BR-1586

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

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 dimensional 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 inspection 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 techniques and analysis methods which are all directed at determining the remaining useful life of these high temperature headers.

INTRODUCTION

HISTORICAL PERSPECTIVE

Hlgh Temperature Headers

In 1982 Babcock & Wilcox (B&W) first began its investigation 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 relatively 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 l/$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 resulted. 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 significant damage are the superheater outlet headers that

operate at temperatures near 1000°F. High temperature headers are generally constructed of 1 1/4Cr-1/2Mo (SA335 Pll) or 2’/,,Cr-1Mo (SA335 P22) steels. The typical operating 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 dependent upon sustained stress at relatively high temperature. Given sufficient time in operation, creep damage will accumulate from exposure to the normal operating temperatures 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 mechanisms of oxide notching and fatigue. These damage mechanisms, 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 frequent 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 normally 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 of 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 deforms 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 indicates 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 comprehensive 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 pressure 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 excessive 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 problems 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 cracking 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 expansion 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 assessment 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 was 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 l/,Cr-l/zMo (Pl 1) and 2 /,Cr-1Mo (P22) headers.

Ligament cracking generally initiates as numerous longitudinal 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

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 temperature 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 off 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, cracking 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 typically 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 effects of the thermal stresses. Second, the bore hole openings 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 temperature variations that can occur across the width of the boiler and superheater. Tube temperatures may vary resulting 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 mismatch is more likely in transient operating conditions, such as load 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 temperature can accelerate creep damage, oxide growth and crack growth rates.

In general, quantifying the remaining life of high temperature headers focuses on analysis and prediction of header crack growth which has been developed using time dependent 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 headers, 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 inspected. 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 thickness 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, was 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 temperatures over 1050°F were found to have experienced ligament cracking, illustrating the significance of temperature and its effect on creep.

Radial Nonradial RadiaVNonradial

Figure 6 Header bore hole penetration patterns.

Babccck & wlkox 5

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 distribution 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 superheater and reheater experience differences in the steam flow in individual tubes within the bank. A tube carrying greater steam flow will experience less of a steam temperature increase than a tube with reduced flow, 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 occur 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 distribution will vary from that shown in Figure 7.

Figure 7 Steam temperature variation in a header.

As a consequence of the through-wall temperature differences 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 increasing 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 transient 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 difference 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.

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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 combined 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 follows 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 consider in the phase I review includes the material and design type. Is it 1 l/&r or 21/qCr alloy? Is the header made of seam welded pipe? Does it have radial, nonradial or a combination 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 support failures or cracks? Have steam leaks been experienced? 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 occurred 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 as how 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 operation. 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 selecting 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 particle testing (MT), liquid dye penetrant testing (PT), dimensional 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, internal inspections (normally with video probe), dimensional analysis and ultrasonic testing for volumetric examination 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 was 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 problems. 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 identify macroscopic cracking associated with creep or fatigue. Overheating of the header or of the outlet legs can sometimes 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 surfaces 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 magnetic 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 examinations, 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, therefore, 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 performed 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 consequence 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 larger 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 shallow 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 important. 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 available then the outlet leg temperatures measured during operation can be used to guide the selection. If thermocouples data is not readily available then B&W recommends 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 temperature. 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, off 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 information 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 considerations. 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 excursions are associated with the highest outlet leg temperatures. 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 arrangement 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 environment 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 primary 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 assessment 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

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 measurement. 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 applicability 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 volumetric 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, regardless 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 was 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 consider 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 between 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 volumetric, 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

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 effectiveness 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 different 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 specially 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 analysis 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 presented in the discussion that follows.

DAMAGE MECHANISMS

As previously discussed, there are several damage mechanisms that contribute to ligament damage in elevated temperature components. These mechanisms include creep, fatigue and oxidation. The damage process consists of two phases: crack initiation and crack propagation. The following 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 first 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 measurements 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 engineering 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 recoverable; 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 deformation 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

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 temperature. In general, all measures of strength decrease as the test temperature increases. The elastic modulus decreases 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 sensitive 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 general, the ductility increases as the test temperature increases.

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 operating transients, the strain rate effects are generally considered insignificant.

Long term exposure to elevated temperatures, e.g., experienced 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

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 temperature 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 deformation, 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 specimen cross-sectional area can also be reduced by the formation of grain boundary voids and microcracks, thus contributing 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 illustrated in Figures 15 and 16. As an exam

P le, at a stress level of 10 ksi, the minimum creep

rate of 2 /4Cr-1Mo is increased by approximately 50 percent 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 minimum 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’ krrm-77 I 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 characterize 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

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 behavior 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 illustrated 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 temperatures 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 selected 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 temperatures. The oxide grows during periods of sustained operation 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 temperature steam environment.

14 Babcock 8 Wilcox

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, eventually 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 significant 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 temperature. 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

TIME

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 temperature 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 temperature 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 cr2 at 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 damage incurred during any one of these time intervals is determined as the time fraction. The time fraction is defined as the time interval, At, divided by the time-to- rupture, T,, at the corresponding stress and temperature.

Babcock 8 Wilcox 15

Tz

Tl I

I I f I I

‘1 ‘2 ‘3

TIME

I

T ‘2

T r3

T r1

LOG TIME-TO-RUPTURE

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 combination) 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, fatigue 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 generally 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 FATIGUE DAMAGE D, = x($ 5 1

JIl

Df =3+1+4 “I “2 4

Figure 22 Miner’s Linear Damage Rule.

16 Baboook 8 Wilcox

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 operation 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

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 relaxation. 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 will occur as the metal temperatures reach into the creep regime during the start-up depicted in Figure 24.

CRACK PROPAGATION

Once a crack has been initiated in a high temperature header by either oxide notching or creep-fatigue, it can propagate under fatigue or creep conditions, with the potential to cause leaking or failure. The crack driving force for fatigue is the cyclic stress intensity factor, AK, and for creep it is the C, parameter of SaxenaW

Macroscopic crack growth in a creeping material occurs by local failure resulting from nucleation and coalescence of micro-cavities in the highly strained region ahead of the crack tip. When the fracture process zone ahead of the crack tip is small, a detailed accounting of the fracture process is not necessary for predicting creep crack growth. Creep crack growth has been shown to be governed by a time-dependent loading parameter that characterizes the

geometry and applied loading of a flawed component. In linear elastic fracture mechanics (LEFM)-controlled fatigue crack growth, the governing parameter is the range of the stress intensity factor, AK. In elastic-plastic fracture mechanics (EPFM)-controlled ductile crack growth (tear- ing), the governing parameter is the J-integral. In time- dependent fracture mechanics (TDFM), the analogous crack tip parameter is the energy release rate (power) parameter, C,, which correlates creep crack growth rates through the relationship:

da/dt=bC; (7)

Figure 27 illustrates a typical correlation between C, and da/dt for 1 */qCr- l/ZMo steel. The remaining service life of an elevated temperature header, with an existing crack, can be estimated by the numerical integration of the above equation. In order to do this, a methodology for determining C, must be established.

JOULEShn’HR

I,-,3 1 114 Cr - l/Z MO STEEL 536°C (lCOO°F)

0.01 0.1 1.0 10.0 C, (IN LBWlN’HR)

Figure 27 Typical correlation between C, and da/dt.

18 Babcock 81 Wilcox

The TDFM crack tip parameter, C,, which correlates creep crack growth rates is dependent on the level of creep deformation at the crack tip. Creep crack growth occurs under small-scale creep characterized by a creep zone which is small relative to the overall dimensions of the cracked ligament and the crack length. In the steady-state condition, this creep zone spreads over the entire untracked ligament. The transition creep conditions lie between the small-scale creep and steady-state regimes. Both these small-scale and transition creep regimes are under non- steady-state conditions because the crack tip stress varies with time. Under steady-state creep, where the crack tip stresses no longer change with time, the crack growth behavior can be characterized solely by the path-independent energy rate line integral, C*W By interpreting strains and displacements in the definition of the line integral, C*, as their rate counterparts, C* can be thought of as the analog of the line integral, J.

Both C* and C, can be interpreted as the difference between energy rates between two cracked components with incre- mentally differing crack lengths “a” and “a + Aa”. Further- more, C* characterizes the strength of the crack tip stress singularity commonly known as the Hutchinson, Rice, and Rosengren (HRR) singularityti3~i41 in the same manner as the J-integral characterizes the elastic-plastic stress singularity. Experimental studies have shown that C* characterizes the creep crack growth under large-scale (steady-state) creep conditions.

The problem of determining C, is analogous to determining the magnitude of the J-integral under elastic-plastic conditions.ti51 From Reference 16, accurate estimates of J, over a wide range of elastic-plastic conditions, can be made by adding the J values, obtained from small-scale yielding expressions in terms of K,, to J plastic values from expressions for fully plastic loading. Motivated by the analogous J-integral formulation, Saxenatil established a general formulation for estimating C, based on applied loading (stress) for a wide range of creep conditions. An alternate, much simpler, expression for C, originally proposed by BassanW and later verified by Bloomtl*l as giving similar (to within 2%) life predictions to Saxena’s general formulation, has been used in BLESS for creep crack growth rate calculations.

The alternate formulation for C!,, for the special case in which primary creep (i.e., the first term of Equation 1) may be considered negligible , is

where tr is the transition time for extensive secondary creep conditions to develop from small scale creep, and

(9)

Note that C, can be estimated from the remotely applied unrelaxed stresses and the knowledge of the elastic and the creep behavior of the material, the stress intensity factor,

K,, of the flawed structure and the C* expression for the geometry of interest. The stress intensity factors, K,, can be found in handbooksti91 while the C* expressions can be found in Reference 16 for limited geometries. The second stage creep properties can be obtained from uniaxial creep teStS.

While formulations based on secondary creep alone are adequate for the evaluation of base load operation, primary creep must be considered in the evaluation of cycling operation. The high strain rate associated with primary creep contributes to rapid crack growth rates at short times. This detrimental effect of primary creep can be sustained throughout the life of a cracked component under cyclic loading conditions. The local plastic straining that occurs during start-ups and shut-downs can negate the effects of prior creep relaxation of the stresses at the creep tip so that the initially high values of C, are reinitiated during each operating cycle as schematically illustrated in Figure 28. The local plasticity can similarly wash-out prior creep strain hardening so that the creep response is once again characteristic of primary creep at the start of each cycle. The significant effect of primary creep on the service life of flawed components subjected to cyclic elevated temperature service was demonstrated in Reference 18. The inclu- sion of primary creep in the creep crack growth driving force parameter, C, or C(t) is thus highly desirable.

Figure 28 Schematic of pressure, temperature, and C, versus time for a simple on/off cycling.

If the uniaxial primary creep strain can be represented by the first term of Equation 1 then[zJp211

C(t) = [l+(tTPlt)+(t2/t)P’(1+p)]C* (10)

This expression is used in the BLESS Code as the approximate creep crack growth driving force parameter. The transition time, trr, in Equation (10) is the time for extensive primary creep conditions to develop from small scale primary creep and is given by the following equation

t, = [Kgj”“31”p’ . & (11)

Babcock 8 Wilcox 19

This result is due to analyses by Riedel.[**l

Figure 28 schematically illustrates a plot of the pressure and temperature of a typical header during a start-up and shutdown cycle. For this situation, the value of C, decreases with time during normal operation and goes to zero at shutdown. At the next start-up, the value of C, starts from its initial value as shown in the figure. For start-up and shutdown (cyclic) operation, the crack growth rate is best estimated on a per operating cycle basis and creep-fatigue crack growth (CFCG) can be determined by partitioning the crack growth into a cycle-dependent part and a time- dependent part such that

The time-dependent crack growth occurs only under constant amplitude loading during the hold period of the cycle, t,,. Thus, C* and C, can be used in characterizing the crack growth rate during this hold period. Due to experimental limitations, it is difficult to obtain instantaneous values of da/dt and C, during the hold period. However, average values of the crack growth rate and the C, parameter can be accurately measured. The average da/dt and C, are obtained as follows:

and

(13)

(14)

The (da/dN),, is the crack growth during the hold period aud is obtained by subtracting the cycle-dependent crack growth rate from the total crack growth rate. Note that using Equation (14) for cyclic loading along with Equation (8) or (10) would be a problem because of the singularity in those equations at t=O. However, following the application of the load, there is a finite incubation period, fplr during which crack growth is very slow. The incubation time is related to the time required for generating the creep zone within the cyclic plastic zone. This time, t,,,, has been calculated by YoorW for several material types and is on the order of 0.03 hours for cyclic loading conditions with primary creep. Allowing for instantaneous plasticity during cyclic loading using the C, expression modified by replacing t by t+$, eliminates the problem of the singularity at t=O.

The cyclic term can be written as a!u L-1 (w’cYde = C(AK)” which has the same dependency for fatigue crack growth rates under conditions where time-dependency is insignificant.

The total crack growth rate is then given by

g = c,fw + C,[CJ,j%* (16)

The significance of using Equation (1) versus Equation (16) is that experiments of creep crack growth (CCG) have been used to determine b and q of Equation (7) while creep fatigue crack growth (CFCG) experiments have been used to determine C, and q, of Equation (16). Data from these two different type fatigue experimenW31 have shown (Fig- ure 29) that time-dependent crack growth behaviors under CFCG and CCG conditions can be expressed as a single trend if (da/dt), is characterized by (C& or da/dt is characterized by C,. This is an important conclusion for applications because material behavior measured by CCG testing can be used to predict component life under creep- fatigue conditions or material behavior measured by CFCG testing can be used in CCG conditions. Even though C, and (CJ, are equivalent parameters, their exact numerical values may differ slightly in the small scale creep regime for a given material. However, in application, this difference will not cause any significant problem, as C, is used for steady-state continuous operation, while (CJ,, is used onrn cyclic situations.

KJ/m’-hr

W2 lo” 1 10’ 10’ lU’@ 1111111, , ,,,,,,,, , ,,,1 ,,,, , , ,,,,,,, , ,,,1 ,,,, 4

l CCG 1.25cr - OSMO A CFCG 98 SECOND HOLD TIME 539% (1000°F)

lo” H CFCG SIX SECOND HOLD

1o-61 ’ ’ 111111’ ’ 0 llllld ’ 0 lrlllc’ ’ ’ llltll’ 1 ’ 0 ~J 1CP lo” lo“ d 1lS’ 1

(C,h”,~ klpsiln-hr

Figure 29 Comparison between CCG and CFCG data in terms of the measured (C,),.

A solution is not available for header ligament cracks that grow simultaneously from the header ID toward the OD and between adjacent bore holes as was shown schematically in Figure 5. As a result, the BLESS Code bounds the ligament cracking problem with the models illustrated in Figures 30 and 3 1. BLESS models the ID-to-OD cracking, illustrated in Figure 30, using the solution for a double- edge cracked plate under uniform stress. The BLESS model assumes that the header wall thickness represents one-half of the width of the double-edge cracked plate. The header OD is then representative of the line of symmetry of the double-edge cracked plate model. BLESS models the bore hole-to-bore hole cracking, illustrated in Figure 31, using the same double-edge cracked plate model. In this case, the width of the ligament (i.e., the distance between bore holes) represents the width of the double-edge cracked plate model.

For header cracking away from ligaments, and for piping, models for both longitudinal aud girth cracking are avail-

20 Babcock 8 Wilcox

Figure 30 ID-to-OD cracking in header ligaments.

able. Continuous longitudinal cracks can be modeled at the ID, OD, and subsurface. Semi-elliptical longitudinal cracks can be modeled at the ID. Continuous girth cracking can be modeled at both the ID and OD.

BLESS Code calculations of crack growth are terminated, and failure is considered to occur, when the crack depth reaches a user-specified fraction of the section thickness (in the direction of crack growth), typically 50 to 70 percent. At this depth, the crack is usually growing so fast that the remaining time to become through-wall is negligible. As a result, an accurate definition of critical crack size is not an important aspect of the problem[“Jsl.

HEADER TEMPERATURES AND STRESSES

As illustrated in Figure 32, a boiler header is a complicated three-dimensional (3D) structure. The detailed calculation of time-dependent temperature and stress distributions is a formidable. time-consuming task, requiring 3D finite element analysis techniques. As discussed earlier, it is also required to address both the time-independent (plasticity) and time-dependent (creep) inelastic behavior of the header material in order to obtain realistic damage estimates. Thus, a detailed finite element analysis further requires that nonlinear material behavior be included. Furthermore, it is generally not practical to perform those detailed 3D nonlinear finite element analyses for a complete, complex

LIGAMENT B

Figure 31 Bore hole-to-bore hole cracking in header ligaments.

(although typical) boiler service history. In order to make the solution of this problem practical requires that some simplifying assumptions be made. Appropriate simplifying assumptions were made in the development of the BLESS Code.

The major simplification of the BLESS Code is that the transient temperatures and stresses (both pressure and thermal) are based on a simple thick-walled cylinder model. The temperatures and stresses vary only through the wall in this one-dimensional (1D) model. Stress and temperature predictions of this 1D model are representative of the conditions far-removed from the local effects of the bore holes. However, the temperatures and stresses from the 1D model are modified in BLESS to represent the important features of the actual temperatures and stresses at critical areas of the specific header geometry. The 1D temperatures and stresses are modified using closed-form empirical equations that were developed using the results of detailed 3D linear finite element analyses of several headers subjected to various transient and steady-state operating conditions. The header geometries and operating conditions studied are representative of secondary superheater outlet headers designed for steam temperatures of 950°F to 1050°F and pressures of 2400 to 3800 psi. A typical finite element model is illustrated in Figure 33. That model is representative of a slice of a header as shown in Figure 34.

Babcock 8 Wilcox 21

CIRCUMFRENTIAL

Figure 32 Schematic representation of a header illustrating bore holes and ligaments.

The development of closed-form equations, capable of appropriately modifying temperatures and stresses from a 1D model to approximate the transient conditions in header ligaments, becomes practical when only those features essential to the crack initiation and propagation models are considered. There is no attempt, or necessity, to approximate the temperature and stress distributions throughout the entire ligament area. Specifically, the oxide notching model requires the time history of the local metal temperature at only that interior location (i.e., in contact with steam) experiencing the most severe temperature history. The creep-fatigue (time and cycle fractions) initiation model similarly requires the time history of the local metal temperature and local stress/strain at only that location experiencing the most severe conditions. In both cases, the location of interest is on the surface of the bore hole, either near, or at, the intersection of the bore hole and interior surface of the header main cavity. The most severe location can actually traverse up and down a limited length of the bore hole as the boundary conditions vary with boiler operation. However, there is no attempt, nor is there thought to be a necessity, to predict and record this limited move- ment of the most severe location. As a result, a generic, static location is used for all boundary conditions throughout the operating history.

The creep crack growth and fatigue crack growth models are designed to focus on the behavior of an entire ligament rather than on the behavior at a single surface location. The creep crack growth model is provided the time history of the metal temperature and pressure stress, both averaged over the ligament surface that is defined by the postulated crack growth path. Hence, only pressure-induced membrane forces on the ligament are considered in the calcula-

22

tion of the creep crack driving force. The fatigue crack growth model is provided the time history of the pressure- plus-thermal stress, again averaged over the ligament surface defined by the postulated crack growth path.

The transient 1D thermal solution is obtained by implicit finite difference procedures that consider an insulated outer surface and time-dependent steam temperature and

Figure 33 Typical finite element model of a transverse slice of a header.

Babcock 8 Wilcox

Figure 34 Transverse slice used in header modeling.

convection heat transfer coefficient on the interior surface. The convective heat transfer coefficient is based on standard heat transfer correlations for single phase steam.

Once the temperatures, which depend only on time and radial position, are obtained by the finite difference procedures, stresses are calculated. The pressure stresses are calculated using the Lame solution for thick cylinders, and the elastic 1D thermal stresses are evaluated by numerical integration of the temperature field.

The simplified 1D temperature and elastic stress solutions for the idealized smooth cylinder are then modified to account for the specific geometric details of the numerous bore holes in the header body. As described earlier, the 1D temperature and stress solutions are modified by using closed-form empirical equations that were developed using the results of detailed 3D finite element analyses. These empirical relationships separately address the effects of the geometric discontinuities, the additional heat transfer surfaces of the bore holes. and the potential (in outlet headers) for the steam temperature in the bore holes to be significantly different than that in the main cavity of the header. This latter effect is important since there is generally a significant variation in outlet leg temperatures across the boiler.

As mentioned earlier, the time and cycle fractions crack initiation model requires an accurate estimate of the stress- time history at the local point of interest. The direct use of an elastically calculated stress-time history to predict creep damage would result in grossly inaccurate damage estimates. To obtain more realistic damage estimates, the elastically calculated stress-time history is modified to reflect the inelastic deformation characteristics of the material prior to evaluating the creep-fatigue damage. Pure strain control is assumed as the basis for modifying the stress-time history to account for the effects of both plasticity and creep.

In the case of header seam welds away from ligaments, the 1D temperature and stress solutions are adequate without empirical modifications. In the case of header girth welds, stresses due to nozzle and support loads can be important and are not calculated by the BLESS Code. In piping systems, the restraint of thermal expansion stresses generally dominate the failure of girth welds. The BLESS Code does not perform a piping analysis to determine the system stresses. Evaluation of restraint of thermal expansion (and deadweight) bending moments in piping must be made outside of the BLESS Code, and serve to define an equivalent load controlled membrane stress input to BLESS. The influence of creep on piping bending moments should be consideredt~l. Elastic follow-up in piping systems is capable of concentrating considerable creep strains at bends and other compliant locations. These calculations are not required to determine the growth of cracks in longitudinal seam welds.

ANALYTICAL LIFE ASSESSMENT

BLESS MODELING CONSIDERATIONS

The BLESS Code greatly facilitates the life assessment of elevated temperature headers and piping by eliminating the need for finite element thermal and stress analyses and utilizing recent developments in nonlinear creep-fatigue crack growth. Finite element stress analysis is, however, required for piping for evaluation of the forces and bending moments that result from the restraint of thermal expansion. Used in conjunction with its preprocessor, INBLESS, the code requires temperature and pressure history, material type, and geometry inputs to evaluate lifetime. The evaluation includes both crack initiation and crack growth. Provisions are also made for using inspection results to define current conditions, and then evaluating remaining life based on these current conditions. The code can be run in either a deterministic or probabilistic mode, and a set of &fault material properties is provided for 1 ‘/Jr- f;Mo and 21/,,Cr-1Mo base metal, weld metal, and heat-affected zone (HAZ)W

The estimated remaining life is calculated by BLESS either as a single value (when run in the deterministic mode) or a statistical distribution. This distribution is obtained when BLESS is run in the probabilistic mode and defines the probability of failure as a function of time. Such infotma- tion can be useful in making run/repair/replace and reinspection decisions for aging or cracked headers and piping.

The BLESS Code permits the evaluation of the effects of extremely detailed thermal and mechanical load histories on headers with very complicated geometric details. The INBLESS preprocessor was developed to simplify the specification of the geometry details and the loading history. The INBLESS preprocessor permits the description of a complex header geometry using relatively few geometry parameters. Once a header is described, any of the ligaments may be selected for evaluation. The INBLESS preprocessor also assists in the description of complicated operating histories in the familiar terms of steam tempera-

Babcock 8 wllcox 23

tures and pressures and flow rates. INBLESS guides the definition of a collection of individual operational procedures, such as, start-ups, shutdowns, load changes and periods of steady operation. The BLESS analysis module then assists in the definition of an operating history by linking together individual operational procedures in a user-prescribed sequence. The effects of different operating scenarios are then quite easily evaluated.

A header may be penetrated by several hundred tubes. The tubes are generally in several well-defined tube rows, with each row penetrating the header at a different circumferential location, as was shown in Figure 32. The header material between bore holes, of adjacent tube rows, is defined as a circumferential ligament. Note that a circumferential ligament supports the axial stress. The circumferential penetration pattern is repeated along the longitudinal axis of the header. The centerline-to-centerline distance between tubes, in the axial direction, is defined as the axial pitch. The header material between bore holes of adjacent tubes, within the same tube row, is defined as an axial ligament. An axial ligament supports the circumferential, or hoop, stress as illustrated in Figure 32. As previously discussed, the closed-form empirical equations that were developed to modify the 1D temperature and stress solutions were based on the results of 3D finite element analyses of header slices as was illustrated in Figure 34. The BLESS Code thus assumes that the geometry and boundary conditions are constant along the length of the header. However, the BLESS Code can be used to separately model different axial positions. The basic geometry parameters, required for the analysis of ligament cracking, are shown in Figures 35 and 36.

ION

-BORE HOLE DIAMETER

HEADER OD HEADER ID AXIAL PITCH TUBE OD TUBE ID BORE HOLE DIAMETER

\ 1

‘\I +

Operating data is input as the time histories of the steam temperatures, pressures and flow rates. These time histories are input as discrete time points, as illustrated in Figure 37 for a typical time history of the steam temperature during a start-up. In addition to specifying the time history of the bulk steam temperature, it is also possible to specify the time history representative of the steam temperature of a typical outlet tube leg. As discussed earlier, this feature is important because of the significant effect, on ligament damage, of outlet tube leg temperatures across a boiler.

Figure 35 Parameters required to define header ligament geometry.

EXAMPLES

Several example problems are provided to clearly demon- strate the effect of operational parameters on the crack growth rates of ligament cracks. The basis for these com- parisons is the typical P22 secondary superheater outlet header illustrated in Figure 38. In all cases, the results are reported for crack growth from the ID toward the OD within one of the two small circumferential ligaments shown in Figure 38. The BLESS representations of average base metal material properties are used.

Figure 39 illustrates the effect of temperature unbalance on the remaining life as a function of the initial crack depth. It is seen that a 50°F unbalance reduces the remaining life by a factor of approximately 3 for this case of steady operation with no cycling. The profound effect of on/off cycling

Figure 36 Parameters required to define circumferential penetration pattern.

24 Babcock 8 Wilcox

Figure 37 Steam temperature during start-up represented by 10 time points.

STUH R

STUB C

\ /

\\\ /

‘\ “\

‘--. ,A

------- _-’

Figure 38 Header cross-section for BLESS example

300,000

problems.

STEADY OPERATION ID-TO-OD CRACKING

2 1/4Cr-1 Mo

CT) 5 200,000

P

E J 150,000

f 100000 o! '

1050/l 100 0

105011075 <r --

1050/l 050 -x-t ~-

50,000

1 .oo 1.20 1.40 1.60 INITIAL CRACK DEPTH, INCHES

1.80 2.00

Figure 39 Effect of temperature unbalance during steady operation (no cycling).

(approximately an order of magnitude in this example) is demonstrated in Figure 40 for a case involving only 12 cycles per year. Note that the comparison of Figure 40 does not include any effect of temperature unbalance. Figure 41 illustrates the strong effect that temperature unbalance can have on the remaining life under on/off cyclic operation. Finally, Figure 42 demonstrates the expected effect that the bulk outlet steam temperature has on the remaining life. In this case, a 50°F difference in the outlet connection tem-

perature is considered. A 50°F temperature unbalance was included in the example involving 12 on/off cycles per year. Note that in all of the above cases the initial crack depths were quite significant, being from one to two inches in depth.

An excellent example of the probabilistic capabilities of the BLESS Code was provided in an earlier paper by H~sP'l.

Babcock & Wdcox 25

300,000

NO UNBALANCE ID-TO-OD CRACKING \

250,000 ~‘. j_ 2 ID-1 MO j<

j ‘..~ j\_\ v) ‘_ !zj 200,000

P j .~ Steady Operation -..

t ‘L

3 150,000 - -\ “,_

1

z 2 100000 oc '

50,000

105011050 STEADY d-

1050/l 05OF CYCLIC --ft-

I--- _~__.__ i -

01 1 .oo

12 Cycles/Year

I I I 1.20 1.40 1.60

INITIAL CRACK DEPTH, INCHES

‘y

f I

1.80 2.00

Figure 40 Effect of cycling on crack growth, 12 cycles/year, no unbalance.

r 50,000

40.000 . k

CYCLIC OPEFWTlON ID-TO-OD CRACKING

2 114Cr-1 MO

2 s I 30,000

e i

10,000

0 I I I I I I I I I

1.00 1.20 1.40 1.60 1.80 2.00


1050/l 100 CYCLIC &. *

1050/1075F CYCLIC A ”

1050/l 05OF CYCLIC b-

+

--\ _

50F Unbalance ---.

Figure 41 Effect of temperature unbalance, cyclic operation, 12 cycles/year.

26 Babcock & Wilcox

60,000

50,000

t/l 5 40,000

8

if 3 30,000

z Z a 3 20,000

10,000

0

‘, .,

CYCLIC OPERATION ID-TO-OD CRACKING

2 1/4Cr-1 MO

--. -- . .._-.

--..

105OF OuGt - --

-----i-___ -----+---,

I .oo I .20 I .40 I .60 I .80 2.00


Figure 42 Effect of outlet temperature on cracking, 12 cycles/year, 50°F unbalance.

A reality check is always a good idea when conducting remaining life analyses. One of the most relevant checks is to use the modeling technique to predict the damage incurred from the date of commissioning to the most recent inspection, or to predict the change in damage between successive inspections.

HEADER UPGRADES

The problems experienced in high temperature headers has led to upgrades and improvements which make the headers less susceptible to the problems of creep, thermal fatigue and creep fatigue. B&W offers design improvements for high temperature headers to minimize the areas where problems initiate. Additionally boiler owners are more likely to anticipate future changes in the operation of the boiler, such as cycling, so that these changes can be taken into account in the design.

As discussed previously, the standard tee nozzle design with its large saddle weld has been prone to cracking as a result of excessive loads placed on the outlet nozzle by the piping system. The large welded nozzle is eliminated by upgrading to a forged design in which the tee section has no large saddle weld (Figure 43).

The ligament area is a critical area in the header. Past designs had both radial and nonradial outlet stub arrange- ments which were relatively close spaced. The ligament area can be made more resistant to crack growth signifi-

cantly extending the header life by increasing outlet leg spacing. Upgraded header designs have wide spaced outlet

,Weld

Old Welded Design

Reinforced area

New Forged Design

Figure 43 Forged outlet nozzle replaces tee section weld.

Babcock 8 Wilcox 27

Present Upgrade

Lower Ligament Stress

Figure 44 Increased ligament size for increased life.

legs with the resulting larger ligaments (Figure 44). This reduces the localized hot spot on the header associated with specific outlet leg upset temperatures and reduces thermal stresses. The larger ligament is also more resistant to creep crack growth. In addition to wider spacing, an enhance- ment can be made to the design of the stub weld itself. Instead of a standard socket weld arrangement a full penetration tube-to-header weld will provide less stress concentration and eliminate the notch associated with lack of fusion in the stub design (Figure 45).

Large Lack of Chamfer Smaller Lack of Fusion Notch \ Fusion Notch

tw Present

Upgrade Lower Stress

Figure 45 Header upgrade with redesigned tube penetration.

When unique circumstances dictate special retrofits can be made to extend header life and delay the need for header replacement. An example is the B&W superheater thermal sleeve stub design which was specially engineered for a customer so added protection could be provided when changing out stubs (Figure 46).

The greatest increase in the service life of high temperature headers can be achieved by material substitution. More and more boiler owners are opting for upgrading header material to modified 9Cr-lMo-V (SA335 P91). The modified 9Cr-1Mo steel has considerably greater creep and rupture strength properties than conventional 21/4Cr-1Mo, offering the capability to dramatically increase header life. The ASME Boiler and Pressure Vessel Code allowable stresses for P91 are compared to the allowables for P22 in Figure 47. At a design temperature of lOOO”F, the allowable stress is 14.3 ksi for P91 and 8.0 ksi for P22. As a result of that difference in the allowable stresses, the required wall thickness of a P9 1 header is only about 55% of that required for a P22 header. At a design temperature of 1050”F, the

Thermal barrier

Superheater outlet leg

Figure 46 Retrofitted thermal sleeve reduces stress.

. .

0 0 0 .

0 . 0

.

0 5

SA335 P91 SA335 P22 0

. 0

0 I /

700 800 %I0 IO00 1100 TEMPERATURE, DEGREES F

Figure 47 Comparison of ASME allowable stresses for SA335 P91 and SA335 P22.

700 Rlnl 900 low II00 TEMPERATURE. DEGREES F

Figure 48 Comparison of yield strengths for SA335 P9 1 and SA335 P22.

28 Babcock 8 Wilcox

required wall thickness of a P91 header is only about 45% of that required for a P22 header. The much thinner wall thickness results in a similar reduction of transient thermal stresses due to steam temperature changes during start-ups, shutdowns and load changes. The significant reduction in transient thermal stresses results in much greater resistance to fatigue cracking. The yield strength of P91 is compared to that of P22 in Figure 48. At temperatures of 1000°F to 1100°F the yield strength of P91 is more than 50 percent greater than that of P22. The greater yield strength affords a further increase in the resistance to fatigue damage in P9 1 headers. This additional resistance to fatigue damage is due to the increased elastic stress range. That is, P91 can be subjected to larger thermal stresses before resulting in damaging plastic strains (i.e., exceeding the yield stress). However, far more important than the resistance to fatigue damage, is the increased resistance to creep damage in P91 headers. Even with the reduced wall thickness, a significant increase in the margin against creep damage is achieved with P91 headers. This increased margin is possible since the allowable stress for P91 is controlled by the ultimate tensile strength at temperatures up to, and including, 1050°F. while the allowable stress for P22 is controlled by the minimum creep rupture strength at temperatures of 900°F and above. The margins against creep damage would be comparable for the two alloys only if the allowable stresses

for both were controlled by the rupture strengths at the specified design temperature.

As the discussion above illustrates, high temperature headers have been an important component in condition assessment and life extension programs implemented by the electric utilities during the past decade. Most of the headers that have been given base line assessments remain in service today. Consequently the reinspection and assessment of these headers will continue to be an important aspect of a plant’s predictive maintenance program.

As explained, through research, material testing, finite element analyses and continuing inspection activity, the failure mechanisms in these high temperature headers have come to be well understood. Reinspection is essential to monitoring the initiation and growth of header cracking. With the development of analytical tools such as the BLESS program the boiler owner can obtain quantitative predictions for remaining life for these components. And finally, the lessons learned by the various manufacturers have led to improved designs to help mitigate the development of these header problems in replacement headers and in the new boilers built today.

References

1. Grunloh, H. J., Ryder, R. H., Gattuso, A., Bloom, J. M., Lee, D. R., Schultz, C. C., Sutherland, D. D., Harris, D. O., and Dedhia, D. D., 1992, An

of Boiler Pressure Parts. Volume 4: BJSS Code I . .

User s w J.ife at Gtud&&$g Report to Electric Power Research Institute on Research Project 2253-10, Palo Alto, CA.

2. Carden, A. E., McEvily, A. J., and Wells, C. H., “Fa- tigue at Elevated Temperatures,” ASTM Special Technical Publication 520, The American Society for Testing and Materials, 1973.

3. Skelton, R. P. (editor), w 9 Applied Science Publishers, 1983.

4. “HighTemperature Creep-Fatigue,” edited by R. Ohtani, M. Ohnami and T. Inoue, -se Material

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30 Babcock 8 Wilcox

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