
planned. High temperature headers with their tenacious
oxide layer and irregular geometries can be difficult on
which to perform some NDE methods. Surface preparation
to assure a bare metal finish is particularly important for
ultrasonic testing and surface techniques such as MT and PT.
Magnetic Particle Testing (MT) is an effective technique
for evaluation of surface indications associated with welds
where the geometry of the weld allows proper placement of
the magnetic yokes. Effective MT requires that the mag-
netic field be applied at two orthogonal axes such that
accessibility of the weld areas is a factor. In general, MT is
performed on all of the major welds, fittings, and most
branch connection welds on the header including: outlet
nozzle welds, girth (circumferential) welds, long seam
welds if present, support welds, hand hold cap welds, and
welds in the drain and vent lines. In most header examina-
tions, the outlet tube stub welds on the header are too
closely spaced to allow effective MT. For stub welds,
liquid penetrant testing is normally preferred.
Wet Fluorescent Magnetic Particle Testing (WFMT) is
more sensitive than conventional dry MT. WFMT is, there-
fore, preferred for magnetic particle testing of girth welds
and long seam welds. It can also be used in lieu of dry MT
on the other welds. WFMT may be required in some
locations where the orientation does not allow use of a dry
medium, such as overhead test locations.
Liquid Dye Penetrant Testing (PT) is used for detection of
flaws or cracks which are open to the surface of the
component. Unlike MT, dye penetrant testing can be per-
formed in locations with limited access, provided the
component surface can be properly prepared. For surface
NDE of high temperature headers, PT is generally used
when MT or WFMT are not possible. ET is used on welds
where limited access prevents placement of the MT yokes,
the most common being tube stub-to-header welds. PT is
also commonly used during intermediate steps in a repair.
When grinding a header to “chase out” a crack or defect, ET
is used to verify that all the indication has been removed.
Surface preparation for PT is particularly important in that
surface preparation methods must not have the effect of
closing potential cracks. For example shot blasting should
not be used on headers as it can mask damage and make ET
ineffective. Because PTrequires multiple steps - apply dye,
allow period for capillary action of the dye, followed by
removal of excess dye and applying of a developer - it
requires more time than other NDE methods. As a conse-
quence it is common to target a partial sampling of the
outlet stubs for PT rather than testing 100 percent of the
welds. NDE of stubs is then expanded only if problems
warrant further testing.
Header Bore Holes Examination. The most important
inspection for early detection of bore hole and ligament
cracking is direct examination of the header bore hole.
B&W strongly recommends that high temperature oxide
scale be removed from the ID of the bore hole before bore
hole examination. Without oxide removal, cracks would
have to advance to a arger size for them to be found reliably
with internal inspection (Figure 10). The larger the cracks
when detected, the less the remaining life of the header; the
owner will have less time to make decisions regarding the
header and boiler. B&W developed the Hone & Glow@
technique to effectively remove oxide scale and allow
examination of the header base material. Hone & Glow@
has been in use since early 1985. Hone & Glow@ is done by
removing the oxide scale layer from the bore hole ID and
then performing dye penetrant testing (Figure 11). This
maximizes the effectiveness of bore hole inspection so that
cracking is detected early in the degradation of the header.
For increased sensitivity, fluorescent dye penetrant may be
used. It is important that care be taken when removing the
oxide scale such that any damage in the bore is not removed
in the cleaning process. Early bore hole cracking can
appear as broad or wide shallow linear indications. This
characteristic may be the effect of oxide notching as a
mechanism of crack initiation. Because of their wide shal-
low features these indications can be removed by excessive
bore hole cleaning when removing the oxide scale. Bore
hole inspection requires that outlet tubes be cut to provide
access into the header and bore hole. Normally, the tube
stub is cut a couple of inches from the OD of the header such
that rewelding of the tube following inspection does not
impact the header itself.
Location selection for bore hole examination is very impor-
tant. As emphasized in the earlier discussions of creep-
Figure 10
Header bore holes with oxide removal to reveal
damage.
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High temperature oxide layer
Superheater
outlet tube stub
Figure 11 Header bore hold exam with fiber optic.
fatigue, bore hole cracking is related to tube leg outlet
temperatures, i.e. the greater the temperature difference
between the outlet leg and the header the greater the
thermal stresses and the greater the impact on creep in the
header. Consequently, the bore holes that have the highest
temperatures are targeted for testing. Various methods are
used in selecting locations. If thermocouple data is avail-
able then the outlet leg temperatures measured during
operation can be used to guide the selection. If thermo-
couples data is not readily available then B&W recom-
mends tubes be selected on the basis of oxide thickness
data. NDE methods such as the B&W NOTIS@ test allow
accurate measurement of the ID oxide thickness from the
OD of the tube stub. As mentioned earlier, oxide scale
grows at a rate that is dependent upon operating tempera-
ture. Oxide thickness measurements are taken along a row
of header stubs to determine those with greatest past
operating temperature, i.e. those with heavier (thicker)
oxides indicate the hottest locations along the header. If
neither thermocouple data nor oxide measurements are an
option then the locations are selected on the basis of
experience. For front wall and or rear wall fired boilers,
tubes will normally be chosen at quarter points and near the
mid-point, of f the header reinforced area if present. For
tangentially fire units, locations will include tubes nearer to
the header end where higher steam temperature is expected
to occur.
Metallographic Replication is the NDE method used for
the evaluation of grain structure in both high temperature
headers and piping. Specifically, replication is the NDE
method relied upon to provide microscopic material infor-
mation needed for assessment of creep. A replica is essen-
tially a “fingerprint” of the surface under examination and
can be used to detect cracking, creep cavitation, porosity,
inclusions, and other similar defects that are undetectable
by other nondestructive techniques. Replication can thus
provide an early warning of an active failure mechanism.
Replication is a technique that complements other NDE
methods when evaluating the high temperature headers.
Because replica information is obtained from discrete
locations, other NDE is needed to accurately assess the
entire header.
Replica location and replica quality are important consid-
erations. Replication should be directed to the locations
where fatigue and creep are most likely to occur. Locations
subject to temperature excursions and/or higher stresses
should therefore by chosen. Site-specific temperature ex-
cursions are associated with the highest outlet leg tempera-
tures. At least one replica location is selected on a tube
stub-to-header weld where temperatures are expected to be
greatest. The options for determining this location are the
same as described previously for selecting the bore hole
inspection site. Damage found in headers is associated with
the weld locations. Selecting the best weld locations on the
basis of higher stresses is done primarily from experience
and a knowledge of typical problem areas. Locations are
also chosen on the basis of other NDE where damage may
have been found indicating a problem or high stress. The
outlet nozzle with its susceptibility to high stresses from the
piping loads is always included; replicas are taken on the
outlet nozzle at various locations which include both the
header-to-nozzle weld and the nozzle-to-outlet pipe weld.
Other welds typically included are girth welds and, if
present, long seam welds. In general, the arrangement of
the header, its interconnecting piping and support arrange-
ment will dictate where replication is done. The replica
tape itself should include the weld metal, heat affected zone
(HAZ), weld fusion line, and the transition between the
HAZ and the base metal. Depending upon the type of
replica made this may require multiple replicas at each
location selected. Replication is sensitive to airborne con-
taminants which can scratch prepared surfaces. The envi-
ronment in which replication is to be performed must be as
dust free as practical to prevent this contamination. Exces-
sive moisture and humidity can also lead to poor replication
and must be considered when planning the NDE work.
Dimensional Analysis. As noted previously, dimensional
analysis of high temperature components is done in an
attempt to assess creep damage by correlating growth in
component diameter to plastic creep deformation. Dimen-
sional analysis along with replication have been the pri-
mary methods of evaluating components for creep. Dimen-
sional analysis has been relegated to a secondary tool for
high temperature headers, primarily because creep-fatigue
at ligaments will not necessarily correlate to a swelling in
the overall header diameter. Today it is felt that bore hole
examinations are more reliable in header assessment. Di-
mensional analysis has greater applicability to piping as-
sessment where ligament cracking is not a factor.
Regardless of the application, for dimensional analysis to
have any value, data accuracy and repeatability are critical.
The actual measurements must be documented in sufficient
detail to exactly locate the points during subsequent
reinspections. The following criteria should be part of data
gathering for measurements on headers.
l
Locations should be permanently identified by punch
marks or by exact position reference measurements from
components on the header, i.e., distance from support
plates or nozzle connections, stub locations, etc. Data
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must be complete for both the axial and circumferential
locations.
l
Circumferential as well as diametral data should be
recorded at sites selected for dimensional checks. This
data can help evaluate the amount of swell and provide
back-up to diameter measurements.
l
At each axial location along the header or pipe, diameter
measurement sites should be cleaned prior to measure-
ment. Surface preparation should be consistent and should
remove oxide scales. Subsequent reinspection should
also ensure data is taken from base metal.
l
Diameter measurements should only be made using
appropriate size micrometers. Outside calipers and tape
measurements have been found to give inconsistent
results.
Multiple locations are selected for swell measurements. In
general, measurements are taken in at least three axial
locations along the header and one location on the outlet
nozzle(s).
A second technique that has been used for dimensional
analysis in headers is bore hole ovality measurements.
Header analysis has shown that creep deformation will
occur more rapidly in the circumferential direction versus
the axial direction in the header. Since the header bore
holes are machined during manufacturing, it was felt that
header swelling due to creep would result in a measurable
ovality of the header bore holes. This technique might have
greater sensitivity to the localized creep associated with
headers. The disadvantage is in the fact that this can only be
done at bore hole inspection locations such that applicabil-
ity is limited to the scope of the bore hole inspections for the
specific header assessment. Not enough data has been
obtained to validate this method. Dimensional analysis is
considered secondary and complementary to other NDE
methods and should not be used as an exclusive condition
assessment technique.
None of the NDE methods discussed above provide for
volumetric examination of the weld. When major welds are
to be examined such as girth welds and especially if the
header has a long seam weld to be evaluated, then volumet-
ric inspection methods must be included. For girth welds
and long seam welds ultrasonic shear wave testing is
pelfOIllld.
Ultrasonic Testing (UT) has been shown to be the most
sensitive technology for the nondestructive volumetric
examination of welds in piping. The EPRI sponsored work
done to investigate techniques for evaluation of seam
welded steam piping established UT as the most reliable
NDE method for detection of small flaws in welds, regard-
less of orientation. EPRI’s CS-4774 Guideline for the
Evaluation of Seam-Welded Steam Pipes has evolved into
the standard for inspection of long seam welds in hot reheat
piping. EPRI’s research wa s targeted toward the relatively
thinner wall hot reheat piping where catastrophic failures
had occurred. These guidelines are also applicable to seam
welded hot reheat headers and should be referred to for long
seam weld inspection in headers. For girth welds found in
higher pressure piping and headers the EPRI criteria for
seam welds is too sensitive due to the thicker materials
involved. The ASME Boiler and Pressure Vessel Code
Section V, Nondestructive Examination, is often cited as
the criteria for ultrasonic examination of girth welds. The
key requirements defined by the code in article 5 include
the following:
l Calibration standard will have a notch depth that is 10%
of thickness. (This is the major difference between ASME
and the EPRI seam weld standard. The EPRI method
requires a calibration on a notch of l/s3 inch depth which
is approximately 2% of typical reheat pipe wall thickness).
l
The UT shear wave examination shall be done with a
nominal angle beam of 45 degrees or others, as needed,
based upon component geometry.
l
Scanning must ensure the entire volume of the weld is
covered; the search unit (transducer) shall overlap a
minimum of 10% of the previous pass; the search unit
scanning speed shall not exceed 6 inches per second; a
straight beam 0 degree UT scan must be performed; and
angle beam scans must be made in two directions -
parallel and perpendicular to the weld.
l
Evaluation must be made of all indications in excess of
20% DAC (Distance Amplitude Correction curve).
Criteria for evaluation of indications is directed back to the
referencing code section. For components such as headers,
the referencing section is the ASME Boiler and Pressure
Vessel Code Section I, Power Boilers. Acceptance criteria
for Section I established the requirements for construction
and manufacturing of new components and does not con-
sider aged or creeped material. Since creep crack growth
analysis relies upon time dependent fracture mechanics
and considers the case of aged (partially creeped) material,
this approach attempts more accurate determination of
critical flaw size. A full discussion of the analysis with
examples is given later in this paper. The most recent
analysis tool developed as part of EPRI sponsored research
project 2253-10 is called the BLESS Code. This is a PC
based program with algorithms to estimate time-to-crack
initiation as well as crack growth and propagation.
Ultrasonic detection of flaws in areas of complex geometry
are not well established. In the past, attempts to detect flaws
or cracking in complex components, particularly high
temperature headers, have had mixed results at best. Since
the geometries that may be encountered vary greatly be-
tween the headers in different boilers, no one technique can
be developed that is guaranteed to be effective in each case.
Once a flaw is detected in the header information is needed
regarding its size and orientation. Accurate dispositioning
of the flaw by nondestructive methods is difficult and
highly dependent upon flaw location in the header, as well
as the experience and knowledge of the technician. Knowl-
edge of flaw size, flaw geometry, i.e., planar versus volu-
metric, flaw orientation, flaw location and flaw depth are
critical to the analysis.
Occasionally other NDE methods are needed in the header
assessment. Normally other methods are used to help
evaluate damage found by methods described above.
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Radiographic Testing (RT) is used sparingly as an NDE
method during level II condition assessment and is not
recommended by B&W for header assessment programs.
Significant research was done to investigate RT as an NDE
tool for heavy section components, particularly seam welded
piping - reference EPRI CS-4774. However, RT effective-
ness was found to be too sensitive to flaw orientation and
flaw size to be a reliable NDE method. RT as part of a
header assessment is more likely to be used as part of weld
repair certification than for detection of damage.
Eddy Current Testing is a common technique used for
inspection of small, thin wall components such as tubing in
heat exchanges and steam generators. Eddy current has
limited applications in the field testing of heavy wall
components such as headers. Evaluations done with eddy
current techniques have included seam weld detection on
headers and piping and crack sizing of bore hole ligament
cracks. Welds in ferritic steel can have appreciably differ-
ent electrical properties compared to the base metal that
they join. These differences vary and are related to the
combined effects of chemistry, fabrication process, and
effective heat treatment. Properly designed eddy current
instrumentation has been shown to have the ability to detect
material changes associated with the header welds. Typi-
cally an eddy current technique is use for scanning and
weld detection followed by an acid etch test to verify the
presence of the weld.
B&W developed an eddy current device for the sizing of
small bore hole ligament cracks. The technique uses spe-
cially designed probes which are inserted into the header
bore hole through an external access. The eddy current
signal response to known ID notch sizes in a calibration
standard is used to provide the data needed for interpreting
and estimating the sizes of bore hole cracks. The inherent
characteristics of eddy current limit this crack sizing ability
to relatively shallow cracks (l/8 inch or less in depth).
Alloy Analysis is sometimes done in the field if there is
question regarding the exact material that was used in
manufacture of the component or weld. Although this can
be a problem in piping with the many spool pieces and
numerous field welds, it is rarely a problem with headers.
The most likely area where field analysis would be needed
would be in verification of a field weld at the outlet
connection. Testing is usually done using one of the com-
mercially available nuclear alloy analyzer instruments.
Field alloy verification is not normally required in the
typical header assessment program.
Data acquired during the outage inspection is next used for
assessment of the header in phase III of the condition
assessment program. Header assessment may include analy-
sis to quantify remaining life. As stated earlier, quantifying
remaining life for high temperature headers is based upon
time dependent fracture mechanics and considers crack
initiation and creep crack growth. A full discussion of the
mechanisms of crack initiation and crack growth, as well as
the analyses for predicting header remaining life are pre-
sented in the discussion that follows.
DAMAGE MECHANISMS
As previously discussed, there are several damage mecha-
nisms that contribute to ligament damage in elevated tem-
perature components. These mechanisms include creep,
fatigue and oxidation. The damage process consists of two
phases: crack initiation and crack propagation. The follow-
ing discussion of the header damage mechanisms is based
on the approach used in the EPRI developed BLESS
(Boiler Life Evaluation and Simulation System) Code. The
deterministic version of the BLESS Code was developed,
for EPRI, by B&W as a subcontractor to General Atom-
icstll. Prior to discussing the damage mechanisms, it is
appropriate to firs t review basic material behavior concepts
and test methods used to characterize material behavior.
MATERIAL BEHAVIOR
Plasticity
The tensile test is used to determine the time-independent
inelastic, or plastic, behavior of materials. The tensile test
involves subjecting a specimen (generally a polished solid
cylindrical bar) to a monotonically increasing elongation
(i.e., stretching) while simultaneously measuring the
uniaxial tensile force required to maintain a constant strain
rate. The test is conducted at a well controlled constant
temperature and constant strain rate and is continued until
the specimen fractures (i.e., complete separation). The
measured load and corresponding elongation measure-
ments are used to construct an engineering stress-strain
curve similar to that depicted in Figure 12. The engineering
stress is determined by dividing the measured load by the
original cross-sectional area of the specimen. The engi-
neering strain is determined by dividing the measured
elongation of the gage length by the original gage length.
The load and elongation are linearly related during the
initial elastic deformation. Elastic deformation is recover-
able; i.e., the specimen will return to its original length if
the load is removed. Plastic deformation will occur as the
elongation continues. This deformation is characterized by
the non-linear load-elongation curve. Plastic deformation
is not recoverable. The specimen will not return to its
original length when the load is removed. The unloading
curve is parallel to the elastic portion of the loading curve,
indicating that the elastic deformation is recovered. The
deformation remaining after load removal represents the
plastic deformation. The initiation of plasticity is often
accompanied by a slight load plateau (or even a drop in
load) at the end of the elastic deformation. This behavior
identifies the yield point. The load, required to sustain
further deformation, continually increases to a maximum
value. The plastic deformation is uniformly distributed
over the specimen length prior to achieving the maximum
load. The plastic deformation becomes localized, and un-
stable, resulting in specimen “necking” as evidenced by the
achievement of the maximum load. Subsequent deforma-
tion is sustained with less and less load. However, the
material continues to strain harden (i.e., becomes stronger,
or more resistant to deformation) throughout the test.
Localized necking occurs when the specimen area de-
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Englnwlng Stress: o = PIA o
L - L.
Engineorlng Strain: E 7
P
Elastic
Modulus
T
L
1
Uniform
SIrSill
\I
--it-
I
0.002 ltllhl
Strain
Figure 12 Typical engineering stress-strain curve.
creases more rapidly than the material strain hardens. This
results in the appearance that the material is becoming
weaker, since less load is required to continue deformation.
The important features of the engineering stress-strain
curve are summarized as follows:
Proportional Limit: The stress level at which the curve
first deviates from linearity.
Elastic Modulus: The slope of the initial linear
portion of the curve, i.e., up to the
proportional limit.
Yield Strength: The stress level associated with a
small amount of permanent, or
plastic deformation; usually 0.2%
strain.
Ultimate Strength: The stress level associated with the
maximum load.
Uniform Strain: The strain (expressed as a percent)
corresponding to the maximum load.
Fracture Strain: The strain (expressed as a percent)
corresponding to fracture.
The stress-strain curve is very dependent on the test tem-
perature. In general, all measures of strength decrease as
the test temperature increases. The elastic modulus de-
creases as the test temperature increases. The modulus is
insensitive to material conditions and minor variations in
alloying additions and thus varies very little from lot-to-lot.
The yield strength and ultimate tensile strengths are very
sensitive to material condition and minor variations in
alloying additions and thus exhibit significant lot-to-lot
variations.
The fracture strain is a measure of the ductility of a
material. However, this measure of ductility is very sensi-
tive to the the gage length as a result of the localized
straining that occurs during necking. The percent reduction
of area is a more useful definition of uniaxial tensile
ductility since it eliminates the effect of gage length. The
reduction of area is defined as the ratio of the decrease in
specimen cross-sectional area to the original area. In gen-
eral, the ductility increases as the test temperature in-
creases.
At very high strain rates, the stress-strain curve can be
significantly affected by the strain rate at which the tensile
test is conducted. However, at the low strain rates that
characterize the response of boiler components to operat-
ing transients, the strain rate effects are generally consid-
ered insignificant.
Long term exposure to elevated temperatures, e.g., experi-
enced during normal boiler operation, results in a decrease
in the short-time tensile properties as determined by the
tensile test. The effect of service time and temperature on
the subsequent yield strength of 2$Cr- 1Mo steel is shown
in Figure 13.
As discussed earlier, at loads less than the ultimate tensile
strength (UTS), the load must be continually increased in
order to sustain continued deformation in a low-temnera-
1.0
0.S
%
0.6
>
i
E
S
% 0.7
i
f
0.6
0.5
10'
10' 10'
10'
10'
nmr, nours
Figure 13 Effect of service time and temperature on the
yield strength of 2 l/&r-1Mo.
12
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ture tensile test. That is, elongation (or straining) will cease
if the load is held constant at some point below the UTS. At
high temperatures, elongation will continue to fracture,
even if the load is held constant. This time-dependent,
elevated temperature, deformation is called creep. The
creep test requires subjecting a specimen (similar to the
tensile test specimen) to a constant, uniaxial load at a well-
controlled constant temperature, while simultaneously
measuring the elongation. If this test continues to rupture
(fracture), it may be referred to as a creep-rupture test. The
primary objective of this type of testing is frequently to
establish only the time to rupture (fracture). With that
objective, the elongation measurements may be made at
longer intervals, and the test may be referred to as a stress-
rupture test.
A classical creep curve is shown schematically in Figure
14. The specimen is heated and stabilized at the test
temperature prior to loading. The specimen elongates as
the load is gradually applied. Depending on the test tem-
perature and stress level, the initial elongation (or loading
strain) may have elastic and plastic components or it may
be entirely elastic. The creep curve generally consists of
three stages of creep deformation: the primary, secondary
and tertiary stages. Primary creep is characterized by a
relatively rapid, yet decreasing, strain rate (or creep rate).
The decreasing creep rate (at a constant stress) indicates
that the material is becoming more resistant to deforma-
tion, i.e., it is strain hardening. Secondary creep is a period
of nearly constant creep rate that results from a balance
between the competing processes of hardening and recov-
ery. Secondary creep is usually referred to as steady-state
creep. The average value of the creep rate during secondary
creep is called the minimum creep rate. Tertiary creep is
characterized by an increasing creep rate. This increasing
rate is, in part, due to an increasing stress, especially at the
higher test temperatures and stresses. The stress increase,
during the constant load test, is the result of the specimen
cross-section being reduced during elongation. The speci-
men cross-sectional area can also be reduced by the forma-
tion of grain boundary voids and microcracks, thus contrib-
uting to the increase in creep rate.
Period of
1
-Primary
creep
Period of
-Initial Extension
0
1
I
I
I
Time
Figure 14 Classic (diagrammatic) creep test at constant
load and temperature.
The test temperature has a very significant effect on the
results of these tests, as llustrated in Figures 15 and 16. As
an exam
P
e, at a stress level of 10 ksi, the minimum creep
rate of 2 /4Cr-1Mo is increased by approximately 50 per-
cent when the test temperature is increased from 1000°F to
1010’F. The rupture life is decreased by a similar ratio.
Figures 15 and 16 also illustrates the strong effect of stress.
As an example, at a test temperature of lOOO”F, the mini-
mum creep rate of 2$Cr- 1Mo is nearly doubled when the
stress level is increased from 10 ksi to 11 ksi. This same
increase in stress level results in a loss of about half of the
rupture life.
100,
I I I 1
(‘36’rrm-77
I I I
,l.~,O.Ol
I I I IllIll
I I I111111 I I IIIIILJ
0.10 1.0 10
Creep Rate, %/loo0 h
Figure 15 Creep rate curves for 2*/&r-1Mo steel,
Figure 16
Steel.
Typical creep rupture curves for 21/,Cr-1Mo
The BLESS Code uses the following equation to character-
ize the creep strain as a function of stress, temperature, and
time.
EC = [Bt(p+ l)] A (a / 1000)m + A(o / 1000)?
(1)
where: EC = creepstrain
t = time
0 = stress
p,m,n
= constants
A3
= functions of temperature
The first term characterizes the primary creep and the
second represents the secondary, or steady-state creep. The
form of the creep equation is dictated by the requirements
of the crack growth model.
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Parameter methods have been developed to assist in the
interpolation and extrapolation of creep rupture tests. The
Larson-Miller parameter is probably the most frequently
used. The Larson-Miller parameter, P, is defined as:
P = T(C+log f)
(2)
where: T
= temperature in degrees Rankine
C = a material constant, often equal to
approximately 20.
f
= time to rupture in hours.
Data obtained over a limited range of test conditions is used
to generate a master rupture curve. The parameter method
then allows the interpolation and extrapolation of the
limited data to conditions for which data does not exist. The
BLESS Code uses the Larson-Miller parameter method to
represent the time-to-rupture behavior of the Pl 1 and P22
materials.
Fatigue
Repeatedly subjecting a material to either load-controlled
or strain-controlled cycling may result in a fatigue failure.
Strain-controlled fatigue tests are used to study the behav-
ior of boiler materials, since boiler component cracking
often results from low cycle, strain-controlled thermal
loading. The fatigue test specimen is generally hour-glass
shaped and is subjected to uniaxial push and pull at a
constant temperature. The tests are usually conducted at a
constant strain rate and constant strain range, with zero
mean strain as llustrated in Figure 17. A strain cycle occurs
as the strain goes from an initial value through an algebraic
maximum and an algebraic minimum and then returns to
the initial value. The number of strain cycles required to
I
OxemNT J
CCWSTANTlEMPERWJRE
STRAIN RATE
SrRAlN -co
t4TlCUED FATKXIE TEST DESCFWTDN
I
LOG -NUMBER OF CYCLES TO FAILURE
Figure 17 Typical representation of fatigue data.
produce a failure is referred to as the fatigue life. The
applied cyclic strain range is the principal variable govem-
ing the number of cycles to failure in a strain-controlled
fatigue test. Data from several tests run at the same constant
temperature and same constant strain rate, but each with a
different constant strain range, allows construction of a
fatigue curve for the test temperature and strain rate. The
fatigue curve is generally presented as log-strain range
versus log-number of cycles to failute, as illustrated in
Figure 17. At low temperatures (i.e., temperatures at which
creep is unimportant), the effects of temperature and strain
rate are insignificant and usually ignored. As a result, a
single fatigue curve provides an adequate representation of
low temperature behavior. Both the temperature and strain
rate can significantly affect the fatigue behavior at tem-
peratures at which creep behavior is important.[~l
CRACK INITIATION
The initiation phase is generally considered to be the result
of two competing processes: oxide notching and creep
fatigue. The time & cycle fractions model is usually se-
lected as the basis of the creep-fatigue initiation model.
Oxide Notching
The oxidizing potential of steam results in the formation of
predominately magnetite (Fe,O,) on the surfaces of Pll
and P22 headers at their usual boiler operating tempera-
tures. The oxide grows during periods of sustained opera-
tion at elevated temperature. The oxide grows initially at a
rapid rate with the growth rate decreasing with time, i.e., as
the oxide thickness increases. The oxide growth is usually
represented as parabolic, as illustrated in Figure 18. The
relatively rapid decreases in outlet leg steam temperatures
that accompany decreases in boiler load (Figure 8) result in
tensile stresses at the interior surfaces of the header and
bore holes. The tensile stresses are sufficient to crack the
relatively brittle oxide. When the oxide is cracked during a
load decrease, the base metal is again exposed to the steam,
allowing the initial high rate of oxidation to be re-estab-
Figure 18 Oxidation of low alloy steel in high tempera-
ture steam environment.
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I
0.0
I
I
0
200
400
600
TIME. HOURS
Figure 19 Oxidation during cyclic operation.
lished. As this process continues over time, the header is
preferentially oxidized along the crack in the oxide, even-
tually forming a notch. Figure 19 schematically illustrates
how boiler load changes can accelerate the formation of
oxide notches. The local steam temperature is also a sig-
nificant contributor since the growth of the oxide is a strong
function of the steam temperature. For example, the BLESS
Code defines the oxide thickness for the P22 material as a
function of time and temperature as follows:
Oxide thickness = 1.23 1 exp (-8496.5/T) P
(3)
where: Oxide thickness is in inches
T = Temperature in degrees Kelvin
t = Time in hours
The bore holes of the outlet legs that operate at the highest
steam temperatures have the most significant formations of
magnetite and thus the highest probability of significant
oxide notching. Figure 20 illustrates the basis used to
extend the above equation to conditions of variable tem-
perature. The curve labeled T, represents the growth of the
oxide at a constant temperature of T,, while curve T,
represents the growth at a higher temperature, T? Assume
that a temperature of T, is sustained for a ttme of t,,
allowing the oxide to grow as illustrated by line segment O-
1 of curve T,. If the temperature is then changed to T,, and
held for a time duration of dt, the oxide will grow as
represented by line segment 2-3 of curve Tz .
11
T I M E
Figure 24 Accumulation of oxide at variable temperamre.
Creep-Fatigue
The phenomenological time & cycle fractions model views
the damage process as being composed of separate rate-
dependent and rate-independent damage processes. The
rate-dependent part is termed creep damage and is based on
Robinson’s Linear Life Fractions RuleP That rule states
that the creep life has been expended when the sum of the
life fractions, or time fractions, equals unity, as:
De=
(4)
where: D, = Accumulated creep damage
9
= number of time intervals (each with a unique
stress-temperature combination) needed to
represent the specified elevated temperature
servie life for the creep damage calculation.
At = time duration of the load condition, k.
Tr
= the time-to-rupture for the temperature and
stress combination of load condition, k.
Determined from constant temperature and
constant load, uniaxial, stress rupture tests.
The time fractions model is thus seen to provide a method
to estimate creep damage, for variable stress and tempera-
ture service conditions, using the results of constant load,
constant temperature,
Stress rupture t&t%
An
example Of
the application of this rule, for a very simple loading
history, is illustrated in Figure 21. The time histories of the
stress and temperature are shown in that figure. The tem-
perature is held constant at T, from time zero to time, b. The
temperature is increased to T, at time, tr, and held at that
temperature until time, ts. The stress is increased from 0, to
cr2at time, t,, and subsequently decreased to o, at time, b.
The creep damage is calculated as the sum of the incre-
ments of damage incurred during each of the three intervals
of constant stress and temperature. The incremental dam-
age incurred during any one of these time intervals is
determined as the time fraction. The time fraction is de-
fined as the time interval, At, divided by the time-to-
rupture, T,, at the corresponding stress and temperature.
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Tz
T
I
I
I
f I
I
‘1
‘2
‘3
T I M E
I
T
‘2
T
r3
T
r1
LOG T I ME -TO-RUP TURE
CREEP DAMAGE D, =,&$ 5 1
D, I ($j + (y) + te)
Figure 21 Robinson’s Life Fractions Rule.
The time-to-rupture, T,, is determined from the stress
rupture curve for the appropriate temperature, as shown in
Figure 21.
The rate-independent part of the damage is termed fatigue
damage and is based on the Miner linear damage model[*l.
That model states that the fatigue life has been expended
when the sum of the cycle fractions equals unity, as:
Df t[#] s l
j=l aj
(5)
where: D, = Accumulated fatigue damage.
P
= number of load conditions (each with a
unique strain range-temperature combin-
ation) needed to represent the specified
elevated temperature service life for the
fatigue damage calculation.
= number of cycles of loading condition, j.
= allowable number of cycles for the strain
range and temperature of loading
condition, j .
The cycle fractions model is thus seen as a method to
estimate damage for variable service conditions using the
results of constant strain range, constant temperature, fa-
tigue tests. An example of tire application of this model is
illustrated in Figure 22 for a very simple cyclic strain
history. The assumed strain-time history consists of three
strain cycles of strain range Ae,, two cycles of strain range
AF+ and four cycles of strain range AQ The increment of
fatigue damage attributable to cycling at any one of the
strain ranges is defined as the number of applied cycles, n,
of that strain range divided by the allowable number of
cycles, N,, at that strain range. The allowable number of
cycles, N,, is determined from a fatigue curve of log strain
range vs. log cycles to failure, as shown in Figure 22. That
fatigue curve is constructed from the data of several fatigue
tests, each run at a constant, yet different, strain range.
Determining strain ranges and counting fatigue cycles for
the actual operating history of a boiler component is gen-
erally not as straight forward as the example of Figure 22.
To accomplish this task in an orderly manner requires what
is commonly referred to as a cycle counting method. The
BLESS Code uses the Range Pair MethodW The basis of
the method is that a strain cycle, or fatigue cycle, is defined
as complete when tensile-going strain is reversed by an
equal amount of compression-going strain, and vice-versa.
The initiation process is assumed to be completed when the
sum of the creep damage and fatigue damage exceeds the
allowable damage, D, asW
A2 Nl
LOG-CYCLES TO FAILURE
P
FATIGU E DAMAGE D, = x($ 5 1
JIl
Df =3+1+4
“I “2 4
Figure 22 Miner’s Linear Damage Rule.
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D,+D, ID
03
The allowable damage is usually defined by a bilinear
damage diagram, or damage envelope, similar to that of
Figure 23.
0.1
1.0
FATIGUE DAMAGE
Figure 23 Damage diagram used with the time and cycle
fractions creep-fatigue model.
As a result of the strong dependence of creep damage on
stress (Figure 16), it is quite important that the stresses be
accurately predicted. For example, the direct use of an
elastically calculated stress-time history will generally
provide grossly inaccurate estimates of creep damage. It is
thus necessary that the stress calculations capture the
important features of the inelastic response of the material.
As an example, consider the behavior of a thick-walled
high temperature header during start-up. The inside surface
of the header is subjected to large compressive thermal
strains as the temperature of the steam rapidly increases
during a start-up. The compressive strain at the inside
surface occurs since that surface is heated more rapidly
than the rest of the thick section. The thermal expansion of
that warmer surface is then restrained by the rest of the
section, resulting in compressive stresses at the inside
surface. As the operating temperature is approached, the
rate of heating is decreased and the temperatures, through
the thickness, begin to equalize. As the metal temperatures
equalize, the thermal strains and stresses are dissipated.
However, as a result of plastic straining, large residual
stresses may remain. These residual stresses may be quite
damaging as the header begins a period of sustained opera-
tion at elevated temperatures. This type of loading history
is illustrated with the aid of a simple bar subjected to strain
controlled axial loading as shown in Figure 24. The bar is
initially loaded, beyond the yield stress, to a strain level of
Aa,. This is representative of the compressive strain at the
inside of a thick-walled header during a start-up. The
elastically calculated stress is represented by point 1, while
the actual stress, represented by point 2, lies on the stress-
strain curve. Note that the stress of point 1 is considerably
in excess of the yield stress. Since the creep damage is a
strong function of stress level, the use of the elastically
calculated stress (point 1) would greatly over-estimate the
creep damage. If the bar is then returned to near its original
strain level (i.e., zero), an elastically calculated solution
would indicate that the stress also returned to zero, as
represented by point 3. However, as a result of the plasticity
incurred during the initial loading, the actual unloading is
along line 2-4, resulting in the residual stress represented
by point 4. This unloading is similar to that in a header as
the temperatures tend to equalize following the start-up. In
this situation, the use of the elastically calculated stress
(point 3) would incorrectly indicate zero creep damage.
The use of the residual stress, represented by point 4,
captures the effect of the plasticity that occurred during the
thermal transient associated with the start-up. That residual
stress is an important contributor to creep damage since it
exists when the unit begins sustained operation at elevated
temperature.
Creep strain may also significantly influence the stress-
strain response. For example, the residual stress of the
above example (i.e., point 4 of Figure 24) will relax to a
lower level as a result of creep strain incurred during
sustained elevated temperature operation. This relaxation
behavior, at constant strain, is illustrated in Figures 25 and
26. Figure 25 illustrates the effect that relaxation can have
P
&
4 RESIDUAL
1
- YIELD STRESS
Figure 24 Residual stress after a boiler start-up.
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Figure 25 Relaxation during sus
quent to start-up.
RELAXATION
I
STRAIN
INITIAL STRESS
POINT 4 OF FIGURE 4.16
YIELD STRESS
Figure 26 Stress-time history during relaxation.
ined operation subse-
on the stress-strain history during a strain-controlled cycle.
Figure 26 illustrates the stress-time history during relax-
ation. It is seen that the sustained stress, and thus creep
damage, would be significantly over-estimated if the creep
relaxation were ignored. That is, the use of the residual
stress (point 4 of Figure 24), throughout the period of
steady operation, would be overly conservative. It should
also be realized that stress relaxation and creep damage
occur during periods of transient operation, as well as
during steady operation. For example, relaxation and creep
damage w