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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


3/31

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

Babcock 8 Wilcox


4/31

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.

3


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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

Babcock 8 Wilcox


6/31

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.

Babccck & wlkox 5


7/31

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

Babcock &Wilcox


8/31

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

Babcock 8 Wilcox 7


9/31

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.

8 Babcock & Wilcox


10/31

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

Babcock

8 Wilcox

9


11/31

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.

10 Babcock 8 Wilcox


12/31

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-

Babcock&Wilcox

11


13/31

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

Babcock 8 Wilcox


14/31

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.

Babcock & Wilcox

13


15/31

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.

14

Babcock 8 Wilcox


16/31

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.

Babcock 8 Wilcox 15


17/31

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.

16 Baboook 8 Wilcox


18/31

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.

Babcock & Wilcox

17


19/31

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

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