:II

'NUREGICR4731

EGG-2469 Volume I

June 1987

Residual Life Assessment of Major Light

Water Reactor Components-Overview

Volume i

Editors

V. N. Shah

P. E. MacDonald A.

ELIEGn Idaho

Work performed under

DOE Contract No. DE-AC07-761D01570

for theU.S. Nuclear

Regulatory Commission

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EGG-2469

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I

RESIDUAL LIFE ASSESSMENT OF

MAJOR LIGHT WATER REACTOR

COMPONENTS-OVERVIEW

VOLUME 1

V. N. Shah and R E. MacDonald, Editors

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Authors

R. L. Cloud

J. F. Cook

M. A. Daye

W. G. Hopkins

L. J. House

V. Malhotra

H. Mantle

W. R. Mikesell

G. R. Odette

R. 0. Ritchie

W. L. Server

V. N. Shah I

Published June 1987

EGtG Idaho, Inc.

Idaho Falls, Idaho 83415

Prepared for the

Division of Engineering Technology

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, D.C. 20555

Under DOE Contract No. DE-AC07-761D01570

FIN No. A6389

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CONTRIBUTORS

Several specialists, engineers, and scientists have contributed to the dil

names, affiliation, and contributions are:

Name Affiliation

V. N. Shah EG&G Idaho

P. E. MacDonald

NV. L. Server Robert L. Cloud Associates

R. L. Cloud

NV. R. Mikesell

G. R. Odette

R. 0. Ritchie

S. A. Traiforos

WV. G. Hopkins

M. A. Daye

P. B. Lindsay

H. Mantle

V. MaIhotra

C. E. Jaske

L. J. House

J. F. Cook

Robert L. Cloud Associates

Robert L. Cloud Associates

University of California, Santa Barbara

University of California, Berkeley

Bechtel National

Bechtel National

Bechtel National

Bechtel National

Bechtel National

Bechtel National

Battelle Columbus

Battelle Columbus

EG&G Idaho

fferent sections of this report. Their

Contribution

Ranking of major PWR and BWR

components

PWR RPV, BWR RPV, PWR

primary coolant piping

PWR primary coolant piping

BWR RPV

RPV radiation damage

RPV fatigue

Task Leader (Bechtel)

RPV supports

PWR containment structures

Recirculation piping, steam

generator

Recirculation piping

Steam generator

Degradation mechanisms, life

assessment methods, emerging

NDE methods

Standard and emerging NDE

methods

ASME Section Xl NDE activities

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environment and cyclic loading. Likely degradation mechanisms include hydrogen embrittlement

of the anchorages; corrosion of the liner, tendons,;

reinforcing steel, and metal shells, including possible microbiologically induced corrosion; and

chemical reactions in the concrete. The potential

failure modes are loss of prestress in the tendons,

leakage of radioactive material caused by linerconcrete interactions, and loss of structural integrity, mainly because of corrosion: The ISI includes

a tendon surveillance program, integrated leak rate

tests, and visual inspection of surfaces. Major

unresolved technical issues are a lack of aging

related data for reinforced'concrete and tendons,

the need for improved inspection programs to identify and quantify degradation, and a better understanding of the potential impact of concrete-liner

interactions in aged containments. A comprehensive and standardized 1SI program is needed to

identify and quantify degradation in reinforced

concrete. Integration and evaluation of the available research results and information from older

facilities on the degradation of reinforced concrete

subjected to long-term exposure to elevated temperatures, radiation, and cyclic loading will support

the development of an appropriate 1SI program.

PWR Reactor Coolant Piping

The key degradation sites for PWR reactor coolant piping are the main coolant pipe nozzles, dissimilar metal welds, and cast stainless steel

components. The reactor coolant piping is sub-,

jected to thermal and pressure loading caused by

system operating transients. The major degradation mechanisms are low-cycle fatigue and thermal

aging. The potential mode of failure is through wall

leakage. IS methods include surface and volumetric inspection. Two unresolved technical issues are a

lack of accurate accounting of operating transients,

and the development of a better understanding and

assessment of the embrittlement of cast stainless

steels because of thermal aging.

PWR Steam Generators

The key degradation sites in the recirculating

steam generators are the inside tube surfaces at the

U-bends, tube sheet, and tube supports; the outside surfaces at the tube-to-tube sheet 'crevice

region; and ihe girth welds in the upper shell

region. The corresponding sites for once-through

steam generators are the outside surface of the

tubes in the upper-tube sheet region, and the welds

attaching the thermal sleeves to the auxiliary feedwater inlets. The major degradation mechanisms

include intergranular stress corrosion cracking

'(IGSCC), intergranular attack (IGA), pitting,

wastage, denting, fretting, and thermal fatigue.

The likely modes of failure are cracking, localized

or uniform tube thinning, and wear out'of tube

.material, which will ultimately lead to leakage of

primary coolant to the secondary system and possibly to the outside environment. 1SI methods

include eddy-current testing and leak detection

methods. Some of the steps taken to mitigate the

effects of the various degradation mechanisms are

using AVT chemistry on the secondary side, shot

peening the U-bend and roll-transition regions of

the tubes, using thermally treated Inconel 600 tube

material and a quatrefoil design for tube support

plate, and using 12% chromium ferritic stainless

steel as tube support material. Additional work is

needed to understand and model the corrosion

mechanisms, and to, better monitor the status of

steam generator tube degradation.

.Reactor Pressure Vessel Supports

The potential degradation sites for the neutron

shield tanks and column supports are at the core horizontal midplane elevations, and for the cantilever supports they are in the active height of the core. These

*RPV supports are subjected to neutron irradiation,

tensile stresses, operating temperatures, and a corrosive

environment because of the presence of water. The

major degradation mechanisms include neutron

embrittlement, corrosion, and radiation damage to the

lubricant used in the sliding foot assembly: The potential failure mode of the neutron tank, cantilever, or

column supports is brittle fracture. Two remaining

types of supports, i.e, the skirt and bracket types; are

not likely to undergo catastrophic brittle failure

because they are exposed to very little irradiation.

However, the skirt support is subjected to fatigue

because of the thermal- and pressure-induced expansions and contractions of the RPV during startup and

shutdown. Currently, there are no requirements for ISI

of RPV supports. Additional Fork is'needed to

develop a fracture toughness data base for RPV support steels irradiated at <2320C (4500F), determine

the range of radiation environment conditions for support structures, and investigate the effects of actual

radiation levels'on lubricants used in the sliding foot

assemblies and between RPV nozzles and supports.

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BWR Reactor Pressure Vessels

The key degradation sites for BWR RPVs are

nozzles, safe end welds, closure studs, and beltline

region. BWR vessels are subjected to mechanical

and thermal loads and neutron irradiation. The

major degradation mechanisms are low and high

cycle fatigue and neutron embrittlement. The

potential mode of failure is ductile overload leading

to a leakage. A surveillance program similar to the

PWR programs is required to assess irradiation

damage. ISI methods include volumetric inspection of weldments, studs, and threads. An unresolved technical issue is the need for close

monitoring of nozzle fatigue usage.

BWR Recirculation Piping

The key degradation sites for BWR recirculation

piping are the dissimilar metal welds at the safe

ends, the cast austenitic stainless steel components,

and the crevices at the shaft sleeves. The BWR recirculation piping is subjected to cyclic tensile

stresses, an oxygen environment, high temperatures, and has sensitized heat-affected regions. The

major degradation mechanisms are IGSCC, thermal fatigue, thermal embrittlement, and crevice

corrosion. The potential mode of failure is a leakage through a crack in the piping. ISI methods

include ultrasonic examination and use of

moisture-sensitive tape. Two unresolved technical

issues are the lack of an accurate accounting of

operating 'transients, and the development of

improved assessments of embrittlement because of

thermal aging. Additional work is also needed to

develop a better understanding of the effects of the

hydrogen added to the recirculation piping loop to

reduce the oxygen level in the coolant.

Current In-service Inspection

Methods

Many of the standard NDE methods employed to

satisfy ISI requirements were developed for the detection and qualitative assessment of fabrication-related

defects. These methods are not entirely adequate for

residual life assessment. Inspections for life assessment

generally require greater detection reliability and a

more quantitative determination of defects and accumulated damage than traditional ISI. The ISI methods

generally practiced by the nuclear industry are visual

examination, 'penetrant and magnetic particle testing,

x-ray radiography, eddy-current testing, ultrasonic testing, and occasionally acoustic emission monitoring.

Visual examination is the most widely used NDE

method. In most cases, it provides an indication

that damage may have occurred but cannot directly

quantify the amount of material damage. Penetrant and magnetic particle testing are used to

improve the visibility of surface-connected flaws,

and are therefore an extension of the visual examination rhethod. X-ray radiography measures density variations that may be due to cracks,

inclusions, porosity, voids,' lack of bonding, and

dimensional changes. The application of x-ray

radiography is restricted because of the slow rate of

examination. The use of single- and multifrequency

eddy-current testing is generally limited to inspection of near-surface cracks in simple geometries

such as PWR steam generator tubes. However, the

standard eddy-current methods are not adequate to

detect and characterize circumferential flaws

caused by intergranular stress corrosion attack.

Ultrasonic testing has been accepted as the most

useful volumetric examination method, especially

for inspection of welds and adjacent base material.

The limitations of standard ultrasonic test methods

are due to deficiencies in the available technology

and human factors, such as operator boredom.

Acoustic emission techniques are used to detect

growing flaws in pressure vessels. The main advantage of the acoustic emission method is its use as a

precursor to impending failure (rather than a

method for flaw sizing required for residual assessment). Another useful application of acoustic

emission monitoring is leak detection in pressure

vessels. Among all the standard NDE methods, the

eddy-current and ultrasonic methods are the most

promising for making quantitative damage related

measurements needed for residual life assessment.

The ISI of major nuclear power plant components is

controlled by USNRC regulations and Section Xl of

the ASME code. Improvements in the ASME code

NDE methodology are being made to detect flaws in

piping, RPVs, containments, and steam generators.

The current ASME code methodology is not adequate

to assess the residual life of the major LNVR components. More efforts are needed to develop field-usable

NDE techniques and equipments. The major unresolved issues associated with current ISI methods are

(a) the need for quantitative sizing of flaws for use in

fracture-mechanics analyses, (b) the need for methods

to inspect cast stainless steel components, and (c) the

need for techniques to measure the degradation in

mechanical properties during long-term service exposure.

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Current Life Assessment

Methods

Life assessment techniques include testing of surveillance specimens, monitoring of operational

parameters, evaluating samples removed from

service-exposed components, and predicting damage accumulation processes. The data from the surveillance programs used to assess the irradiation

embrittlement of the RPV material show a good

correlation between the brittle-to-ductile transition

temperature shifts at 30 ft-lb (41 J) and the tensile

yield strength. However, a correlation with a more

relevant material property such as fracture toughness of the RPV steel is not well established. Monitoring of the operational pressures and

temperatures may be used to determine the fatigue

usage factors at critical locations in the primary

loop. Metallographic and fractographic examinations and fatigue testing of samples removed from

service-exposed components have been used to

assess the structural integrity of LWR piping systems. Analytical prediction methods are required to

estimate accumulation of damage during anticipated future operations. A fracture-mechanics

approach based on probabilistic life assessment

methodology can be used to estimate the reliability

of pressure vessels and piping. Two major unresolved issues are (a) lack of confidence in the adequacy of the current models for material damage

accumulation processes, such as irradiation

embrittlement and fatigue crack initiation and

growth under a spectrum of loadings, and (b) the

unavailability of archival mechanical property data

for comparison with the properties of the same

material after service exposure, so that the degree

of degradation can be assessed.

Emerging Methods for Inspection

and Life Assessment

Recent interest in LWR plant life extension is

encouraging the development of new methods for

inspection and life assessment. Inspection methods

are needed to accurately determine the size, shape,

location, orientation, and type of both surface and

internal flaws, (including microstructural phase

changes), so that fracture mechanics approaches

may be used for life assessment. Current NDE

methods are not capable of detecting timedependent changes in microstructural features,

such as changes in the ferrite phase in austeniticferrite stainless steels, and the precipitate phases in

stainless steel; or the radiation embrittlement of

RPV steels. Such changes can be detected by replication, extraction, indentation hardness testing,

x-ray diffraction, and electrochemical testing at

assessable locations. Another potential method is

the use of eddy-current to measure residual stress.

The synthetic aperture focusing technique

(SAFT) for ultrasonic testing (UT) has been developed to provide enhanced visual images of flaws

detected during inspection. While the initial results

of the SAFT-UT application in the nuclear industry

are encouraging, additional work is needed to fully

qualify this technique. Another emerging method,

a dc potential drop method, may be used to detect

surface cracks on the inside of a pipe. Computeraided techniques for eddy-current testing are also

being developed to detect flaws in austenitic stainless steel pipe.

The USNRC is sponsoring programs using

acoustic emission technology for on-line monitoring of crack growth in pressure boundaries, and

leak surveillance in LWVR systems. On-line monitoring appears to be a promising technique to detect

fatigue-crack growth in RPVs and stress-corrosion

crack growth in stainless steel piping. Initial evaluation results of the acoustic emission leak-detection

technique are encouraging, but the method requires

additional field validation.

The emerging methods for life assessment employ

miniature test samples that can be removed from a

component-with negligible damage to the component, on-line procedures for the calculation of damage,

and improved models of material degradation and

damage accumulation. Miniature specimen testinig

(MST) can provide a direct measure of the degree of

aging. Miniature specimens can be valuable for surveillance testing where only a limited amount of test material is available and where space available for material

irradiation is restricted. The main constraint to the

application of MST is that the specimen must be large

enough to be representative of the material from which

it has been removed. MST may be employed to measure stress-strain response, fracture toughness, and

fatigue cracking. MST also may be used to measure the

through-thickness properties of thin steel plate, and to

characterize the material properties near welds.

On-line damage and remaining life calculations

may be performed for some component locations

by monitoring key operating parameters on-line.

For example, this approach can be applied to calculate fatigue usage factors for .RPV nozzles during

startups, shutdowns, and major operating transients when the nozzle temperatures are monitored

during operation.

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Finally, effective engineering models are required for

the successful implementation of life assessment strategies. These models must provide a balance between

engineering sophistication and practical utility so that

good assessments can be made economically. Using

simplifying assumptions and the results of the emerging improved inspection and monitoring methods,

miniature specimen testing and on-line monitoring,

these models should be capable of predicting the

major, important features of material degradation.

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-ACKNOWLEDGMENTS

The editors and authors of this report acknowledge the important role that

J. P. Vora of the U.S. Nuclear Regulatory Commission has played in providing programmatic guidance, review, and encouragement for this effort.

- The editors and authors also sincerely appreciate the efforts of M. E. Lapides of

the Electric Power Research Institute and D. E. Hostetler, Vice Chairman of the

Nuclear Utility Plant Life Extension Steering Committee, in arranging an industry

review of this report. In particular, the critical review performed by the technical staff

of the Grove Engineering and the Multiple Dynamics Corporation and other consultants is acknowledged. The editors and authors also thank A. B. Johnson, Jr., and his

co-workers at the Pacific Northwest Laboratories and D. M. Eissenberg,

K. V. Cook, R.WV. McClung, and C. E. Pugh of the Oak Ridge National Laboratory for providing detailed and useful reviews of this report. All these reviews have

helped make this report more complete and accurate. *

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CO NTENTS

CONTRIBUTORS ............................ ............................ ii

ABSTRACT ......................................................... iii

EXECUTIVE SUMMARY ....................................................... iv

ACKNOWLEDGMENTS .. ix

1. INTRODUCTION . ...................................................... I

2. RANKING OF MAJOR PRESSURIZED WATER REACTOR COMPONENTS 3

3. PRESSURIZED WATER REACTOR PRESSURE VESSELS .12

4. PRESSURIZED WATER REACTOR CONTAINMENTS AND BASEMATS .32

5. PRESSURIZED WATER REACTOR COOLANT PIPING ...... ...................... 55

6. PRESSURIZED WATER REACTOR STEAM GENERATORS .66

7. REACTOR PRESSURE VESSEL SUPPORTS FOR PRESSURIZED WATER

REACTORS AND BOILING WATER REACTORS .79

8. RANKING OF MAJOR BOILING WATER REACTOR COMPONENTS . .95

9. BOILING WVATER REACTOR PRESSURE VESSELS. 100

10. BOILING WVATER REACTOR RECIRCULATION PIPING ............ .............. 107

11. NONDESTRUCTIVE EXAMINATION METHODS .............. .. ................. 113 r

12. ADEQUACY OF ASME CODE IN-SERVICE INSPECTION METHODOLOGY .. 122

13. CURRENT LIFE ASSESSMENT TECHNIQUES ............... .. .................. 129

14. NEW OR EMERGING METHODS FOR INSPECTION AND LIFE ASSESSMENT 143

15. SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS ...... .................. 167

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RESIDUAL LIFE ASSESSMENT OF MAJOR LIGHT

WATER REACTOR COMPONENTS-OVERVIEW

VOLUME 1 I

1. INTRODUCTION

The United States was one of the first nations to

use nuclear power to commercially generate electricity and, therefore, has some of the oldest operating commercial reactors. As U.S. light water

reactors have matured, problems associated with

time- or cyclic-dependent degradation (aging)

mechanisms such as stress corrosion, radiation

embrittlement, fatigue, and other effects haveoccurred and have raised questions about the continued safety and viability of nuclear plants and, in

particular, about the integrity of the primary coolant pressure boundary. These problems have

included cracked piping at boiling water reactors

(BWRs), steam generator degradation at pressurized water reactors (PWRs), defective valves and

relays and inadequate means -for detecting and

characterizing flaws.

At the same time, with a continually increasing

demand for electricity and limited new generating

capacity under construction, the U.S. electric utilities are motivated to keep their existing plants operating beyond the original design life at as high a

capacity factor as possible. The economics of plant

life extension are clearly favorable. Studies cosponsored by the U.S. Department of Energy (DOE)

and the Electric Power Research Institute (EPRI)

show that replacing any single nuclear plant component can easily be justified, if the life of the plant

can be extended for a number of years.-Extending

the life of a 1 ,000-MWV plant by 20 years is expected

to realize a net present'worth of about $1 billion.

Therefore, the potential problems of managing

aging in older plants and the resolution of technical

safety issues in consideration of the development of

appropriate life extension criteria have become a

major focus for the research sponsored by the U.S.

Nuclear Regulatory Commission (USNRC). This is -

reflected in the following Policy and Planning

Guidance (PPG NUREG-0885) provided to-the

NRC staff in 1986:

"Requests for an operating license renewal

are to be anticipated and will require

advanced planning and analysis. The

Commission intends to continue develop-

, ment of the policies and criteria to define

requirements for operating license extensions to help assure that industry's efforts

in this area are focused on the primary regulatory concerns."

An important part of the USNRC research effort

is the Nuclear Plant Aging Research (NPAR) Program that is being conducted at several national

laboratories, including the Idaho National Engineering Laboratory (INEL). One of the NPAR program tasks at the INEL is to develop the

appropriate technical criteria for the USNRC to

assess the residual life of the major LWR components and structures. These assessments will help

the USNRC resolve certain safety issues and

develop policies and guidelines for making operating plant license renewal decisions. Most of the

effort for this residual life assessment task is

focused on integrating, evaluating, and updating

the technical information relevant, to aging and

license renewal from current or completed NRC

and industry research programs. A five-step

-approach is being pursued to accomplish the residual life assessment task: (a) identify and prioritize

major components, (b) identify degradation sites,

mechanisms, stressors, potential failure modes,.

and evaluate current in-service inspection (ISI)

"methods, (c) assess current and advanced inspection, surveillance, and monitoring methods and

evaluate maintenance programs, (d) develop residual life assessment models, and (e) develop criteria

for-license renewal. This report addresses the progress made toward gaining a qualitative, understanding of the degradation mechanisms active in

several light water reactor components and it represents a completion of the first step and partial completion of the second step of. the approach

described above.

- Virtually all the major components and structures

within a nuclear plant complex must be evaluated in a

life-extension program. However, from the USNRC's

perspective of ensuring the health and safety of the

-public, the first step in the residual life assessment task

is to identify those major components that are critical

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to nuclear power plant safety during normal operation,

off-normal conditions, design-basis accidents such as a

hypothetical large-break, loss-of-coolant accident or

the design-basis earthquake, or a severe accident. This

report presents appropriate criteria for selecting the

components and structures of interest and then ranks

these components in order of importance. The next

step is to integrate the available technical information

so as to identify the degradation sites, stressors, degradation mechanisms, failure modes, and ISI methods

associated with each of these major components. This

report identifies the degradation sites and corresponding processes for seven major components mentioned

later in this chapter. The remaining components will be

addressed in a later edition of this work. Future work

will be concerned with developing recommendations

for quantitative residual life models'

A major factor affecting the life of the component is the amount of aging, i.e., degradation of

performance over time, in the materials from which

the component is fabricated. This is critical for

those components that have a relatively small original safety margin. The degree of aging is often

referred to as the damage state of the material.

Material that has not been degraded by service

would have zero damage, whereas a failed material

would have- I00%o damage. Knowledge of the current damage state of the material is essential to

assess the residual life because it establishes the initial conditions for the remaining life. ISIs are performed to measure the current state' of damage.

This report discusses the current nondestructive

examination (NDE) methods used to detect and

characterize defects in materials. This report also

discusses certain new or emerging examination

methods that are being successfully used in other

industries and that could be used to quantify damage states and measure material properties in

nuclear power plant components. With an adequate characterization of the damage state and a

quantitative model, an accurate assessment of the

residual life of selected major structural components will be possible.

In addressing this task on residual life assessment, major emphasis has been placed on integrating the available technical information on the

relevant degradation mechanisms and applicable

NDE methods. The sources for the technical information are technical reports from the USNRC and

EPRI, technical workshops and conferences,

American Society of Mechanical; Engineers

(ASME) Section XI meetings, technical seminars,

technical journals, technical publications in

Europe and Japan, technical discussions with

experts, and contributions from subcontractor specialists with expert knowledge of specific major

equipment or NDE activities. Several additional

sources of information will be added to this list.

One source that is being currently pursued is the

plans and results of certain ongoing research activities at the national laboratories and other research

organizations. Also, operating experience data,

i.e., licensing event reports and NPRDS, including

information on operating transients and NDE

activities taking place at nuclear power plants will

be used.

This report is organized into three major parts. The

first part, Chapters 2 through 7, discusses major PWR

components. The second part, Chapters 8 through 10,

discusses major BWR components. The third part,

Chapters 11 through 14, discusses NDE methods.

Chapter 2 presents the criteria used to select the

major PWR components for this residual life assessment task. These criteria are also used to select the

major BWR components. Chapter 2 ranks the selected

major components according to their relevance to

plant safety. Chapters 3 through 6 discuss the aging of

the four highest ranked PWR components: the reactor

pressure vessel, the containment, the reactor coolant

piping, and the steam generators, respectively. The discussion for each component includes a description of

the component, the stressors acting on the component,

the degradation sites and their rankings, degradation

mechanisms active in the component, potential failure

modes of the component during normal operation and

transient and accident conditions, and current ISI

methods used for the component. A prioritized list of

degradation sites and associated degradation mechanisms, potential failure modes, and ISI methods are

presented in a table at the end of each section. The

discussion for the other components follows a similar

sequence. Chapter 7 discusses the aging of the reactor

pressure vessel supports in both PWR and BWR

plants. The major BWR components are identified

and prioritized in Chapter 8. Chapters 9 and 10 discuss the aging of the BWR reactor pressure vessels and

recirculation piping, respectively. Chapter 11 discusses

the standard ISI methods. Chapter 12 discusses the

current NDE activities in the ASME Section Xl committees. Chapter 13 discusses the current life assessment methods. Chapter 14 evaluates the applicability

of advanced NDE and life assessment methods

employed in the fossil power plant and petrochemical

and aerospace industries as well as those being developed for LWR applications. Finally, Chapter 15

presents the conclusions and recommendations for further work to assess the residual life of selected LWR

components.

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2. RANKING OF MAJOR PRESSURIZED

WATER REACTOR COMPONENTS

V. N. Shah

I Residual life assessment of the major components is

of primary interest for continuing safe operation and

life extension of a nuclear power plant. This chapter

identifies the major components of interest in commercial pressurized water reactors (PWRs). In identifying

these components, it is assumed that the lifetime of the

sensors, controls, batteries, certain types of pumps,

valves, motors, etc. is much' shorter than 40 years and

these smaller, less expensive components will be main-.

tained, refurbished, and/or replaced frequently. Therefore, the residual life of these small components' will

not significantly affect the life extension of the plant

and is not addressed in this report.

The major components reported here are identified and prioritized according to their relevance to

plant safety'only. The industry-sponsored Surry-l

pilot plant project has also identified components

for plant life extension studies.I However, the identification and prioritization criteria used by industry pilot project were based on plant safety,

reliability, and cost.

2.1 Key PWR Components

The selection of the key components is based on

two main safety criteria: to contain the release of

fission products that may take place during an accident, and to maintain an acceptable'low'level of

radioactivity 'in the-containment'during normal

operation. There are several barriers to the release

of fission products. Two'of iheii barriers are associated with the fuel rods. Because these rods are

replaced several times''du'ring the lifetime of a reac5

tor, they are not-included in'the selected components. The 'remaining two'barriers are the pressure

boundary components and the' containmenit.

Therefore, the pressure boundary components,'-

containment, and their supporting 'structures are

included in-the list of selected major components.

The control 'rod driv&e mechanism (CRDM) is

included in the list because its failu'e'mrnay lead to an

anticipated transient 'without scrarm "(ATWS) or

reactivity insertion accident." The s~afety-related

cables and connectors are included beeause they

provide sigials related 6t plant co'nditini't6 ihe'sen'

sors that control the' operation 'of the safety systems. The failure of the cables' and connectors

because of aging may lead to common-cause failure

of a number of systems. The emergency diesel generator is included in the list because it is required

for plant safety during a station blackout event..

-The reactor pressure, vessel (RPV) internals are

included in this list because their failure may prevent control rod insertion or cause fuel failure. The

.RPV supports and biological shields maintain the

level of radiation in the containment at an acceptable level and provide support to the' RPV and primary coolant piping, and therefore they are

included in the list.

Major PWR Components

1. Reactor pressure vessel

2. Containment and basemat

3. Reactor coolant piping and safe ends

4. Steam generators

5. Reactor coolant'pump body

6. Pressurizer

7. Control rod drive mechanisms

8. Cables and connectors

9. Emergency diesel generators

10. RPV inte'rnals

11. RPV supports and biological shields.

The RPV is ranked Number I because it is the

most' critical major component 'as far as plan''

safety is concerned. It is one of the major components of a nuclear power plant for which there is no

redundant member. The catastrophic failure of a

pressure vessel could lead to a rapid core'meltdowvn,

generating high pressure and temperature loadings

for'which light water re'actor containments ar 'not'

designed. Therefore, the RPV is ranked higher than

Ithe c'ontainrnient. The' vessel is"subjected to high'

system pressures,' temperatures, neutron irradiation, and cyclic fatigue. The most likely degradation' sites 'are' the circumferential rand axial

'weldments in the beltline region of the vessel.' Other

possible degradation siies'of concern are the hot-'

leg' and cold-leg nozzles, the weldment at the lower

shell to bottom head junction, the vessel flanges'

'and studs, and'the instrument penetrations in the

bottom head. The'most important potential degradation mechanism' is neutron embrittlement'. The

factors affecting neutron embrittlement of RPV

steel are fast' neutron fluence, irradiation

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temperature, and the chemical composition of the

steel and weld metal (mainly Cu and Ni content).2 '

3 The fast neutron fluences and the chemical composition of the weld metal vary

circumferentially and through the thickness of the

RPV wall. The irradiation effects on the RPV steels

are well documented, required to be considered in

operation and test limits, and monitored by surveillance programs during service. A less critical degradation mechanism is fatigue. However, the damage

caused by fatigue is irreversible, while irradiation

damage of pressure vessel steels may be reversible.

The containment and basemat are ranked higher

than the remaining components because they are

the major barriers protecting the public during an

accident from released fission products. Under a

severe accident scenario, it is possible that the containment internal pressure can rise to a value that

may cause leakage through the containment wall.

LWR containment structures may be divided into

three types: prestressed concrete containments,

reinforced concrete containments, and steel

cylinder/steel sphere containments. The most

likely degradation site for a prestressed concrete

containment is the posttensioning system, consist-,

ing of tendons and anchors. The potential degradation mechanisms for the posttensioning system are

time-dependent relaxation of tendon material, and

hydrogen embrittlement and stress corrosion cracking of the heat treated anchor heads.4 Corrosionr of

the tendon material is also possible. The liner constitutes the weakest element in the prestressed and

reinforced containment structures during a design

basis accident (DBA) and severe accidents. 5 A

liner-concrete interaction may take place at locations of major concrete cracks and distortions and

introduce highly localized strains in the liner. This

strain concentration may cause localized rupture of

the liner that will lead to leakage from the containment at a relatively early time in a pressurization

transient. This leakage to the environment represents the most likely failure mode for the dry and

subatmospheric types of containment but not for

the dual containment, and it is more likely in the

reinforced structures than in the prestressed structures. Prestressed structures see relatively minor

cracking until nearly twice the design pressure, but

reinforced structures develop widespread cracking

at about 75%. of the design pressure. The relevant

degradation modes affecting the concrete material

properties are deterioration because of aggressive

environments and internal chemical reactions. The

aggressive environments include nuclear heating,

extensive cycling of wetting/drying and freezing/

thawing of the exterior walls, acid rains and sulfatebearing ground water, and carbonation. 6 '

7

Carbonation is the process by which the alkaline

calcium hydroxide present in the concrete is neutralized by carbon dioxide and atmospheric moisture.8 Cracked concrete can provide a direct path of

water through or near to the rebars that may lead to

corrosion of reinforcing steel and subsequently to

additional cracking and spalling of the concrete.

The internal chemical reactions include alkaliaggregate, cement-aggregate, and carbonateaggregate reactions. The most likely degradation

sites for steel cylinder/steel sphere containments

are the base metal and the weldments in the containment wall. The major degradation mechanism

is the corrosion of the steel walls.

The four components representing the pressure

retaining boundary are ranked right below the containment and vessel because they constitute the first

barrier to the release of the fission products following the failure of fuelrod cladding during an accident. These components are the reactor coolant

piping and safe ends, the steam generator, the reactor coolant pump body, and the pressurizer. The

failure of any of these four components may lead to

a large-break or small-break loss-of-coolant accident (LOCA) that may challenge the integrity of

the containment by imposing high thermal and

pressure loadings on it. The reactor coolant piping.

and safe ends are ranked third, i.e. the highest

among these four components, because the impact

of a failure is severe. A break of the reactor coolant

piping will impose significant mechanical loads on

the component supports and other structural members and may damage them. The weldments at the

safe ends have dissimilar metal welds, and are the

most likely locations for degradation. The branch

nozzles and the cast stainless steel components, i.e.

elbows and T-connections, etc., in the reactor coolant piping are also possible locations for degradation. The potential degradation mechanisms are

thermal fatigue and thermal embrittlement.

The steam generator is ranked next to the reactor

coolant piping because the rupture of its tubes will provide a passage for primary coolant to outside the containment by way of the relief valves. In addition, a

major steam generator tube rupture event is of serious

nature because of the rapid depressurization of the primary coolant system and the complications that result

in terminating the pressure transients that ensue. In

fact, relatively small tube rupture accidents have taken

place at several nuclear power plants, 9 and the steam

generators in the operating PWRs have been plagued

by failures because of mechanical and chemical

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problems. The majority of problems have been associated with the recirculating type units manufactured by

Westinghouse and Combustion Engineering. However,

the once-through units manufactured by Babcock &

Wilcox have also'experienced some operational difficulties. 1

I The locations of the degradation sites within -

a recirculating type steam generator include: the outside surfaces of the tubes near the support plate in the

region of the cold leg, I the inside surface of the Inconel

tubing in the region of the hot leg, 1 1 the divider plate,

the tube support plates, the feedater nozzle, the

J-tubes attached to the feedring,12 and the upper shell

to transition cone girth welds. The J-tubes are normally fabricated from carbon steel and'have been

observed to degrade through wall thinning and perforation by a corrosion-erosion mechanism. The J-tubes

in the recirculating type units manufactured by

Westinghouse have been replaced,'and therefore they'

probably will not have a significant impact on the life

extension of those steam generators. A potential degradation mechanism for the girth welds is -corrosion

fatigue. 13 The normal operating water level for the

steam generators is in the vicinity of those welds and

the normal oscillations of the water level may impose

cyclic thermal loads on the welds. Cracks have been

discovered in the girth welds of the Surry-2 and Indian

Point-3 steam generators.

The locations of the degradation'sites within a

once-through steam generator (OTSG) include the

secondary face of'the upper tube sheet and the

uppermost tube support plate in tube rows adjacent

to the inspection lane.lU Another likely degradation site is the weld at the auxiliary feedwater nozzle

thermal sleeve in the OTSG. The cracking of this

weld will increase the potential for thermal shock to

the OTSG shell. The potential degradation mechanisms for steam generator tubes are wastage, denting, intergranular stress corrosion cracking,

intergranular attack,' erosion corrosion, 'pitting,

and fretting. The use of all-volatile treatment of the

secondary coolant has significantly reduced the

damage due to wastage and denting in the recirculating type steam generators. 14 The potential degradation mechanism for the divider plate and

feedwater nozzle is thermal fatigue. '

The massive size'of the reactor coolant pump

casing provides primary protection to the pressure

retaining components within the reactor coolantsystem and containment building from missiles

generated by internal failure of'pump elements

(e.g. impeller parts, shafts, fasteners). Therefore, it

is ranked higher than the pressurizer. The reactor'

coolant pump casings in the PWR systems are thick

(8 in. or greater) and constructed of cast austenitic

steel. Because of the'casing size, the pumps manufactured by Westinghouse are welded by the electroslag process that introduces high residual

stresses in the heat-affected zones near the weldments.' 5 The pump is powered by a large electric

'motor (4000 hp and larger) to maintain the coolant

flow through the reactor vessel, and the pump casing is subjected to the high vibratory forces of the

.rotating pump impeller assembly.' The vibratory

forces, in 'combination with operating 'stresses,

fluid dynamic cyclic loadings, and the primary

coolant environment, may. contribute to'flaw

growth. The potential degradation mechanisms for

the pump casing are corrosion fatigue and thermal*

embrittlement.L6 The other possible locations for

degradation'sites are flange and seal housing boltings. Corrosion fatigue is the potential degradation

mechanism for the boltings.

The pressurizer is the last key component that

represents the'primary coolant pressure retaining

boundary. A' surge line connects the pressurizer to

the hot leg, while a spray line connects it to the cold

leg. The surge and spray nozzle connections-at the

pressurizer vessel are'subjected to cyclic temperature changes resulting from the transient conditions

in the reactor. The spray nozzles'are subjected to

larger temperature changes than the surge nozzle.

Thermal fatigue is the potential degradation m'echanism for these two nozzles. The'spray head, upper

shell barrel, and seismic lugs are the other possible

locations for degradation. The potential degradation mechanism for the spray head is erosion, and

for the shell barrel and seismic lugs it is fatigue.

Another potential degradation mechanism for the

,spray head is thermal embrittlement, if it is made of

cast stainless steel.

The next three components play critical roles in mitigating LWR operational transients and accidents.' The

control rod drive mechanism is ranked highest among

these three components because its failure may lead to

an ATWS, or'a reactivity initiated accident (RIA), or a

small-break LOCA. The key subcomponents that may

degrade are the drive r6d assembly and the control rod

pressure vessel. Electrical failure caused by closed contacts in a switch may result in a control rod withdrawal

event leading to a RIA. Mechanical binding of the

latch mechanism may result in the failure of the scram

system and lead to an ATWS accident. The pressure

vessel houses the latch assembly and drive rod assembly, and provides a pressure barrier for'the CRDM

between the primary coolant and containment. Failure

of the control rod drive pressure vessel may result in a

rod ejection accident along with a small-break LOCA.

Eleven failures of pressure vessels and two failures of

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drive rod assemblies are reported in the data base that

covers the information from 31 Westinghousedesigned, operating plants up to and including 1983.17

The potential degradation mechanism for those pressure vessels having cast stainless steel latch housings, is

thermal embrittlement. I Further research is required to

determine the impact of the thermal embrittlement on

the residual life of the CRDM. Wlear is the most likely

degradation mechanism for the drive rod assembly.

Safety-related cables and connectors are active during normal operation of the reactor, while the emergency diesel generators are needed if offsite power is

lost. Therefore, the safety-related cables and connectors are ranked higher. The key degradation sites are the

cable insulation, especially in the higher temperature

regions of the containments, and the connectors.

Cable insulation becomes brittle because of thermal

aging, radiation, and oxidation, and may break during

a seismic event causing electrical shorts. Cables may

also fail by creep shortout, which occurs when a

mechanically-stressed cable is bent over a corner and

the conductors are pulled toward each other or toward

a grounded electrode through insulation softened by

aging.18 Certain types of connectors have inserts that

are made of carbon, which can become embrittled and

fracture because of aging. The Westinghouse-designed

reactor, San Onofre-l, which has been in operation

since January 1968, has experienced the problem of

embrittled connectors. 1I Moisture intrusion has

caused corrosion of some connector material.

The emergency diesel generators must provide

backup power to operate critical reactor safety

equipment in the event of a loss-of-offsite power to

the emergency bus bars. The failure of the diesel

generators to provide backup power will lead to a

station blackout transient. A station blackout transient followed, by the failure of the steam-driven

emergency feedwater pump to provide cooling

water to the steam generators and the absence of

operator intervention, is identified as a dominant

core melt sequence by the recent USNRC reactor

risk study. 19 The diesel generators have instrumentation and a control system that should start them

automatically on loss-of-offsite power, low reactor

water level, high contamination level, or other

emergency condition signals, and automatically

apply the required load. A review of the various

failure modes of the diesel generators indicates that

the governor in the instrumentation and control

system is the subcomponent most susceptible to

aging degradation. 20 The other subcomponents

susceptible to aging degradation are the-cooling

system pumps and piping, the fuel system injector

pumps and engine piping, and the turbocharger.

The potential degradation mechanisms are metal

fatigue and loosening of the fasteners.

The main function of the reactor internals is to

provide orientation and support for the reactor

core, and guide and protect the reactor control rod

assemblies. It also provides a passageway, support,

and protection for any in-vessel instrumentation.

One of the potential failures of the RPV internals,

i.e. baffle jetting, has led to fuel rod cladding degradation and disbursement of fuel into the coolant

in certain reactors. Failure of the RPV internals

also may relocate fuel away from the control rods

or prevent the control rods from inserting properly

and lead to an operational transient without scram.

The key RPV internal components susceptible to

aging degradation are the lower core plate, the

baffle-former assembly, the upper support column

bolts, the control rod guide tube sheaths and support pins, the thermal shield bolts, the core barrel

bolts, the in-core instrument nozzles, and the flux

thimble tubes. I Thermal shield and core barrel bolt

failures have occurred in a number of older

Westinghouse plants. Twenty-one out of 54 in-core

instrument nozzles in the Oconee-l RPV had been

cracked and broken off in the region of the weld.

The loose broken pieces of the in-core instrument

nozzles subsequently caused extensive damage to

the tube ends and to the tube-sheet welds in the

upper head of one of the two steam generators. 21 It

should be noted that some of the past RPV internal

problems were caused or aggravated by original

design errors. The potential degradation mechanisms are neutron embrittlement, stress-corrosion

cracking, mechanical I wear, low- and high-cycle

fatigue, and stress relaxation. The low-cycle fatigue

is caused by the loads due to changes in power levels, vessel inlet and outlet temperature differences,

and coolant pressures and flow rates. The highcycle fatigue is a result of the flow-induced vibrations. The control rod guide tube sheath and

support pins, and flux thimble tubes are locations

that may experience significant mechanical wear.

The last major components are the RPV supports and biological shields. There are several different designs of RPV supports and biological

shields employed in the commercial nuclear power

plants. One designed by Stone & Webster includes a

sliding foot assembly and a neutron tank. The sliding foot assembly supports both the vertical and

tangential loads of the RPV and its permanently

lubricated foot slides radially to allow free thermal

expansion of the RPV. The neutron tank supports

the reactor and, in addition, stores water in the

annular space between the RPV and the tank wall

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and provides shielding against neutrons. The sliding foot assembly is made of a maraging steel and

coated with Heresite, an air-drying phenolic coating. The potential degradation mechanism for the

sliding foot assembly is the stress-corrosion cracking of its threaded parts. 1,9 The potential degradation mechanisms for the neutron tank are neutron

embrittlement and corrosion at the inside surface

of the tank walls.

2.2 Summary, Conclusions, and

Recommendations

The major PWR components are selected and

ranked such that the release of the fission products

that may take place during an accident is contained. The RPV has been identified as the most

important component in PWR plants. This means

that the failure of a reactor pressure vessel would

have a very significant impact on the safety of the

power plant. The other key components, according

to their ranking, are containment, reactor coolant

piping, steam generators, reactor coolant pump

body, pressurizer, control rod drive mechanisms,

cables and connectors, emergency diesel generator,

RPV interna,''ahd RPV supports. The insights

and results presented in this chapter are based on

past experience as reported in the available literature. These results will be updated as more data are

made available.

Table 2.1 summarizes the aging related information on the key PWR components discussed in this

section. It lists the components and gives reasons

for their ranking. Table 2A identifies the most

likely degradation site and other degradation sites

for each component. Table 2.1 also lists the most

likely degradation mechanism and other potential

degradation mechanisms.

The major components reported in this section

were selected and prioritized according to their relevance to plant safety only. These criteria are different than the ones used in the industry pilot study.

Therefore, the ranking of the major components

presented in this section is somewhat different than

the one in the pilot study. However, all major components identified in this section were also ranked

high in the industry pilot study. Detailed discussions of the possible degradation processes of

PWR reactor pressure vessels, PWR containments

and basemats, reactor coolant piping, steam generators, and RPV supports are presented in

Chapters 3 to 7, respectively.

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Table 2.1. Key PWR components for residual life assessment

Rank Component Reasons for Ranking

Degradation Sites

(most likely, others)

Degradation Mechanisms - (most likely, others)

I I. Reactor pressure

vessel

2. Containment

and basemat

3. Reactor coolant

piping and safe

ends

4. Steam generator

Severe safety impact

of catastrophic failure

of an embrittled

vessel

Public protection

during an accident

Severe safety impact

of a large break

Thbe rupture will

provide a passage from

the primary system

directly to the

environment for the

primary coolant,

relatively poor

operating experience

Weldments in the beltline

region, hot-leg and

cold-leg nozzles, weldment

at the lower shell to bottom

head junction, vessel flanges

and studs, instrument

penetrations in lower head

Posttensioning system

(tendons and anchors),

concrete, metal liner,

reinforcing steel

Weldments at the safe

ends, branch nozzles,

cast stainless steel

components (elbows and

t-connections), complete

primary loop piping in new

Westinghouse plants

Inside tube surfaces

at U-bends and tube

sheet, outside surfaces

at tube-to-tube sheet

crevices, divider plate,

tube support plate,

feedwater nozzle, girth

weld

Neutron embrittlement,

corrosion fatigue, thermal

fatigue

Hydrogen embrittlement

and SCC of anchor heads,

corrosion and relaxation

of tendon material,

environmental degradation of

concrete, corrosion of

reinforcing steel and liner

I

Fatigue, thermal

embrittlement

I.

Wastage, denting, IGSCC,

pitting, fretting, IGA,

thermal fatigue, corrosion

fatigue

5. Reactor coolant

pump casing

6. Pressurizer

7. Control rod

drive

mechanism

Primary protection

from any internal

failure of pump

elements (impeller

parts, shafts), primary

coolant pressure

boundary

Primary coolant

pressure retaining

boundary

Failure may lead to a

reactivity initiated

accident, an operational

transient without scram.

or a loss-of-coolant

accident

Casing wall, flanges,

seal housing bolts

Surge and spray nozzles,

spray head, upper shell

barrel, seismic lugs

Drive rod assembly, control

rod pressure vessel

Corrosion fatigue, thermal

embrittlement

Thermal fatigue, erosion,

thermal embrittlement

Wear and thermal embrittlement

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Table 2.1. (continued)

Rank Component

8. Safetyrelated

cables and

connectors

9. Emergency

diesel

generator

10. Reactor

internals

11. RPV supports

Degradation Sites

Reasons for Ranking (most likely, others)

Active during normal Cable insulation, inserts

operation in mitigating in connectors

operational transients

and accidents

Needed to operate Governor in the

critical safety instrumentation and

equipment in the event control system, cooling

of a loss of offsite system pumps and piping,

power fuel injector pumps,

turbocharger

Failure may lead to Lower core plate,

dispersement of fuel baffle-former assembly,

into the coolant, or a upper support column

reactivity-initiated bolts, control rod bolts,

accident control rod guide tube

sheaths and support pins,

thermal shield bolts, in-core

instrument nozzles, flux

thinible tubes bolts, in-core

instrument nozzles, flux

thimble tubes

Failure will challenge Inside surface of the

the integrity of the support structure (neutron

primary coolant piping tank or column support)

and the small lines at the core horizontal

connected to the RPV. midplane level; lubricant

The neutron shield in sliding foot assembly

tanks maintain of neutron tank support

acceptable levels of

radioactivity in the

containment

Degradation Mechanisms

(most likely, others)

Thermal aging and creep of

insulation, thermal

embrittlement and

corrosion of connectors

Fatigue and vibrations

Neutron embrittlement, wear,

high cycle fatigue, SCC,

relaxation

rNeutron embrittlement,

corrosion, degradation of

lubricant

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L.

9

172
2.3 References

1. Seminar on Nuclear Power Plant Life Extension, Seminar Proceedings, cosponsored by EPRI, Northern States Power, U.S. DOE and Virginia Power, Alexandria, Virginia, August 25-27, 1986.

2. J. Ahlf, "Irradiation Experiments Related to Reactor Safety in Particular Pressure Vessel Steels,"

Nuclear Engineering and Design, 81, 1984, pp. 231-245.

3. U.S. Nuclear Regulatory Commission, "Effects of Residual Elements on Predicted Radiation Damage on Reactor Vessel Materials," Regulatory Guide 1.99, Revision 2, February 1986.

4. "Farley Tendon Problem Blamed on Water as NRC Mulls Generic Impact," Nucleonics J'eek, 26, 10,

March 7, 1985, pp. 1-2.

5. R. S. Dunham et al., Methodsfor UltimateLoadAnalysisof ConcreteContainment, EPRI NP-4046,

June 1985.

6. C. A. Negin et al., Extended Life Operation of Light Water Reactors: Economic and Technological

Review, Volume 2, EPRI NP-2418, June 1982.

7. D. J. Naus, "Concrete Material Systems in Nuclear Safety Related Structures-A Review of Factors

Relating to Their Durability, Degradation Detection and Evaluation, and Remedial Measures for

Areas of Distress," Appendix B in The Longevity of Nuclear Power Systems, EPRI NP-4208,

August 1985.

8. G. H. Lewis, Degradation of Building Materials Over a Lifespan of 30-100 Years, Volume 1, Commission of the European Communities-Nuclear Science and Technology, EUR 10020 EN, 1985.

9. V. L. Sailor et al., Summary of Barrier Degradation Events and Small Accidents in U.S. Commercial

Nuclear Power Plants, NUREG/CR-4067, BNL-NUREG-5 1842, March 1985.

10. H. R. Carter, M. T. Childerson, and T. E. Moksal, Model Tests ofa Once-Through Steam Generator

for Lane Blocker Assessment and THEDA Code Verification, EPRI-NP-3042, June 1983.

11. G. Taylor, "Maintenance Practices for Longer Plant Life," Nuclear News, 29, 3, March 1986, pp. 50-59.

12. J. D. Roarty, "Corrosion-Erosion of Steam Generator J-Tubes," International Meeting on Nuclear

Power Plant Maintenance, Salt Lake City, Utah, March 23-27, 1986.

13. C. J. Czajkowski, "Investigation of the Shell Cracking on the Steam Generators of Indian Point 3,"

Proceedings of the 9th International Congress on Metallic Corrosion, Toronto, 1984, pp. 229-234.

14. 0. S. Tatone and R. S. Pathania, "Steam Generator Tube Performance: Experience with WaterCooled Nuclear Power Reactors During 1982," Nuclear Safety, 26, 5, September-October 1985,

pp. 623-640.

15. D. T. Umino and A. K. Rao, Long-Term Inspection Requirements for PWR Pump Casings, EPRI

NP-3491, May 1984.

16. J. F. CopelandandA. J. Giannuzzi, Long-Term Integrity ofNuclear Power Plant Components, EPRI

NP-3673-LD, October 1984.

10
I

17. G. K. Roberts and G.'R. Ahdre,' A 'Component Wiear-Out Analysis'on Control Rod Drive Mechanism, Westinghouse Electric Corporation, July 1985.; -.

18. 0. M. Stuetzer, Correlation of Electrical Cable Failure with Material Degradation, Sandia National

Laboratories, NUREGICR-4548, March 1986.

19. U.S. Nuclear Regulatory Commission, ReactorRiskl"ReferenceDocument, Volume1, NUREG-1150,

February 1987.'

20. J. V.' Vause, D. A. Dingee, and J. F. Nesbitt, Aging of Nuclear Station Diesel Generators: Evaluation of Operating and Expert Experience, to be published by the Pacific Northwest Laboratories for

the U.S. Nuclear Regulatory Commission.

21. H. W. Bertini, Descriptions of Selected Accidents that Have Occurred at Nuclear Reactor Facilities,

ORNL/NSIC-176, Ap'ril 1980.

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3. PRESSURIZED WATER REACTOR PRESSURE VESSELS

W L. Server, G. R. Odette, and R. 0. Ritchie

3.1 Description

The design of pressure vessels for commercial

pressurized water reactors (PWRs) is very similar for

the different domestic nuclear steam supply system

vendors: Westinghouse (W), Combustion Engineering, Inc. (CE), and Babcock & Wilcox Co.

(B&W). Both CE and B&W manufactured their own

vessels, while W procured vessels from CE, B&NV,

Chicago Bridge and Iron (CB&I), or Rotterdam

Dockyard Co. The vessels for Prairie Island, Units 1

and 2, were manufactured by Societe des Forges at

Ateliers du Cresot. Also, at least one B&W vessel

was completed by Rotterdam Dockyard Co. for W,

and this vessel is part of the Surry-l plant that is

being studied in detail for life extension under Electric Power Research Institute (EPRI)/Department

of Energy (DOE) funding. Vessels were designed

and built in accordance with the applicable

American Society of Mechanical Engineers (ASME)

Boiler and Pressure Vessel Code, Section IIII at the

time of fabrication, except for the earliest vessels

that were constructed before Section III existed.

These earlier vessels were built to ASME Boiler and

Pressure Vessel Code, Sections I or VIII. Different

versions or updates to the Code apply to different

vessels depending upon the vintage. The structural

integrity of the Sections I and VIII vessels has been

judged to be comparable to Section III vessels of

approximately the same era.2 Section III requires

more limiting nondestructive examination of welds

so that the probability of manufacturing defects is

small.

The CE, early B&W, and CB&I production vessels used steel plates formed and welded to produce

the vessel structure. The earliest commercial reactor was Yankee-Rowe (operational in 1961), which

had a W vessel made by B&W of SA302B steel.

SA302B was used in several early vessels until

SA533B-I steel plate became the industry standard.

(Nickel-modified SA302B is essentially the same

composition as SA533B-1). Automatic submerged

arc welding was typically employed for both longitudinal and circumferential welds. The materials

consumed in this welding process are a manganesemolybdenum-nickel filler wire and a granulated

flux that minimizes atmospheric contamination

and provides ingredients to form a slag to remove

oxides during the welding process. The type of flux

material is important because the mechanical properties of the weldment can differ depending upon

what flux is used. B&W welds employed a Linde 80

flux that typically gives lower values of upper-shelf

Charpy V-notch properties than other fluxes. CE

and CB&I used Linde 0091, 1092, and 124 fluxes;

those. three fluxes produce similar mechanical

properties. The Linde flux designations are proprietary codes used by the Linde Division of Union

Carbide Corporation. Some shielded metal arc

(SMA) welding may be used for complex geometry

welds (nozzles), fit-up, backwelding, and repairs.

The SMA electrode is a wire coated with a bonded

flux (which provides a protective arc environment);

the welding current is automatically controlled, but

the arc voltage and welding speed are manually

controlled. An E8018 electrode is typically used for

SMA welds by all fabricators. Minimum preheat

and interpass temperatures (%3000F or 1491C) are

controlled, and intermediate postweld heat treatment may be employed. All vessel welds are

postweld heat treated to reduce residual stresses.

For later B&W vessels, an SA508-2 forging steel

was used for fabrication. The main advantage of

the forging is that the entire ring section is made of

one piece, eliminating the need for longitudinal

weld seams. The W vessels are only slightly modified versions of the CE and B&W vessel designs.

W has three slightly different versions of vessels,

depending on whether a two-, three-, or four-loop

plant was built; the number of loops dictate the

overall size (diameter) of the vessel as well as the

number of nozzles. B&W has three slightly different versions with the main difference being the

change in fabrication from plate to forging material. The CE vessels change slightly in terms of

diameter, but a more significant change is that the

newer System 80 design is larger to reduce neutron

exposure and has fewer welds.3 Additionally, the

nonsystem 80 CE-design vessels have instrumentation nozzles on the top head; all others (including

the System 80) have bottom-head instrumentation.

A typical reactor vessel section is shown in

Figure 3.1 to illustrate the important areas of concern. These regions will be discussed in more detail

later. All vessels are lined with a stainless steel cladding (usually Type 308 or 309 stainless steel) that is

applied in one or two layers by either multiple wire

or strip cladding SMA processes. 4 The cladding

12
I

, :' .,l ., SE->;

studs

flange

rVessel wall

thickness transition

II In-core - Instrumentation

pentrations

7.3042

Figure 3.1. Typical PWR vessel showing important degradation sites. U.-

13
I

thickness can vary, but is typically around 0.2 in.

(5 mm).

Many of the issues described in the next subsections are also directly applicable to the discussions

of Chapters 5 and 9 on primary system piping and

BWR RPVs, respectively.

3.2 Stressors

The radiation environment is a major stressor

because it causes a degradation in mechanical

properties. The other. primary stressors are

mechanical pressure loads during operation, periodic thermal transients that induce through-wall

bending moments, dead weight loads, residual

stresses, and other abnormal loadings that can

cause damage (such as pressurized thermal shock).

Other stressors that can be important are water

chemistry and mechanical contact that result in

wear.

The effects of neutrons from the nuclear reaction

process bombarding the inside wall of the reactor

vessel are a serious issue for maintaining the integrity of the vessel. As a result of irradiation exposure, the ferritic steel strength increases with an

attendant decrease in ductility. In terms of the

Charpy V-notch (CVN) energy properties, the

ductile-brittle transition temperature increases to

higher temperatures as shown in Figure 3.2 (below

the transition temperature, the material behaves in

a brittle manner and its fracture requires only little

energy, whereas above the transition temperature,

the material behaves in a ductile manner), and the

upper shelf level decreases in magnitude as the irradiation exposure increases (upper shelf level in the

CVN energy versus temperature curve represents

the fracture energy that corresponds to approximately 100°1o ductile shear and remains approximately a maximum for the upper range of test

temperatures). The fracture energy of the pressure

vessel steel thus changes continuously with radiation exposure at a given temperature. Therefore,

10 Code of Federal Regulations (CFR) 50 has

established a criterion that defines the transition

temperature as the temperature at the fracture

energy level of 30 ft-lb (41 J). Monitoring of the

condition of the reactor beltline region is conducted through the use of surveillance specimen

testing, which is covered later in Section 3.4.1.

Operating transients that occur with regularity

were used in the original ASME Section III fatigue

design analysis. Typical normal, test, and upset

transients are given in Table 3.1. The actual numbers of occurrences of these transients are usually

much less since conservative numbers were estimated for the design. Other transients that are

abnormal also are important since the fatigue

design did not consider their occurrence. These

events, such as a main-steam-line break or a smalll4reak loss-of-coolant accident, can create a potential challenge to the vessel integrity (but not

because of frequent occurrences).

Not only are many of the number of design transients unlikely to occur, but the magnitude of actual

events are not generally as severe as those designed

against. For example, plant operating procedures

generally limit the heatup and cooldown rates to a

I

I

1.

II

Temperature (OF)

-400 -200 0 200 400

C)

CD

C.)

0

Cd

0.

D

.L

160 >.

0)

C)

a)

80 0

c.

30 co

C ,

I

Temperature ("C) 7.3197

Figure 3.2. Effect of irradiation on the Charpy impact energy for a nuclear pressure vessel steel.

14
Table 3.1. Typical plant transients and assumed design occurrences5 6

Transient

Plant heatup at 100l F (38QC)/h

Plant cooldown at 1000F (380C)/h

Plant loading at 5 full . power/mi

Plant unloading at 5%yo full power/min

Step load increase of 10010 full power

Step load decrease of 10% full power

Reactor trip from full power

Loss of flow and abnormal loss of load

Loss-of-secondary pressure ' - :

Hydrotest to 3125 psig (21.55 MPa), 400'F (204'C)

Operating basis earthquake

Normal plant variation 1100 psi and 100F (-120C)]

Number of Events

500

'500

15,000

15,000

2,000

2,000

400

80,

5

10

200

>106

maximum of 50'F/h (281C/h), and actual practice

reduces the rate even lower. Therefore, the original

fatigue design is generally very conservative.

The primary transients leading to fatigue damage (in terms of the cumulative fatigue usage factor

used in Section III of the code) are plant heatup/!

cooldown, plant loading/unloading, reactor trips,

loss of flow, and abnormal loss of load. Other transients, such as inadvertent safety injections, can

also contribute to fatigue damage even though they

are not explicit design transients. If these transients

can be controlled and properly monitored to indicate the actual level'of loadings, estimates of the

actual fatigue usage can be obtained: This issue is

discussed further in Section 3.4.2.

The other issues of water chemistry and mechanical wear are less important. The water chemistry

issue pertains 'mainly to' the possibility of

corrosion-assisted fatigue-crack 'propagation if a

crack were to exist tha' penetrates the inner

cladding layer' and extended into the ferritic steel

vessel material. Fatigue-crack growth rates for ferritic steels in water environments at operating temperatures may be an order of magnitude higher

than those measured in air environments at the

same temperatures. 7 The mechanical wear problem

concerns surfaces that contact and periodically

move relative to each other. The action of the forces

between the surfaces creates potential wear of the

surfaces that can lead to eventual inoperability.

3.3 Degradation Sites

The sites of highest degradation depend on the

degradation mechanisms. Obviously, the sites of

degradation relative to radiation embrittlement are

in or near the reactor beltline region. Therefore, the

welds within that region become possibly the weakest link because the welds are more likely to contain

defects that can become cracks. Additionally, the

problem exists of higher copper (and nickel) content in many of the older vessel welds leading to

higher radiation damage sensitivity. The steel filler

material used in the fabrication of older submerged

arc welds was copper plated to enhance electrical

contact during welding and resistance to corrosion

during storage. The copper content in the welds

results directly from the, copper plating and the

residual copper in the consumable steel filler material. It is unfortunate that the effects of copper

(and nickel) were not recognized to be detrimental

to neutron exposure resistance earlier. Appendix G

to 10 CFR 50 defines the beltline region as including all shell material directly surrounding the effective height of the fuel element assemblies plus an

-additional volume of shell material both below and

above the active core, with a predicted reference

temperature shift (ARTNDT) at the end of service

life of 50°F (>280C). In some earlier CE vessels,

(especially considering license renewal) the

10 CFR 50-defined beltline region could even

include the nozzles, the nozzle welds, and the

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15

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I

nozzle shell. The base metals should not be ignored

because the copper content in older plates and forgings was not controlled to a minimum level; however, there appears to be less radiation

embrittlement in base materials as compared to

welds with the same copper/nickel concentrations.

Also it is possible that flaws can exist in the vessel

cladding which could potentially lead to flaws in

the ferritic vessel wall.

In the case of fatigue, most of the problems

involve geometric discontinuities or changes.

Informal review of CE-designedvesselsa reveals the

following order for the areas'of relatively higher

usage factors: closure head studs (inside bottom

surface), outlet nozzles (two per vessel), inlet nozzles (four per vessel),' and instrumentation nozzles

(penetrations).

The instrumentation nozzles for the older CE

design (at the top) did not have high fatigue usage

factors, but the System 80 design (at the bottom)

does. Other areas considered include the closure

head itself and the vessel flange, control rod drive

mechanism (CRDM) housings, vessel supports (at

the nozzles), vessel wall transitions, veisel'wall to

bottom head juncture, surveillance holders, core

stabilizing lugs, and a flow baffle. The CE design

for integrally welded surveillance holders is unique

because none'of the other designs allow attachment

to the vessel inner wall (cladding surface). The CE

surveillance holders are made of' a nickel-based

alloy welded to the stainless steel cladding. Core

stabilizing lugs are located on the bottom: shell as

well as core stop lugs. A flow baffle is located at the

bottom head 'shell course near the core stop lugs. A

vent tube through the top head also is present' but

does not present a fatigue problem. The support of

the vessels for all W and CE vessels is at the nozzles,

but the usage factors at these support areas are relatively low.

The W_ design is somewhat similar to CE design

and yields similar results.b The CRDM housing in

the W design also appears to have relatively high

fatigue usage. The CRDM nozzles in some cases

are made'of cast stainless steel that has the potential to degrade because of thermal aging. The same

areas of potential degradation turn up regardless of

whether a two-, three-, or four-loop plant (i.e.,

number of nozzles) is considered. Also, no difference between vessel fabricators is seen.

For the B&W design, a list of high usage factor

locations for a typical B&W system were obtained.c

Those regions for the reactor pressure vessel were

the closure studs, instrumentation nozzles, and the

core flood nozzle venturi sleeve. Interestingly, the

inlet and outlet nozzles were not included, nor were

the CRDM housings.

Therefore, in general, the regions of greatest concern for fatigue damage are:

1. Closure studs

2. Outlet nozzles

3. Inlet nozzles;'

4. Instrumentation nozzles (penetrations)

5. CRDM housings/penetrations (nozzles).

Comparing fatigue usage factor values for different

regions of a component or calculations between

different vendors can be misleading because different simplifying assumptions are made in their computations. The ASME Code, Section 111, I specifies

that the cumulative fatigue usage factor cannot

exceed a value of one. Therefore, if a value less than

unity is obtained by making simplifying assumptions to allow an easy calculation, there is no need

to continue. Refined calculations can significantly

reduce the usage' factor values reported. These

refinements include extensions from simple elastic

calculations to simplified' elastic-plastic calculations and to highly detailed fully plastic calculations. Other refinements in the way that transients

are grouped and counted can also make a significant change.

Corrosion fatigue can occur anywhere fatigue in

general is a consideration. Therefore, the sites indicated above are appropriate.

For wear problems, EW has identified thimble

tubes as an area of concern.8 Evidently, this concern stems from the fact that some French reactors

have experienced wear problems induced by vibration. Replacement appears to be an economically

viable option if and when problems arise, although

_W indicates that there is no significant effect on life

extension. The other vendors did not indicate any

problems relative to their equivalent of thimble

tubes. Thus, wear is not considered to be a significant problem.

a. Private communication with E. Siegel, Combustion Engineering, 1986.

b. Private communication with S. L. Abbott, Westinghouse

Electric Corporation. 1986.

c. Private communication with A. D. McKim, Babcock &

Wilcox Company, 1986.

16
I

3.4 Degradation Mechanisms

The primary degradation mechanisms covered in

this section are irradiation damage (embrittlement)

and fatigue (both classical design fatigue and crack

growth using fracture mechanics). The effects of

environmentally assisted fatigue growth also are

discussed. The list of potential degradation mechanisms can be very large, because many mechanisms

are always present but specific conditions suppress

their direct influence. General corrosion and stresscorrosion cracking are not a problem in PWR vessels because the water is kept very clean (and

essentially free of oxygen). Erosion and cavitation

are not a problem because flow velocities are within

acceptable limits. High-temperature creep is not a

problem because the temperatures in the vessel are,

well below one-half the melting point (on an absolute temperatures scale) of the materials. Thermal

aging of cast stainless steels is not considered a

problem (CRDM housings), because the

temperatures are not sufficiently high- to promote

large changes in properties. The following discussions reflect only those mechanisms that the

authors feel to be most significant.

3.4.1 Irradiation Embrittlement. Irradiationinduced increases in yield strength and decreases in

the fracture toughness of low alloy steels may limit

the operating life of PWR pressure vessels. The

major variables controlling such "irradiation

embrittlement" are the copper and nickel content

of the steels and the total neutron exposure or

fluence. Other variables believed to be important

include the irradiation temperature, neutron spectrum and flux, phosphorous content, thermomechanical history, concentrations of some other

impurity, and minor alloying elements. Indeed,

embrittlement may be controlled by the particular,

combination of these first- and second-order variables; this .combination greatly complicates the

problem of developing reliable embrittlement pre--

dictions based on a limited experimenital data base.'.

Embrittlement is normally monitored by testing

samples of base metal, heat-affected zone '(HAZ),

and weld in the formn of CVN and tensile specimens

that have been irradiated in surveillance capsules

located near or adjacent to a vessel wall. This represents an effort to come as close as possible t6 siniu'-

lating the actual conditions experienced by the,

vessel itself. Federal regulations 9 require that the

samples be'selected from the mosI limiting'material. There are usually several surveillance capsules

which are tested at specified intervals over the life

of a vessel;'the neutron flux in the surveillance cap-

'sules is greater than that in the actual vessel by a

lead factor that is intended to provide an 'early

warning of unacceptable levels of property degradation. The CVN specimens provide two measures

of embrittlement: changes in the' upper-shelf

energy (USE) and shifts in the transition tempera-

-ture at the 30 ft-lb (41 J) energy level,' AT. The

former is used as a signal of significant degradation

in the upper-shelf ductile toughness level based on

'a criterion of the USE falling below 50 ft-lb (68 J).

The latter is used to shift RTNDT, the temperature

used to reference a lower bound toughness curve,

i.e., AT = ARTNDT.

Excellent descriptions of most of the vendors'

surveillance programs are presented in

Reference' 10. Certain nuclear power plants do not

have any active program because of the structural

failure of capsule holders. One approach to dealing

with this situation is to use data from other plants,

including irradiation of extra capsules in selected

host reactors. The use of host reactors is an

approach used by B&W after the damage of all sur-

'veillance capsule holder tubes in their reactors in

operation in 1976.11 There are two categories of

-plants having surveillance programs of limitedapplicability because the materials in the program

do not adequately represent the limiting physical

properties and chemical compositions of the RPV

materials. In the first category of plants, the material in one representative utility surveillance program is actually part of the material from another

utility's RPV. In the second category of plants,'

there may be an undesirable weld material in a

number of reactor vessels that is reflected in the

surveillance program material of only one of the

vessels. The plants in the above two categories may

benefit by the industry sponsored programs that

provide the unified data from the pre- and postirradiation testing of the specimens.

In general, however, estimates of USE changes

and shifts in such circumstances are based on correlations (of the surveillance data base) that are contained in the USNRC Regulatory' Guide 1.99.12 A

second revision of this regulatory guide has been

proposed and iscurrently out 'for' public com-'

ment. 13 The differences between the two versions

of the Reguiatory Guide are- copper and 'nickel

concentrations are now taken into account for both'

weld and base metal, a distinct margin is used to

account for uncertainties in the data base, and the

damage change through the wall 'uses a displacements per atom (dpa) approach rather than taking

I

I

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r.LI

17
I

fluence for energies > I MeV. Regulatory

Guide 1.99, Revision 2, provides methods for calculating numerical estimates of changes in USE as

a function of copper and fluence, and shifts in transition temperature as a function of copper and

nickel content and fluence. Extrapolations of radiation damage into the vessel wall are based on a

prototypical fluence and spectral gradient and the

assumption that calculated dpa provides an appropriate estimate of spectrum weighted radiation

damage exposure. Limitations of the application of

the numerical estimates are specified for classes

and-grades of steels, temperatures, and composition ranges.

The consequences of the new regulatory guide

affect at least two areas: the pressurized'thermal

shock screening criteria and heatup/cooldown

curves. With regard to the pressurized thermal

shock issue, the higher magnitudes of the new

shifts could potentially lead to a reduced residual

life in some utility vessels depending upon their

exact chemistry and fluence conditions. The lower

attenuation through the vessel wall calculated using

Regulatory Guide 1.99, Revision 2, than that calculated using Revision I means that the heatup

curves can now become more limiting.

A major source of uncertainty in current correlations is the limited accuracy and variable range of

the surveillance data base. For example; there are

little or no data for high-fluence and low-flux combinations that may characterize the end-of-life or

life extension conditions of some vessels.

One method'of recovering CVN toughness and

tensile properties is to perform postirradiation heat

treating (annealing). This technique has been successfully applied to two vessels at temperatures less

than or equal to 650 0F (3430C). 1

3 For higher temperatures [up to 850'F (455 0

C)J, the feasibility of

this technique has been shown analytically 14 but it

requires a major engineering effort to actually perform the annealing and sophisticated calculations

of the postannealing reembrittlement rate. The

complexities in estimating irradiation' damage are

increased when the effects of an annealing are

included. 15 Adjustments in surveillance testing

also are required. Even though annealing may be

physically possible, the economics have only dictated two attempts: the Army SM-IA vessel in

1967 and the Belgian BR-3 vessel in 1984. 14 Both of

these successful annealings 'were conducted at a

lower temperature than is' needed for commercial

United States reactor pressure vessels. Therefore,

more work leading to a demonstration of a highertemperature annealing for a full-size vessel is

needed.

There have been a very large number of accelerated materials test reactor (MTR) embrittlement

experiments over the last several decades. It is now

generally accepted that because of the high flux levels in these studies, the MTR data should be used as

a supplement rather than as a direct substitute for

surveillance data. Most of this research has been

aimed at generating engineering data (e.g.,

CVN shifts) for establishing selected effects such as

the influence of copper content on shifts or the'

degree of recovery because of postirradiation

annealing treatments.

Sustained efforts to establish'the underlying

mechanisms of embrittlement have been initiated

only recently, primarily by groups in the

United States, Great Britain, and Germany. However, substantial improvements in our understanding of embrittlement have been achieved through

these efforts. 16 -

27 Indeed, this research has

resulted in significant improvements in the reliability of data correlations. For example, based on the

guidance of early theoretical models, correlations

of weld surveillance data were developed with

standard errors of only .1607o of the mean shift; 15

more recent analysis has improved the physical

basis for the models while maintaining a similar

correlation accuracy for an expanded surveillance

data base for weldments. The basic insights from

the recent studiesl7-19 can be summarized as follows:

I. Embrittlement is primarily caused by

irradiation-iniduced increases in the yield

strength, which can be quantitatively

related to CVN USE changes and temperature shifts.

2. Yield strength increases are because of

irradiation-induced, fine-scale microstructures (.l nm) that act as obstacles to dislocation motion.

3. The most likely candidate microstructures

are precipitates, vacancy clusters (microvoids), and interstitial clusters (dislocation

loops). Copper-rich precipitates, possibly

alloyed with either elements such as manganese and nickel or vacancies, are

believed to be a dominant hardening featu're in sensitive steels containing significant 'concentrations of this impurity

element. Other, possible types of

irradiation-induced or -enhanced precipitates are phosphides and small carbides.

l

18

I'
I

4. The major effects of, irradiation' are

enhanced diffusion rates' and defect clustering.

5. These microstructural evolutions are

kinetic phenomena and are functions of

flux and temperature, as well as composition and microstructure. In general, the

changes in the microstructure can be

understood from basic principles of alloy

thermodynamics and precipitation kinetics, coupled with rate theories'of radiation

damage. 17-20-

A number of detailed issues remain to be

resolved, including: the composition of copperrich phases and the balance between defect and precipitate microstructures, the mechanisms of nickel

enhancement of embrittlement, quantitative treatment of neutron flux effects, the role of factors

such as other elements and thermomechanical

treatment history, and the kinetics and mechanisms

of postirradiation anneal recovery and reembrittlement. Also, the effects of such variables as flux,

composition, 'and annealing' conditions on

recovery/reembrittlement need further resolution.

In principle, some other mechanisms may contribute to embrittlement in a synergistic,' additive,

or competitive fashion. These include: thermal

aging of ferritic materials caused by precipitation

(which is not irradiation enhanced), segregation

(temper embrittlement and strain aging), and

hydrogen embrittlement/hydrogen attack. For

example, there is some evidence that fine-scale copper precipitation can take place at vessel operating

temperatures over periods on the order of 25,000 to

250,000 h.17 Hydrogen effects could be promoted

by circumstances that may or may not be present in

PWR environments, 'related to 'nonequilibrium

partial pressures or local chemical reactions that

yield high local hydrogen fugacities.

There are'a number of'other unresolved questions. For example,' the accuracy of using CVNbased RTNDT shifts to reference toughness curves

(ASME Code' Section Ill, Appendix G) has not

been demonstrated: Theoretical consideratioiis'

suggest that in'some sensitive'steels this practice

may be nonconservative, and that relatively low.

cleavage fracture toughness levels may persist up to

elevated temperaiures in steels subject to very large

irradiation-induced yieldI strength inicreases.16 In'

other words, the actual RT1DT shifts in some pressure vessel steels may be significantly larger than

the CVN-based RTNDT shifts. Indeed, the serious-'

ness of this'issue may be mitigated only by the'

uncertain conseri'atisrri of using RTNDT-referenced

lower bound toughness curves. (It should be noted

that some work is being pursued within the Heavy

Section Steel Technology (HSST) Program spon-''

sored by NRC to actually measure shift and shape

change in the Kic and Kia curves from largespecimen fracture toughness tests).28 Further, no

reliable basis exists for characterizing irradiation

effects on ductile upper-shelf static fracture toughness and crack arrest toughness from surveillance

data. All these properties may be required in a comprehensive integrity analysis.

Finally, this section has not covered the topics of

defect monitoring and evaluation or application of

the material properties in fracture mechanics analyses, including the treatment of elastic-plastic

effects an'd multidimensional part-through flaws

subject to combinations of primary and secondary

stresses (which may be rapidly time varying). Information on the effects of irradiation on additional

properties of ferritic steel, such as fatigue crack

growth rates (including the effect of the environment), stress corrosion cracking, strain rate effects,

strain hardening exponents, and various measures

of ductility are generally required to carry out such

analysis. Indeed the effect of radiation damage on

the overall constitutive behavior and macroscopic/

microscopic strain distribution may be important.

It is noted that if large increases in yield strength

are correlated with large decreases in cleavage fracture toughness, a greatly expanded regime of brittle

vessel operation might be anticipated. 17

3.4.1.71 Uncertainties. The prior discussion

sets the context for uncertainties in characterizing

the material state as a function of service life and

,history for use in analysis of pressure vessel safety

margins. Summarizing, uncertainties are related

to:

1. The accuracy of either plant specific or

correlation based estimates of CVN transi-'

tion temperature shifts for specified values

of thedominant variables

2. The existing surveillance programs do not

include the effects of high-fluence and'

low-flux conditions, 'and treat the effects

of spectrum and irradiation temperature,

approximately.

3. The actual versus nominal value of significant variables in the most critical location

in the vessel

4. The questionable validity of using reference toughness curves and shifting these

1.

II

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19
I

curves by transition temperature increases

measured in CVN tests

5. The essential lack of information on a

number of properties, such as static uppershelf toughness (which may be needed in

an integrity analysis in general);' and an

even greater lack of information on the

effects of in-service irradiation on these

properties. In addition, there are uncertainties in the initial properties such as the

unirradiated RTNDT.

6. Gaps in basic understanding and detailed

empirical characterization of many aspects

of embrittlement, postirradiation annealing, and reembrittlement phenomena.

While this admittedly incomplete list of technical

uncertainties is imposing, it may not be all that significant. Currently, it is simply not possible to

assess the combined effect of the issues listed above

on the general reliability of the integrity assessments. However, a change in attitude about how to

treat the aggregated uncertainties,'or discovery of

new issues (e.g., an unanticipated accident

sequence) could change the assessment of residual

life far more drastically than'would be expected

from further resolution of some technical issues

listed above (e.g., the accumulation of additional

CVN surveillance data or the identification-of a

new second-order variable or embrittlement mechanism). Clearly, the magnitudes of the technical

uncertainties increase with the length'of the projected time frame; and may be paramount to the

issues of relicensing and life extension.

3.4.1.2 Practical Issues. There are a number

of practical issues that constrain the range of

approaches for resolving the uncertainties indicated above. Some of the technical issues are being

addressed by the USNRC research29 and others,

but it is not totally clear that all issues'are integrated

relative to life extension considerations. Some of

the pertinent practical technical issues include:

1. The limited range of potentially important

variables and properties contained in the

existing data base and the reliability of

subsets of that data base.' ' ''

2. The limited amounts of critical material

and limited numbers of surveillance capsuies. (For example, if plant life extension

is anticipated, how should this influence

the schedule and approach to testing surveillance capsules? Or, what can be done

after all the capsules are used up?)

3. The limited amount of property information that can be obtained from current

CVN-based surveillance programs relative

to the data needed for comprehensive

integrity assessments.

4. The flexibility (or inflexibility) of existing

regulation and ASME code requirements.

5: Questions regarding the suitability of the

use of MTR irradiations to simulate radiation damage under service conditions; and

the validity of using (prototypical) surrogate alloys to represent the behavior of vessels materials (e.g., if archival samples are

not available).

6. Lack of a pertinent data base to assess the

potential for a long-time, irradiationinduced aging effects in ferritic steel.

3.4.1.3 Issues to be Resolved. Clearly, the

most pressing need is the development and timely

execution of a comprehensive, integrated, and balanced plan for dealing with the array of issues associated with residual life assessment and life

extension of reactor pressure vessels. The NRC

Materials Branch has such a plan underway for the

reactor pressure vessel, i.e., the HSST Program.30

Timely execution of this plan will allow resolution

of several issues associated with aging of the reactor

pressure vessels. It is not' appropriate to attempt to

design such a plan here; however, it is useful to elaborate on the approach and to identify a few key

elements that can be expected to emerge from a

detailed study;

1. Comprehensive-After identifying the key

variables, develop a means of obtaining

and validating the necessary array of properties for alloys and conditions (i.e., the

irradiation, thermal and chemical environment) reasonably representative of the

realities of long-term in-service behavior.

2. Integrated-The experimental effort to

obtain appropriate properties in suitably

conditioned alloys should be closely tied to

the overall basis for integrity analysis. The'

integration with the mechanics component

of the integrity analysis should probably

anticipate significant advances in computational capabilities. On one hand, such

development of computational capabilities may permit the further exploitation of

I

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20

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I

miniature specimen testing and reconstituted specimen testing techniques '(see

Chapter 14 for additional discussion),

greatly extending the potential sources of

information about embrittlement; and, on

the other hand, these advances will permit

more direct and reliable application of

fundamental property data to practical

structural analysis. Significant improvements in the basic theoretical understanding of the mechanisms of embrittlement

and techniques to characterize the

embrittling microstructures (as well as

property changes) should also be anticipated.

3. Balanced-The efforts should be carefully

balanced, because the impact of a few

weak links in the sequence of analysis will

far outweigh high precision in one or

another step in the process. For example, it

would do little good to develop methods to

account for each and every neutron

impinging on a surveillance capsule or vessel if the irradiation temperature or alloy

microstructure and composition were still

poorly characterized.

Some of the important specific elements in this program are listed below (the HSST program has some

activities planned in many of these elements):

1. Rapidly resolve critical outstanding issues,

e.g., the potential for high-temperature

cleavage fracture in sensitive steels with

high levels of irradiation strengthening -

2. Comprehensively define the property array

needed to carry out reliable integrity analysis, including ranking of the overall sensitivity to uncertainties in the properties

3. Develop improved ways of extracting an

array of mechanical property information

from material in' surveillance-programs,

e.g. miniature specimen techniques

4. Validate available RT:NDT shift data from

accelerated MTR irradiations'as a' basis for

establishing long-term in-service 'behavior

for RPV steels. These data will supplement

the surveillance data from the power reactors

5. Hold additional surveillance specimens in

power reactor internals above'or below the

core as well as on the sides to get the appropriate flux and fluence levels, and temperatures

6. Initiate`sorme long-time studies of other

environmental effects (e.g. thermal aging),

and their potential synergistic interactions

with irradiation embrittlement

7. Resolve postirradiation annealing

recovery/reembrittlement kinetics and

mechanisms

8. Optimize surveillance programs with longterm objectives clearly defined

9. Continue efforts to develop fundamental

understanding of embrittlement and postirradiation annealing recovery/

reembrittlement phenomena in terms of

- physically based models and techniques to

characterize the state of the 'alloy (such as

small-angle neutron scattering).

I

3.4.2 Fatigue. The fatigue analysis of Class I

components in nuclear plants is 'defined by the

ASME Boiler and Pressure Vessel 'Code, specifically in Section Ill for the fatigue design evaluation

procedurel and in Section XI for the flaw acceptance standards and evaluation procedures. 31 The

Section Ill procedures are based on the classical

stress-strain/life (S/N) approach, using an experimentally determined relationship between the elastically calculated "stress range and fatigue life.

Under variable amplitude cyclic loading, a Miner's

rule approach of accounting for damage accumulation in the component at different stress ranges is

used by linearly summing the fraction'of life consumed at each stress range with the cumulative

usage factor set at less than, or equal to, unity. Section XI procedures, conversely, rely on fracture:

mechanics-based, damage-tolerant analyses,' where

continued operation of components containing

defects can be permitted in-service, provided predictions of fatigue-crack growth show that the

defect will remain subcritical until the next inspection. .

-From the perspective of improved life-prediction

; rocedures, the Section III nd Xl approaches may

be open to question. First, they represent fundamentally different'fatigue approache's (one based

primarily on crack initiation and the other based on

crack growth)' which'may cause problenii of consistency wheri applied to' the same'.component. 2

'Second, with regard to the S/N analysis, no allowance is made for the sequence of loading cycles,

'with'the Miner's 'linear damage' law, and further

separate analyses are required for high- and lowcycle fatigue. Third, the fatigue-crack-growth,

damage-tolerant analysis does not fully account for

'such factors as near-threshold behavior, variable

1.

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amplitude loading, and the anomalous behavior of

small flaws (typically of a size < I mm).33

Described 'next are various corisiderations that

may be used to improve the life-prediction capability of the Section 111 and Xl fatigue design analyses. (Note that the NRC Pressure Vessel Research

Program is conducting work on both environmental fatigue-crack-growth behavior and cumulative

fatigue-damage modeling.) Such advanced lifetime

computation procedures are designed for life extension and rely on:

I. A modified S/N procedure for crack initiation based on the local-strain

approach, 3 4 which incorporates the

effects of both stress concentrations at

notches and the sequence of variable

amplitude cycles

2. An updated version of the damagetolerant procedures that incorporate an

improved appreciation of environmental

factors, low growth rate behavior, and the

role of small flaws.

3.4.2. 1 Modified Life Prediction Approach

for Crack Initiation. The local strain-approach

represents an advanced version of the classical S/N

approach, and provides a more consistent

approach to fatigue life prediction, based primarily

on crack initiation. Specifically, it incorporates

improved analyses of the roles of mean stress and

stress concentrations and in the summation of

cumulative'damage. 3 4 The analysis is essentially

based on total strain, and so as such, does not

require any distinction between high- and low-cycle

fatigue. Typical S/N analyses generally define lowcycle fatigue (LCF) as failure in the'ran'ge of'

102-105 cycles' (depending upon the maiterial) at

stresses near or above yield, and high-cycle fatigue

(HCF) as failure at a high number of cycles under

nominally elastic stresses. The Coffin-Manson law

can be used for life prediction in LCF, and the Basquin relationship can be used in HCF.3 5

The first commonly used equation for the localstrain approach is the experimentally-determnined

cyclic stress-strain curve, which relates cyclic stress

range (Ao) to 'strain range (Ae) in terms of Young's

modulus (E) and the cyclic strain hardening exponent (n'), and is written

Ae AO /aa \In'

_2 = _2E + ~Vk') (3.1)

where k' is the cyclic-strength coefficient representing the slope of the log stress/log plastic strain

curve, and n' is typically 0. 15. Use of this flow rule

allows a far better estimate of actual material

behavior (e.g., cyclic hardening or softening). The

second commonly used equation is the basic stressstrain/life equation, which relates total strain range

(Ae) to life (Nr) by incorporating modified versions

of the Coffin-Manson law for LCF, the Basquin

relationship for HCF, and the Goodman relationship for mean stress (or) into a single equation that

gives 3 4 '

3 5

1.

Ae = ar_ a 2__ +Er_ 2

2 Eo + f) (3.2)

where af and ef are the fatigue-strength and fatigueductility. coefficients, respectively (approximately

equal to the true fracture strength and fracture

strain in a tensile test), and b and c are the fatiguestrength and fatigue ductility coefficients, respectively (with approximate values of -0.1 and -0.6). It

should be noted here that no fatigue limit is

assumed.

Equation (3.2) is also used for the prediction of

fatigue life in the presence of stress concentrations,

where the values of Aa and AE now define the local

stresses and strains at the notch root. Instead of

employing the classical approach of computing

these local stresses merely in terms of the theoretical stress concentration factor (K,) or the fatigue

notch strength reduction factor (Kr), Neuber's

rule,36 or equivalent numerical calculations, can be

used to compute the local strain history. For example, Neuber's rule gives

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Kr = (KOKE )I/ (Aa Ae E) AS (3.3)

where K0is the actual stress concentration (ratio of

local to nominal stress, Aav/AS) and K, is the actual

strain concentration (ratio of local to nominal total

strain, Ae/Ae). Such an approach can result in

much improved life prediction, particularly in

lower strength steels where local yielding at the

notch results in significant strain concentration as

well as stress concentration.

For irregular cyclic loading conditions, a modified Palmgren-Miner linear-cumulative-damage

rule can be employed to allow for consideration of

the sequence of variable amplitude loading cycles.

(Sequence effects, overstressing, and understressing are not accounted for in classical S/N

22
analyses.) The basic linear-cumulative-damage'la'w' 3.4.2.2 Modified Life Prediction Approach

can be stated as ' ' ''for Crack Propaigation. The flaw-evaluation

-procedures in Appendix' A of Section XI of the'

.' ASME Code3 ' use linear elastic fracture mechan-

= usage factor (3.4) ics to predict the growth of existing flaws under

i "'e-cyclic loading. For these procedures to be effective,

crack extension rates (da/dN) must be predicted on

where n1 is the number of cycles of a given stress the basis of a realistic crack growth 'relationship,

range (Aat), Nr, is the number of cycles to failure 'which generally has the form39

had all the cycles been of magnitude (au1), and the'

usage factor is set not to exceed unity. In the modi-' da ' ' F '- (35

fied approach, which has been shown to yield supe- dN = f (ak, Kmax* (

rior fatigue-life'estimates with'complex loading'

histories,37 the primary input is to account for' - where Kmax and AK'are the maximurn value and

mean stress and sequence effects by using a moret' range of he stress intensity factor ( - 'a

appropriate means for' cycle counting. There are F is the frequency, and C and m are'experimentally

several procedures that have been used with '' -"determined constants appropriate to thermaterial

success, but the so-'called rainflow method is gener- - the frequency, and C and m are experimentally

ally the most widely used.37 This method involves: 'determined constants appropriate to the' material

3 ' ' ' '' '- ' ' ' '''' intermediate cyclic crack growth rates and can' be

1. The determination'of the 'number of corm- used with altered (realistic) constants' for some porplete hysteresis loops in the loading spec '. 'tions outside the intermediate'crack growth range.

trum - .' ' ' ' ' .The approach of James and Jones40

2. Measurement of'a mean-and alternating stainless steels represents an ideal example. These

stress (or strain) for each loop authors employed an extensive data base for stain3. Computation of the value of Nf for each less steels in order to apply least squares regression

loop from Equation (3.2) '' techniques to derive a power-law, crack-growth

4. Calculation of the usage factor, using relationship [Equation (3.5)] for use in' life calculaEquation (3.4) in the usual manner. tions. However, because in this particular case no

'attempt was' made to allow for environmental

Because these procedures require a great deal of effects, the"equation can only be used for crack

tedious bookkeeping to implement, several'- 'growth behavior in air." '

algorithms have beerf published to reduce computa- The approach of integrating an appropriate.,

tion time.38 , crack growth'relationship 'from initial to final

Finally, as with all fatigue analyses, it is vital to defect size41 is' generally considered to be in'herconsider the effect of service environment,-as it is ently conservative because it assumes negligible ini-'

well known that fatigue life can be drastically tiation life. However, based on recent studies of

altered by the presence of an active environment variable-armiplitiide, near-th eshold, 'and small-

(such as PWR water). This generally entails defin- crack behavior (e'j., Refereii6es 33,' 42, 43), 'there

ing the life equation [Equation (3.2)] in simulated are several factors that should be considered addiservice environment with loading applied at a real- tionally to improve the accuracy of life prediction. [

istic frequency, 'if possible.'' ' ''; ' '' 'These factors are discussed later.'

Compared to classical S/N analyses,' the local-"' *

strain approach has 'two principal 'advantages. ' '; -

First, life predictions f6r a wide ivariety'of situa- '' "3.4.2.3 'ita Base' 'Establishmnent. Crack

tions can be made'from a limited amount'of small- ''pr6pagation' data should be generated experimenspecimen test data. Secorid,' it provides a rationial' 'tally in as realiitienvironments and frequencies as

basis for considering various material'belhaviotr 'possible, over 'i' ikide spectrum of growth rate's;.

related interaction 'effects that 'are -importanti in ' Sufficient data must be obtained to g'uaranitee con- 'i

metal fatigue."For' these ieasons,-the approach' 1 " fidenceintheres~Iltingeqiations. Dataintheultra-' ' '

yields superior life pOedictions compared with clas- low 'growth-rate regime below 10.6 mm/cycle,

sical analyses;, 'and'could be successfully used to 'approaching the so-called fatigue threshold, 'are

better define usage'factors for life-extension con-'' pairticularly important as'fatigue' lifetimes; espesiderations. ''. i!: ' " - ' cially those of practical importance to' life

23
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extension, can be dominated by crack extension in

this region. For example, using a linear extrapolation of higher growth rate data (above 10-5 mm/

cycle 4 l may lead to heavy penalties on predicted

life and indicate component retirement long before

its safe useful end-life is reached). The need for

ultralow growth rate data is further amplified by

estimates of reactor component service loads. For

example, service loads at the Zion-I pressurized

water reactor plant have been estimated to include

over 2 x 1010 cycles of vibrational stress cycles [of

magnitude typically 1000 psi (6.89 MPa)1, 3 2

which over the 40-year life of a particular component could lead to extensive crack growth at the

seemingly insignificant 'near-threshold growth

rates. Because near-threshold growth' rates can be

severely retarded by the phenomenon of crack closure42 in laboratory tests (e.g., using comjpact-type

specimens) at low load ratios, it is essential that

such data: be determined at high' load'. ratios

above 0.5 to generate a conservative'crack-growth

relationship (see also small crack discussed later)..

3.4.2.4 Variable Amplitude Behavior.

Unlike crack initiation under irregular'cyclic load-'

ing, as described by Miner's law [Equation (3.4)],

crack growth under variable amplitude loads can be

subjected to severe load interaction effects, particularly for long cracks under broad band spectrum

loading.4 4 Such interaction effects generally take

the form of transient' crack-growth retardations,

following, say, single tensile overload'cycles or

high-low block loading sequences, which contribute significantly to enhanced fatigue lifetime.45

Although physical explanations for suclibeliavior

are still mechanistically uncertain, the aerospace

industry (in particular) has been successful in devising several modified damage-tolerant analyses that

can account, (albeit empirically) for the increased

life caused by such transient behavior. The' most

widely used models are those of Wheeler 46 and

Willenborg et al.,4 7 although-nore sophisticated

approaches based on numerical modelin'g'ofcirack

closure have recently been developed.4 8 The

approach in all these models is essentially to. use

constant amplitude data to predict variable ampli-.

tude behavior through modification of the integra-

- tion of the crack growth relationship, e.g., through

determining a retardation factor on crack growth

while the crack is propagating within the: plastic

zone of the overload cycle. Incorporation of such

procedures in the life prediction analyses may, well

contribute to increased, life (which corresponds

essentially to decreased usage factors).

3.4.2.5 Small-Crack Effects. One further.

factor, which should be considered with advanced

damage-tolerant life prediction procedures, is an

appreciation of the possible anomalous behavior of.

small flaws. Cracks are considered small when they

are of a size comparable with the scale of the microstructure, or the scale of the local plastic zone size,

or when they are simply physically small (i.e.,

<I mm).33 In general, their behavior is to display

growth rates in excess of those of conventional long

cracks (e.g., over 10 mm in length), when exposed

to the same stress intensity range (although mild

reverse effects have been reported). In addition,

small crack growth can occur below the fatigue

threshold, below which long cracks are presumed

dormant. Moreover, under severe environmental

conditions, the accelerated growth behavior of

small cracks may be further enhanced. 49 Short of

generating experimental data on such small cracks,

damage-tolerant procedures to account for such

behavior have yet to be developed. However, since

mechanistically the behavior of small flaws is primarily influenced by a diminished role of crack closure, by generating long crack data at high load

ratios, some of the discrepancy between long- and

short-crack growth rates can be diminished.

In summary, it is felt that by incorporating

advanced S/N analyses to replace the typical

approaches outlined in Section III of the code, and

by incorporating revised damage-tolerant, crackgrowth procedures to supplement the Section XI

procedures, significantly improved life-prediction

capabilities can be obtained. Such approaches

could further be embellished by consideration of

the small-crack issue, which represents the interface between crack initiation and crack propagation analyses.

3.5 Potential Failure Modes

The eventual failure modes resulting from degradation during service life will generally emerge when a

severe abnormal transient or event occurs. For example, in the reactor beltline region, if a pressurized thermal shock event were to develop and the material

properties were already severely degraded by irradiation embrittlement (i.e., the RTNDT was raised to high

temperatures), the potential for failure of the vessel in a

brittle manner is possible. However, there must also be

a defect or crack of critical size present to produce this

failure scenario. The possibility of an, irradiationinduced low upper-shelf material in the beltline region

could lead to a low-energy ductile tearing overload

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situation, if again a crack is present. Therefore in the

beltline region, the necessity of having a defect is paramount for producing failure. However, the regions of

interest in the beltline have extremely low fatigue usage

factors indicating the srmall probability of having a

growing defect or crack. Defects are possible'as evidenced by the concern over underclad cracks in the

1970s;50,51 these manufacturing-induced cracks can

result from high heat input multipass strip cladding

techniques (at the overlap points), especially for

SA508-2 forging steel. The United States processes

typically use a single- or double-layer submerged arc

pass;- and there is no embrittlement of the HAZ, leading to cracks. The United States vessels have not shown

in-service evidence of underclad cracking. These small

cracks, even if present, would not pose an integrity

problem because a series of them would have to link

together to form a continuous defect of any significant

size. 6

In the areas where fatigue is a potential problem,

the type of final failure will almost always result

from a plastic overload situation caused by the

stress concentration and reduced ligament load carrying ability (i.e., the resultant material left from

the fatigue crack tip to the surface.) These ductile

tearing conditions are not necessarily failires, but

are potential through-wall leakers. However, the

inlet nozzles could be exposed to cold water during

some accident conditions that could lead to a brittle fracture there (especially if some irradiation

damage also was present). The CRDM housings

and instrumentation penetrations would probably

be exposed to only ductile overload situations, and

the closure studs also would be exposed to only

overload conditions resulting in ductile failure for

conditions caused by internal pressure, boltup

loads, and some thermal transients.

In planning for extended life, the closure studs

should be replaced near the end of the design life of

the vessel because of the high fatigue usage. The

CRDM housings, instrumenitation penetrations,

and outlet/inlet nozzles should be, monitored

closely and will most likely limit the vessel extended

service life before repair or replacement is required.

The beltline region may be most limiting, but each

individual vessel will require independent assessment depending on material chemistries, weld

types, and fluence conditions.

3.6 In-service Inspection and:

Surveillance Methods

In-service inspection (ISI) is. required by the

ASME Code, Section XI. 31 The general philosophy lof Section XI ISI and the status of the

improvements in the code requirements for the

RPV are presented in Chapter 12. There are basically four inspection intervals, and certain welds'

must be examined during those intervals. All shell,

head, shell-to-flange, head-to-flange, and repair

welds must be subjected to a lOO1o volumetric

examination during the first inspection interval

(over 3 to 10 years). Successive inspection intervals

require fewer beltline region, head, and repair weld

examinations. The nozzle-to-vessel welds must all

be subjected to a volumetric examination during all

four inspection intervals. Twenty-five percent of

the partial-penetration nozzle welds (CRDM and

instrumentation) are required to have a visual,

external surface examination during each inspec-.

tion period (leading to total coverage of all nozzles). All of the nozzle-to-safe end butt welds with

dissimilar metals (i.e. ferritic steel nozzle to stainless steel or Inconel) must be subjected to volumetric and surface examinations at each interval. These

welds are covered in Chapter 5 (primary system

piping) of this report. All studs and threaded stud

holes in the closure head studs need surface and

volumetric examinations at each inspection interval. Any integrally welded attachments are required

to have surface (or volumetric) inspections of welds

at each inspection interval. A brief description of

the currently used volumetric examination methods

is given in Chapter 11.

Thus, the inspection plan is very complete and

results in close monitoring of potentialfatiguecrack formation and growth. The additional monitoring and recording of transients are usually done

in accordance with the plant technical specifications; however, this area could be greatly enhanced

by more accurate monitoring and assessment of the

transients (identified earlier) leading to fatigue

usage.

The surveillance program for monitoring

changes in the vessel wall and weld properties

caused by neutron embrittlement has been discussed in Section 3.4. Coupling these surveillance

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results (and existing data) with the accurate characterization of flaws or defects: through ISI would

allow prudent calculation of vessel integrity. If better logging of the critical fatigue transients also is

undertaken, even further confidence in the reliability of the pressure vessel is possible. More confidence now in the safety of PWR vessels will extend

directly into later life extension.

3.7 Summary, Conclusions, and

Recommendations

A summary of the important degradation sites,

mechanisms, and consequences for PWVR vessels is

shown in Table 3.2. Note that the ranking of sites is

based on the importance of that area, not on the possibility of that site having severe degradation. For example, the closure studs are probably the most likely site to

have significant damage (except for possibly the reactor beltline region), but the studs are listed at the end

because they are easily replaced. Other conclusions and

recommendations are:

I. Radiation embrittlement is the primary

safety concern that may limit vessel life

extension; flux reduction programs used to

reduce fluence levels need to be

coordinated with overall life extension

plans; Charpy referencing of toughness

needs to be resolved, fundamental studies

of irradiation damage and further work on

phenomenological correlation models

should be developed for both transition

temperature shift and upper shelf energy

drop (rate effects need to be better understood and correction factors developed);

surveillance program schedules may need

adjustment and revision to meet lifeextension goals.

2. Thermal annealing needs to be pursued as a

contingency plan (a long-term option);

annealing mechanisms and kinetics need to

be better understood and models developed.

3. Application of miniature specimen testing

and reconstituted specimen testing need to

be evaluated. Use of these testing techniques is valuable because the available

space to irradiate the test specimens and

the available test material are limited. This

is discussed further in Chapter 14.

4. Other synergisms associated with irradiation may be present and need to be studied

along with some simple long-time thermal

aging studies, the total package of information needed for structural integrity

analysis should be reviewed and uncertainties quantified.

5. Nozzles' and penetrations represent.,the

major fatigue problems associated with

life extension; better monitoring of fatigue

transients is needed to properly assess

actual usage factors; more sophisticated

fatigue usage factors can be developed

using the local-strain approach presented

earlier; as actual usage factors become

large (i.e,'approach unity), fatigue-crack

propagation calculations using the state of

the art should be used to predict usable

life.

6. Flange closure studs should be replaced

when considering life extension; also, the

thimble tubes in W plants should be

inspected and assessed for replacement.

7. Inspection and transient logging records as

well as' all material properties (unirradiated and surveillance results), design

data, and other key information, should

be maintained for future reference; early

collection of these records will allow life

extension plans to be developed.

8. It is probable that the reactor pressure vessel, in most cases, can be used for longer

than the 40-year license period dictated by

federal regulations. 9 Actual ,decisions are

plant specific, however, due to slightly different designs, different limiting materials,

and operating conditions (fluence levels

and transient history). Continued close

monitoring of the material condition in

the vessels is required for life extension.

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Table 3.2. Summary of degradation processes for PWR reactor pressure vessels

Rank of

Degradation

Site

Degradation

Site Stressors Degradation Mechanisms Potential Failure Modes

,I , ;

I Beltline region Neutron

irradiation,

mechanical

and thermal

stresses

* Irradiation embrittlement

(degree is dependent on

individual vessel

materials and flux

spectrum history) I , fI ,

Ductile high-energy

tearing leading to

leakage (net section

over-load)

ISI Surveillance Methods

100% volumetric during

first inspection; one

weld for subsequent

inspection

Surveillance program for

assessing irradiation

damage is required by law

if

Brittle fracture (i.e..

- ' pressurized thermal

shock)

, Ductile low-energy

tearing (low uppershelf toughness)

Environmental

fatigue

Ductile overload leading

to a leak; possible brittle

fracture if PTS occurs

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2 Outlet/inlet Mechanical

nozzlesa and thermal

stresses

Fatigue crack initiation

and propagation

, Fatigue crack initiation

and propagation

Ductile overload leading

to a leak; possible

brittle fracture if

pressurized thermal

shock occurs'with

some irradiation

embrittlement

Ductile overload leading

to a leak

3 Instrumentation

nozzles

(penetrations)

and control rod

drive mechanism

housing nozzles

Mechanical

and thermal

stresses

All nozzle welds inspected

volumetrically at each

interval

Visual, inspection of

external surface; 25%o of

nozzles inspected at first

interval; remaining 755o

spread out over next three

intervals

Volumetric and surface

inspection of all studs

and threads in flange stud

holes at each interval

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4 Flange closure Mechanical

studs ' Iand thermal

stresses

Fatigue crack initiation Ductile overload failure

and propagation (possibly (can be replaced)

corrosion assisted)

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a. Welds at the vessel to safe end are covered in Chapter Son primary system piping; the comments here are also applicable except

that volumetric and surface'examinations ire required. '

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3.8 References

1. American Society of Mechanical Engineers, ASMDE Boiler and Pressure Vessel Code, Section 111,

Nuclear Power Plant Coinponents, Division 1, New York, 1983.

2. K. Hoge, Evaluation of the Integrity of Reactor Vessels Designed to ASME Code, Sections I and/or

VIII, NUREG-0031, NRC-I, U.S. Nuclear Regulatory Commission, 1976.

3. J. J. Koziol, "Radiation Surveillance Program for Combustion Engineering, Inc. Reactor Vessel

Materials," in Status of USA Nuclear Reactor Pressure Vessel Surveillance for Radiation Effects,

L. E. Steele (ed.), ASTM Special Technical Publication 784, Philadelphia: American Society for

Testing and Materials, 1983, p. 216.

4. NV. R. Corwin, R. G. Berggren, and R. K. Nanstad, "Charpy Toughness and Tensile Properties of a |

Neutron-Irradiated Stainless Steel Submerged-Arc Weld Cladding Overlay," Effects of Radiation on

lMaterials: Twelfth International Symposium, WVilliamsburg, Virginia, June 18-20, 1984, ASTM Special Technical Publication 870, American Society for Testing and Materials, 1985, pp. 951-971.

5. G. T. Yahr, R. C. Gwaltney, A. K. Richardson, and NV. L. Server, "Case Study of the Propagation

of a Small Flaw Under PWR Loading Conditions and Comparison with the ASNIE Code Design

Life: Comparison of ASME Code Sections I11 and Xl," ASMEPressure Vessel and Piping Conference, Chicago, July 20, 1986. C

6. T. J. Griesbach, "Section XI Future Trends in Equipment Requalification of Nuclear Components,"

ASMfE Pressure Vessels and Piping Conference, San Antonio, Texas, June 17-21, 1984, PVP

Volume 90.

7. NV. H. Bamford, "Technical Basics for Revised Reference Crack Growth Rate Curves for Pressure

Boundary Steels in LWR Environment," Journal of Pressure Vessel Technology, 102, 1980, p. 433.

8. Seminar on Nuclear Power Plant Life Extension, Seminar Proceedings, cosponsored by EPRI,

Northern States Power, U.S. DOE, and Virginia Power, Alexandria, Virginia, August 25-27, 1986.

9. Code of Federal Regulations, Title 10-Energy, Part 50-"Domestic Licensing of Production and

Utilization Facilities."

10. L. E. Steele (ed.), Status of USA Nuclear Reactor Pressure Vessel Surveillancefor Radiation Effects,

ASTM Special Technical Publication 784, Philadelphia: American Society for Testing and Materials, 1983.

r

11. A. L. Lowe, K. E. Moore, and 1. D. Aadland, "Integrated Reactor Vessel Material Surveillance

Program for Babcock and Wilcox 177-FA Plants," Effects of Radiation on Materials: Twelfth International Symposium, Williamsburg, Virginia, June 18-20, 1984, ASTM Special Technical

Publication 870, American Society for Testing and Materials, 1985, p. 931.

12. U.S. Nuclear Regulatory Commission, "Effect of Residual Elements on Predicted Damage to Reactor Vessel Materials," Regulatory Guide 1.99, Revision 1, 1977.

13. U.S. Nuclear Regulatory Commission, "Effects of Residual Elements on Predicted Radiation Damage on Reactor Vessel Materials," Regulatory Guide 1.99, Revision.2, February 1986.

14. NV. L. Server, In-Place Thermal Annealing of Nuclear Reactor Pressure Vessels, NUREG/CR4212,

EGG-MS-6708, 1985.

28
15. W. L. Server and A. 7Iboada, "An Approach for Estimating Post-Anneal Reirradiation Embrittlement of

Reactor Vessel Steels," Proceedingsof the Second International Symposium on Environmental Degradation of

Materials in Nuclear Power Systems- Water Reactors, Monterey California, September 9, 1985, American

Nuclear Society, 1986, p. 496. -

16. G. R. Odette and P. M. Lombrozo, Physically Based Regression Correlations of Embrittlement Data

From Reactor Pressure Vessel Surveillance Programs, EPRI NP-3319, 1984.

17. G. R. Odette, "Recent Advances in Predicting Embrittlement of Reactor Pressure Vessel Steels,"

Proceedings of the Second International Symposium on Environmental Degradation of Materials in

NuclearPowerSystems- WaterReactors, Monterey, California, September9,1985, American Nuclear

Society, 1986, p. 121.

18. G. R. Odette, "On the Dominant Mechanism of Irradiation Embrittlement of Pressure Vessel

Steels," ScriptaMetallurgica, 17, 1983, p. 1183.

19. G. R. Odette and G. E. Lucas, "Irradiation Embrittlement of Reactor Pressure Vessel

Steels: Mechanisms, Models and Data Correlations," in Radiation Embrittlement of Reactor Pres-

,, sure Vessel Steels: An International Review, ASTM Special Technical Publication 909, L. E. Steele

(ed.), American Society for Testing and Materials, 1986, p. 206.

20. G. E. Lucas and G. R. Odette, "Recent Advances in Understanding Radiation Hardening and

Embrittlement in Pressure Vessel Steels," Proceedings of the Second International Symposium on

Environmental Degradation of Materials in Nuclear Power Systems-Water Reactors, Monterey,

California, September 9, 1985, American Nuclear Society, 1986, p. 345.

21. G. R. Odette, P. M. Lombrozo, and R. A. Wullaert, "Relationship Between Irradiation Hardening

and Embrittlement of Pressure Vessel Steels," Effects of Radiation on Materials: Twelfth International Symposium, Williamsburg, Virginia, June,. 18-20, 1984, ASTM Special Technical

Publication 870, American Society for Testing and Materials, 1985, p. 840.

22. G. E. Lucas, G. R. Odette, P. M. Lombrozo, and J. W. Sheckherd, "Effects of Composition,

Microstructure and Temperature on Irradiation Hardening of Pressure Vessel Steels," Effects ofRadiation on Materials: Twelfth International Symposium, Williamsburg, Virginia, June 18-20, 1984,

ASTM Special Technical Publication 870, American Society for Testing and Materials, 1985, p. 900.

23. G. E. Lucas, G. R. Odette, R. Maiti, and J. W. Sheckherd, "Tensile Properties of Irradiated Pressure Vessel Steels," Proceedings of the Thirteenth International ASTM Symposium on the Effects of

Radiation on Materials, Seattle, Washington, 1986. - .

e-b~,'ThEe *rr of a resr

24. G. R. Odette, G. E. Lucas, and T. Griesbach,"TheEffectof Irradiation and Aging on Pressure

Vessel Steel Embrittlement," Proceedings of the International Conference on Nuclear Plant Aging,

Availability Factor, and ReliabilityAnalysis, San Diego, California, July 7,1985, American Society for

Metals, 1986, p. 375.

25. G. E. Lucas and G. R. Odette, "Application of Small Specinene Test Techniques to Power Plant

Reliability and Life Extenision 'Pr6blemis," Proceedings of the International Conference on Nuclear

Plant Aging, Availability Factor, and Reliability Analysis, San Diego, California, July 7, 1985,

American Society for Metals, 1986, p. 385. .

29

I
26. J. T. Buswell, E. A. Little, R. B. Jones, and R. N. Sinclair, "Analysis of Microstructural Changes

in Irradiated Pressure Vessel Steels Using Small Angle Neutron Scattering," Proceedings of the Second International Symposium on Environmental Degradation of Materials in Nuclear Power SystenmsWHater Reactors, Monferey, California, September 9-12, 1985, pp. 139-152.

27. P. A. Beaven, F. Frisius, R. Kampmann, and R. Wagner, "SANS/TEM Studies of the Defect Microstructures of Test Reactor Irradiated Fe-Cu Alloys and Cu Containing RPV Steels," Proceedings of

the Second International Symposium on Environmental Degradation of Materials in Nuclear Power

Systems- Water Reactors, Monterey, California, September9-12, 1985, pp. 400-408.

28. Vessel Integrity Review Group Meeting, Review of NRC Pressure Vessel Research Program, Oak Ridge

National Laboratory, December 1986.

29. M. Vagins and A. Taboada, Research Program Plan: Reactor Vessels, NUREG-I 155, Volume 1, U.S.

Nuclear Regulatory Commission, 1985.

30. Engineering Technology Division, Heavy-Section Steel Technology Programn-Five-Year Plan,

FY 1985-1989, NUREG/CR-4673, ORNL/TM-10108, August 1986.

31. American Society of Mechanical Engineers, ASME Boiler and Pressure Vessel Code, Section XI,

Rulesfor In-Service Inspection of Nuclear Power Plant Components, New York, 1983.

32. G. T. Yahr, R. C. Gwaltney, A. K. Richardson, and W. L. Server, Case Study ofthe Propagation ofa r

Small Flaw under PWYR Loading Conditions and Comparison with the ASME Code Design Life,

NUREG/CR-3982, ORNL-6099, 1984.

33. S. Suresh and R. 0. Ritchie, "Propagation of Short Fatigue Cracks," International Metals Review,

29, 1984, pp. 445476.

34. N. E. Dowling, NV. R. Brose, and W.- K. Wilson, in Fatigue under Complex Loading-Analysis and

Experiments, Warrendale, Pennsylvania: Society of Automotive Engineers, 1977, p. 55. E

35. M. R. Mitchell, "Fundamentals of Modern Fatigue Analysis for Design," in Fatigue and Microstructure, Metals Park, Ohio: American Society for Metals, 1979, pp. 385437.

36. H. Neuber, "Theory of Stress Concentration for Shear-Strained Prismatical Bodies with Arbitrary

Nonlinear Stress-Strain Law," Journal ofApplied Mechanics, Transactions of theAmerican Societyof

Mechanical Engineers, 28, 4, 1961, pp. 544-550.

37. N. E. Dowling, "Fatigue Life and Inelastic Strain Response under Complex Histories for an Alloy

Steel," Journal of Testing and Evaluation, 1, 4, 1973, p. 271. L

38. R. M. WMetzel (ed.), Fatigue under Complex Loading: Analyses and Experiments, Warrendale,

Pennsylvania: Society of Automotive Engineers, 1977.

39. P. C. Paris and F. Erdogan, "A Critical Analysis of Crack Propagation Laws,"Journal of Basic

Engineering, Transactions of the American Society of Mechanical Engineers, 85, 4, 1963, pp. 528-534.

40. L. A. James and D. P. Jones, "Fatigue Crack Growth Correlations for Austenitic Stainless Steels in M

Air," in Predictive Capabilities in Environmentally Assisted Cracking, Proceedings of the Winter U

Annual Meeting of the American Society of Mechanical Engineers, Miami Beach, Florida, November

17-22, 1985, PVP Volume 99, p. 363.

30
41. AU. A. Van Der Sluys, Corrosion Fatigue Characteristics of Reactor Pressire Vessel Steels, EPRI NP-2775,

1982.

42. S. Suresh and R. 0. Ritchie, "Near-Threshold Fatigue Crack Propagation: a Perspective on the

Role of Crack Closure," in Fatigue Crack Growth Threshold Concepts, D.L. Davidson and S. Suresh F

(eds.), Warrendale, Pennsylvania: The Metallurgical Society/American Institute of Mechanical L

Engineers, 1984, pp. 227-261.

43. J. Schijve, in Fatigue Crack Closure, ASTM Special Technical Publication, J. C. Newman and

Nv. Elber (eds.), American Society for Testing'and Materials, in press.

44. W. Elber, "Significance of Fatigue Crack Closure," in Damage Tolerance in Aircraft Structures,

ASTM Special Technical Publication 486, American Society for Testing and Materials, 1971,

pp. 230-242.

45. C. M. Ward-Close and R. O. Ritchie, "On the Role of Crack Closure Mechanisms in Influencing Fatigue

Crack Growth Following Tensile Overloads in a Titanium Alloy: Near-Threshold vs. Higher AK Behavior,"

ASTM International Symposium on Fatigue Crack Closure, Charleston, South Carolina, May 1986.

46. 0. E. Wheeler, Spectrum Loading and Crack Growth, New York: American Society of Mechanical

Engineers, 1971.

47. J. Willenborg, R. M. Engle, and H. A. Wood, A Crack Growth Retardation Model Using an Effec- L

tive Stress Concept, AFWPL Report No. AFFDlTM-71-1-FBR, Wright-Patterson Air Force Base,

Dayton, Ohio, 1971.

48. J. C. Newman, "A Crack-Closure Model for Predicting Fatigue Crack Growth under Aircraft Spec-'

trum Loading," in Methods and ModelsforPredictingFatigue Crack Growth under Random Loading,

ASTM Special Technical Publication 748, Philadelphia: American Society for Testing and Materials, 1981, p. 53.

49. R. P. Gangloff and R. O. Ritchie, "Environmental Effects Novel to the Propagation of Short [

Fatigue Cracks," in Fundamentals of Deformation and Fracture, Eshelby Memorial Symposium,.

B. A' Bilby, K. J. Miller, and J.'R. Willis (eds.), Cambridge, England: Cambridge University'

Press, 1985, pp. 529-558.

50. A. G. Vinckier and A. W. Pense, A Review of Underclad Cracking in Pressure Vessel Components,

WRC Bulletin 197, Welding Research Council, New York, 1974.

51. A. Disooge et al., "A Review of Work Related to Reheat Cracking in Nuclear Reactor Pressure Vessel 'r

Steels," International Journal of Pressure Vessels and Piping, 6,5, 1978, p. 329.

., ., ,, : .~~~~~~

. . . . . . . .~~~~~~~~~~~~~~~~~~~~~~~~

31
I

4. PRESSURIZED WATER REACTOR CONTAINMENTS AND BASEMATS

M. A. Daye

4.1 Description of Containment

Structures

The majority of the pressurized water reactor

(PWR) containments in the United States are prestressed concrete containment vessels (PCCV) with

steel liners covering the inside surfaces. Reinforced

concrete containment vessels (RCCV) with steel liners are also commonly used. Other types of containments constructed as steel cylinders or steel

spheres were used for some PWRs.

The function of the containment structure is to

house the nuclear steam supply system (NSSS) and

limit the release of radioactive material to the environment during normal operation and in the event of accidents. In particular, the containment structure provides

the much needed shielding against radiation in the

event of an accident so that necessary emergency measures can be provided to minimize the consequences

and deal with the causes in a protected environment.

The structure also provides protection to the reactor

system from the effects of severe and extreme environmental loads such as tornadoes, external accidents,

and penetration by moving objects.

Containment structures are designed to withstand design earthquakes, loss-of-coolant accident

pressures and temperatures, and operating loads

and temperatures for a continued, predesignated

life span of normally 40 years.

4.1.1 Prestressed Concrete Containment Vessels.The principal components of a PCCV are

reinforced concrete, steel liner plate, and posttensioning wire or strand tendons..

In general, the containment structure consists of a

flat reinforced concrete slab, an upright posttensioned

concrete cylinder and posttensioned concrete dome.

All internal surfaces are lined with steel plates anchored

to the concrete. Figures 4.1 (a) and 4.1 (b) represent two

configurations for the PCCV containments.

4.1.2 Reinforced Concrete Containment Vessels. The RCCVs are basically similar to the prestressed concrete containment vessels except that

the cylinder and dome are not posttensioned and

rely completely on the reinforcing steel and concrete for load resistance (Figure 4.2).

4.1.3 Steel CylinderlSteel Sphere Containment

Vessels. Steel cylinder containments consist

mainly of a freestanding steel cylinder with a steel

dome on top and an ellipsoidal steel bottom that is

anchored to a concrete foundation. The inside of

the steel bottom is filled with concrete to approximately the transition line between the ellipsoidal

bottom and the base of the cylinder. Within the

bottom concrete inside the containment, the reactor cavity is formed to house and support the reactor vessel. The freestanding cylinder and dome are

enclosed inside a reinforced concrete reactor building that forms the biological shield around the containment (Figure 4.3). The PWR containments

with ice condensers, i.e., Sequoyah-I and -2, and

Watts Bar-I and -2, have this type of containment.

Some of the other plants having this type of containment are Indian Point-I, St. Lucie-l and -2,

Prairie Island-I and -2, and Davis-Besse-l.

The steel spherical containment structure is similar in concept to the freestanding steel cylinder

except that the cylindrical portion is omitted and a

complete sphere is used. The bottom of the sphere

is anchored to a concrete foundation. The inside is

filled with concrete to a height sufficient to provide

structural stability and form the reactor cavity

within. There is one commercial PWR plant that

has this type of containment structure in the United

States, San Onofre-I. Yankee-Rowe, a 175-MIVe

demonstration plant, also has a steel spherical containment and it is supported by concrete/steel

columns through the lower hemisphere.

4.1.4 General Discussion. Exterior and internal

walls and floors of the containment structure are

invariably of thick concrete primarily for shielding purposes. Occasionally, high-density aggregate such as

ilmenite, is used to provide additional shielding without increasing the thickness of the concrete members.

The base slabs for both PCCV and RCCV containment structures are essentially the same and

constructed of reinforced concrete. Some variations in the geometry of the base slabs may exist

depending on the NSSS system used. The slab is

either completely flat or has a discontinuity where

the reactor vessel is located in a cylindrical reactor

cavity compartment. formed into the slab.

Posttensioning systems consist of high-strength

steel wire or strand tendons, end anchorages, and

corrosion protection filler material. The tendons

L

L_

7,

L.

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I

200 ft.

I

I

L

: r,

Figure 4.1(a). PWR prestressed concrete containm'ent Type l and Typel11.

:~~~~~~ I

i - .

i,

L

33
I

210 ft.

U

rU

Secondary shielding

and pressure boundary

S 11 BPimay shieldin

' (~~~~~~ - D ~~~~~~It.

~~a~: -10 ft.

7-3054

Figure 4. 1 (b). PWR prestressed concrete containment Type I I l-three buttresses and hemispherical dome. Li

34

E'
I

..

-Secondary shielding

and pressure boundary

V'O ft.

-10 ft.

7-3076

I

Li

- . I . l | } , l j .; .~~~~~.

-. ~~Figure 4 2. PWR reinforced concrete containrnent.

.. , . ^ , .A. , ,,i.

35
I

Primary shielding,

Figure 4.3. PWR steel cylinder containn

are placed inside ducts within the containment wall

and dome, which were formed during construction. The tendons provide prestressing forces in two

orthogonal directions within the curved shell structure. For types I and II prestressed containments,

[both types are shown in Figure 4.1 (a)], the tendon

anchorages are provided at the top and bottom of

the cylinder for the vertical tendons, at six (Type I)

or three (Type 11) equally spaced buttresses on the

circumference of the cylinder for the hoop tendons

and around a ring girder at the top of the cylinder

4 ft annular space

ndary Secondary shielding

v85 f t.

'v 0 ft.

7.3041

with reinforced concrete reactor building.

for the dome tendons. In Type Ill containments,

[Figure 4.1 (b)], the tendon anchorages are provided

at the bottom of the cylinder and the vertical tendons extend from one side of the cylinder over the

dome and down the opposite side. Tendon anchorages also are provided at three equally spaced buttresses on the circumference of the cylinder and

extending vertically into the dome for the hoop tendon in the cylinder and the dome. All tendons and

anchors are encased in a corrosion protective material and tightly capped at the ends.

7.

lent

Li

36
I

The PWR containment structures are designed in,

accordance with the American Society of Mechanical Engineers (ASME) Section III, Division 2 code

or equivalent requirements approved by the

USNRC.I Material used in the construction of the'

containment structure is controlled by' specific

requirements specified by the American Society for

Testing and Materials (ASTM) and other industry

standards to guarantee a quality product. Quality

control and quality assurance programs are set to'

control all steps of design, production, and construction of the containment.

The typical values of design/capability pressures

and free volumes for the three PWR containments

are as follows:

Design/Capability Free Volume

Containment Pressure (10 E+03

(plant) (Psig) cubic feet)

PCCV (Zion) 47/134 2,600

RCCV (Surry) 45/119 1,800

Steel cylinder (Sequoyah) 12/50 1,200

4.2 Containment Environment

and Degrading Factors

Containment structures are subjected to a number of stressors that influence the characteristics

and physical properties of the structural and nonstructural components of the containment. In' general, exposure can be categorized as external,

caused by the outside environment, and internal,

caused by the environment created by the operating

systems inside the containment building. An additional deteriorating factor unique to concrete is the

adverse chemical reactions 'in concrete that are'

related to certain characteristics of the concrete.

mixture and its constituents.

4.2.1 External Environment. Exposure to 'the

external environment affects mainly the concrete

and in advanced cases of cracking and spailling it'

can possibly affect the embedded reinforcement.

The adverse effects are due to wetting and drying,

and freezing aind thawing cycles.'Acid rain, carbonation, salt spray on containment structures located

near the sea, and harmful sulfates in ground water

affecting the concrete foundation are the major

degradation factors for both prestressed and reinforced concrefe containments, and for concrete

reactor buildings and steel shells in the case of steel

containments.

The degree 'of deterioration caused by each of

these factors varies, depending on the intensity of

each and on the quality of the concrete and the corrosion 'protection system for the steel.

The' external environment also affects the posttensioning system anchorage components and tendons mainly in' the form of corrosion or hydrogen

',embrittlement caused by-either rain or ground

water seeping in, through the' concrete' or end

anchorage caps, to the tendon ducts. The different

degradation mechariisms are in most cases related

to moisture conditions and accumulation of free

water near the anchors.

These effects and degradation mechanisms will

be discussed in detail in Section 4.4 of this report.

4.2.2 Internal Environment. 'The effects of the

internal environment can be categorized as

mechanical deterioration because of vibration

'loads, high temperatures, and fatigue stress and

deterioration resulting from exposure to moisture

and nuclear heating'. The internal environment

stressors affect both steel and concrete components

in different degrees, depending on the severity of

the stressor and the proximity of the components to

the source of deteriorating factor. The posttensioning system' is only slightly affected by the internal'

environment because of its location, embedded in

the concrete wall. However, the steel liner may suffer from corrosion caused by the hot, moist atmosphere' inside the containment, coupled with an

acidic environment.

4.2.3 Chemical Reactions in 'Concrete. Aside

from the 'exterril and internal environmental

effects, concrete can be subject to adverse internal

chemical reactions in the form of alkali-aggregate

reactions, carbonate-aggregate reactions -or

cement-aggregate reactions resulting from adverse

characteristics of the concrete mix ingredients. 2

Outside factors such as moisture and temperature

conditions can enhance such reactivities. Reactivity

in concrete may not occur or be noticeable for a

long time'after concrete' placement and 'culd 'vary

in the degree of seriousness from insignificant scaling to complete spalling and deterioration.

4.2.4 Degradation Environment During Containment Service. Degradation factors'also can

be categorized as continuously acting stressors during normal operation of the plant, stressors

L

L

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E.

37

r
I

occurring for short durations during accidents or

malfunction of certain components, and as stressors naturally acting during the life of the plant.

Stressors during normal. operation consist, of

radiation and nuclear heating, causing strength

degradation and changes in physical and chemical

properties of steel, concrete, and other materials

such as gaskets and wires. Thermal stressors caused

by elevated temperature, vibrating loads causing

fatigue stresses, moisture and condensation causing corrosion, and acidic environments causing

concrete and metal deterioration also exist during

normal operation. In addition, normal operating

procedures and scheduled pressure testing and refueling shutdowns can result in load and temperature

cycling leading to fatigue and ratcheting effects.

Naturally occurring stressors such as freeze and

thaw, wetting and drying, acid rain, etc., may be

considered as stressors occurring during normal

operation. The degree of deterioration that may

occur depends on how well concrete and other

materials are constructed. The stressors that occur

during accidents or malfunction of certain equipment can cause similar effects to those present during normal operation but at an accelerated rate and

high magnitude that can possibly result in overstress conditions.

Structures are designed to a specified service life.

Under normal conditions, most structures meet

such a design criteria. However, beyond the design

life span, degradation may alter the function, characteristics, and physical properties of components

of the structure. Thus, reevaluation and possibly

reanalysis of the load-bearing capability of the

structural members becomes necessary. using the

properties of the members at the time of the reanalysis. Main stressors affecting the structures beyond

the design life are those causing concrete deterioration, metal fatigue, and corrosion.

4.3 Degradation Sites

The degradation sites are best identified by examining each PWR containment type separately.

However, some of these sites may be common to

more than one containment type. The degradation

sites for each containment type are identified in

order of importance to the functionality and safety

of each structure.

4.3.1 Prestressed Concrete Containment Vessels. The major structural components for the

PCCV are the posttensioning system for the wall

and dome, and the reinforced concrete wall, dome,

and base slab as discussed above. The steel liner

plate covering the internal surfaces of the containment, although it is not a structural load-bearing

element, is important in providing a leak-tight barrier that will prevent radioactive releases.

Because the wall and dome are prestressed and

prestressing induces a state of compression in both

the steel liner and concrete, stress cracks in the steel

liner are less likely to occur in this type of containment, and the possibility for release of radioactive

material through the shell under normal operation

is highly unlikely. Corrosion may develop in the

liner or liner welds resulting in through-wall cracks.

However, as long as the concrete is still in compression, the danger of radioactive material release is

small. Therefore, degradation in the shell steel liner

cannot be considered to be as important as loss of

stress in the prestressing tendons. On the other

hand, corrosion or cracks in the steel liner over the

base slab can be important because radioactive

water or other liquids may seep through the base

slab and contaminate the ground water in the vicinity of the containment structure. Degradation in

the base slab liner could become more significant if

cracks were to have developed in the slab as a result

of a seismic event or any other reason.

Loss of the prestressing stress (below the limits

accounted for in the design) can be caused by broken

wires or strands, cracked anchor heads resulting in partial or complete release of the tendons, excessive wire

or strand relaxation or creep, and/or shrinkage in the

concrete. Although the possibility of anchorage failure

is remote, it has happened at various plants, after varying lengths of service times and different environments.

This type of degradation can have important consequences because accident pressures inside a containment may cause the concrete shell and steel liner to

enter a state of tension, at which time any cracks in the

steel liner may possibly permit leakage of radioactive

gases to the outside environment. 3 Therefore, the following degradation sites for prestressed containment

are identified and ranked in the following order of

importance:

I. Loss of prestressing force caused by:

a. Loss of tendon anchorage

b. Loss of tendon wires or strands

c. Excessive tendon wire or strand relaxation

d. Excessive concrete creep and shrinkage

2. Corrosion of the reinforcing steel embedded in the concrete

r

38
I

3. Cracks or corrosion in the steel liner ovej

the base slab

4. Deterioration of the tendon corrosion pro.

tection material

5. Cracks or corrosion in the steel liner ovei

the walls and dome

6. '.Deterioration in the concrete caused by the

outside environment or internal reactions.

4.3.2 Reinforced Concrete Containment Ves

sel. RCCVs rely on reinforcing steel in the con.

crete to resist tensile stresses. With the absence of a

prestressing system, the state of stress in the stee'

liner and concrete is not always in compression.

Tension zones exist both in the wall, dome, and ir

the base slab. Therefore, cracks in the steel linei

can be significant. Although seepage of radioactive

material through cracks in the base slab is of firsi

importance; leakage of radioactive gases through

cracks in the steel liner on the wall and dome are

also important. Because the tensile loads are car.

ried by the reinforcing steel, corrosion of the rein.

forcement steel is of concern. Cracks and 'high

porosity in the concrete cover over the reinforcing

steel may permit moisture and oxygen to permeate

and enhance the corrosion process. Therefore,

good quality concrete is essential for both protec.

tion and strength. The possible degradation sites

for the reinforced concrete containment can, there.

fore, be ranked as follows:.

l. Corrosion of the reinforcing steel in the

wall and dome

2. Corrosion of the reinforcing steel in the

slab

3. -Cracking and deterioration in the wall anid

dome concrete caused by'the outsidelenvi.

ronment and internal chemical reactions

4. Cracks'or corrosion in the steel liner'ovei

the base slab

5. Cracks or corrosion in the steel liner ovei

the walls and dome

6.' Deterioration inthe base slab cbncrete

' caused by groundw'ater sulfate and internal chemical reactions.

4.3.3 Steel CylinderlSteel Sphere Containmen,

Vessel. For this type of containment, the prime pro

tection against leakage of radioactive gases'is achievec

by the same steel shell that also provides the structura

barrier' against pressure buildup' inside the contain

ment. Therefore, the most important degradation sites

are associated with the shell itself. Degradation can be

- in the form of cracks or corrosion, either in the base

metal or weld seams of the steel shell. Because the bottom part 'of the containment is embedded in concrete

and the rest of the shell is freestanding, there exists a

region of discontinuity atfthe top of the concrete where

' stress concentrations and vibrating loads in the'steel

shell may cause fatigue cracking. In addition, the same

region is also susceptible to either stress-enhanced or

galvanic corrosion. Other stress concentration zones

exist in this type of containment in all other areas of

discontinuities such as hatches and penetrations. Corrosion of the bottom portion of the containmenf that is'

embedded in the concrete is likely to occur and might

* eventually lead' to escape 'of radioactive fluids to the

ground outside the boundaries of the containment.

The corrosion of this embedded portion could possibly

e be the result 'of attack by sulfates in the ground water,

which reach the steel by penetrating through the outside concrete.

Elevated temperatures affecting the steel shell,'

either in areas near high-temperature equipment or

at penetrations for high-temperature mechanical

piping, can cause deterioration in the strength and

other physical properties of the shell material.

The degradation sites for the steel cylinder/steel

sphere containments can be ranked as follows:

I . Cracks'or corrosion in the welds or base

metal of 'the 'shell because of vibrating

equipment and aggressive environments'

2. Stress cracks or corrosion in the shell at the

concrete' embedment interface at the base

of the containment

3. Stress cracks or corrosion in the shell at discontinuities near'h'atches and penetrations''

4. Corrosion of the bottom steel of the shell

embedded in concrete

5.'- Deterioration and cra'cks in the base slab

concrete inside the containment

6. Regions of elevated temperature 'on the

steel shell.'

4.4 Degradation Mechanisms

The discussion of the 'potential degradation

mechanisms is separated into 'three parts.

Section 4.4.1 deals exclusively with the 'effects of

i ; radiation and nuclear heating on the physical propI' ' erties of concrete. This 'is -the only phenomena

unique to structures that house nuclear reactors

L

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39
I

and associated systems. Several other phenomena

that affect all concrete structures and are not

unique to nuclear power plants are described in

Section 4.4.2. For this reason, there is an abundance of research and literature on degradation

caused by temperature, freezing-thawing cycles,

sulfate attack, corrosion of the reinforcing material, etc. The potential degradation mechanisms

associated with the various steel components are

discussed in Section 4.4.3.

4.4.1 Effects of Radiation on Concrete.

Because of the dual role of nuclear power plant

concrete as a structural support and radiation

shielding material, an evaluation of extended service life must address the continued effectiveness of

the concrete in both its functions. From the standpoint of shielding, water is the most desirable ingredient in concrete, because water's high hydrogen

content makes it an effective neutron moderator

and attenuator. Because concrete generally serves

as a structural support as well as a shield, the initial

water content is maintained at a minimum because

the structural strength of the concrete decreases

with increasing initial water content: Thus, in the

initial selection of the concrete constituents, shielding design considerations must be balanced with

structural considerations. Similarly, the effects of

aging and the various concrete degradation mechanisms must be evaluated in light of both shielding

and structural requirements..

The most significant radiation degradation

aspect of concrete results from nuclear heating.

Nuclear heating produces an increase in concrete

temperature causing evaporation of the free water

in the concrete, 5 that is, water that is not chemically bound to the hydrated cement. Both the structural and shielding properties of concrete are

somewhat. affected by increased temperatures and

loss of the free water.

As concrete temperatures increase, decreases in compressive strength, tensile strength, bonding strength,

and the modulus of elasticity have been observed. 6'

7

For gamma-ray shielding, the water loss as a result of

nuclear heating has little impact because of the relatively small change in density associated with the loss.

Because of the role hydrogen plays as a neutron moderator, the water lost at high temperatures can alter the

neutron attenuation properties of concrete.. However,

the amount of the water loss during a 40- year operating,

life is not expected to be substantial enough to impair

structural or shielding properties during extended operation.7

In contrast to the potential degradation of concrete resulting from nuclear heating and water loss,

direct radiation damage is a less serious concern.

This is because the fluxes that can cause direct radiation damage to concrete are higher by orders of

magnitude than those that cause damage from

nuclear heating. Hungerford7 suggests that even a

fast neutron exposure of 1021 nvt (n/cm2) has little

apparent effect on concrete aside from reducing its

rupture strength by a factor of 2. Figure 4.48

shows the axial distribution of the fast neutron flux

at the boundary of the reactor pressure vessel for a

typical PWR. Because of attenuation, only the

concrete closest to the surface directly exposed to

the source is expected to approach a fast neutron

exposure of 1018 nvt, which is well below expected

damage levels.

Evidence of radiation-induced dissociation of the

water (in small concrete samples) into hydrogen and

oxygen, accompanied by gas evolution has also been

reported. 7 This phenomenon is not expected to deplete

the concrete's water content significantly, and, is most

likely to occur near (i.e., within A.l-1/2 to 2 feet of) the

surface adjacent to the source.

Concrete samples obtained from the Nuclear

Waste Tank Farm built in 1953 (Rockwell Hanford

Operations, Richland, Washington) have been

tested for the effects of long-term exposure to temperature and radiation. 9"0 ' 1 1 The results indicated limited reduction in strength and modulus of

elasticity but did not indicate any reduction in the

integrity of the concrete. Thus, even in extended

plant operation, degradation of concrete because

of radiation is not expected to be life limiting.

4.4.2 Effects of Other Concrete Degradation

Mechanisms. The degradation of reinforced

concrete can be described in terms of elements

affecting the deterioration of the concrete and the

subsequent corrosion of the reinforcing steel after

its protective cover is degraded. The potential degradation of concrete can be divided into two categories. The first category of potential degradation

is related to the initial material selection, proportioning and mixing procedures, and placement and

curing of the concrete. The second category of

potential degradation is related to the environment

to which the concrete will be subjected during the

life of the plant.

4.4.2.1 Initial PotentialDegradation of Concrete. This category of potential degradation is

influenced by such factors as physical properties of

I

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40

U'
I

C.a)

0

C

0,

U)

0

IL

r~

L

0 50 100 150 200 250 300

Distance (cm) 7.3053

Figure 4.4. Isoflux contours for neutrons with energy > I MeV in the PWR (R,Z) model (neutrons/cm2 s).9

41
I

the aggregate, water-cement ratio, type and quantities of cement, shrinkage and settlement cracking,

temperature effects related to heat of hydration,

and adverse chemical reactions. These potential

degradations can be controlled by the proper choice

of materials, mix ratios, and curing procedures.

However, if any of the key parameters affecting the

concrete integrity are not handled properly during

mixing and placing, the quality of the concrete can

be reduced which will increase the potential for

later degradation.

These degradation modes usually result first in

surface cracking of the concrete. Cracking of concrete can occur while the concrete is in either a plastic or hardened state. The presence of cracks

normally does not affect the ability of the structure

to carry load, but can affect structural durability by

exposing the embedded reinforcements to a corrosive environment. Excessive shrinkage resulting in

cracks of unacceptable sizes usually results from

use of materials and techniques that are not in

accordance with good engineering practice and that

would not pass standard construction inspections.

4.4.22 Environmental Degradation of Concrete.

1. Chemical Reactions. A number of deleterious chemical reactions can produce concrete cracking. These generally are related

to the concrete aggregate and

include: alkali-aggregate reactions,

cement-aggregate reactions, and the reaction of carbonate aggregates. Sulfatebearing waters also present problems, but

by attack of the cement paste rather than

involvement of the aggregate.

a. Alkali-Aggregate Reactions., Cracking

can occur as a result of expansive reactions between aggregates containing

active silica and alkalies usually derived

from cement. Problems with structures

built since about 1950 have been significantly reduced through proper selection

of aggregate materials (petrographic

examination to identify potentially active

materials), use of low alkali cements

(<0.6qo equivalent NaO), and addition

of pozzuolanic materials that restrict

expansion.

b. Cement-Aggregate Reactions. Concrete deterioration attributable to

cement-aggregate reactions has been

observed with some sand-gravel aggregates. These materials are highly siliceous and produce map cracking in

concrete. This type of distress is not

prevented by the use of pozzuolana or

low-alkali cements. Tests have been

devised to indicate potential damage

from this phenomenon. 2 A recommended remedy, if these materials

must be used, is to dilute the aggregate

with crushed limestone.

c. Carbonate-Aggregate Reactions. Certain dolomitic limestone aggregates

containing some clay react with alkalis

to produce an expansive reaction. The

alkali-carbonate reaction can be controlled by keeping the alkali content of

the cement low (<0.4%) or by diluting the reactive aggregate with a lesssusceptible material.

d. Sulfate-Bearing Waters. The sulfates

of sodium, potassium, and magnesium have caused deterioration of

many concrete structures though

chemical reaction of the sulfates with

the hydrated lime and hydrated calcium aluminate in the cement paste.

Calcium sulfate and calcium sulfoaluminate are formed along with

considerable expansion that disrupts

the concrete. Resistance of concrete to

sulfate attack can be improved by the

use of admixtures such as pozzuolana

and blast furnace slag or the use of

sulfate-resisting cements (Types V,

VP, VS, IP, or IS).12

2. Elevated-Temperature Effects. Elevated

temperature has an important effect on

concrete structures because it affects concrete's compressive strength, stiffness, and

durability.

A review of research and testing performed on concrete at elevated temperatures indicates that temperature effects are

a function of a number of factors, some of

which are related to the type of aggregate

and mix proportions and others of which

are related to the limits and duration of the'

temperature exposure. Other important

factors influencing the results of tested

concrete are the temperature of the concrete at the time of test and the condition

of the concrete during heating (whether

sealed or open). Thermal cycling is also

important because it can cause cracking

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and separation between' the concrete

matrix and aggregate and between the concrete and the reinforcing steel.

The effects of the type of aggregate on

compressive strength do not become significant until the temperature reaches

%600'F (3160C). Neville2 reported that

the compressive strength exhibits a slight

increase between 400 and 600'F (204 and

3160 C) (Figure 4.5). The modulus of elasticity is reported by Neville to drop with

increased temperature, Figure 4.6.

Bertero and Polivkal3 concluded that

moisture in sealed specimens has a detrimental effect on concrete exposed to high

temperature. They reported a 71 % and

5SIo reduction in compressive strength

and elastic modulus, respectively in sealed

specimens exposed to 3000F (1490C).

However, other specimens from the same

concrete experienced insignificant reductions in mechanical properties when moisture was allowed to escape during the

thermal exposure.

3. Wetting and Drying. In structures-with

cracks, improperly treated construction

joints, or areas of segregated or porous

concrete, water may enter and exit. As

water passes through the concrete, it dissolves (leaches) some of the calcium

hydroxide and other solids so that in time

serious disintegration of the concrete may

occur. As the concrete dries, salts (sulfates

and carbonates of sodium, potassiun, or

calcium) are deposited through evaporation of the water or interaction with carbon dioxide in the atmosphere. Extensive

leaching' causes an increase in concrete

porosity, which leads to reduced strength

and increased vulnerability to other corrosive environments.

4. Freezing and Thawing. Repeated cycling

of wet concrete between freezing and thawing can lead to severe damage or destruction of concrete structures. The principal

force responsible for concrete damage

resulting'from this phenomenon is internal

hydraulic pressure created-by the expand--

ing ice-water system during freezing. Construction of highly freeze-thaw resistant

structures requires attention to these factors: (a) use of air entrainment, (b) selection of aggregate with durability adequate

for the' proposed exposure, and (c) use of

initial low water-to-cement ratio concrete.

5. Cracking. Cracks in concrete surfaces can

be indicative of normal processes such as

drying shrinkage or of unusual (and

unwanted) processes such as foundation

settlement, structural overload, chemical

reactions, and other degrading processes.

Because the opening of cracks allows

access by hostile environments to reinforcing steel and other embedded items, it is

important to investigate cracks of significant size.

6. Aggressive Chemical Environrnent.

Aggressive environments of high temperature, moisture, and acid-saturated fluids

or gases are common in power plants.

Such environments are detrimental to both

concrete and steel containment types.

When concrete is exposed to an acidic environment, a reaction occurs between the

acids and the calcium hydroxide of the

hydrated portland cement. Some of the

acid reactions result in water-soluble calcium compounds that are then leached

away by aqueous solutions. In the case of

sulfuric acid, the deterioration is more

pronounced because the reaction product

occupies a larger volume than the acid and

the calcium hydroxide, resulting in spalling and cracking in the concrete.

7. Corrosion of Embedments. Cracking and

spalling of concrete can result from the corrosion of embedments, primarily mild steel

reinforcing bars. Concrete cracking and spalling results from the buildup of corrosion

products that can act as a Wedge to delaminate and spall the adjacent concrete.

Continuous exposure of aluminum to

moist concrete produces severe corrosion

because of the destruction of its passive film

at high alkalinities. Lead and zinc also 'corrode in moist concrete, with zinc being more

resistant than aluminum or lead. In all these

cases, the corrosion process is aggravated in

the presence of chloride ions. -

8. Vibration and Structural Loads. Concrete

and steel can both fail in fatigue. After a fewv

hundred thousand reversals at a significant

- stress level, the concrete strength can be

reduced to half its original value. The effects

of fatigue on concrete are more pronounced if

the stress reversals alternately place the concrete in tension and compression. 2 ,

4 The

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Temperature (OF)

1000 1200 1400 1600 1800

'a

La)

C

0

a)

Cq

a)

U

a)

Cu

-c

Cn

I

L

20 100 200 300 400 500 600 700 800 900 1000

Temperature (@C) 7-3052

Figure 4.5. Compressive strength of concrete after heating to high temperatures. 3

Temperature (OF)

-300 .200 -100 0 100 200 300 400 500 600 700

151 1 1 | I I I I I I |

C.)

c)Cu

'5 1

a)

-a

0

E 05

at

:0co

a:

-200 -100 0 100 200 300

Temperature (*C)

Figure 4.6. Influence of temperature on modulus of elasticity of concrete. 3

400

7-3051

Fn

Li

44

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I

durability of the concrete' can be improved lated at the lower end of the vertical tendons in the

significantly if the stress reversals are such area of the anchorage. This problem has been

that the concrete always remains in compres- encountered in Types I and 11 postiensioned PWR

sion. containments, Figure 4.1(a), where the configuraExcessive 'vibration can lead to cracking tion of the vertical tendon anchorage possibly

of the concrete. The presence of such allows some rain water ponding around the dome

cracks may expose the reinforcing steel to to seep through the top and travel to the bottom

corrosive elements. anchorage. Another reason for pitting corrosion of

*the wires and anchor heads may have been a poor

construction practice where the tendons were either

4.4.3 Steel Degradation Mechanisms and Their stored at the site without proper protection for long

Effects. This section will discuss the potential periods or were not properly protected after instal- Effects. Ths s n wlation and before application of the permanent cor- degradation mechanisms associated with reinforc- rosion protection material.

ing steel, posttensioning system steel, steel liner A few anchor heads on different containments plate, and steel material used for the steel cylinder/ a eweancho hea onadint ontainme

steel sphere containment shells. Unlikec have experienced either cracking or partial failure themajr.dgraatin.mchac et r e - and in one case a complete failure because of the major degradation mechanism 'for steel is cor- hydrogen embrittlement. The time at which these

rosion. Elevated temperatures and radiation also

have the potential to degrade steel. ' failures occurred ranged from a few days to a fewv

years after stressing. The chemical composition of

the anchorhead material and heat treatment proce4.4.3.1 Corrosion of Steel. The potential for dures used are thought to have contributed to the

corrosion of the reinforcing steel is highly depen- failures. It has not been positively confirmed

dent on the quality of the concrete in'which it is whether such failures are time dependent and

embedded and on the presence of cracks and poros- whether they are related to a specific environment

ity in the concrete that may allow moisture and near the anchorage.

other harmful substances to reach the steel. Corrosion, when it occurs, will result in iron oxides that 4.4.3.2 Effects -of Elevated -Temperatures on expand In volume causing spalling of the concrete expand saigofthecoSteel. in volume causing In general, elevated temperature decreases the and further exposure,~ allowing more corrosion to' andcfurtWhera eosureallcowing'mren corxosin to an yield strength and modulus of elasticity of the steel and

occur. 'Whenvrne san ntainmens otee meosured th ans increases its ductility. However, these effects are only acidic environment and protective measures such as

coatings or a dry environment are not provided (or pronounced when the steel is exposed to temperatures

have worn out), the acid attacks the steel. This of-3000F (1491C) and above.14,15 One characteristic resuls in cof special importance, even at temperatures of 100 results in chemical corrosion. Scaling of the corro- to 150'F (38 to 660C), is the relaxation of the posttension product will result in a reduced structural sec- sionin(g tendon steel that increases with increased temtion leading to increased stress 'intensities -and reduc.d leak-tightness cap.bility. Similarly, 'co- perature resulting in loss of the prestressing force. reduced leak-tightness capability. Similarly, corro- --- sion linr of the cnt seel en Fortunately, the tendons are located in the containment sion of the steel liner m i concrete containments.. could' j e ishell within the concrete which acts as an insulator and could'Jeopardize its 'leak-tightness capability,whih iprotects the tendons from exposure to harmful temper- which is the primary function of the liner.

Corrosion in posttensioning system components atures.

may be either localized or uniform. Mosi prestressing, corrosion-related failures have been the result 4.4.3.3 Effectof Radiation on Steel. This is not

of localized corrosion in the form of pitting, stress a potentially 'damaging mechanism for reinforcing

corrosion, or hydrogen embrittlement. A review of' - bars, 'liner, structural steel, anid posttensioning tenthe prestressing tendon surveillance reports indi-t dons. Reinforcing steel and posttensioning tendons in

cates that corrosion incidents are extremely'rare' the containment shell are normally exposed to limited

However, corrosion has been detected at a number ' amounts of radiation compared to the structural steel

of plants during'the tendon system inspection. The' inside the 'containment. Exposing test 'specimens to

most common incidents of corrosion are pitting on '- radiation flux levels higher than that which posttenthe tendon wires or strands. The pitting in most sioning tendons are normally exposed to has resulted in

cases has been limited and attributed to the press harmless effects.7 Therefore, this mechanism is not

ence of small amounts of water that have accumu-' " important.

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4.5 Potential Failure Modes

Possible containment failure modes during normal operation, accidents, and postdesign life are

discussed in this section.

4.5.1 Prestressed Concrete Containment Vessels

4.5.1.1 Failure Modes During Operation.

Structural failure under load is not anticipated during normal operation because load-carrying members are designed with considerable margin to carry

such applied loads. Therefore, any failure modes

are expected to be the result of exposure to adverse

environmental conditions. The failure modes are

ranked in order of importance as follows:

I. Loss of prestressing force. There are a

number of factors that can lead to different levels of losses in the prestressing force.

The most severe occurrence is the failure of

an anchor, resulting in either partial or

complete loss of the tendon stress. In addition, dealing with the situation of a partially failed anchor presents a safety hazard

to workers replacing the damaged anchor.

In most cases, failure of anchors has been

caused by hydrogen embrittlement resulting from either improper chemistry or

improper heat treatment of the anchor

materials. Failure of individual wires or

strands in the tendons is also of concern.

This type of failure can be caused by either

overstress in individual wires or strands,

localized inclusions of weak material,

localized brittle fracture because of inclusion of an embrittling type of material, or

reduction in cross section because of pitting and corrosion. Corrosion in posttensioning components has occasionally

occurred and may be the result of breakdown in the effectiveness of the corrosion

protection material (grease) around the

tendons and anchor components or the

presence of moisture and water in areas

around the tendons and anchors. Microbiologically influenced corrosion of the carbon steel tendon wires has been reported at

some PWR sites. 16 .Prestressing losses

caused by creep and shrinkage of the concrete and relaxation of the tendon material

-* might also result in failure of the posttensioning system, if the losses exceed those

accounted for in the design.

2. Corrosion of the reinforcing steel. This

mode normally occurs as a result of poor

concrete quality or concrete deterioration.

The consequences can be serious and

repairs are complicated.

3. Corrosion of the steel liner over the base

slab. This mode of failure can lead to leakage of radioactive fluids through the liner

into possible cracks or porosity in the concrete slab and out to the outside environment.

4. Deterioration of the concrete. During the

life of the plant, the concrete may be subjected to the degradation mechanisms

described in Sections 4.4.1 and 4.4.2. The

resistance of the concrete to degradation is

related to the quality of the concrete and

the construction procedures used. Deterioration of concrete can have a number of

side effects, ranging from serious structural consequences to superficial scaling.

The primary shield wall is among the concrete structures that should be inspected.

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4.5.1.2 Failure Modes During Accidents.

Accidents which could adversely affect containment structures are those which result in large

increases in pressure or a hydrogen burn explosion

inside the containment. In the case of prestressed

concrete containments, the hoop stresses in the vessel control the failure modes. Resistance of the containment to this type of loading is dependent on the

adequacy of the prestressing system. When the

internal pressure generated by the accident exceeds

twice the design pressure, it could cause sizable,

cracks in the concrete while structural stability, is

still maintained. A liner-concrete interaction may

take place at locations of major concrete cracks and

distortions and introduce high strain concentrations that may lead to localized rupture of the steel

liner plate provided the liner is firmly bonded to the

concrete. The ruptured liner plate will then allow

some leakage of radioactive gases to the outside

environment. This failure mechanism has been proposed by Rashid and is sometimes referred to as the

Rashid mode of failure.17 Other failure modes

have been identified for containment failure at

pressure levels higher than design pressure. 18

4.5.1.3 Failure Modes Beyond Design Life

of Containment. Most of the possible failure

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s ~ ~~~~ fal 'cno es modes for this stage'are similar to the failuremodes

possible during normal operation: However, problems associated with the loss of the prestressing

force because of creep and shrinkage of the concrete and relaxation of the tendon' material may

become more important because these' changes,

with time, will eventually consume all or most of

the design margin provided. Load cycling resulting

from scheduled and forced'shutdowns, and inte-'

grated leak rate tests, may also develop fatigue

cracks that jeopardize the integrity of the containment components.'

4.5.2 Reinforced Concrete Containment Ves

sels. 'The potential failure modes for RCCVs are

similar to those for prestressed conicrete containments except for the failure modes related, to the

posttensioning system (that does not exist in the

reinforced concrete containment). However, reinforced concrete containment failure in the hoop

direction under accident pressures can be more pronounced than in the prestressed concrete containment. RCCVs develop widespread cracking 'at

about 75% of the design pressure because of the

absence of the prestressing forces. Therefore, localized strains in the steel liner plate 'due to linerconcrete interaction can' be large and may lead to

leakage of radioactive gases.

,capability. The SIT and ILRT consist of pressurizing the containment to 15% above its design pressure. Associated with the SIT are measurements of

containment growth under pressure for evaluation

of actual containment behavior'in comparison with

design predictions. The ILRT includes instruimented and visual examitiation' of surface-crack

initiation and growth at critical locations such as

hatches, the ring girder, and the posttensioning systern anchorage regions. After the ILRT, the surface

of the steel liner plate also is visually inspected for

any cracks, separation, hollow sounding or inward

buckling regions, and corrosion that might indicate.

any degradation in the system. A brief description

of the visual'examination is given in Chapter 11.

The ILRT is repeated every five years to ensure

structural integrity and continued safety of the containment. 'Additi6nal'discussion on the visual

examinations and leak testing of containments is

given in Chapter 12.'

In addition to the ILRT tests, ISI of the posttensioning system in the prestressed concrete'contain'-

ment vessels is performed one, three, and five years

after the initial SIT-ILRT in-spection and every five

years thereafter.-This inspection is required by the'

operating license and ASME Section XI and is conducted to monitor the level of prestressing forces in

the postternsioning tendons and 'inspect the condition of theltendons, anchorages, 'shims, bearing

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4.5.3 Steel Cylinder/Steel Sphere Containment 'plates, and corrosion protection materials. In addiVessels. The loads in the steel containment tion, wire or strand samples' are obtained during

structures are entirely resisted by the structural steel each ISI from one or more tendons and tested for

of the shell. Cracks and corrosion in welds and base tensile strength The concrete surface condition

metal of the shell dominate the possible failure ' * near the anchorages, where it is stressed the most, is

modes in this type of containment structure during visually examined for the presence of any cracks or

normal 'operation. Containment cracks present .) ,deterioration. The ISI surveillance is also expected

both structural and safety hazards because of stress to identify any'adverse effects related to tendon

concentrations and leakage of radioactive gases, relaxation or creep of the concrete.

respectively. Corrosion and fatigue are the poten -' mProblems identified during these'tests' are' nortial failure modes for operation beyond the design y corrected immediaely. Gaskets, seals, welds, '

life. Of special concern are the shell areas near dis- etc. are provided when measured leak rates exceed

continuities such as the base of the shell, hatches, the standard design allowables. Prestressing levels

and 'enetraiions. ' are restored when tendon losses exceed those pre- penetrations. - ' -''''''''dicted. Tendons, anchors, and corrosion protection

'material are replaced when' corrosion or cracking is' 4.6 In-service Inspection ' ' observed on' the tendons or anchors" and when

' ' ' ''''chemical ' and physical test results indicate any degIn-service inspection (ISI) of the critical compo, radation in the corrosion 'pro materials.

nents of the containment structures area manda-'' In addition'to the 'SIT, ILRT, and ISI, a

tory part of the operating licenses of nuclear power walkdown and inspection of all 'components,' supplants.1Eachl'contairinnent is subjected to a struc-' ports, structural elements, and system irains is pertural integrity test (SIT) and an-integrated leak rate formed during each refueling outage. A punch list

test (ILRT) before operation of the plant to ensure is generated and executed for correction of deficomplete leak tightness and pressure resistance ciencies discovered during the walkdown.

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A great deal of information is collected on the

condition of each containment as a result of all the

above described inspection and testing programs.

However, in light of the degradation mechanisms

described in Section 4.4 and the failure mode's postulated in Section 4.5, additional monitoring may

be required to maintain the' containment in a reliable condition beyond the design life.

General monitoring and ISI procedures for extended

life operation should be established to identify the

damage due to any likely and significant degradation

mechanisms in a timely manner. Such procedures may

include both nondestructive and destructive test methods. Nondestructive methods may consist of visual

inspection, ultrasonic inspection and testing, radiography, infrared thermography, rebound method for' surface hardness, etc.19 20 Destructive 'methods should

be used only when the nondestructive methods are not

sufficient to quantify the extent of the: degradation

damage. Destructive methods may consist of obtaining

core samples from the in situ concrete for compressivestrength tests and petrographic 5caminationsA or the

use of penetration techniques sich as'the Windsor

Probe and pullout testing for determination of in situ

concrete strength. Core samples from the degraded

regions would also provide a direct means to evaluate

the width and depth of cracking or the extent of any

voids. Steel samples from both reinforcing steel and

structural steel components also may be required for

laboratory testing.

Establishment of inspection procedures to cover critical areas where adverse environmental conditions such

as high temperature, humidity, and/or radiation, and

locations subjected to an acidic environment, will be a

necessary measure to determine the exient of degradation. Representative samples. of steel 'and concrete

should be obtained from critical areas'and tested to

determine the existing chemical and physical properties

of the materials. Specifically, inspection of concrete

against deterioration from'environmental factors'and

underground water attack'and monitoring of concrete

compressive strength by testing some core samples may

be necessary. Specific examination of steel members

for corrosion, cracking, and distress because of cyclic

loading conditions may also be necessary.

In summary, by implementing a well planned,

meaningful, and comprehensive inspection program followed by a firm program of repairs and

maintenance, the service life of containment structures can be extended, probably to at least'double

the initial licensed term and possibly much longer.

4.7 Summary, Conclusions, and

Recommendations

The nuclear power plant containment degradation sites include areas and components where degradation may cause either structural damage or a

safety hazard or both. Degradation mechanisms

unique to nuclear power plants such as those resulting from radiation and nuclear heating and elevated temperature have been 'discussed in detail.

Other degradation mechanisms that affect general

structural components have also been discussed.

The available research and test results indicate

that radiation and nuclear heating have little effect

on the shielding and structural capability of concrete. However, activation of primary shield concrete surfaces in the reactor vessel cavity by

neutrons, even to a few inches in depth, can be a

real radiation protection' problem after 40 years of

operation.2 2 The effects of radiation on the steel

used in PWR containmen'is appear to be insignificant within the dose range experienced during normal operation of the plant. Although it appears

that elevated temperature does affect concrete, its

degradation effects are unique to each plant based

on the concrete material and quality used.

I Degradation caused by freezing and thawing, wetting and drying, and chemical reactions is a condition

that is related to the quality and durability of the concrete selected for the"plant. Careful testing, evaluation,

and selection of the concrete material with proper mix

proportioning and construction procedures could minimize or eliminate degradation related to these factors.

However, some remedial mneasures may be taken to

limit or slow the degradation process if it occurs. Such

measures can be in the form'of actions to prevent moisture and free water from entering and leaving the concrete.

Corrosion of the structural or reinforcing steel

can be a major degradation mode; however, it

should be noted that'corrosion is initiated as the

result of adverse environmental conditions. Distress caused by corrosion whether in structural steel

or reinforcing steel can have a great structural

impact. Cracks and fatigue caused by vibrating

loads and dynamic loads can be serious in terms of

the structural integrity of the affected members.

Failure modes'are closely related to the described

degradation mechanisms. The failure modes postulated are general because they are presented on a

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generic basis. More unique failure modes such as

liner-concrete interaction are expected to be identified when evaluations are performed on a plant-byplant basis.

The importance of ISI has been recognized and

reflected in the presently imposed inspection programs for the containment structures and components, in particular, the posttensioning system for

the prestressed concrete containment vessels. It

should be realized that great safety and economic

benefits can be derived if an expanded ISI is implemented to cover identified degradation sites that

may not be frequently inspected.2 3 Immediate

repair or remedial action is of great importance in

arresting any degradation before it progresses to a

point where other components begin to degrade or

other degradation mechanisms start that otherwise

would have not started. A simple example of this is

deterioration of the concrete cover over the reinforcement. If deterioration of the cover material is

promptly detected and repaired, degradation of the

reinforcing steel will not occur.

Tables 4.1,- 4.2, and 4.3 provide a summary of

the degradation sites, stressors, mechanisms, failure modes, and necessary inspection methods for

the three types of the PWR containments discussed

in this report.

There are several issues associated with aging and

life extension of the PWR containments that need

to be resolved. A comprehensive and standardized

ISI program is needed to identify and quantify deg-.

radation in reinforced concrete. The state-of-theart methods used in Europe, Japan, and the U.S.A.

for evaluating the condition of reinforced concrete

need to be assessed and included in the program. A

standardized program implies the same ISI pro-'

gram for all LWRs. The ASME Section XI committees on metal and concrete' containments are

evaluating the current ISI methods for reinforced

concrete and developing specific requirements for

thorough visual inspections.

Additional information about the long-term

degradation of reinforced and prestressed concrete

containments is needed. Data should be collected

from the older LWR containments. Also, facilities

that have been shut down after an extended period'

of service (such as Shippingport, Dresden, or

Humbolt Bay) could be used to obtain additional

aging-related data on reinforced concrete materials. Accelerated aging techniques could also be

'.investigated and, if found to be appropriate, used

to obtain posttensioning system and concrete materials aging data. The results from these investigations will contribute to the development of an

aging-related material property data base and support the development of quantitative models of the

degradation:of reinforced concrete subjected to

long-term exposure to elevated temperatures, radia-

:tion, and cyclic loadings. Two items in the postten-

:sioning system that need particular attention are

tendon relaxation and degradation of the grease

used as the corrosion inhibitor. Also,' a monitoring

method' is needed to detect degradation of the

anchorages because of hydrogen embrittlement

and aging of the grease caused by impurities. Other

less important recommendations include evaluation of the potential interactions between chemical

reactions in concrete and radiation, and determining the extent of the deterioration of the shielding

properties of concrete caused by radiation heating.

The structural integrity of aged metal containment

and analysis'of concrete-liner interactions in 'aged

RCCVs and PCCVs during severe accidents are two

other issues that need to be resolved. The results from

the testing of containment models subjected to pressure and thermal loadings and the results from the

EPRI-sponsored research'programs to determine the

significance of concrete-liner interactions in aged containments will be helpful.

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Table 4.1. Summary of degradation processes for prestressed concrete containment

vessel

Degradation

Rank Site

I Posttensioning

system

anchorage

2 Posttensioning

. tendon wire or

strand

Stressor

Material

properties and

trapped water

Moisture, trapped

water, or

breakdown of

grease material

Moisture, acidic

environment, and

stress

Degradation

Mechanisms

Hydrogen

embrittlement

Pitting,

microbiologicalinduced

corrosion

Potential

Failure

Modes

Loss of stress

Loss of stress

1SI

Method

Tendon

surveillance

program

Tendon

surveillance

program

.

3 Steel liner . dome and wall

Corrosion Liner-concrete

interaction,

leakage of

radioactive

gases

Leakage testing

(10 CFR 50,

Appx. J)

4 Steel liner

over base slab

Moisture, acidic

environment, and

stress -

Corrosion Leakage of

radioactive

material

Leakage testing

(10 CFR 50,

Appx. J)

5 Dome, wall, and

base slab

reinforcing

steel

Aggressive

environment

Corrosion Loss of

structural

integrity

Visual

6 Concrete Aggressive

environment and

internal chemical

reactions

Cracking and

spalling

Loss of

integrity,

corrosion of

reinforcing

steel

Visual, rebound

methods, core

samples if

required

Li

50
a

Table 4.2. Summary of degradation processes for reinforced concrete containment vessel

Degradation

Site

- Degradation

Mechanisms

Potential

Failure

Modes

ISI

Rank Stressor Method

I Dome and wall

reinforcing

steel

2 Base slab

reinforcing

steel

Aggressive

environment

Aggressive

environment

Corrosion - Lossof

structural

! integrity

Visual

Corrosion Loss of Visual

structural

integrity .

3 Steel liner

over dome and

wall

Moisture, acidic

environment, and

stress

Corrosion Liner-concrete

interaction,

leakage of

radioactive

gases -

Leakage of

radioactive

material

Leaking testing

(10 CFR 50,

Appx. J) [;

4 Steel liner

over base slab

*Moisture, acidic

environment, and

stress

Corrosion Leakage testing

(1OCFR 50,

Appx. J)

5 Dome and wall

concrete

Aggressive

environment,

internal chemical

reaction

Cracks and

spalling

Loss of . integrity,

corrosion of

reinforcing

steel

Loss of

integrity,

corrosion of

reinforcing

steel

Visual, rebound

methods, core

samples if

required

I

6 Base slab

concrete

Aggressive

environment,

internal chemical

reactions

. Cracks and

spalling

Visual, rebound

methods, core

samples if

required

[

51
a

Table 4.3. Summary of degradation processes for steel cylinderlsteel sphere containment

vessel

Degradation

Rank Site Stressor

Degradation

Mechanisms

Potential

Failure

Modes

5l1

Method

I

I Shell welds and

base metal

2 Interface

between shell

and concrete

slab at base

of shell

3 Discontinuities

in the shell

such as hatches

and penetrations

Stresses,

vibration, cyclic

loading,

aggressive

environment

Stresses,

vibration,

cyclic loading,

aggressive

environment

Stresses,

vibration,

cyclic loading,

aggressive

environment

Corrosion

Corrosion

Corrosion

Loss of

structural

integrity,

leakage of

radioactive

gases

Loss of

structural

integrity,

leakage of

radioactive

gases

Leakage of

radioactive

gases

Visual, leakage

testing

Visual, leakage

testing

F

Leakage testing

(10 CFR 50,

Appx. J)

4 Steel bottom

of shell

embedded in

concrete

Aggressive

environment

Corrosion Leakage of

radioactive

material

Leakage testing

(10 CFR 50,

Appx. J)

5 Base slab

concrete

Aggressive

environment,

internal chemical

reactions

Cracks and

spalling

Corrosion of

reinforcing

steel,

corrosion of

steel bottom

of

containment

shell

Visual, rebound

methods, core

samples if

required

I

52
I

4.8 References

I. American Society of Mechanical Engineers, ASME Code for Concrete Reactor Vessels and Containments, Section ill, Div. 2, New York.

2. A. M. Neville, Properties of Concrete, third edition, London: Pitman Publishing Limited, 1981.

3. N. Hanson, D. Schultz, and J. Roller, "Results of Larger Scale Test of a Discontinuity Region in a

Prestressed Concrete Containment," Construcdion Technology Laboratories, Third Workshop on

Containment Integrity, May 21-23 1986, Washington, D.C.

4. American Concrete Institute, Considerationsfor Design of Concrete Structures Subjected to Fatigue

Loading, ACI 215R-74, 1981.

5. G. E. Troxel and H. E. Davis, Composition and Properties of Concrete, New York: -McGraw-Hill

Book Company, Inc., 1968.

6. American Nuclear Society, Guidelines on the Nuclear Analysis and Design of Concrete Radiation

ShieldingforNuclearPowerPlants, Americani National Standard, ANSI/ANS-6, 4-1985, La Grange

Park, Illinois, 1985.

7. H. E. Hungerford et al., "Concretes, Cements, Mortars and Grouts," in Engineering Compendium

on Radiation Shielding, New York: Springer-Verlag New York, Inc., 1975. X

8. M. L. Gritzner, G. L. Simmons, T. E. Albert, and E. A. Straker, PWR and BW'R Radiation Environmentsfor Radiation Damage Studies, EPRI NP-152, September 1977, pp. 30, 49.

9. Construction Technology Laboratories of the Portland Cement Association, Strength and Elastic

Properties of Concretes from Waste Tank Farms, report prepared for Rockwell Hanford OperationsEnergy Systems Group, Richland, Washington, RHO-C-22, December 1978.

10. Construction Technology Laboratories of Portland Cement Association, Effects of Long-Term Exposure to Elevated Temperature on Mechanical Properties of Hanford Concrete, report prepared for

Rockwell Hanford Operations-Energy Systems Group, Richland, Washington, RHO-C-54,

October 1981.

11. "Strength and Elastic Properties of Concrete Exposed to Long-Term Moderate Temperatures and High

Radiation Fields," American Concrete Institute, Fall Convention 1984, New York, RHO-RE-SA-55P.

12. American Society for Testing and Materials, Standard Specification for Portland Cement, ASTM C 150. r

13. V. V. Bertero and M. Polivka, "Influence of Thermal Exposures on Mechanical Characteristics of

Concrete," in Concrete for Nuclear Reactors, Volume 1, ACI Special Publication SP-34,

Detroit: American Concrete Institute, 1971.

14. T. T. Lie, "Fire Resistance of Structural Steel," Engineering Journal/AISC, fourth quarter, 1978.

15. M. Fujimoto, F. Formura, and T. Ave, "Primary Creep of Structural Steel (SM50A) at High Temperatures," Transactions of the Architectural Institute of Japan, 306, August 1981. E

16. D. H. Pope, A Study ofMicrobiologically lnjluenced Corrosion in Nuclear Power Plants andaPractical Guidefor Countermeasures, EPRI NP-4582, May 1986, pp. 1-2, 1-3.

53
17. R. S. Dunham et al., Methodsfor Ultimate LoadAnalysis of Concrete Containment, EPRI NP-4046,

June 1985.

18. 'M. A. Daye and B. L. Meyers, "Progression of Stress Development in Pressurized Containment

Structures," Transactions of the 8th International Conference on Structural Mechanics in Reactor

Technology, Volume J, Brussels, Belgium, August 19-23, 1985, North-Holland for the Commission of

the European Communities, pp. 7-12.

19. American Society for Testing and Materials, Aarnual Book of ASTM Standards, Section 00,

Volume 00.01.

20. D. J. Naus, Concrete Component Aging and its Significance Relative to Life Evtension of Nuclear

Power Plants, NUREG/CR-4652, ORNL/TM-10059, September 1986.

21. American Society for Testing and Materials, Standard Recommnended PracticeforPetrographicExarmination of Hardened Concrete, ASTM C 856.

22. R. G. Jaeger et al. (eds.), Engineering Compendium on Radiation Shielding, Volume 11, Shielding

Material, New York: Springer-Verlag New York, Inc., 1975.

23. J. Cook, "Review of Containment Examination and Testing Requirements," 3rd Workshop on Containment Integrity, Washington, D.C., May 21-23, 1986.

54
I

5. PRESSURIZED WATER REACTOR COOLANT PIPING

R. L. Cloud and W L. Server

5.1 Description coolant piping for the CE and B&W plants is

wrought ferritic steel clad with stainless steel. The

materials for all piping are shown in Table 5. 1. - 7hr^ ferriti- nini;npn iicpr! hy nRurWnnri (7F in rindj - ~ . ~ ~~~~~~~~~~~~~~~~~~~ ua*u .31 tsssJz u 2rps ~ o *svwLa sla t.

5.1.1 Geometry, Design, Vendors. Reactor with austeritic stainless steel, as noted in Table 5.1.

coolant piping systems for the three principal pres- B&W uses weld-deposited cladding, and CE 'uses

surized water reactor (PWR) types are shown in ' roll-bonded cladding. B&W uses explosively

Figures 5.1, 5.2, and 5.3. The PWR designers are bonded cladding in the piping elbows. The elbows

Babcock & Wilcox (B&W), Combustion Engineers - are fabricated with two half elbows (clam shells)

ing (CE), and Westinghouse (W). and the clad is applied before the two halves are

As can be seen, the B&W and CE designs are welded together.

similar in that there are two heat 'exchange loops. There is no need for cladding in the W pipe because

Each loop has a' hot leg from the reactor to the it is entirely stainless steel. As noted in Table S.1, howsteam generator, and two cold legs that return the ever, there are special circumstances with the weld

coolant from the steam generator to the reactor. between the austenitic pipe and the ferritic reactor vesThere are four pumps, one in each cold leg. sel and steam generator. Because any weld to the ferThe only major conceptual difference between ritic steel would require stress relief in these wall

the CE and B&W design is that the B&W steam thicknesses, a layer of Inconel 600 is deposited on the

generator (SG) is a once-through straight tube type, nozzle ends before final vessel stress relief. Then, the

and the CE design is a recirculating U-tube type. stainless steel pipe can be welded to the Inconel and no

Circulation of primary coolant in the U-tube SG is additional stress relief is required. The Inconel was not

from the bottom hot-leg side up to the top of the thought to become sensitized in the vessel stress relief.

tube bundle, then down to the bottom of the cold- CE piping is constructed of roll-bonded cladded

leg side. The B&W straight-tube SG path for the plates. The plates that are cladded during the rollprimary coolant flow is from the top headv'ertically ing process are formed into rounds and seam

downward to the bottom head. ' welded. Careful weld preparations permit austenConsistent with the SG designs, the primary sys- itic weld material to be used on the cladding and

tem hot-leg piping for the B&W design goes from ferritic weld material on the ferritic plate forming

the reactor vessel outlet nozzle upward to the top of the main pipe wall.

the SG. The hot-leg piping for the CE design goes

directly across the compartment from the reactor

vessel outlet nozzle to the bottom' head of the SG. 5.2 Stressors

The overall arrangement of the cold-leg piping for - -

these two designs is quite sirnilar. " - ' t ~~~Nuclear plants are most economical when they

The third major PWR design, .xV. is conceptually -* - .. operate at steady state,' full power from one refueldifferent from the CE and B&W designs 'in that - v~-ing to the next. Certainly, all plant owners strive for

each heat exchange loop has a SG and a 'reactor this type of operation. From a technical point of

coolant pump. A given reactor plant may have twiO, view; this mode of operation is also optimum. The

three, or four heat exchange loops. The number of only significant source of stress is the internal presloops in any given plant is dependent on a number sure. In'the steady-state condition, thermal stresses

of factors, but the most important is the capacity of do exist because of minor degrees of restrained

the plant. Each loop increases the amount of power expansion, but they are minimal.

that may be extracted from the reactor. However, for one reason or another, nuclear

plants rarely operate for a complete fuel cycle at

5.1.2 Materials and Fabrication. The later 3Y full power in an undisturbed condition. The conplants have main coolant piping made from centrif- - cern about safety is so great that the plants are freugally cast austenitic stainless steel, and the fittings quently shut down and examined for off-normal

are statically cast stainless steel. However, the main situations. Equipment failures often require shutcoolant piping in the early plants was wrought aus- down to cold conditions. For various reasons,'

tenitic stainless steel (such as in Surry-l). The main reductions in power are common. Testing of

1.

I

I

55
Reactor

coolant

pump/motor---,

Steam

generator

Reactor

coolant

pump/motor --

Reactor

coolant

pump/motor

._Steam

)- generator

- Reactor

coolant

pump/motor

Steam

generator

Reactor

coolant

pump/motor

7-3068

Figure 5.1. Babcock & Wilcox system.

56
Pump 1B Pump2A

Steam

generator

No.2 I

I

I

7-3069

Figure 5.2. Combustion Engineering system.

fL

57

I
Two loops

Three loops

Four loops

Reactor

7-3067

Figure 5.3. Westinghouse system.

58
Table 5.1. 'PWR main coolant pipe material . I .; ,. i,. . *I

Vendor'

Bib

CE

Piping ':

CF8AC

CF8M ;- i

SA516GR70

Fittings

CF8A

*- - SA516GR70

! Type 309L SS

Claddinga

N.A.

Type 308L SS

B&W ;. . SA 106GRC . l: SA 516 GR 70

Type 308L SS

Type 309L SS

Type 304L SS

a. Welds between austenitic stainless steel (SS) and ferritic material normally have an Inconel 600 layer to overcome the need for

sensitizing stress relief heat treatment.

b. Early W PWR plants had wrought Type 304 SS main coolant piping.

c. CF8A and CF8M are cast grades of l'x 304 and Type 316 SS respectively.

systems and equipment is a major cause of tran,-,,

sients. These transient conditions contribute addi-_-

tional thermal stresses to the main piping.

In addition, the plants are designed to be able to

accept unusual conditions, such as severe earthquakes and dynamic effects resulting from cata-;

strophic pipe failure in connected systems; These

various conditions are discussed in the following

paragraphs. : -

5.2.1 Steady-State Stress. All PWR primary

systems are designed to freely expand as their temperature is increased from ambient to operating

temperature, a rise of "500'F (280'C). Inall cases,

the reactor vessel is rigidly attached to the reactor

building and serves as a fixed point when the sys-..

tem heats up and expands. The hot-leg piping

attached to the reactor vessel nozzles expands along

its axis to the steam generator. All three PWR--

designs have SGs that move freely to accommodate.

the pipe expansion. L : -

The W SG is supported on columns with pinned

connections that rotate freely and permit the SG to

assume the natural or unconstrained hot position,-.

of the piping. The CE and B&W designs have SGs lasupported on skirts with sliding bases that accomplish the same result. The SGs for all three vendors

move '.2 in. (50 mm) during the plant heatup from

ambient to operating temperature. The reactor

coolant pumps also are supported in such a manner

that cold-leg expansion is essentially unrestrained.

In the steady-state condition then, the state of

stress in the main coolant piping is dominated by

the internal pressure contribution. There are small

amounts of thermal stress that result from secondary effects because perfect free expansion cannot

be achieved. In addition, there are small amounts

of gravity-induced stress that are inconsequential

because of the short, stiff nature of the primary

piping. Because the thermal expansion is unconstrained, there is little or no steady-state bending

stress. The piping itself is of relatively large diameter with very short lengths..

For the reasons given above, the main coolant

piping may be more accurately thought of as a

series of pressure vessels than piping. This concept

is important from the viewpoint of life extension,

because the measures to be taken with main coolant

piping will be'the san'xe measures to be taken with

the connected pressure vessels, in contrast with

those measures needed for long runs of relatively

flexible piping.

As discussed above, the steady-state stress is essentially given by the 'ressure6 contribution. The stress is

uniform tension in the pipe wall and is written as

I

I

I

I

Sb=pr Sh = Pr (5.1)

59

1'
I

S2 = Pr (5.2) 2t

S. = 2 (5.3)

where

p = pressure

r = pipe mean radius

t = pipe thickness

Sh = circumferential or hoop stress

SI= longitudinal or axial stress

Sr = radial stress.

The American Society of Mechanical Engineers

(ASME) code stress intensity (Pm) is equal to twice

the maximum shear stress, shown as

p pr + p (5.4)

t 2

and is limited to the ASME code design stress Sm.

Sm in turn is limited to the smaller of 2/3 S. or

1/3 Su

where

Su = specified minimum, ultimate tensile

strength

Sy = specified minimum yield strength

and these quantities are the minimum specification

values for the steel in question. Because of the high

strain hardening nature of some austenitic stainless

steels, Sm may be >2/3 Sy but never >90% of S.

nor 1/3 Sul, where S., and Su, are the yield and ultimate strengths, respectively, at temperature.

The foregoing paragraphs describe the steadystate stress on the pipe wall. As mentioned, there

may be small amounts of additional stress caused.

by secondary constraints, but these are not significant. The stress picture is quite different, however,

at discontinuities or changes in section. Stresses

will be much higher at openings in the pipe wall

where auxiliary pipe nozzles are attached and also

at terminal ends where the pipe is welded into the

primary vessels. In general, the elevated stresses at

these locations are the result of purely geometrical

effects that tend to produce incompatible deformations under the pressure load. The continuity of the

structure requires compatibility, however, and

results in locally higher stresses at geometrical discontinuities. As will be discussed later, these

stresses are secondary in nature and are important

mainly from the viewpoint of fatigue.

5.2.2 Transient Stresses. Each PWR primary

system of the current generation is designed for a

specific set of transient conditions. That is, the*

stresses resulting from a specific set of transients

satisfy the limits of the ASME code. Earlier plants

that did not fall under the requirements of

Section III of theASME code of 1971 or later or of

American National Standards Institute B31.7 were

not specifically required to be analyzed for transient stresses.

In this segment, nuclear plant transient conditions and the stresses resulting therefrom are discussed. As mentioned earlier, different generations

of plants had to meet increasingly stringent analytical requirements, but as a practical matter the physical configurations of the plants changed but little,

and the actual transient conditions were also very

similar. Therefore, little or no generality is lost by

discussing transient loads and stresses in currentgeneration plants.

5.2.2.1 Transient Conditions. The manufacturer of the PWR system is required by Section III

of the ASME code to compile a design specification that lists all' transient'design conditions. The

design specification contains a description of the

transient conditions, usually in the form of a

temperature-time relation, and lists the number of'

occurrences for which each transient condition

must be considered. Design conditions and transients have historically been grouped into the following categories:

0

0

0

Design'

Test

Normal operation

60

I'
0

Upset conditions ! 7

Emergency conditions' ' -

Faulted conditions.

Faulted conditions normally include the safe

shutdown earthquake and loss-of-coolant 'accidents. If a' nuclear plant is afflicted by either of

these events, there is no doubt that it will receive

considerable special attention, probably far exceeding any life extension program procedures. In view

of this consideration, and the extremely low probability of occurrence of faulted plant conditions,

these conditions will be omitted from this discussion on life extension.' .

All the system transients in the remaining catego-'

ries are considered in the plant analysis. Most of

the transients arc in the category of normal and

upset conditions: To understand the effect of the

transient condition, a specific transient, the system

heatup and cooldown is discussed in some'detail.'

This transient is especially appropriate for several

reasons: it certainly must occur for the plant to

operate, it is one of the simplest transients, and it'

contributes significant fatigue damage.-

Ideally, in a plant designed for 40 years, the primary

system would need to be shut down only during refueling, say conservatively once a year or on the order of

40 times during life. For design purposes, usually 200'

(or even '500) heatup, cooldown cycles are specified to

account for nonideal operation.

Normally the heatup would begin with the cool-'

ant at ambient temperature and pressure. In PWR

plants, the 'heatup 'is controlled by a pressuretemperature relation that limits the ' pressure

according to the temperature elevation. This

pressure-temperature limit ensures that the reactor

vessel will not be pressurized in a cold condition'

when the fracture toughness'of the 'feiritic steel is'

depressed, particularly later in life after some radi-!

ation embrittlement has occurred. Heat is added to'

the'system" gradually, usually the heat-loss from

operation of the 'reactorr'coolant pumps and the

temperature and pressure' increase gradually, usually at a rate'on the'order of 50'F/h'(280C/h)

although specification limits' may be higher, say

l00F/h'(380C/h). Once the temperature is elevated sufficiently,' 'the control rods can be !pulled

and the heatup completed with reactor heat. In a'

cooldown, the p'rocess is reversed.'} ''"

Temperature gradients 'develop throughout the

primary system in the c6urse'of the heatups.'From'

a stress viewpoint, the most significant temperature

gradients are those'from the inside diameter to the

outside diameter of the'piping and nozzles. The

warmer inside' diameter "seeks 'to :expand and is

restrained by the cooler outside diameter. The situation is exacerbated at nozzles and other locations

of changing' geometry and wall thickness. After

steady state is reached,' the gradients even out and

the transient stresses vanish. Each change in system'

operation produces 'these effects in a greater or

lesser degree. For example, once the system is at full

power, a change in power level lowers the hot-leg

temperature, which develops gradients in the pipe

and associated transient stresses.'

The significance of the transient stresses is that

they, being of a cyclic nature, can cause fatigue.

This is further discussed in Section 5.3.1. As can be

seen from the foregoing discussion, the damage can

be limited by reducing the number of transients and

the rate of change of temperature. ''

I

5.3 Degradation Mechanisms

and Sites

The primary-coolant piping consists of the hot-leg

piping, cold-leg piping, pressurizer surge-line piping,

and the several nozzles in the pipe connected to, for

example, the charging line, safety-injection line,

residual-heat-removal line, etc. In this section, various

mechanisms are discussed that tend to reduce the life of

piping and the nozzles in the general case. For the sake'

of completeness, an attempt has been made to discuss

the applicability of most such mechanisms to the primary coolant piping of the PWR. No particular order

is followed.

I

5.3.1 Fatigue. There is an entire list of mechanisms that can cause failure of pressure vessels and

pipings and have done so in the past. As has been

discussed in previous paragraphs, very few of these

mechanisms are applicable to the main coolarit'piping of PWRs. However, one of the most widespread

and best known, namely fatigue, is applicable. So

far as is known, there have been no failures, not

even 'discovered cracks, in existing PWR 'reactor

coolant piping. That does not mean, however, that':

there will not be fatigue problems in the future.'

The significant fatiguie mechanism is low-cycle

fatigue caused by 'a combination of, pressure and

transient 'thermal stresses. The points' of highest

stress throughout the systems are the most vulnerable. These sites include the nozzles welded into the

primary-coolant piping and the'terminal ends of

the piping. Each transient event (that is,' each

heatup :cycle,'each test of the safety injection

I

F

r

61
I

system, each operation of the charging pump, etc.)

produces a cycle of stress throughout the system.

At the geometrical irregularity of a nozzle or metallurgical irregularity of a. dissimilar, metal weld,

these stress cycles may be significant. It is perhaps

worthwhile to digress briefly to describe fatigue

damage and the means of calculating it.

At a given point subject to a cyclic stress of

sufficient magnitude, a crack will form. Then, during succeeding cycles the crack will grow until it

progresses through the wall. It was shown by

Coffinl that the fatigue life of a small, smoothpolished specimen was proportional to the range of

strain imposed on the specimen in a cyclic constant

strain range test. Since most of the fatigue life of a

small smooth specimen is in the precrack phase, it

may be inferred that the range of strain determines

the onset of crack initiation.

Later it was shown by Paris2 that the amount of

crack growth of a given crack under a given cyclic

stress correlated almost exactly with the applied

fracture mechanics stress intensity range, AK,

which is a function of the crack size, shape, stress

level, and geometry. Detailed discussions of crack

initiation and growth during low-cycle fatigue are

provided in Section 3.4.2 of Chapter 3 and in

Chapter 13 of this report.

The primary system piping is designed and analyzed on the basis of Section III of the ASME code.

The ASME code provides a fatigue design curve

derived directly from smooth, small specimen failure data by imposing safety factors of 2 on stress

or 20 on cycles, whichever is lower. These factors

are intended to account for size, environment, and

other differences between laboratory specimens

and actual equipment.3 Large-scale vessel fatigue

tests have been performed for the express purpose

of checking the ASME fatigue design curves.4 It

was shown by these tests that cracks, may be

expected to initiate below the fatigue curve, but

that, wall penetration is not expected until fatigue

usage exceeds the design curve.

The ASMtE code also employs Miner's linear

damage rule to combine fatigue damage that accumulates at different stress levels. In this way, fatigue

is accumulated according to a cumulative usage

factor that is equal to unity when the accumulated

fatigue usage is just equal to the ASME code design

fatigue curve. See Section 3.4 for a detailed discussion on this topic.

Referring then to the tests of Reference 4, it may

be inferred that for those locations in the primary

system that: have design fatigue usage factors near

unity and in-service actually experience the design

usage, then it would not be unreasonable to find

fatigue cracks initiating near the end of design life.

Of course, there are mitigating factors such as variable sequences of load cycles, etc. that could

improve the situation. Nevertheless, it is clear that

low-cycle fatigue cracking is definitely one of the

most important damage mechanisms that will have

to be dealt with in any life extension program for

primary piping.

5.3.2 Thermal Aging. Cast austenitic-ferritic

(duplex) stainless steels with significant amounts of

delta ferrite will experience a reduction of toughness when aged at elevated temperatures. The maximum effect occurs at 8850F (4750 C) and the

general phenomenon is often identified as 8850F or

4750C embrittlement. This temperature is well

above the maximum piping temperatures of PWR

main coolant piping; nevertheless, the phenomenon obeys a time-temperature relationship such

that at lower temperatures the toughness properties

are still affected, but a longer time is required. The

general phenomenon and the effect on PWR maincoolant piping is discussed at length in

References 5, 6, and 7.

In the references cited, substantial investigations

have been carried out and it has been suggested that

in the design lifetime of existing PWR plants, the

safety and integrity of the materials affected has

not been compromised. The reason for this statement is that the cast stainless steels are very tough

initially, and the amount of aging that occurs at the

reduced temperatures of PWR operating systems

does not lower the properties enough to render the

materials susceptible to brittle fracture, nor to

affect the crack-growth rates significantly to compromise the design capability.

It turns out that the degree of aging is related to

the volume percentage of delta ferrite in the duplex

material. It would definitely seem that piping or

components such as pump or valve casings that

have high delta ferrite content that are in the hot-leg

side of the plant would become susceptible to brittle fracture at some period during an extended lifetime. These components will in the future require

careful monitoring. Just as the reactor vessels are

monitored today, both as to property changes and

the appearance of cracks.

The NRC Office of Research is sponsoring a program at the Argonne National Laboratory to study:

the long-term aging effects on cast duplex stainless

steels. 8 Service-aged, as well as laboratory-aged,

specimens are being examined in this program.

Close monitoring of this program is recommended.

62

I'
Additional information on the thermal aging of

cast stainless steel components may be found in

Chapter 10, Boiling Water Reactor Recirculation

Piping, of this report.,

5.3.3 Stress-Corrosion Cracking. As is well

known, intergranular stress-corrosion 'cracking

(IGSCC) occurs in metals when tensile'stress'is

present, the environment is conducive to stress corrosion, and the material has been sensitized. There

are different conducive environments but the halogen salts promote stress corrosion cracking in austenitic stainless steels, and caustic 'environments

promote stress-corrosion cracking in ferritic steels.

In the PWR primary piping, to date, there have

been no known cases of stress corrosion cracking.

Cracking is suppressed by good control of the environment including scavenging of oxygen to very

low levels. Also, the boric acid used for reactivity

control is believed to contribute to suppression of

stress-corrosion cracking. In any event, unlike the

boiling water reactor (BWR) plants, those. few

PWRs with wrought stainless steel piping have

shown no evidence, so far, of susceptibility to

stress-corrosion cracking.

Later W plants with main coolant piping of cast

stainless have also been impervious to stresscorrosion cracking; however, this resistance to*

cracking was definitely expected. The cast stainless

steel is a duplex alloy with significant percentages

of delta ferrite that renders these alloys greatly

resistant to stress-corrosion cracking.

* The B&W and CE plants have ferritic steel piping

and therefore are free from the threat of stresscorrosion cracking because there is little or no possibility of caustic or related compounds entering

the primary system. Conceptually, the stainlesssteel cladding would be vulnerable to cracking in

some circumstances, but there has not been a problem; nor is it expected in the wrought stainless piping of the early 3M plants.

In summary, stress-corrosion cracking does not

seem to be a problem in the primary piping system

of any of the PWRs. A caveat on this conclusion,

however, is that the piping should nevertheless continue to be inspected closely, since it is to be remembered that no stress corrosion was expected in the

Inconel steam generator tubes, but after a few years

it appeared. Similarly, for BWR recirculation piping, IGSCC was not expected but occurred.

5.3.4 Corrosion. Ordinary metal wastage

because of any of the well-known electrochemical

mechanisms of corrosion has never been a problem

in the primary system piping of any of the PWRs;

It is not expected to be in the future either. The'

main reasons for this good experience are that oxygen is completely removed from the system, and

other impurities that could lead to electrochemical

action are likewise removed. Primary PWR coolant

is extremely pure water.

5.3.5 Brittle Fracture. Brittle fracture is the most

feared!of all mechanical-failure mechanisms

because these types of failure are invariably sud-

- den, complete, and catastrophic. This mode of failure is not considered possible in PWR primary

piping because all the piping materials that have

ever been used in this application are more than

sufficiently tough. Indeed, recent calculations for

all plants have proved that leaks will occur well

before fracture is possible.

5.3.6 Erosion and Cavitation. Flow velocities in

primary systems are well within acceptable limits

for austenitic stainless steel, whether used for piping or cladding. As a result, erosion of main coolant piping has never been a problem, nor is it

expected to be one in the future.

Cavitation erosion is not an uncommon problem

in the process industries. In the PWR primary systems, however, there are operating restrictions to

ensure net pump suction head limits are never

exceeded. Similarly, the static overpressure is more

than enough to control cavitation at other flow' disturbances in the system.

5.3.7 Creep. In light-water nuclear. plants, all

system temperatures are below the creep range, i.e.,

less than 0.3 Tm, where Tm is the melting point of

the material. Creep and other time-dependent

deformations are insignificant.

5.3.8 Summary. In summary, the most significant degradation mechanism in the primary PWR

piping is low-cycle fatigue. The locations at which

this occurs first are the primary piping nozzles. In

general, the larger-diameter nozzles tend to be most

susceptible, as well as those nozzles with the most

severe transients (such as the safety-injection nozzles). However, the true ranking of the nozzles is

difficult to determine because the results are dependent on the type and exactness of the analyses

(which is variable) and the actual transient usage,

which is somewhat unknown. In addition, each

manufacturer analyzes the system somewhat

1.

I

I

I

63
differently. The general results for all significant

degradation sites are shown in Table 5.2.

5.4 Recommendations for Future

Research

The fundamental and most significant damage

parameter is fatigue because of system transient operation. However, nuclear plant operators rarely keep a

careful accounting of the number, type, and severity of

transients. The truth is that very few nuclear plants

knowv with any certainty how much of their design life

has been consumed at any given time.

Therefore, it is strongly recommended here that

one or two plants, at least on a pilot basis, institute

an in-depth study of past operations in order to

gain some insight into the rate that the design life is

being consumed. The issue sounds simple, but it is

not. The plant was qualified to idealized descriptions of the transients. Actual plant transients do

not conform to the idealization and are rarely as

severe as the design descriptions.

The second recommendation is that programs be

put in place now to monitor the actual degree of

thermal-aging embrittlement. This issue is not seen as a

current problem, but it could turn out to be a problem

if very long-term operation were to be achieved.

Table 5.2. Summary of degradation processes for PWR primary system piping

Rank of

Degradation Site

Degradation

Site

Degradation

Mechanisms

Potential

Failure

Modes

IS'

Stressors Methods

I Main coolant

pipe nozzlesa

Terminal end

dissimilar

metal weldb

System

operating

transients

System

operating

transients

Fatigue crack

initiation and

propagation

Fatigue crack

initiation and

propagation

Through

wall

leakage

Through

wall

leakage

Volumetric

inspection

for diameter

:-4-in.

Surface

inspection

for diameter

<4-in.

2

3 Cast

stainless

steel

Temperature Thermal

embrittlement

Through

wall

leakage

Volumetric

a. The nozzles in the primary coolant piping are the highest ranking degradation sites. Additional specificity is avoided because the

most severely loaded nozzles are difficult to determine. Reported results are heavily dependent on type of analysis performed. Actual

damage will completely depend on the actual transient usage that occurs in the plant. This is the basis for the recommendation.

b. In W plants, the dissimilar metal welds are at the reactor vessel and steam generator nozzles.

64
I a. ! ~5.5 References' r

1. L. F. Coffin, "A Study of the Effects of Cyclic Thermal Stresses on a Ductile Material," Transactions

of the American Society of Mechanical Engineers, 76, 1954, p. 931.

2. P. C. Paris, "The Fracture Mechanics Approach to Fatigue," in Fatigue-An Interdisciplinary

Approach, J. J. Burke, N. L. Reed, and V. Weiss (eds.), Syracuse, N.Y.: Syracuse University Press,

1964, pp. 107-132.

3. "Criteria of the ASME Boiler and Pressure Vessel Code for Design by Analysis in Sections III

and VIII, Division 2,". in Pressure Vessels ,and Piping:, Design' and Analysis, Volume 1, New

York: American Society of Mechanical Engineers, 1972, pp. 61-83.

4. L. F. Kooistra, E. A. Lange, and A. G. Pickett, "Full Size Pressure Vessel Testing and Its Approach

to Fatigue," Journal of Engineering for Power, Transactions of the Amnerican Society, of Mechanical

Engineers, 86, 4, 1964, pp. 419428.

5. WV. H. Bamford, E. 1. Landerman, and E. Drar, "Thermal Aging of Cast Stainless Steel and Its

Impact on Piping Integrity," Journal of Engineering Materials and Technology, Transactions of the

American Society of Mechanical Engineers, 107, 1985, p. 53.

6. E. 1. Landerman and W. H.; Bamford, "Fatigue Characteristics of Centrifugally Cast Type 316

Stainless Steel Pipe after Simulated Thermal Service Conditions," in Ductility and Toughness Considerations at Elevated Temperature Service, American Society of Mechanical Engineers, WinterAnnual

Meeting, 1978, MPC-8.

7. G. Slama et al., "Effect of Aging on Mechanical Properties of Austenitic Stainless Steel Castings

and Welds," SMIRTPostconference Seminar 6, Monterey, California, 1983.

8. M. Vagins and J. Strosnider, Research Program Plan: Piping, NUREG- 1155, Volume 3, U.S. Nuclear

Regulatory Commission, July 1985. , .

65.

[
6. PRESSURIZED WATER REACTOR STEAM GENERATORS

V Malhotra

6.1 Description

The steam generator is an important part of the pressurized water reactor (PWR) primary system pressure

boundary. Failure of the steam generator tubes will

provide a passage for the primary system coolant to the

outside of the'containment by'way of the relief valves.

The recirculating and the once-through steam generators shown in Figures 6.1 and 6.2, respectively, are the

two types of steam generiators commonly used. In the

recirculating type steam generator, the secondary system water is fed into the downcomer (Figure 6.1)

where it is mixed with recirculating water that is draining from the moisture separators. The downcomer

water flows to the bottom of the steam generator,

across the tube sheet, and then upward through the

tube bundle where steam is generated. The primary

system coolant flows through U-tubes. It enters the

steam generator at 599 to 621OF (N-315 to 3270C) on

one'side and leaves at 5501F (\2881C)'on the other

side.

A variation of the above recirculating design is

the use of a section of the secondary side as a feedwater preheater to achieve'greater thermal efficiency. In this, a section of tubes on the outlet or

cold-leg side of the U-bends is partitioned off from

the rest of the secondary side to form a preheat

section. Water is then fed to the preheat section

rather than into the downcomer.

In the once-through steam generator, the

secondary-system water enters a feed annulus (Figure 6.2) above the ninth tube support plate level.

Here it is mixed with steam aspirated from the tube

bundle area and preheated to saturation. The saturated water flows down the annulus, across the

lower tube sheet, and upwards into the tube bundle

where it becomes steam. This superheated steam

flows radially outward and then down the annulus

to the steam outlet connection.

Operating experience has demonstrated that

steam generators have had and will probably continue to have aging-degradation problems. These

degradation problems are mainly associated with

the recirculating type of unit that is manufactured

by either the Westinghouse Electric Company or

the Combustion Engineering Company. Some of

these steam generators have developed thousands

of degraded tubes after just seven or eight years of

operation. I The five- to ten-year span of operation

appears to be the critical timeframe in which many

nuclear plant steam generators begin to exhibit

severe tube degradation.

In the existing steam generators, the tubes are

made of Inconel 600 with carbon steel as the support material. Over the years, various defects have

caused tube failure as shown in Figure 6.3. A defective steam generator tube is defined as any tube

which has been plugged for whatever reason. A

large number of tubes have been plugged, however,

the tubes that actually leaked are a small proportion of the tubes plugged. Phosphate wastage was

the major cause of steam generator defects in the

early 1970s. 2 From 1976 to 1979, denting was the

major cause of tube failures in PWR steam generators. In 1980, stress corrosion cracking/

intergranular attack (SCC/IGA) was the major

cause of tube defects (31qo), followed by denting

(29To), and wastage (9016). In 1982, the major

causes of tube'defects were primary side SCC

(54.5%), pitting corrosion (11.7%), and secondary

side SCC/IGA (10.3%). The major causes of tube

defects in 1983 and 1984, were primary side SCC

and secondary side SCC/IGA.3

6.2 Stressors

Stressors in steam generators are any load or.

environment that tends to affect the functional

capability of a part or component. Stressors

include chlorides, copper compounds, air leakage

into the system, type of water treatment, applied

stresses during fabrication, tube support design,

pH, temperature, velocities, and crevices. These

factors, either in combination or separately, can

produce intergranular stress corrosion cracking

(IGSCC), SSC/IGA wastage, erosion-corrosion,

pitting, fretting, or denting.

6.3 Degradation Sites

The type of degradation acting on steam generator

tubes varies; however, it can be divided into the two

broad categories of chemical and mechanical degradation. The following is a short description of the various

chemical and mechanical degradation mechanisms

and the sites at which they have been observed:

66

I'
viw) iFE L11 F- ~ aa

~~~ts L]~~~~~~~~~F Q ~~Deformation

~~of tube U-bend

rged) ~~~~~~~~Denting[

Local

Ae ~~~~~~~~~~~~~~~concentration

of Impurities

_,. _To blowdown

(removal of

Impurities)

rng T~ube

ng lntergranular attack

Primary Inlet Primary outlet

7-3079

Figure 6.1. Sketch of recirculating steam generator with indicated problem areas. 4

i

67

I
Upper span

lane region

circumferential

cracks - Upper

tube

sheet

-Auxiliary

feedwater

inlet

-Steam

outlet

-Feedwater

inlet

Aspirating

steam

First

support

plate

Lower tube

sheet

7-3066

Figure 6.2. Sketch of once-through steam generator with indicated problem areas.4

68
C

IU

0)

CL

76 ' 78 ' '80 ' 82 '84

Year ' ' ' 7-3050 F H u 63 t y r t g r t f l m a

Figure 6.3. History of steam generator tube failure me'chanisms.3

* Wastage (thinning): General corrosion of

the outside surface of tubing in the region

above the tube sheet when phosphate

chemistry is used. Sludge provides a mechanism for the concentration of chemicals

that attack the Inconel 600 tubing.

* Denting: The plastic deformation of

tubes resulting from buildup of carbon

steel support-plate corrosion products

(magnetite) in the tube-to-tube support

plate annuli.

* Pitting: A localized attack on the outside

surface of tubing resulting from nonuniform corrosion rates caused by the formation of local cells.' Pitting can be severe in

the area of the sludge pile or where scale

containing copper'deposits is'found. Pitting of Inconel '600 tubes takes place in the

temperature range'' 212 -to 392IF

(100 to 200 0

C), hence, pitting is 'found

mainly on the cold-leg side of the steam

generator. However, Pacific Northwest

Laboiatories (PNL) has reported the presence of pitting'on the hot-leg side in the

original Surry steam generator.

Intergranular Attack (IGA): Grain

boundary attack on the outside surface of

Inconel tubing with no stress orientation.

This attack is noted in the hot-leg tube-totube sheet crevice region and tube-to-tube

support plate-crevice region.

* Intergranular Stress Corrosion Cracking

(IGSCC): The primary side IGSCC

sometimes initiates on the inside surface of

the tubing at the U-bends of tubes in the

inner rows with small bend radiu' and the

roll-transition region where the tube is

'expanded 'into 'the tube sheet. The

'secondary side IGSCC initiates on the outside'suiface'of the iubing'in the crevice

"formed by the tube and the tube sheet, and

tube and tube support plate. '

* Fretting: The loss of tube material caused

by excessive rubbing of tubes'against their

support structure. This can be caused by

either primary side or secondary side flow-

'induced vibrations 'of the tubes.' Severe

fretting has occurred at baffles in the inte-

''gral preheater nearjthe feedwater inlet, and

at antivibration bars in the U-bend region.

69

.
* Erosion-Corrosion: The mechanical

damage caused by a corrosive environment

and impingement of liquid or solid particles in a system.. The carbon steel J-tubes

in the early Westinghouse-designed steam

generators have experienced wall thinning

and perforation caused by an erosioncorrosion mechanism. The carbon steel Jtubes have been replaced by the J-tubes

made of improved corrosion resistant

materials to avoid this type of degradation, and therefore they will probably not

have a significant impact on the life extension of these steam generators.

* Environmental Fatigue: This degradation is

a result of synergistic effects of corrosion and

fatigue. The locations of the degradation sites

within the once-through steam generators

include the secondary face of the upper tube

sheet and the uppermost tube support plate

near the open lane. The upper shell to transition cone girth welds in a recirculating steam

generator have also been degraded by corrosion fatigue.

* Thermal Fatigue: The welds at the auxiliary feedwater nozzle thermal sleeve in

once-through steam generators are subjected to thermal fatigue. The cracking of

this weld will increase the potential of thermal shock to the steam generator shell.

6.4 Degradation Mechanisms

and Failure Modes

Corrosion in recirculating steam generators

tends to occur in stagnant regions, where impurities

can concentrate and attack the tubes. These regions

include the tube-to-tube sheet crevices, the tube

support crevices, and the sludge zone above the

tube sheet. The following sections describe the

major types of degradation mechanisms and their

failure modes.

6.4.1 Wastage. Wastage, or tube thinning, or

corrosion of the tube material primarily takes place

on the outside of the tubes in the area of a sludge

pile. The sludge collects in the steam generator and

causes chemicals to concentrate.4'

5 The result is

that the tube wall thinning occurs in the region

above the tube sheet when phosphate chemistry is

used. It is believed that the rate of thinning, in part,

is a function of the chloride and oxygen concentrations, as well as the sulfate concentration. Resin

leakage from the condensate polisher bed is postulated to provide an acidic sulfate environment contaminated with chlorides. The progression of tube

wall thinning appears to be a serious problem only

in those plants that are still on phosphate water

chemistry control. About 200 tubes were plugged

in. 1982 because of the wastage problem. Closer

control of the phosphate water treatments was

attempted by a number of utilities in order to avoid

steam generator tube wastage and cracking, but

these efforts were unsuccessful. Thus in order to

halt these problems, many utilities changed the secondary water treatment from phosphate to ammo-

-nia. Hydrazine chemistry [all-volatile treatment

(AVT)h, in essence, eliminates phosphate problems.

6.4.2 Denting. Oxygen, chloride, and copper

oxide and their compounds can form a voluminous

(nonprotective magnetite) corrosion product on the

carbon steel support plate and subsequently

deform some of the steam generator heat transfer

tubes at the tube-to-tube support plate intersections. This reduction in tube diameter was subsequently termed denting. The forces that cause the

indentation result from the accelerated corrosion of

the carbon steel support material because of chloride contamination of the steam generator water.

Chloride ingress from condenser leaks concentrate

at the tube/support structure intersection and lead

to fast linear growth of nonprotective magnetite.

Several nuclear power plants experienced steam

generator tube-denting problems in the early 1970s.

The incidents of denting were observed both in

plants that were originally on phosphate chemistry

control and then switched to an AVT and plants

that were on AVT from start-up. However, the

plants experiencing tube denting generally used sea

water or brackish water for condenser cooling and

were therefore susceptible to higher and more

acidic levels of chloride contamination from condenser inleakage than were plants that used fresh

water. Some steam generator tube denting has been

observed in recent years in plants that are on AVT

control and fresh cooling water; however, only

plants with sea water cooling have shown extensive

tube denting to date.

Research has been undertaken to understand the

role of certain parameters such as chemistry, temperature, heat flux, crevice condition, and materials in the tube-denting mechanism, and to

determine the threshold concentration levels at

which denting occurs. Devices such as the

70
isothermal capsule, 'single-tHbe-mode[ boilei

heated-crevice apparatus, and inleakage concentrator were used in these studies. Additionally,'existing

data from various operating plants were analyzed

to understand 'the tube-denting mechanism. 5 A

summary of some of the conclusions derived 'fr6m'

the above work follows. - '' '

An acidic chloride environment is required at the

crevice surface of the carbon steel support 'for the

occurrence'of denting'or nonprotective magnetite -

formation (NPM). The threshold concentration of''

strongly acid-forming chloride solutions (copper

chloride or hydrogen chloride) for the formation of

NPM is between 175 and 700 ppm 'chloride.

Sufficient NPM was produced in solutions with

concentrations 'of 700 to' 1500 ppm chloride 'to

cause tube denting. The threshold concentration of

a weak acid-forming chloride solution for the for-"

mation of NPM is also between 175 and 700 ppm

chloride. The threshold chloride concentration in

sea water for NPM formation is between 2000 and

10,000 ppm and for tube denting is 20,000 and

90,000 ppm chloride. Oxygen in the makeup water

made a sea water environment more aggressive with

respect to the production of NPM. ' ' ': -'-'

In heated-crevice tests with prepacked crevices in

a I1% sea water (200 ppm chloride) and copper

oxide environment, a crevice superheat of only

0.3O F (0.17'C), which yields'a concentration of

2000 ppm chloride (based on boiling point elevation) produced 'significant 'denting.'The 'denting'

rate increased with initial crevice superheat in the

range of 0 to I IF (0.56 0C). Above this, the denting

rate decreased with increasing superheat up 'to a

value of 23IF (12.81C). As the size of an initial

open crevice was reduced in these tests, a width of

4 mils' was reached -for isothermal capsules 'ahd

2 mils for heated crevice tests at which tube denting

is reduced or stopped.'

The rate of tube denting increased in isothermal

capsules containing 5800 ppm chloride 'as the

temperature increased from 450 to 600°F (2320C to

316°C). Also, when the pH of the chloride solution

was above 4, NPM was not formed even with chloride levels of 35,000 ppm chloride.

Cooling tower water containing 0.34 ppm chloride

and silicon dioxide, which' is mildly alkaline-forning

when heated to ste'am generator temperature, pro-'"

duced denting-when copper oxide was added as an oxidant. Silicon'dioxide by itself produced slight denting

in a nondenting environment."'" -' ' '

Increasing the chromium content of "steel

decreased 'the tendency for NPM formation.

Type 405 stainless steel in place of carbon steel did

not produce'NPM. Therefore the support plates in

the new' steam 'generators are made' of Type 405

stainless steel.'Quatrefoil designs for support plates'

are currently used to reduce further the damage due,

to denting. ' ' '

Large quantities of hydrazine (40 ppm) in the

blowdown were not immediately effective in stopping denting in either a copper chloride or an oxy-'

gen environment.

Corrosion rates are a function of available superheat, but they'also depended on the concentration of

chloride 'present. The existing plant sludge "data

revealed that'plants with a copper-to-iron ratio'greater

than' I developed accelerated denting, while those with'

a copper-to-iron ratio less than I had little or no indication of denting. Current trends 'are for isea-water-'

and brackish-water-co6led plants to use AVT uith con-

.densate demineralization and titanium instead of copper as the condenser tube material.3

6.4.3 Pitting. Pitting'is an extremely localized

attack that results in -holes in the metal (pitting

occurs because of a localized breakdown in passiv:'

ity).' In' most cases these holes are relatively small

with a surface diameter about the same or less than

the depth. It is difficult to detect pits because of

their small size and because they are often covered

with corrosion products.'Pitting usually requires

several months to years to show up. Once initiated,

however, it progresses 'at an accelerating rate, and

failures often occur with extreme suddenness'.- For

this'reason, it is difficult to detect pitting''and the

best fix for life extension is to eliminate the coriditions that lead to pitting:

With'the exceptions of Millstone-2 and-lndian

Point-3,' only shallow pitting has been observed in

PWR steam'generator tubes removed from service.

The maximum pit'depthsnireasured in'tubes

removed were "s0.005-in. Pitting 'was sufficiently

minor that these pits were not detected by'eddy-'

current testing nor'did the existence of these pits

promote 'concern about'possible primary-tosecondary leakage. .

As a result of the extensive pitting; in the

Millstone-2 and Indian'Point-3 steam generators

(40%b throughwall pits), the Electric Power'-

Research'Institute (EPRI) conducted a workshop 6

on steam generator tube' pitting. It 'was reported

that'both affected 'plants had pits'of similar morphology (chromiuiri'oxide and copper metal-filled

undercut pits). These pits were found under scale'

and in similar locations (e.g., the cold-leg side near'

the sludge ' pile). 'PNL has also 'reported the

71 ;"

________________________ I
presence of pitting in the hot-leg side in the original

steam generator removed from the Surry plant.

It has been shown in the laboratory that it is possible

to pit Inconel 600 under a variety of conditions but

only a fewv of the laboratory-grown pits have the same

morphology that is found in operating plants.

Electricite de France has produced pits in low-oxygen,

high-chloride (740 ppb), high-sulfate (760 ppb) water.

These pits had chromium-rich, copper-included scabs

similar to those found in the Millstone and Indian

Point steam generators. There is no universal agreement as to whether steam generator pit initiation and

growth is primarily an operating temperature. (hot)

phenomenon or a layup (cold) phenomenon. However,

chromium oxide-filled pits are seen in hightemperature tests, but not in low-temperature tests.

Also, laboratory tests indicated that the chloride content of water and pitting are closely related, and that

pits grow under sludge or scale and are associated with

copper deposition. PNL reports that oxygen inleakage

will also cause pitting of steam generator tubes. There

is no straightforward explanation about why pits generally appear on the cold-leg side and what role, if any,

resin ingress may play in the initiation and growth of

pits. Also, no mention is made of the effect of pitting

on corrosion fatigue in the literature.

6.4.4 Intergranular Attack. IGA has been

observed in Inconel 600 tubing in the hot leg and

tube-to-tube sheet region of some of the operating

steam generators. From analytical determinations,

it appears that for IGA to occur, a highly alkaline

condition must exist in the crevice caused by the

concentration of alkaline species present in the secondary side water. It was noted that a combination

of IGA and SCC is often present close to each other

in failed tube samples, with the: IGA being more

extensive than the SCC. Tubes removed from existing plants7 indicate that both the outside diameter

surface and the intergranular fracture face had in

addition to. the three major elements,(nickel,

chromium, and iron), the presence of, sodium,

potassium, calcium, phosphorus, sulfur, aluminum, and chloride. . .

Tests. were conducted; using high-temperature

electrochemical measurementsto identify conditions leading to this kind of IGA.8. The results of

these tests indicated that in 10° Oocaustic media at

6080F,,(3200 C), IGA is commonly observed in

Inconel 600. in the mill-annealed: condition.

Thermally-treated material at 1292°F, (7000C)

showed. definite improvement, over mill-annealed

material in both resistance to IGA and SCC.

The presence of various anions strongly influences the corrosion attack. Carbonates and sulfates, and to a lesser extent phosphates, are very

deleterious and can develop deep IGA and SCC,

depending on the value of the electrochemical

potential. The electrochemical potential of the

Inconel 600 tubes is governed by the composition

of the secondary water during operation. When

AVT control is used, reductive conditions are

encountered, and in the case of caustic pollution,

IGA should preferentially occur in crevice regions.

On the other hand, when oxygen enters the steam

generator (either in the form of oxygen or metallic

oxides), the potential is raised and favors an SCC

mechanism. For this reason, the composition of the

sludge and, in particular, the oxidizing potential

could be the deciding factor in whether IGA or SSC

will occur.

Occurrence of SCC/IGA at or within the tube

sheet can be treated by sleeving affected tubes. The

large-scale sleeving at Point Beach-2 and

San Onofre-l included many tubes with SCC/

IGA. The possibility of SCC/IGA at the tube sheet

may also be reduced by expanding the tubes for the

full depth of the tube sheet and, thus, eliminating

the crevices.

Secondary side SCC/IGA has also been found,

recently, in the crevice formed by the tubes and tube

support plates.3 This degradation mechanism at

the support plates caused several failures in 1983.

The SCC/IGA at the support-plate intersections

cannot be eliminated by closing the crevices and it is

of greater concern because it can occur at more

than one level in the steam generator.

6.4.5 Intergranular. Stress Corrosion Cracking.

Increasingly, tubes have been removed from service as a

result of primary-to-secondary leaks at points of high

stress concentration: the Row I U-bends, the rolled

transition regions in the tube sheet, and dented intersections. These cracks appear to have initiated on the

inside surfaces of the tubing. It is not clear whether

inside initiated cracking of nondented units is a. sporadic occurrence related to some flaw in the manufacturing process, which imparts excessive stresses to that

surface, or if it is a generic problem4 characterized by a

long initiation time and slow rate of progression or

both. Inside surface initiation of. tube cracking, is

believed to be related to pure water or the Corion phe-..

nomenon. In the first Corion tests, pure water was

shown to crack Inconel 600 at temperatures well above

the primary side operating conditions. Laboratories

are trying to define quantitative relationships between

72
failure time and variables such as 'strain rate, 'stress,

environment, and temperature.

The outside surface initiation of SCC in the secondary side of the tube-to-tube sheet crevice region w as

discussed in Section 6.4.4.

6.4.6 Fretting. Tube fretting and 'wear is caused by

vibration. Tube vibration can be induced by fluid

crossflow perpendicular to the tubes or by parallel flow

along the tubes.4 In the PWR steam generators, a

clearance exists between the tube and the tube support

plate through which the tube passes. This clearance is

required for manufacturing and design considerations.

Vibration of the tubes can cause'impact and tangential

sliding of the tube against the tube support plate and

adjacent tubes, resulting in local wear.

Fretting regions are sensitive to fatigue cracks.

Under fretting conditions,' 'fatigue cracks can'initiate

close to the material 'surface at very low stresses, well

below the fatigue limit of nonfretted tube material. If

the vibration cyclic stresses in the tube' are sufficient to

cause these fretting-induced cracks to grow and propagate, early failure of the tube can occur. ' ' ' '

Fretting has caused extensive tube damage at the

antivibration bars in several steam generators. This

damage has been blamed on the poor design of the

antivibration bars that were in point contact with

tubes. The design has now been corrected, and the

new design provides a broader region of contact

with the tubes. ' -

6.4.7 Erosion-Corrosion. Impingement of liquid

or solid particles is known to'cause mechanical

damage, especially to protective surface films.

This can happen to a variety of materials for certain

impinging velocities, sizes, shapes, and hardness of

particles' and impinging angles.' When the' erosion

factor is' combined with a' corrosive environment,

the mechanical damage can be accelerated.'

The impingement of solid particles on steam generator tubes under operating conditions can'ncause

mechanical damage. The probable sequence of

events is that solid particles entrained in the bulk

flow impact against the tube; promoting'erosion.

In a noncorrosive environment, the erosion process

removes metal from the tube wall. In a corrosive

environment, the erosion process may first remove

a protective' film from the -tube, thus making the

tube susceptible to corrosion.- In both cases, wall

metal loss occurs, either directly or by accelerated

corrosion of the tube surface. Sources of solid particles can include any particles entering the 'steam"

generator and those particles entrained from within

the steam generators.

Erosion-corrosion has occurred in 'once-through

steam generators'around the fourteenth support

plate. This process has caused more than 40% wall

reduction in some steam generators.

6.4.8 Environmental Fatigue. Some oncethrough steam generators, particularly early units,

have an open lane, i.e., an untubed lane in the steam

generators. 9 The steam flow carries entrained water

droplets up the open lane to the upper, tube sheet

region. Evaporation of water concentrates any contaminants in these droplets and also carries the chemicals

up the open lane to the upper tube sheet region. Here

the droplets impact the heat transfer tubes and dryout,

thus further concentrating the chemicals on the steam

generatortubes. Even chemicals that are generally

innocuous in the feedwater can cause metal loss in high

concentrations.;

Tube samples removed from alongside the untubed

lane revealed a serpentine band of shallow metal losssometimes containing microcracks on the tube outside

surface just below the secondary face of the upper tube

sheet or near the highest tube support plate. Several

tube failures, attributed to fatigue, are associated with

this band of metal loss. Circumferential transgranular

cracks have also led to primary to secondary leaks. The

serpentine metal loss acts as sites for fatigue crack initiation, since the steam generator, tubes do undergo

cyclic vibration.

Laboratory tests have indicated that an acid solution containing sulfates, silicates, and chlorides can

produce metal loss and microscopic indications

(microcracks), closely resembling those observed

from field-examined tubes.

6.5- In-service rinsp'ectio'n'.

Methods''

A 'host of in-service equipment has been

designed for steam generator inspection, some of

which have been' used 'on existing power plants.

This equipment provides information about tube

denting,- intergranular. 'stress corrosion' cracking,

pitting,'tube-to-tube support gap," etc.'"A brief

description of some of ihe'techniques available to

the industry follows: ' '

* Hydrogen Evolution Monitoring as a Mea-

:sure of Steam' Generator Corrosion.10

High-temperature aqueous corrosion of

carbon'steel in steam generators produces

hydrogen and magnetite. Measuring the

73
hydrogen content of the exiting steam is, at

present, the only available on-line means

of monitoring steam generator corrosion.

It also can provide information about the

effectiveness of the corrosion inhibitors

and the effect of changes in water chemistry. There is no available information as to

whether or not this technique has been

used at any operating plant.

* Eddy-Current Nondestructive Examination for IGA. Routine inspection of steam

generator tubing is conventionally performed using a bobbin-type 'eddy-current'

coil. I I However,- IGA defects have not

been detected successfully in'all the'cases.

Extensive copper deposits on the secondary side of the U-bends and the'hot-leg

side of the tubes interferes with the eddy

current inspection. Successful detection

has been achieved by the use of pancake

coils that are arrayed with their axes parallel to the tube radius and multiplexed to

'permit use of available electronic packages. Field eddy-current data show that

volumetric intergranular attack can be

detected when'the 'flaws are from 20 to

30% through the wall, whereas an oriented

IGA crack can only be detected when the

'flaws are about 60% through the wall. Oriented IGA flaws tend to be orienied'along

the axial direction' within the tube sheet

crevice region, and along the'circumferential direction at dented tube sheet intersections and in the roll transition regions.

(Additional discussions on the eddycurrent examination technology are given

in Section 11.2.4 in Chapter 11 and Section 14.1.1 in Chapter 14 in this report.)

* EMAT System. This method addresses ultrasonic inspection of steam generator tubing

using an electromagnetic acoustic transducer

(EMAT).12 It was developed to detect. flaws

in certain areas of tubes where conventional

single-frequency differential-coil eddycurrent has had difficulty, including circumferential cracks, defects at dents or support

plates, and defects in U-bends.

The EMAT system has a good defectdetection capability for circumferential

cracks and other defects that provide a

fairly wide and sharp circumferentialoriented cross section (e.g., dented areas);

however, the system has difficulty detecting flaws with small cross sections, i.e.,

cracks that are tight. For example, the system can inspect U-bends with its advance

signal but does not effectively detect the

axial cracks that are found in U-bends. In

addition, the system has limited defect

depth-sizing capability.

* Optical Probe for Steam Generator Tube

Dent Measurement. 1 3 The optical profilometer has been tested successfully under labora-.

tory conditions. It can measure inside tubing

profiles in the range of radii from 0.32 to

0.40 in., with an average calibration error of

0.005 to 0.008 in. on nominal inner diameter

and + 0.006 in. on dents. Preliminary investigations suggest that the calibration error is

most probably caused by variation in interior

surface finish.

* Pulse-Echo Ultrasound for Steam Generator Tube-to-Support Plate Gap Measurement. 14 This equipment uses an ultrasonic

technique for determining the condition of

the gap between steam generator tubes and

support plates, which-allows monitoring

corrosion product buildup that might lead

to denting. It also can determine the efficiency of chemical cleaning for removing

this buildup..

Other nondestructive methods include sonic leak

detectors and helium leak testing to search for air

leakage into the system:

* Sonic leak detectors are used to check all

of the potential leak locations during the

period after a vacuum has been drawn on

the condenser but before the unit gets to

load. This technique is used regularly, in

finding steam leaks in boilers and has been

used on industrial vacuum systems. Its use

on power plant vacuum systems is new.

* Helium leak testing is used to search for air

leaks after the unit reaches partial load. A

load of at least 20% is necessary for efficient leak testing.

6.6 Summary, Conclusions, and

Recommendations

I

The major causes of PWR steam generator tubing failures were identified in this chapter along

with the probable causes of the identified failure

74

I
modes. Conclusions were drawn,v here applicable,

at the end of the discussion of each failure mode.

The degradation mechanisms for the steam generator tubes are prioritized based on operating

experience as summarized in Table 6.1. For example, in 1982, IGSCC was the number one cause of

tube defects, followed by IGA, and so on. To control or mitigate each failure mode, the following

recommendations are made:

I. Wastage. Wall thinning of tubes has been

reported in plants that use phosphate water

chemistry. The number of tubes affected by

phosphate water chemistry is small; this problem can' be controlled by periodic lancing of

sludge to minimize tube degradation, affected

tubes may be sleeved, or the plant changed to

AVT control.

2. Denting. Denting is no longer an important cause of tube failures. ln'1982, only

1.4%o of the defective tubes were dented,

mainly because of more stringent control

of the secondary side chemistry by reactor

operators. Some of the rmeasures that have

helped 'to control denting in older' steam

generators include sludge lancing, minimizing condenser and air leaks, eliminating copper alloys from the feed train, using

boric acid, installing deep bed polishers,

and reducing the oxygen in the condensate

storage tank. The steam generators should

be subjected to the soak procedure to

remove as much chloride as possible. In

this method, the steam generators are

soaked at about 300'F (1490C) for at least

8 h, with maximum bleed and feed, and

then the steam generators are cooled to

below 200'F (930C), drained under a

nitrogen blanket and refilled with clean

water. Newer steam generators that are better designed and equipped with more

corrosion-resistant tube supports made of

ferritic stainless steel (SS405) should not

suffer denting if good water chemistry is

maintained.

3. Pitting. Two reactors (Indian Point-3'and

Millstone-2) reported severe secondary

side pitting of their Inconel 600 tubing.

The pitting occurred in the cold leg

between the tube sheet'and the' first support plate. Ingress of brackish water, air,

and copper is believed to'have assisted the

pitting process. To mitigate pitting, remove

all copper material, minimize chloride and

air 'ingress,' change' to"AVT control, and

maintain strict chemistry control.

4. Intergranular Stress Corrosion Cracking.

The only demonstrated method to prevent

inside-surface-initiated IGSCC in an over-'

rolled tube is to peen the inside surface.

This places all of the inside surface in compression where parts of it'after over-rolling

are in tension and therefore subject to

IGSCC. A Westinghouse method called

Rotopeening has not been satisfactory in

steam generators that have been operated;

however, it has been successful in new

steam generators. Shot-peening has been

highly successful in both'new and previously operated steam generators.

The cracks resulting from inside-tubingsurface-initiated stress corrosion, caused by

over-rolling, are difficult and in some cases

impossible to detect by nondestructive'examination of the installed tubes because the tube

sheet provides interfering signals. If cracking

does occur, the tubes should be sleeved or

their lower ends replaced.

1GSCC has also been reported in'Row I

of the U-bends in'several steam generators.

Development of an in situ stress-relief procedure is recommended to mitigate such

cracking. Testing has shown that thermally

treated Inconel'600 provides better resistance to IGSCC than, ,.annealed

Inconel 600. If leaks do develop in the Ubends, they should be plugged. The use of

thermally treated Inconel 690 for tubing

for new steam generators should be considered because the available metallurgical

and corrosion data show that it is the most

corrosion-resistant Inconel tube alloy.15

The French are already using this alloy for

their steam generator" tubings.

Kraftwerkunion AG uses Inconel 800 for

steam generator tubing and shot-peens it

after bending to introduce compressive

stresses in the U-bend that eliminate stress'

corrosion cracking.

'On -the secondary side, -SCC/IGA are

widespread problems that affected 14 reactors in 1982.; This problem led to replacement

of the steam generators at Point Beach-l and

large-scale sleeving at Point Beach-2. Tube

degradation occurred within the deep tube-totube sheet crevice in 11 reactors. The shortterm solution to SCC/IGA is to install sleeves

of either thermally treated Inconel 600

75 -
Table 6.1. Summary of degradation processes for steam generator tubes

Degradation

Ranka Site

Degradation

Mechanisms

Failure

Stressors Modes

Is'

Method

Inside surface of

U-bends and rolltransition regions

2 Outside surface

of hot-leg tubes

in the tube-totube sheet crevice

region

3 Cold-leg side in

sludge pile or

where scale

containing copper

deposits is found

Tube rolling stresses,

Corion phenomenon,

denting

IGSCC Cracking Eddy-current testing

Alkaline environment,

presence of SO4 and

CO3 anions

Brackish water, air,

and copper

IGA, IGSCC

Pitting

May eventually

result in cracking

Local attack and

tube thinning may

eventually lead to

a hole

Eddy-current testing

Eddy-current testing,

optical scanner system,

sonic leak detector

system

4 Outside surface

of tubing above

tube sheet

Phosphate chemistry,

chloride concentration,

resin leakage from

condensate polisher bed

Wastage (thinning) Uniform attack, tube

thinning may

eventually wear out

the material

Eddy-current testing

5 Tubes in the tubesupport regions

6 Contact points

between tube and

- antivibration bar

Oxygen, copper oxide,

chloride, temperature,

pH, crevice conditions

Flow-induced vibrations

Denting

Fretting

Flow blockage in

tubes caused by

plastic deformation

Wear out of material

caused by rubbing

and/or fatigue

Helium leak and sonic

leak testing, optical

probes, hydrogen

evaluation monitoring,

pulse echo ultrasound

method

Eddy-current testing

Once-through steam

generator tubes

Velocities, sizes.

shapes, impact angle,

and hardness of

particles

Chemicals, (localized

corrosion) vibrations

Erosion-corrosion

from impingement

of particles

Wear out of material Eddy-current testing

8 b Once-through steam

generator tubes in

the upper tube

sheet region

Fatigue Primary to secondary

leaks

Eddy-current testing

it a. Based on operating experience for steam generator defects.

b. Denotes once-through steam generator (Items 7 and 8 do not reflect rank order). First six items are for recirculating steam

generators.

or 690. A change to AVT control chemistry

also is recommended.

5. Fretting. Seven reactors in 1982 experienced fretting damage to their steam generator tubes, caused by flow-induced

vibration. In two- cases, severe fretting

occurred at baffles in the integral preheater near the feedwater inlet. Remedial

actions included installation of modified

impingement baffles and plugging of

affected tubes. Fretting caused by antivibration bar movement in the U-bend

region was observed in other units, which

is considered a design problem in the older

I

76

I'
units. In newer units, the problem is eliminated by reducing the gap size between the

antivibration bars and the tubes..,

6. Erosion-Corrosion. Affected tubes should'

be plugged. Stringent control on water

chemistry should also be applied. -

7. Environmental Fatigue. Most of the'previously afflicted plants were partially bypassing their polisher during and shortly

before most failures occurred. Action was

taken to reduce the vibrational loads on the

tubes, and to provide strict control of water

chemistry (including restoration of lO0o

condensate demineralization) to prevent

ingress or buildup of aggressive chemicals.

Blockage of the open lane with a laneflow blocker is used now. This forces the

open lane's steam-water mixture out into

the tube bundle at a lower elevationthus,

dispersing the damaging chemicals before

they can concentrate at a weak point in the

upper tube sheet location.

To mitigate the various degradation mechanisms associated with steam generators, the manufacturers have

taken or recommended the following steps:

Change from drilled tube support plates to

other designs, such as, the quatrefoil

design in the Westinghouse Model F unit

and the trifoil design for B&W steam generators. Additionally, shifting from carbon steel tube-support material to .l2o

chromium ferritic stainless steel.

- Strict control on water chemistry on the

secondary side and a change from copper

to titanium condenser tube material.

,Wesiinghoiuse now recommends 'mostly

the AVT chemistry over phosphate treatment.,

* To control IGSCC, heat-treated

Inconel 600 or 690 now is used instead of

annealed Inconel 600.

; *The use of shot peening has been highly

successful in the U-bend and rolled areas

of the new and operating steam genera-

- tors. The fretting problem is eliminated in

the new steam generators by redesigning

the antivibration bars. The new design has

reduced the gap size between the tubes and

the antivibration bars.

There are'several issues associated with aging and

life extension of steam generators that require followup. Life-assessment models are needed to determine

the degradation of steam generator tubes by the two

' potentially most tstatus of corrosion within nuclear

steam generator tubes. The significance' damaging

mechanisms: IGSCC and IGA. Monitoring methods

' are needed to determine of the changes made in the

designs and material selectio&'of the tubes and tube

support plates needs to be evaluated. Also, useful

results are expected from the EPRI program that is

developing improved nondestructive-evaluation methods for steam generators, the USNRC steam generator

program, and similar programs in France, Germany,

and Japan.

I I

77
6.7 References

1. C. A. Creary and E. L. Houchens, "Economics of Sleeving Steam Generators," Proceedings of the

American Power Conference, 45, 1983, pp. 811-817.

2. 0. S. Tatone and R. S. Pathania, "Steam Generator Tube Performance: Experience with WaterCooled Nuclear Power Reactors During 1982," Nuclear Safety, 26, 5, September-October, 1985,

pp. 623-639.

3. 0. S. Tatone, "Worldwide Tube Performance: Analysis of the 1983-84 Statistics," Nuclear Engineering International, June 1986, pp. 81-83.

4. Steam Generator Owner Group, Steam Generator Technical Status Report, EPRI, January 27, 1982.

5. Causes of Denting, Volumes I through 5, EPRI NP-3275.

6. M. J. Angwin, Workshop Proceedings: Pitting in Steam Generator Tubing, EPRI NP-3574-SR.

7. G. P. Airey and F. W. Pement, "A Comparison of Intergranular Attack in Alloy 600 Observed in the

Laboratory and in the Operating Steam Generators," National Association of Corrosion Engineers,

Houston, Texas, March 22-26, 1982.

8. G. Pinard-Legry and G. Plante, Intergranular Attack of Alloy 600: High Temperature Electrochemical Tests, EPRI NP-3062, May 1983.

9. M. 0. Speidel (ed.), Corrosion in Power Generating Equipment, Proceedings of the Eighth International Brown Boveri Symposium, Baden, Switzerland, September 19-20, 1983, New York: Plenum

Press, 1984.

10. C. R. Wilson, Hydrogen Evaluation Monitoring as a Measure of Steam Generator Corrosion, EPRI

NP-2650, November 1982.

11. S. D. Brown, Eddy-Current NDEfor IntergranularAttack, EPRI NP-2862, February 1983.

12. R. B. Thompson and R. K. Elsley, A Prototype EMAT System for Inspection of Steam Generator

Tubing, EPRI NP-2836, January 1983.

13. D. L. Oberg, Optical Probefor Steam Generator Tube Dent Measurement, EPRI NP-2863, February 1983.

14. E. S. Furgason and V. L. Newhouse, Evaluation of Pulse-Echo Ultrasound for Steam Generator

Tube-to-Support Plate Gap Measurement, EPRI NP-2285, March 1982.

15. C. E. Shoemaker et al. (eds.), Proceedings: Workshop on Thermally Treated Alloy 690 Tubes for

Nuclear Steam Generators, EPRI NP-4665M-SR, July 1986.

78
7. REACTOR PRESSURE VESSEL SUPPORTS FOR PRESSURIZED

WATER REACTORS AND BOILING WATER REACTORS

W. G. Hopkins

7.1 Review of Types of Supports,

the Materials and

Environment

There are five major types of reactor pressure

vessel supports used in light water reactor (LWR)

plants: neutron shield tanks, columns, cantilever

supports, brackets, and skirt type supports as

shown in the Figures 7.1 to 7.5, respectively. The

first four types of supports are used to support

pressurized water reactor (PWR) vessels, while the

fifth type is used to support boiling water

reactor (BWR) vessels. The design responsibility

for the four PWR types of support varies from ven--

dor to architect/engineer, while the design responsibility of the skirt support lies solely with the

reactor vendor because it is an integral part of the

RPV. The major degradation mechanism for the

RPV supports is irradiation embrittlement caused

by neutron irradiation. Therefore,- the designs of

the RPV supports'are discussed with respect to the

susceptibility of their materials, e.g., ferritic steel

and weld metal, to irradiation embrittlement. A

detailed discussion of the damage due to irradiation embrittlement was presented in the Chapter 3

of this report and therefore, is not repeated here.

A major factor in producing nil-ductilitytransition-temperature (NDTT) shifts in support

materials is direct exposure to the reactor core

beltline neutron flux. The neutron shield tank is a

skirt-mounted, annular, water-filled tank formed

by two concentric shells extending above and below

the reactor core (as shown in Figure 7.1), and is

fully exposed to the reactor core beltline neutron

flux. WVater is circulated through the tank for cooling and shielding. Similarly, the column type support is also fully exposed to the reactor core beltline

neutron flux. The cantilever type supports may or

may not experience significant neutron exposure;

depending on a'design's specific location relative to

the core's location. The-bracket type supports are

sufficiently elevated above the region of significant

flux and are, therefore, not likely to experience

great shifts in their NDTT temperatures. The skirt

supports used in BWRs do not experience significant shifts in their NDTT because the BWR reactor

vessels are larger in size than the PWR vessels and

have much more water between the reactor core and

pressure vessel that absorbs neutrons and shields

the pressure vessel, skirt and other hardware outside the vessel. However, the BWR skirt supports

are subjected to fatigue damage when the RPVs

change temperature and pressure during startup

and shutdown.

In order for crack propagation to occur, the support member must' be in tension at some point.

With the neutron shield tank, column land skirt

type supports, tension is experienced only under

dynamic loadings, such as 'safe-shutdownearthquake and loss-of-coolant-accideni (LOCA)

loadings. The bracket and cantilever supports experience tensile forces under their deadweight loads.

The Ndrth Anna units 3 and 4 have neutron shield

tanks. In 1979,'Virginia Electric and Power Company

notified the USNRC (in accordance with 10 CFR 21)

that predicted end-of-life NDTT shifts would exceed

* the operating temperature of the supports. The

USNRC commissioned the Naval Research Laboratory (NRL) to perform a confirmatory study. In that

study, the NRL used displacements per atom (dpa) to

predict NDTT shifts. This parameter was chosen so

that embrittlement or hardening data taken from one

rieutron spectrum could be applied to the same material in a different spectrum. The end-of-life level (reactor life of 30 full-power years) of damage in dpa.was

calculated for the shield tank. The NDTT shift was

calculated based on the assumption that the elevation

in transition temperature is proportional to the square

root of dpa. From this result, it was found that the

estimated end-of-life NDTT exceeded the service temperature of the shield tank. In addition, it was predicted that the Charpy V-notch upper shelf energy level

would be reduced to 30 ft-lb (41 J).

From this study, NRL recommended to the

USNRC that additional investigation's were warranted in the following areas1:

"1. Determination of the range of radiation environment conditions (neutron spectra and flux

l levels) for other shield tank support structures

and for support structures of the open

column design (column support). Comparison of the findings with those for North

Anna Units 3 and 4, relative to potential for

notch ductility reduction.

79
Incotr Inst

Neteon Reactor vessel supporte assembly unnel transfer1 pipe ElI11.9 151161

I 2 tEI.31511816'11

I1~~~~ ~Neutron1

detector Reactor neutron shield lank Plan

sahroud 0 2 4 6 8

1 l _ ~~~Ventilation Primary shield cooler Scale (It)

manlfotd 2.999 5pmn R eanif.... EL14.10'Nom 2.998' O

0.0005 I A

(5 ) 43/8 015. ~ 14.Reactor _t24J\/Realort~ vessel

shbiding 4. 4t 12-* 8' _ 3v J8 \-'~p ~ t

471±0 w 1 UN 3A Del G

Access sil

1tank

X-X \ /support assembly D

0 2 4 6 8 0 1 2

Scat.'4 (It)Req'd Seite tt}

0123456

8 -10 34 nom- Btil o Matertials

r .4 1.2eto - Reactor neutron shield tank

t2 112- No. Name Material

Dust cover 2 * 1mC () Plate A ASTMAA51t GA 60 FBX (normalizedM

ttont ' Plate B ASTM-A516 GR 60 FBX laustemlzd)

2@ _Et8 I 3 5/16 ( . PtateC -ASTMA240Type316L(usefor7 IDtubes)

k V (D~~~~~~~~ Pipe A ASTM.A312type3l8LSST

®( PlpeC ASTM.AlO GR A or B carbon steel Iry L.,, Reactor vessel support assembly .16 1t4 Item

L_-3 .1nom No. Name Material

0D0 DBl Maraged steel 330.00 PSI min Y.R

0 1 2 3 O Ptate Maraged steel 270.000 PSI min Y.P.

Scale (ItI 4 Forging Maraged steel 270.000 PSI min YR sun

Figure 7.1. Neutron shield tank RPV support.

I

80
"6. If possible, conduct experimental irradiations, which will directly compare radiation effects of spectra in far-core versus

near-core irradiation locations to verify

the usefulness of the dpa approach for

making radiation damage projections."

If the above recommendations by NRL are followed, the estimate for the shift in NDTT will be

improved. An improved estimate for the NDTT

shift will most probably result in increased margins

for the plant life extensions.

Table 7.12,3 lists the steels most commonly used

in RPV supports. The NDTT shifts are a'functioni

of irradiation temperature and neutron'fluence

(E > 1.0 MeV). At lower irradiation temperatures

[<4500 F (2320C)] the NDTT shifts are more pronounced because of the inability of the material to

self-anneal. In addition, those steels with relatively

high copper contents will experience greater transition temperature shifts. Table 7.24 shows the computed relationship between copper content and

irradiation temperature on:the' expected NDTT

shift: The effect of neutron fluence levels at a constant irradiation temperature on the expected

NDTT shift can be seen from Table 7.3.

The typical range of service temperatures for the

RPV supports is 120 to 250IF (49 to 1210IC). The

temperature can fall to values as low as 501F (280C)

when the reactor is shut down. This low temperature,

combined with the neutron environment'and copper

content of the support material, can lead to significant

shifts in NDTT. In addition, it has been found that

when using dpa as a neutron exposure parameter rather

than fluence, (in both cases E > 1.0 MeV), greater

NDTT shifts are calculated (see Reference 5, Table 1).

Comparison with test results suggests that dpa might

be a better neutron exposure parameter to use.5

Table 7.4 is taken from examples presented in Reference 5 and illustrates the effect of copper content and

fluence versus dpa calculations of NDTT.

7.2 Review of the Design Basis

for RPV Supports

In considering plant-life extension (PLE) for the

RPV supports a brief look must first be taken at

their design basis. The General Design Criteria of

10 CFR 50 address the minimum design requirements with regard to licensing that must be met for

structures, systems, and components important to

safety. In particular, Criterion 2 states that the

design of a component should be able to withstand

Figure 7.2. Column RPV support.

"2. Development of more detailed information on the upper shelf notch toughness of

support structure materials in the transverse (weak) 'test orientation in the asfabricated condition.

"3. Development of fracture toughness and

strength assessments for low upper shelf

steels irradiated at <4501F and a correlation of fracture toughness with Charpy-V

properties in this condition.

"4. Assessments of the significance of radiation reductions in notch ductility and ele-'

vation in yield strength to support

structure fracture safe reliability.

"5. Develop, for the steels of. application,

more detailed information on the potential

for and sources of variable steel sensitivity

to <450'F irradiation.

81 ':
.0 ~~91 09

o k ~~~~~~~El. 566' L -~~~~~~.

Ad sold' >, 5'2'

@~~~~~~~~~~~~~~~- 57181rll 151!X 

4' grout

<,~~~'6 Section (E

rn support 2

El. 571' AO

tie 1 1/2' web

1 112 El. 568'-7/8'

82~~~~52

L~~~~~~~~~~~~

h-~~~~~~~~16- - 2- 6X

& ~~~~~2'0' 2'0' .

<3 ~~~~~~~~~~Section 2) 7-3071

Figure 7.3. Cantilever RPV support.

82

I'
. I ' I . : . I . i . .I

!  1 ; , -

I : I ,

I ; , i

I ;

I I

/

, Vent

opening

I . ~Reactor ..1;:

\ Anchior ~ ' bolts openings

Shear lugs

.. - -. l . ,,'.1 ' ... * -, , : 7.~~~~~73049

Figure 7.4. Bracket RPV support.

83
Control rod drive

hydraulic lines

7.3070

Figure 7.5. Skirt RPV support.

84

I
Table 7.1 Common RPV support steel characteristics 2,3,a.b

Type of Steel

A572

(normalized)

A302-B

SA533-BI

Classification

High-strength

low alloy

Low alloy (not

quenched and

tempered)

Quenched and

tempered

Initial NDTT

(°F)

-'50

0

0

Irradiationc

NDTT Shift

(OF)

85

85

I .

8.5

Minimum

Yield Strength

(ksi)

40

50

so

SA516-70 Carbon-manganese 10

A572 High-strength 25

(as-hot rolled) low alloy

A36 Carbon-manganese , 60-100

a. Private communication with P. B. Lindsay, Bechtel National, November 4, 1974.

b. Private communication with G. A. McCoy, Bechtel Power Corporation, January 28, 1976.'

c. Fast fluence (E > I MeV) = 1.23 x 101 n/cm2.

100

85

38

40

100 36

Table 7.2. NDTT shifts at a neutron

fluence of 1018 njcm2

(E> 1 MeV) for various copper'

levels, wvith 0.012% maximum '

phosphorus4

550°F Irradiation'' 450°F Irradiation

Copper ND1r Shift! NDTT Shift

(wt°0 o) (°F) (°F)

0.35 !105 .158

0.25 75 113

0.20 58 ; 87 ..,

0.10 : 25 38

Table 7.3. 'Dependency of NDTT change's

: i- ' of A302B steel upon various

neutron fluence'levelsa

.

- Neutron'Fluence ' 

*NDTT.Increase .. at 250°Fr

(O°F) (n/cm2 > 1 MeV)

- 80 1 x 1018

85. 1.25 x 1018 .

100 . 2 x 1018 i

150, I; 4.1 x IV8 ,

' ';: '170 ''''-'x5 XA1018

a. Private communication with P. B. Lindsay, Bechtel

National, November 4, 1974.

85
Table 7.4. Computed NDTT at PWR supports based on neutron fluence (E > 1 MeV) and

dpa (E > 1 MeV) 5

Fluence at

Core Midplanea

(E > I MeV)

(neutrons/cm2 )

(A)

1.08 x 1018

4.86 x 1017

3.08 x 1017

dpa at Core

Midplane

(B)

9.03 x 10-4

3.96 x 104

2.92 x 10-4

Copper

Content

(No)

0.4

0.1

0.05

NDTT `F

(A)

110

23

11

NDTT 0F

(B)

127

26

13

a. Assumed 40-year plant life and 80% capacity factor.

the most-severe-case effects of natural phenomena

recorded for the plant site and not lose the capability to perform its safety function. In addition, the

effects of natural phenomena must be appropriately combined with the effects of normal and accident conditions. Criterion 4 deals with the

environmental- and missile-design bases. It states

that the design of a component should "accommodate and be compatible with the effects of environmental conditions associated with normal

operation, maintenance, testing, and postulated

accidents, including loss-of-coolant accidents."

Also, "the components should be appropriately

protected against dynamic effects that may result

from equipment failures and from events and conditions outside the nuclear power unit."

In any combined event, the design-basis limitation and primary-failure mode is catastrophic brittle failure of the RPV support. Catastrophic brittle

failure is a result of three conditions being met

simultaneously. These conditions are (a) that a

flaw of critical size is present, (b) that there is

sufficient load on the support member to develop a

critical stress level at the crack tip of the flaw and,

(c) that the service temperature is low enough to

promote a cleavage fracture at the crack tip of the

flaw.6 The NRL has developed a method that correlates the above needed conditions for catastrophic brittle failure. The method is known as the

fracture analysis diagram (FAD) 7 and is based on a

generalized stress-temperature diagram for fracture

initiation and crack arrest.

Use of the FAD to determine fracture-safe operation

requires determination of the NDTT, which can be

accomplished by either Charpy V-notch tests or by

drop-weight tests. The NDTT is defined as the highest

temperature at which the stress needed for fracture initiation becomes the same as the stress taken from the

curve representing the' yield strength of the steel. The

nominal stress needed for fracture initiation decreases

in the presence of a flaw. As the flaw increases in size,

the nominal stress required decreases. The FAD predicts that, for a given level of stress experienced above

the NDTT, fracture initiation is possible only with flaw

sizes equal to or greater than a critical flaw size. The

critical flaw size is proportional to the square of the

stress intensity KIc. The crack-arrest temperature

(CAT) curve relates the temperature to the applied

nominal stress at which propagation of brittle fracture

is arrested. For stress levels at one half the yield strength

of a material and at temperatures 30 to 350F above the

established NDTT, it was found that there was a definite cutoff in fracture propagation. The CAT was,

therefore, found to be directly relatable to the NDTT.

Thus, the FAD method can be used to determine

fracture-safe operation based on the NDTT of the

material and expected stress levels. If the service temperature of the material in use never falls below the

NDTT plus some safety margin temperature, then catastrophic brittle failure cannot occur. Should the service temperature fall below the NDTT plus safety

margin under certain loading conditions, the determination of the critical crack size and location is required

for the analysis, using linear elastic fracture mechanics,8 to ensure continued safe operation.

Regarding the stress levels for RPV supports, the

various loading combinations used in the design of

RPV supports should be reviewed. These loading

1.

r

I.

9.1

Li

86
combinations are an explicit interpretation of the previously discussed NRC General Design Criteria 2 and

4 of 10 CFR 50 as implemented in detail by

Section III of the American Society of Mechanical

Engineers (ASME) code. 9 Table 7.510 shows the various loading combinations that must be accounted

for under the four designated design conditions.

Table 7.610 lists the appropriate parts of Section III

of the ASME code for obtaining the appropriate

stress limits allowed for each design condition.

With regard to brittle failure, special attention

must be directed to those design conditions when

the service temperature is at its lowest, i.e., during

shutdown. If the applied loads, such as deadweight

and safe shutdown earthquake, are high enough

when combined with the low temperature, then two

conditions for brittle failure are present.

Also of particular interest are the loadings associated with a postulated LOCA. It was discovered

in May of 1975 that asymmetric loadings on the

reactor vessel supports because of a postulated

double-ended rupture of the reactor-coolant piping'

had not been taken into consideration in the origiTable 7.5. Design loading combinations

for supports for ASME Code

Class 1, 2, and 3 components1 0

.. ...*Sa ...... inal design of North Anna units I and 2 RPV supports. These asymmetric loadings result from

forces induced on the reactor internals by transient

differential pressures across the core barrel and by

forces on the vessel caused by transient differential

pressures in the reactor cavity. I I Because of these

newly identified loads, the USNRC on January 20,

1978,12 required that all PWR plants evaluate the

adequacy of their reactor system components and

supports under the effects of asymmetric LOCA

loadings. Analysis of postulated guillotine breaks

has shown that the additional loads experienced by

the supports are approximately equal to the combined loads of deadweight and SSE.

In addition to the above environmental parameters that influence the fracture toughness of RPV

supports and their design lifetime, one other

important environmental parameter must also be

considered. This parameter is the degradation

effects of neutron radiation embrittlement' on the

RPV support material discussed above and in

'Chapter 3. Only the RPV and the RPV support

structures in a nuclear power plant experience this

unique type of degradation mechanism. Consequently, Section III of the ASME Boiler and Pressure Vessel Code, Part NF-2121, requires that the

materials used in reactor pressure vessel supports

"... shall be made of materials that are not injuriously affected by ... irradiation conditions to which

the item wkill be subjected." A recent report by the

Electric Power Research Institute13 has shown that

there are significant fast neutron fluences (i.e.,

> 1018 neutrons/cm2 at energies' greater -than

> I MeV) in the annulus between the outer surface

of the reactor pressure vessel and the inner surface'

of the primary shield, in the region opposite the

reactor core. As discussed previously, supports are

located within'this region in some pressure-vesselsupport system designs and are therefore subject to

shifts in'their NDTT and subsequent decrease in

fracture toughness. In addition, supports'in this

region are usually maintained'at or below '150 0F

(660C);so thai the concrete in the priniary shield

meets the concrete-codel 4 requirements for maximum temperatures. Therefore,' any inherent

annealing in the supports is nil compared'to the

annealing in ihe reactor pressure vessel that is maintained at temperatures of -5501F (2880C).'

Studies1 5 at the NRL have demonstrated the importance 'of reucing the copper content in steels to suppress both the magnitude of the NDTT shift and the

decrease in upper-level fracture toughness. The NRL

also has developed the drop-:weight test to ascertain the

true NDTT. The USNRC has incorporated the results

Design

Conditions

Normal

Upset

Emergency

Faulted

Design Loading'

Combinationsa

DW + OBE'

DW+ RVC

DW + FVDW + OBE + RVO

'DW + DU

DW + DE : -

DW

DW

DW

+ SSE + RVO

+SSE ,

+ DF

a. Legend:

DW Piping dead weight . . ::

OBE Operating basis earthquake (inertia portion)

SSE Safe shutdown earthquake (inertia portion)

FV Fast valve closure

RVO Relief valve - open system {sustained)

DU Other transient dynamic events associated with the

upset plant condition

DF Dynamic events defined as a faulted condition

87
Table 7.6. Allowable stress limits for Class 1 component supports10

Conditions

Support Type

Plate and shell

design by analysis

Linear type supports

by analysis

Component

standard

supports design by

analysis

Component

supports

design by load

rating

Design

NF-3221

NF-3231

NF-3240

NF-3260

Normal

NF-3222

NF-3231

NF-3240

NF-3260

Upset

NF-3223

NF-3231

NF-3240

NF-3260

Emergency

NF-3224

NF-3231

NF-3240

NF-3260

Faulted

- NF-3225

NF-3231

NF-3240

NF-3260

of this research into Regulatory Guide 1.99,

Revision 216 for prediction of radiation damage in

reactor pressure vessels.

With the emergence of the North Anna syndrome17

and the' additional loads it imposes on reactor vessel

support systenms, it is important that any decrease in

fracture toughness and increase in NDT1 in reactor

pressure vessel supports be assessed in PLE considerations.

Regulatory Guide 1.99, Revision 1,4- was used to

calculate the increase in NDTT and 'assesses the

potential lifetime of a RPV support in the following

example.

1. Deterrnine the actual current NDTT of the

RPV support material by way of sampling

and monitoring techniques as described

below in Section 3 orby,',ay of analytical

estimates as described in the following steps.

2. Assess the fast neutron fluence to which the

RPV support material will be exposed over its

'iormhal design' basis ' lifetime" (e.g.,

1018 neutrons/cm2 've a 40-year lifetime

with an 80% plant capacity factor. This is

equivalent to 'a fast neutron' flux of

109 neutrori/cM2-s). This can be done using

reactor physics 'calculations' as outlined in

American Society' for Testing and

Materials (ASTM) Standard E 482.!

3. Deternine the lowest temperature the RPV

support material will experience [e.g., 50F

(28 0

C)]. This temperature is usually set in a

technical specification in the license of the

facility.

4. Determine the initial NDTr of the RPV support material from the material mill certification records [e.g. -230F (-21.5 0C)J.

5. Determine the maximum loading conditions

at the lowest service temperature. This establishes what safety margin must be used based

on Pellini's FAD procedures, e.g., at half the

yield strength, the FAD indicates that the

material must be at NDTT + 300F

(NDTT + 170C). For this example, it is

assumed that the location of the maximum

tensile stress (half of yield strength) in the

support coincides with the location of maximum fluence. Then, for this example, the

maximum NDTT shift must be given by

I

I

500F - 30 0

F - (-23 0

F)

(initial

NDTT)

(lowest

operating

temperature)

(FAD

safety

margin)

Ii = 43 0

F

(maximum

NDTT shift)

88
6. Using Regulatory Guide 1.99 (see Table 7.4),

assess the NDTT shift of 251F (14'C) for the

normal operating lifetime for -the support

material's weight percent copper. However,

because the. supports are irradiated below

450'F (2320C), the 251F (141C) shift must be

increased by 50%, (i.e., the supports in the

example with 0.10°7 copper and 0.012%

phosphorus, and a fast neutron fluence of

1018 n/cm2 have experienced an NDTT shift

of 380F (20'C) over their 40-year lifetime).

7. The available PLE margin for temperatures

less than 4500F, can now be determined in the

following manner:

NDTT margin for PLE NDTrmamum

':--- ND ;TT4O years

NDTT margin for PLE = 430F - 380F

=50 F

- 8. Referring to Figure 7.6, a 30F (20C) increase

15SF (30C), decreased by 50°7 for irradiation

at temperatures greater than 450'F] in NDTT

would result in an increase of the fast neutron

fluence from the 40 year value of %l.0 x 1018

' ieutrons/cm2, to' a fast'' fluence' of'

1.2 x 1018 neutron/cm 2. Recalling from

Step 2 that the neutron flux was 109

- neutrons/cm 2-s for our example, the PLE

margin is given by i .

1.2 x JOB - I 10X1o8 PLE margin = 1

.^109.

.O- . ; 

._A~kt) ,; , , ,

recent exemption granted by the USNRC when considering loads from post-LOCA guillotine breaks.

Recent advances in fracture mechanics has shown

that there is an extremely low probability of a

double-ended :break occurrinig in PWR primary

coolant loop piping. Instead, it has been found that

the PWR primary coolant pipe is much more likely

to leak before it breaks, so that detection of small

flawvs by either in-service inspection or leakage

monitoring systems is assured. This has led to a

modification of General Design Criterion 4

Requirements by' the'USNRC. 'A statement has

been added as follows:

"... the dynamic effects 'associated with

postulated pipe ruptures of primary coolant loop piping in pressurized water reactors may be excluded from the design basis

when analyses demonstrate the probability

of rupturing such piping is extremely low

under'design basis conditions." 18

The analysis referred to 'above is outlined in

NUREG-1061, Volume 3, dated November 1984.19

Once an extremely low probability (104 per reactor

year) of rupturing is ascertained, redesign of the RPV

supports is possible. In particular, support loadings

because of the dynamic effects of postulated pipe

'breaks can be eliminated. This'resuit, in effect, extends

the lifetime of the RPV supports because the remaining loads (dead weight + earthquake) are about one

half of the total design basis loads caused by the asymmetric loads from a'LOCA. Thus, if a support steel

had been chosen with a yield strength of 100 ksi with

50 ksi ensured for NDTT + 300F (NDTT + 170C),

dropping the loading from 50 ksi to 25 ksi implies that

a larger fast neutron fluence may be tolerated.' In the

previous example, where the final NDTT must be

430 F (60C) because' of the, NDTT + 300F

(NDTT + 171C),' a new -design basis can be used of!

NDTT'+ 150F (NDTT +'91C) (for one-quarter'

yield strength from the FAD) or the final NDTT is 50'-

15°F- .c.(-231;) ='581F (14.50 C) . Compared to the

old value of 430F (61C), an additional 150F (90C) PLE

marginris thereby realized:.Again using the above

example, with a 1018 fast neutron' fluehce at 40 years,

this PLE margin of 151F (91C) can be translated from

Regulatory Guide :1.99 as before into an extended life-,

time.The 150F(90C)isonlyworth 10F(5.50C) recall

the S0%'perialty for <4500F (<2320C) irradiation];

therefore, 'a250F (141C) NDTT 1380F (201C), for

T <4500 F (2320C)] -corresponding' 'to a starting

fluence of 1018 would result in a final NDTT of 35F

(1.70 C). This corresponds'to fast'fluenice from!

Figure 7.6 of 2 x' 1018 neutron/cm2. Proceeding as

before -

= 2.0 x 10' seconds

where '' .,

- - ! ; i, -d' neutronfast flux

Ot ' = neutron fast fluence

PLE margin '= 6.3 years. ' ' '

A reexaminiation of design'-basis loads can also' -

result in a similar PLE gain. This is because of a

89
C,

0

U

a

0

0)

a

Q

C3

0

(C

0

U

0

400 11fir" 140'- 1000 (% Cu 0.08) + 5000(0/0 P -0.008)JIf/10191.1 M 11 W Will Till M11111111111i 11 [M.P111-MM U '4- -1;114114111111111 JiMMU1111i ii !W

300 lit 1.111111 111111 I IM110 11 ININ 4 14" lull III 11MI HIM I IM 6,11 R. URM1111 11 11MIM-1141 iM W1111111iii 1$11 t 11111MIt II

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i Aforl I1. II I III 11111110 0 Wli!' 4- J-44 I II IMIN1104 1111 till :17 : 1:111 1111111Di"llit1h Hill lit 111IM-1- UK 11111") DHnl I 1-1:11 .1 I111p, 11k I-1 t. air, 11111IM111-111 101 ft Lf IWIT NM 41444N 11 I

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419141;11111114= 44441-1-11 a/.p n niq TTM MIR life Lower limit

1 MHH- 1: L-r jtdf Extrapol I I dI I IT.,

2 x 1017 4 6 8 1018 ,. 2 4 6 8 1019 2

Fluence, n/cm2 (E > 1 MeV)

4 6

7.304B

Figure 7.6. Predicted adjustment of reference temperature, "A", as a function of fluence and copper content. For

copper and phosphorus contents other than those plotted, use the expression for "A" given on the figure.

PLE margin = = 2 x 1018- 101

4) 1019

= I x 109 seconds

Additional PLE margin = 32 years.

The above two examples show that certain RPV

supports may have a PLE margin of several years.

However, to develop a valid and more acceptable

approach for calculating RPV' support margins,

more research is needed to determine the NDTT

shifts :due to both low energy neutrons

(E < I MeV) and lower irradiation temperatures

[<4500 F (232 0C)]. The purpose of the above

examples was to show the importance of the loading assumptions. However, it should be noted that

the correlations for the temperature shifts in Regulatory Guide 1.99, Revision- l, are based on high

energy neutrons (E > I MeV), while the neutron

spectra in the RPV supports have a very large component of low energy neutrons (E <I MeV).

Therefore, the NDTT shifts as developed from dpa

gradients (similar to Equation 3, Regulatory

Guide 1.99, Revision 215) may be more appropriate then the use of Regulatory Guide 1.99,

Revision 1.

7.3 Recommended Strategies for

RPV Supports PLE

There are two basic options that can be simultaneously pursued for RPV supports PLE. They are

administrative relief and monitoring and sampling.

The first option, administrative relief by way of

the new GDC4 will result in the elimination of

design basis loadings because of a double-ended

break of the reactor coolant piping. As previously

mentioned, this one tactic alone provides a good

means of extending the use of RPV supports.

The second option of monitoring and sampling

allows an actual NDTT evaluation of the supports

that have experienced operational neutron radiation exposures. Once the neutron environment of

the RPV support has been assessed per the following ASTM Standard E 1035,20 the corresponding

shift in NDTT can be calculated, 5 based on the

type of material employed in the support.

In ASTM Standard E 1035, it is recommended

that all pressurized water reactors whose supports

see a lifetime neutron. fluence (E > I MeV) that

exceeds I x 1017 neutrons/cm 2or3 x 104 dpa consider monitoring as a first step in evaluating any

possible NDTT in the supports.

I

I

90

U
I

The standard outlines state-of-the-art methods

in dosimetry that are now possible to properly characterize the neutron field near the RPV supports.

In particular, it describes how scrapings from the

RPV or the supports themselves can be used as passive dosimeters through the use of a relatively new

and very sensitive technique 21 known as helium

accumulation 'fluence monitors (HAFM). In this

technique, the helium atoms released from the reaction Cu(n,a) are counted and are a measure of the,

neutron field.

Furthermore, the standard suggests the refinements in the reactor physics calculations that are

needed to properly assess the analysis of the integral dosimetry data and the final prediction ofNDTT. For example, if the vessel supports do not

lie within the core's'active height, then an asymmetric quadrature set'rmust be chosen' for discreteordinates calculations' that' will accurately

reproduce the neutron transport in the direction ofthe supports. Care also has to be exercised in constructing the quadrature set to ensure that ray

streaming effects in the cavity air gap do not distort

the calculation of the neutron transport. If the support system is so large or geometrically complex

that it actually perturbs the general neutron field in

the cavity, then the analytic method of choice is a

coupled discrete ordinates/Monte Carlo calculation. The normal coupling for this type of problem

is to perform the two-dimensional discrete ordinates analysis only within the vessel. The neutron

currents generated by this analysis are then used to

create the appropriate cumulative distribution

functions in the final Monte Carlo analysis.

After the monitoring data have been unfolded, estimates of the exposure in dpa can be made using ASIM

practices E 944 and E 1018. The dpa then can be con-'

verted to NDTT using the newest relations outlined by

McElroy.22 If these monitoring results indicate a

potential for the PLE margin, then it also may be prudent to take actual samples of the'RPV support to

confirm the'current NDTT by way of NDTT tests

(cf. ASTM E 208). The structural considerations of

removing suitable samples without 'any subsequent

repair are discussed below.

Four basic RPV support types have potential for

sampling if the reactor cavity is large enough to

allow access. These include the column, bracket,

cantilever, and neutron shield tank type (see

Figures 7.1 through 7.4). Samples are assumed to

be removed from an area adjacent to the mid-plane

of the core usually considered to be the location of

highest neutron fluence. However, it should be

pointed out that this should be confirmed by waykof

reactor physics calculations and by way of the

dosimetry monitoriig outlined above or both. This

is because new low-leakage fuel patterns may cause

the maximum neutron fluence to occur off the reactor core's horizontal midplane. Based on the above

'assumptions, a preliminary review indicates that

permanent removal of a specimen from the neutron

shield tank type is not feasible because it is a fluid

retaining vessel. -,

However, for the bracket, cantilever, or column

type supports removal of a sample may be practical. The design basis for existing RPV supports are,

in general, extremely conservative. This is, in part,

because of the unrealistic loads or load combinations that may have been assumed in their design

bases. Large design margins such as those previ-

'ously explained regarding pipe break exist, as well

as state-of-the-art refinements in analytical techniques used in the original design.:

7.4 Other Factors Affecting PLE

for RPV Supports

The above sections have discussed the primary

area of concern with RPV supports for PLE, i.e.,

the potential catastrophic failure of embrittled supports. This section will address those other. factors

that, although they are not as significant as brittle

failure, are still worth considering primarily

because their synergistic effects in a nuclear power

plant environment have not been the subject of

detailed research programs.-

I

I

* Corrosion. One of the factors present in

all LWR containments is a greater than

average humidity'a an elevated temperature coupled with the presence of a radiation environment. -The effect of the

gamma radiation can be important in

causing increased corrosion rates synergistically with the high humidity and temperatures. The long-term effects of such an

environment on the integrity of the RPV

supports should be examined because the

total gamma dose alone on the supports

will be over 5000 Megarads in the first

40 years of life. RPV support corrosion

during the original license period is not

expected to be a problem; only minor local

pitting has been observed to date.

[

r

91

I
I

* Radiation damage to nonferritic parts of

the RPV support system. Some' support

systems depend on a dry lubricant, e.g.

Lubrite, which is located between the support and the RPV nozzle. This, material

has a radiation threshold dose of

2.2 x 103 Megarads. Therefore, any PLE

plans should evaluate the actual radiation

levels the lubricant has received and will

receive over a PLE period. This evaluation

would determine if supplement or replacement of the lubricant is needed.

* Stress corrosion cracking (SCC) of threaded

parts in the sliding foot assembly. The threads

are coated with Heresite. The threshold stress

for initiation of SCC is 150 ksi while the

applied tensile stresses are calculated to be

'about 20 ksi, therefore SCC of the sliding

foot assemblies is not expected.

* Fatigue. The skirt type supports for BWR

reactor pressure vessels are subjected to

fatigue because the expansion and contraction cycles associated with the

temperature- and pressure-induced expansion and contraction of the vessels during

the plant startup/shutdown cycles. However, the fatigue usage factor during the

first 40 years of operation is expected to be

well below 1. (This 'degradation mechanism for the skirt-type supports is also discussed in Chapter 9 of this report.)

7.5 Summary, Conclusions, and

Recommendations

This chapter of the report has discussed the primary

controlling phenomenon for PLE for reactor pressure

vessel supports. Five basic types of supports used in

LWRs were examined: neutron shield tank, column,

cantilever, bracket and skirt type supports.

The major conclusion reached was that the only

important potential failure mode for the neutron shield

tank, column, and cantilever type supports is catastrophic brittle failure due to irradiation embrittlement. Table 7.7 summarizes these findings and also

points out other possible degradation mechanisms

(corrosion), stressors, and in-service inspection methods. The bracket and skirt type supports will experience no significant NDTT shifts due to their location.

Therefore, BWRs as a class of LWRs will have to

address only fatigue for PLE because they are all skirt

supported. Because some Babcock & Wilcox PWRs

are also skirt supported, they, too, should be considered to be in this class.

A method to predict PLE margins for RPV supports based on a combination of the USNRC Regulatory Guide 1.99, various ASTM standards on

dosimetry for RPVs, and Pellini's FAD was discussed in Section 7.2. In addition, it may be possible to get actual support material for NDTT tests.,

It was demonstrated that by using only the administrative relief for leak-before-break (GDC-4 of

10 CFR 50) that margins of several years could

probably be realized.

As a result of the. review conducted for this

report, the following are recommendations for further study:

1. Develop fracture toughness and strength

assessment data for the RPV support steels

irradiated at temperatures less than 450'F

(2320C). Develop a correlation of fracture

toughness versus Charpy V-notch properties at temperatures less than 450'F

(2321C) and as a function of dpa and neutron fluence (E >I MeV). '

2. Determine the range of radiation conditions (neutron spectra and flux levels) in

and around shield tanks, and cantilever

and column-type support structures.

3. Investigate the'effects of the expected radiation levels due to extended operation on

the lubricants between the RPV nozzles

and supports.

I

I

I

I

I

92

i.
I

Table 7.7. Summary of degradation processes for PWR and BWR RPV supports

Rank Degradation Sitea Stressor

Degradation ' Potential

Mechanism Failure Modes ISI Methodb

I Neutron shield tank at

the core horizontal

midplane elevation

2 Column support at the

core horizontal

midplane elevation

Neutron irradiation, ' Neutron embrittlement, Catastrophic brittle IMonitorinigC

operating temperature, corrosion failure

water chemistry L 1-, :

Neutron irradiation, _Neutr n embrittlement Catastrophic brittle Monioring and sampling

operating temperature ' ' ' I ' failure

I

3 Cantilever support in

the active height of

the core

Neutron irradiation,

tensile stresses, operating

temperature

Neutron embrittlement Catastrophic brittle

failure

Monitoring and sampling

4 Threaded parts in

sliding foot assembly

5 Dry lubricant in

sliding foot assembly

6 Skirt support

Tensile stresses, operating Stress corrosion Binding that causes

temperature cracking possibly excessive

stresses in the primary

coolant system during

heatup and cooldown.

Neutron irradiation, Degradation caused by Binding that causes

operating temperatures neutron irradiation possibly excessive

stresses in the primary

coolant system during

heatup and cooldown

' Mechanical and thermal Fatigue Ductile overload

stresses ' ' - ' ' - ' : -' failure

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a. The bracket and skirt RPV supports will experience no neutron embrittlement. a.. Th c sup ... .

b. There are no national standard in-service inspection methods per se to determine state of degradation.

c. Monitoring of neutron field near RPV support. I_

1; ;'.: "I'

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7.6 References

1. R. E. Johnson and R. P. Snaider, "Radiation Effects to Reactor Vessel Support Structures," U.S.

Nuclear Regulatory Commission memorandum to D. G. Eisenhut and S. H. Hanauer,

November 23, 1979:

2. L. E. Steele, Neutron Irradiation Embrittlement of Reactor Pressure Vessel Steels, Technical Reports

Series No. 163, International Atomic Energy Agency, Vienna, 1975.

3. G. A. Knorovski, R. D. Krieg, and G. C. Allen, Jr.; Fracture Toughness of PWR Components Supports, NUREG/CR-3009, SAND78-2347, February 1983.

4. U.S. Nuclear Regulatory Commission, "Effect of Residual Elements on Predicted Damage to Reactor Vessel Materials," Regulatory Guide 1.99, Revision 1, 1977.

5. W. C. Hopkins and NV. L. Grove, "A Study of the Embrittlement of Reactor Vessel Steel Supports,"

Reactor Dosimetry, Volume 2, J. P. Genthon and H. Rottger (eds.), Dordrecht,

Netherlands: D. Reidel Publishing Company, 1985, pp. 621-628.

6. W. S. Pellini, Principles of Fracture-Safe Design, Pressure Vessels & Piping: Design and Analysis, 1,

American Society of Mechanical Engineers, 1972, pp. 275-293.

7. W. S. Pellini and P. P. Puzak, FractureAnalysisDiagram ProcedurefortheFracture-SafeEngineering

of Steel Structures, NRL 5920, March 15, 1963.

8. G. R. Irwin, Fracture of Metals, Cleveland: American Society for Metals, 1948.

9. American Society of Mechanical Engineers, ASME Boiler and Pressure Vessel Code, Section III,

Subsection NF, July 1, 1974, Summer 1975 Addenda.

10. Bechtel National, Incorporated, Bechtel Standard SafetyAnalysis Report, Chapter 3.9, Tables 3.9-11

and 3.9-12, October 2, 1975.

11. U.S. Nuclear Regulatory Commission, Resolution of Generic Task Action Plan A-2, Asymmetric

Blowdown Loads on PWR Primary Systems, NUREG-0609, January 1981.

12. Victor Stello, Jr., letter to all PWR licensees from U.S. Nuclear Regulatory Commission Director of

the Division of Operating Reactors, Office of Nuclear Reactor Regulation, January 20, 1978.

13. M. L. Griztner, G. L. Simons, T. E. Albert, and E. A. Straker, PWR and BWR Radiation Environmentsfor Radiation Damage Studies, EPRI-NP-152, September 1977. r

14. American Concrete Institute, Code Requirements for Nuclear Safety Related Concrete Structures,

ACI 349-76, Appendix A, June 1977.

15. J. R. Hawthorne, Further Observations on A533-B Steel Plate Tailored for Improved Radiation

Embrittlement Resistance, NRL 7917, September 22, 1975.

16. U.S. Nuclear Regulatory Commission, "Effects of Residual Elements on Predicted Radiation Damage on Reactor Vessel Materials," Regulatory Guide 1.99, Revision 2, February 1986. P

17. ."Docket 50338-207 North Anna Power Station, Units I and 2," Summary of Meeting Held on

September 19, 1975 on Dynamic Effects of LOCAs, September 22, 1975.

94
I

18. U.S. Nuclear Regulatory Commission, "Modification of General Design Criterion 4 Requirements

for Protection Against Dynamic Effects of Postulated Pipe Ruptures," Federal Register, 51, 70,

April 11, 1986.

19. U.S. Nuclear Regulatory Commission, Report of the U.S. Nuclear Regulatory'Commission'Piping'

Review 'Committee,' Evaluation of Potential for Pipe Breaks, NUREG-1061, Volume 3,

November 1984. . - ;

20. American Society for Testing and Materials, Standard Practice for Determining Radiation Exposures

for Nuclear Reactor Vessel Support Structures, ASTM E 1035-85, Volume 12.02, 1985.

21. American Society for Testing and Materials, Method ofA pplication andAnalysis of Helium Accumulation Fluence Monitors for Reactor Vessel Surveillance,' ASTM Standard E 910-82, Volume' 12.02,

1982.. - - 1 - :

22. W. N. McElroy et al.;"Trend Curve Exposure Parameter Data Development and Testing," Reactor

Dosimetry, Volume I, J. P. Gentihon and H. Rottger (eds.), Dordrecht, Netherlands: D. Reidel

Publishing Company, 1985, pp. 113-135.-.

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I -~~~~~~~~~~~~~~~~

8. RANKING OF MAJOR BOILING WATER REACTOR COMPONENTS

V N. Shah

This chapter identifies the major components of

interest in commercial boiling water reactors (BWRs)

for continuing safe operation and life extension. [Pressurized water reactors (PWRs) were discussed in Chapter 2.1 In identifying these components, it is assumed

again that the lifetime of the small, less-expensive components, i.e. sensors, controls, batteries, certain types

of pumps, valves, motors etc., is much shorter than

40 years and these components will be maintained,

refurbished, and replaced frequently. Therefore, the

residual life of these small components is not significant to life extension of the plants and is not addressed

in this section.

The major components reported here are identified and prioritized according to their relevance to

plant safety only. The industry-sponsored

Monticello pilot plant project has also identified

components for plant life extension study. 1,

2 However, the identification and prioritization criteria

used by the industry pilot project were based on

plant safety, reliability, and cost.

8.1 Key BWR Components

The two criteria employed to select the key BWR

components are the same as those employed in

Chapter 2 to select the key PWR components: to

contain any release of fission products that may

take place during any accident, and to maintain an

acceptable level of radioactivity in the containment

during normal operation. The pressure boundary

components, containment, control rod drive mechanism (CRDM), reactor pedestal, and biological

shield are included in the selected key components.

Major BWR Components

Many of the reasons for the above ranking of the

major BWR components are the same as those discussed in Chapter 2 for the PWR components. However, the degradation of BNVR pressure vessels caused

by neutron embrittlement is less severe than that of

PWR vessels, and therefore it is not ranked Number 1.

Also, some of the degradation sites and the potential

degradation mechanisms of the major BNVR components are different than those identified for the PWR

components and are discussed below.

The containment and basemat are'ranked first

because they are the most critical major components,

as far as public safety is concerned. N-ost of the early

BWRs have a Mark I pressure suppression containment system. This 'contairini-nt concept features a reiatively small drywell connected by a system of vent pipes

to another containment compartment commonly

referred to as a wetwell. The other BWR containment

concepts (Mark 11 and Mark III) also have a wetwell

and drywell but are of larger size. The foundation of

the drywell of the Mark I containment is made of concrete. Two demonstration BWR plants, i.e., Big Rock

Point (71-MWe capacity) and Dresden-i (207-MWe

capacity) have spherical metal containments without

water suppression.

The potential degradation mechanisms for the

Mark I drywell metal shells are corrosion of the outside surface and fatigue. The potentially life-limiting

mechanisms for the suppression chamber are corrosion

of inside and outside surfaces and fatigue of the torus

metal shell and vent pipes.

The BWR reactor pressure vessel differs from the

PWR vessel in the way it provides support to the reactor core. In the PWR vessel, the lower core support

structure carries the weight of the reactor core and

transmits it to the reactor vessel head flange. In the

BWR vessel, the control rod drive (CRD) stub tubes

provide vertical support for the CRD housings, guide

tubes and mechanisms, and the fuel assemblies. Intergranular stress corrosion cracking (IGSCC) has been

found at the weldments joining the stub tubes to the

lower head. Also, the degradation of BWR reactor

pressure vessels by neutron embrittlement is less severe

than that of PWR vessels because there is a lot more

water between the BWR fuel and vessel because of the

location of the jet pumps. Most BWR vessels have an

expected end-of-life fluence (E >I MeV) of about

5 x 1O17 n/cm2 , while typical PWR vessels have an

expected end of life fluence of I x 1019 n/cm2 .

Therefore, vessel fluence and pressurized thermal

I

1.

2.

3.

4.

5.

6.

7.

8.

9.

10.

Containment and basemat

Reactor pressure vessel

Recirculation piping, safe ends, and safety

system piping

Recirculation pump body

Control rod drive mechanisms

Cables and connectors

Emergency diesel generators

Reactor pressure vessel (RPV) internals

Reactor pedestal

Biological shield.

96
I

shock are not significant issues in BWRs.3 However,

irradiation embrittlement cannot be totally discounted

when assessing the residual life of BWR vessels. The

new Regulatory Guide 1.99, Revision 2, takes into

account Cu and Ni content in the vessel materials, and

dictates higher transition temperature shifts than those

previously determined by Regulatory Guide 1.99,

Revision 1, for BWR vessels. Fatigue is the most likely

degradation mechanisms for BWR reactor vessels.

The feedwater nozzles of the BWR reactor vessel

constitute an important part of the primary coolant

pressure boundary. Cracks have been found in the

feedwater nozzles at Monticello and 17 other BWR

plants.4 The feedwater nozzles are subjected to relatively severe cyclic thermal loading because of turbulent mixing of incoming cold water 1340 to 4350F (171

to 2240C)] and the hot water [5450F (2850C)] returning

from the steam separators and dryers. The thermal

loading is more severe during startup and shutdown

because the incoming cold water is at a temperature of

about 1000F (381C). The main degradation mechanism is high-cycle thermal fatigue. The stainless steel

cladding on the nozzle surface contributes' to the

fatigue cracking because of the different thermal

expansion'coefficients of the stainless steel cladding

and the carbon steel base metal.

The key BWR piping components that are susceptible to aging are the recirculation piping, the

safe ends, the feedwater piping, and the core spray

spargers. After more than ten years of operation,

several piping components in the Monticello plant

have been repaired or replaced. 1 The core spray

safe end was replaced in 1981, and some recirculation piping welds were weld overlayed in 1982. The

recirculation piping and safe end were replaced in

1984. Similarly, the recirculation piping in the

Pilgrim power station has been replaced.5 Several

other BWRs have experienced similar failures at the

recirculation piping safe ends.6 The potential degradation mechanisms for the recirculation piping

and its safe ends are IGSCC, crevice corrosion, and

thermal fatigue.3 The feedwater spargers distribute

the feedwater such that the flow is evenly distributed to the jet pumps. Sparger heads have experienced cracking at the flow holes because of fatigue

and IGSCC.4 BWR spray headers also have experienced similar fatigue and IGSCC problems. For

example, the Peach Bottom spray header experienced significant cracking and has been repaired. 7

The key locations of the potential degradation sites

and potential degradation mechanisms for the recirculation pumps, CRDMs, cables and connectors, and

emergency diesel generators are likely to be similar to

those discussed for the corresponding PWR components. If there are any variations in the degradation

processes, they will be reported later.

The key RPV internal components susceptible to

aging degradation are the core shroud, jet pumps,

fuel support pieces, steam separators, core plate,

and steam dryers. The CRDM,- feedwater piping

and spargers, and core spray spargers that were discussed above also are key RPV internal components. The reactor internals provide positioning

and support of control rods, fuel rods, and other

components as in PWRs. However, the design of

the BWR internals is quite different than that of the

PWR internals. Their failure may cause binding of

the control blades, thus prohibiting their insertion

into the core. This binding -may lead to an operational transient without scram. Failure of the BWR

internals also may cause fuel rod cladding degradation, fuel relocation, wet steam because of failure

of dryers, and loss of forced recirculation flow

because of failure of the jet pumps. Type 304 stainless steel and Inconiel 600 are 'the principal materials of the reactor internals that see high fluences.

Neutron irradiation of Type 304 stainless steel will

increase its yield and tensile strengths, decrease its

uniform elongation, cause relaxation, lead to

IGSCC, and reduce the Charpy impact upper shelf

toughness. Neutron irradiation also reduces the

low-cycle fatigue life of Type 304 stainless steel

components. Among all the reactor internal components, the core shroud is most susceptible to

thermal shock following a design-basis accident.3

The main degradation locations in the jet pumps

are the holddown beams, restrainer brackets, and

sensing lines.8 The holddown beam is fabricated.

from heat-treated Inconel 600 and has failed in several BWR plants because of lGSCC.9 The restraining brackets have failed because of improper

assembly, and sensing lines have failed because of

vibrations. The jet pumps have been successfully

repaired under field conditions, and therefore, the

degradation of their subcomponents may not have

a significant impact on the life extension of BWR

reactor internals. The failure of the steam separators and steam dryers will result in poor quality

steam passing through the steamlines to the turbines but probably will have no safety significance.

The wet steam may cause erosion and corrosion in

the steam piping and the first stage of the turbine

blades.10 Other potential degradation mechanisms

for the BWR internal components are low-cycle

fatigue, stress-corrosion cracking, and high-cycle

fatigue.

The reactor pedestal and biological shield are made

of concrete. The key locations and potential

TI

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97
I

degradation mechanisms are the same as those for the

concrete in the PWR containment structures. However, the reactor pedestal is subjected to high-cycle,

low-amplitude vibration and thermal cycling. An

additional loading on the biological shield is gammaray heating resulted from the proximity of this component to the core.

8.2 Summary, Conclusions, and

Recommendations

The major BWR components were selected and

ranked such that any release of fission products

during an accident will be properly contained. The

containment and basemat have been identified as

the most important safety components in BWR

plants. This means that a failure of the containment and basemat will have a major impact on the

safety of the power plant. The other key components, according to their ranking, are reactor pressure vessel, recirculation piping and safe ends,

recirculation pump body, CRDMs, cables and connectors, emergency diesel generator, RPV internals, and RPV supports and biological shield. The

results presented in this chapter are based on past

experience, as reported in the available literature.

These results will be updated as more data are made

available.

Table 8.1 summarizes the aging related information on the key BWR components discussed in this

chapter. It lists the components and gives reasons

for their ranking. Table 8.1 identifies the most

likely degradation site and other degradation sites

for each component and also lists the most likely

degradation mechanism and other potential degradation mechanisms.

The major components reported in this chapter

are selected and prioritized according to their relevance to plant safety only. These criteria are different than the ones used in the industry pilot study.

Therefore, the ranking of the major components

presented in this chapter is somewhat different than

the one in the pilot study. However, all the major

components identified in these chapters are also

ranked high in the pilot study. Detailed discussions

of the degradation processes for BWR pressure vessels and recirculation piping are presented in Chapters 9 and 10, respectively. Discussion of the

degradation processes of both PWR and BWR

RPV supports was presented in Chapter 7.

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Table 8.1.. Key BWR components for residual life assessment

Degradation Sites Degradation Mechanisms

Rank Component Reasons for Ranking (most likely, others) (most likely, others)

I Containment and

basemat

Public protection during

an accident

,I

Concrete, metal liner, and Environmental degradation of

reinforcing steel in suppression concrete, fatigue and corrosion

pool and primary containm'ent of metal containments

wall, base metal, welds and vent

pipes in metal containment

2 Reactor pressure

vessel

Primary pressure boundary, Nozzles, closure studs, beltline Thermal fatigue, neutron

poor operating experience region, welds at the stub tubes embrittlement, IGSCC

with RPV nozzles

3 Recirculation

piping, safe

ends, and safety

system piping

Primary pressure boundary, Safetends, austenitic stainless

relatively poor operating steel fittings

experience

IGSCC, thermal aging

4 Recirculation

pump body

Primary protection from

any internal failure of

pump elements, primary

coolant pressure boundary

Heat-affected zones near

weldments in the wall

Thermal aging, corrosion

fatigue, crevice corrosion

5 CRDM

6 Safety-related

cables and

connectors

Failure may lead to a

reactivity-initiated

accident, an anticipated

transient without scram,

or a small-break loss-ofcoolant accident

Active during normal

operation in mitigating

operational transients and

accidents

Drive rod assembly, control rod Wear and thermal embrittlement

pressure vessel

[

Cable insulation, inserts in

connectors

Thermal aging of insulation,

thermal embrittlement and

corrosion ofconnectors

E

7 Emergency diesel

generator

8 Reactor

internals

Needed to operate critical

safety equipment in the

event of a loss-of-offsite

power

Failure may cause fuel

failure or problems in

scraming the reactor

Governor in the instrumentation Fatigue and vibrations

and control system, cooling and

lubrication system pumps and

piping, fuel injector pumps,

turbocharger, generator

windings

Core shroud, top guide plate, Irradiation-assisted stress

core plate, holddown beam in a corrosion cracking,

jet pump, feedwvater and core intergranular stress corrosion

spargers cracking, fatigue

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9 Reactor pedestal Failure will challenge

integrity of RPV

Concrete and reinforcing steel Thermal cycling of concrete,

corrosion and fatigue of

reinforcing steel

10 Biological

shield

Maintains acceptable level

of radioactivity in

containment

Inside surface of the biological Neutron irradiation, gamma-ray

shield at the core horizontal heating, environmental

midplane level degradation of concrete,

corrosion of reinforcing steels

I.

99
8.3 References

1. Nuclear Plant Life Extension, EPRI/DOE Workshop, Dearborn, Michigan, August 1985.

2. Seminar on Nuclear Power Plant Life Extension, Seminar Proceedings, cosponsored by EPRI, Northern

States Power, U.S. DOE, and Virginia Power, Alexandria, Virginia, August 25-27, 1986.

3. Northern States Power Company, Updated Safety Analysis Reportfor the Monticello Plant, April 30, 1982.

4. R. Snaider, BIVR Feedwater Nozzle and Control Rod Drive Return Line Nozzle Cracking, NUREG-0619

(issued for comments), April 1980.

5. L. C. Chan and G. H. Hart, "Experience of Lowering A High Drywell Air Temperature at the

Pilgrim Nuclear Power Station," International Meeting on Nuclear Power Plant Maintenance,

Salt Lake City, Utah, March 23-27, 1986.

6. R. A. Carnahan and R. MN. Horn, Evaluation of Safe-End Weld Materials and Safe-End Replacement

Experience, EPRI NP-4443, February 1986.

7. N. P. Alexakos, "Peach Bottom Atomic Power Station U/3 Core Spray Header TEE Box Bracket

Installation," International Meeting on Nuclear Power Plant Maintenance, Salt Lake City, Utah,

March 23-27, 1986.

8. J. E. Charnley, "Repair of BWR Jet Pumps," InternationalMeeting on NuclearPolverPlant Maintenance, Salt Lake City, Utah, March 23-27, 1986.

9. R. S. Mills et al., Closeout of LE Bulletin 80-07: BWR Jet Pump Assembly Failure, Parameter, Inc.,

Elm Grove, Wisconsin, for the U.S. Nuclear Regulatory Commission, NUREG/CR-3052,

November 1984.

10. G. A. Delp, J. D. Robison, and M. T. Sedlack, Erosion/Corrosion in Nuclear Plant Steam Piping:

Causes and Inspection Program Guidelines, EPRI NP-3944, April 1985.

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9. BOILING WATER REACTOR PRESSURE VESSELS

W. R. Mikesell and W. L. Server

9.1 Description

This discussion for boiling water reactor (BWR)

pressure vessels is not as detailed as that for the

pressurized water reactor (PWR) pressure vessel

presented in Chapter 3. However, the same important degradation mechanisms (namely fatigue and

irradiation damage) are considered for the BWR

vessels, as will be discussed. Primarily, the recollections and experience of one of the authors are presented; therefore, some of the details and specifics

should be considered as close approximations.

BWR vessels have been manufactured throughout

the years with the same basic geometry shownrin

Figure 9.1. They consist of a cylindrical shell and a

hemispherical bottom head containing both the control rod drive (CRD) penetrations and the in-core

instrumentation penetrations. The top head also is

hemispherical and includes nozzles with bolting

flanges attached. One of the nozzles is for venting

(pressure relief) and top head cooling spray, and the

others are spares. The top head is bolted to the cylindrical shell by means of ^.-in. (150-mm) diameter studs.

The flanges are butt welded to the shell and to the head

and are furnished in one piece. The earliest BWRs were

of several sizes/geometries and represented demonstration and small power reactors (i.e.,'Dresden-1,

Big Rock Point, Humboldt Bay, and La Crosse). The

early commercial-generation BWRs consisted of three

sizes (not including Duane Arnold which was smaller).

These sizes were \.201-in. (5.11-nm), 218-in. (5.54-r),

and 251-in. (6.38-m) diameters. For example, the

Vermont Yankee and Monticello vessels were 201-in.

(5.11-m) diameter, Zimmer and Brunswick-I and -2

were 218-in. (5.54-m) diameter, and Quad Cities-I and

-2, Susquehanna-l and -2, and Limerick-] and -2 were

all 251-in. (6.38-m) diameter. These are just examples

and do not include all of the plants of any given size.

The latest BWR model, which was designated as;.

BWR/6, was built in two sizes. These were 218-

(5.54-m) and 238-in. (6.05-m) diameters. Examples of

plants of these sizes are River Bend-l and Clinton-l for

the 218-in. (5.54-m) diameter and Grand Gulf- Iand

Perry-l for the 238-in. (6.05-m) diameter. Quite a few

of the larger diameter vessels were constructed for utilities in this country (many of which have been cancelled), while the other 218-in (5.54-m) diameter vessels

were built for overseas plants (i.e., Laguna Verde and

Kuo Sheng).

The predominant vendors of BWR vessels in this

country have been Babcock & Wilcox (B&W), Combustion Engineering (CE), and Chicago Bridge and

Iron (CB&I). 'One foreign-produced vessel from

Babcock/Hitachi was fabricated for Hope Creek. One

of the vessels finished by CB&I was partially fabricated

by Rotterdam Dockyard. Rotterdam Dockyard did

have a contract to construct one of the vessels for the

Black Fox units that were cancelled. In total, CB&I

built most of the BWR vessels used in the

United States.

All of the major nozzles of the vessel where placed in

the cylindrical shell, but they were 'not located at the

elevation of the reactor core (see Figure 9.1). The nozzles in the lower portion of the shell were, in general,

limited to the primary coolant recirculation inlet and

outlet nozzles which were located in a circumferential

band a short distance above the attachment of the

cylindrical shell to the bottom head. Nozzles above the

core region were designed for core spray, feedwater, and

the steam outlet. The largest nozzles were the two for

the recirculation outlet lines (primary coolant outlet) at

the bottom part of the shell; they were typically 26 in.

(660 mm) in diameter. Few, if any, of the other nozzles

exceeded 12- to 15-in. (305- to 380-mm) diameters.

All the nozzles in the vessels.built by CB&I were

forged and provided a lip at the shell end of the forging

to permit butt welding of the nozzle to the shell. Reactor vessels furnished by at least one other fabricator

included nozzles which were forged but did not contain

a lip so that the attachment weld of the nozzles to the

shell was a corner weld.

The earliest BWR reactor vessel was Dresden-I. The

plates used in its construction were procured to the

requirements of American Society for Testing and

-Materials (ASTM) A300 for fracture toughness

(Charpy V-notch) testing and American Society of

Mechanical Engineers (ASME) Code Case 1280 which

provided requirements for fine grain melting. Before

__the introduction of the SA533 Grade B, Class I

(SA533B-I) plate specification, SA302 Grade B

(SA302B) plate was used. All vessels since about 1965

have been constructed of SA533B-I material (this

includes the plates used in the shell and in both heads).

It appears that all BWR vessels have been constructed

using plates for the shell and the heads. That is, there

have been no forged rings, nor forged head segments,

used in this country.

The forging materials used for the nozzles in

Dresden-l were A105 11 (equivalent to present A105).

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U.

101
I

Vent and

top head

spray _ ,Vessel head

Closure studs

V

, Main steam

outlet

r

Jet pump/

recirculating

water Inlet ,

Skirt

supports -

Control rod

drive -

Recirculating

water outlet [

-Shield wall

E.

1.

7.3045

Figure 9.1. lypical BWR pressure vessel.

102
I

Probably, this material was used for most of the

SA302B vessels, before the introduction of the SA508

specifications. All of the more recent vessels have used

SA508, predominantly Class 2, (SA508-2) although

some Class 3 material (SA508-3) has been used.

Before the introduction of the SA540 bolting specification, the studs were furnished to an A193-B7 specification. Subsequently, the studs have been SA540, B22,

or B23 for the 130 ksi (900 MPa) yield strength level. It

is possible that some studs were furnished at the

140 ksi (965 MPa) yield strength level.

Until about 1974, the CRDs were attached to a stub

tube that was welded to the inside of the bottom head

of the reactor vessel. Until about 1966 or 1967, the stub

tubes were made of stainless steel. Thereafter, they were

made of Inconel (Alloy 600). From about 1974 to the'

present, the CRDs were welded directly to the bottom

head without an intervening stub tube. The in-core

instrumentation penetrations were always welded

directly to the bottom head. The vessel is supported by

a skirt welded to the bottom head just below the

attachment of that head to the cylindrical shell, as is

shown in Figure 9.1. The design of that skirt in terms

of load-carrying capability is governed by the weight of

the vessel and its contents and includes consideration

of the operating basis earthquake loadings.

As mentioned earlier,'all fabricators made the heads

and cylindrical shells of BWRs with formed and

welded plate. CB&I purchased the heat-treated plate

from the mill, and then cold formed and welded it

together. The welding used proprietary welding filler

metals that generally matched the chemistry and properties of the base material. The other primary fabricators of BWR vessels in this country, B&W and CE,

purchased unheat-treated materials from the plate

mills, hot formed the material, electroslag welded the

joints, and heat treated the complete assembly after the

rings were complete. This heat treatment consisted of

austenitizing followed by tempering. All vessel maiiufacturers welded the circumferential seams by the submerged arc process and completely stress relieved the

vessel in accordance with ASME code requirements

after the welds were made. In some cases, an intermediate postweld heat treatment may have been 'made for

individual ring assemblies before joining' them in a

complete vessel.'

The earliest vessels were designed to the requirements of the ASME Code, Section 1, and the 1270

series of code cases. All vessels manufactured after

1963 were furnished to the 'requirements 'of

Section III of the ASME Code. The mechanics of

making the design and satisfying the code analytical requirements are pretty much uniform and conventional regardless of the vendor. Of course, the

analytical methods did change as available computer programs became more developed. An example of this evolution would be the evaluation of the

discontinuity analysis'from the old beam on an

elastic foundation analysis to computer programs

for closed form numerical integration of the shell

equations and finite clement methods.

Problems that have occurred throughout the years

have resulted in changes to both design and materials.

For instance, the safe ends on the nozzles of BWRs

were originally forged of austenitic stainless steel and

stress relieved after welding to the vessel. The occurrence of intergranular stress corrosion cracking

(IGSCC) in the primary coolant lines of BWR plants

led to the removal of these postweld heat-treated austenitic safe ends' and replacement with safe ends made

of austenitic materials that had not been stress relieved.

Additionally, some safe ends were fabricated of

Inconel. Similarly, all of the nozzles of BWRs were

originally stainless steel overlayed. However, thermal

fatigue cracking as 'experienced at Millstone-I as a

result of bypass flow and a loose thermal sleeve'in the

feedwater line. Subsequently, the BWR vessel nozzles

have not been overlayed.

There also has been an evolution in the design of the

attachment to' the inside of the nozzles for thermal

sleeves and the recirculation inlet lines. As mentioned

above, there was cracking at Millstone-l because of

bypass of feedwater flow around the safe end of that

nozzle. The thermal sleeve was not integrally attached

to the nozzle. All of the lines in these vessels have since

been repaired to provide for a welded attachment to

preclude'this problem. Similarly, the attachment to the

inside of the recirculation inlet nozzle provides flow

from the inlet line to the jet pump risers. Initially, these

jet pump riser lines were close-fitted to the inside of the

nozzle and welded at the end (a detail that was a combination of a groove and a fillet weld). :However, stress

corrosion cracking of the safe end (that is actually

where these lines are welded) at the Duane Arnold

plant created a concern about IGSCC of Inconel anrd

crevice conditions. Therefore, most of the safe ends

have been replaced, eliminating this crevice condition.

9.2 Stressors

The basic design requirements for the BWR vessel were'1250 psi (8.62 MPa)'pressure and v5501F

(2880C) temperature. The'actual operating conditions at steady 'state are around 1,000 psi

(6.89 MPa) and 5401F (2820C). The design specifications included the transients necessaiy to heatup

and cooldown the vessel from normal operating

I

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conditions and a variety of normal and abnormal

conditions that might be expected throughout the

life of the vessel. The magnitude of the changes

were provided as were the rate at which they

occurred. Pressure loading in excess of 1250 psi

(8.62 MPa) design level occurs during safety valve

discharge when the pressure is permitted to rise to

1375 psi (9.48 MPa).

The skirt support for BWR vessels is shown schematically in Figure 9.1. One of the key design considerations here is that of earthquake loadings.

Additionally, the attachment of this skirt to the bottom

head requires a degree of refinement beyond that normally encountered in pressure vessel construction

because of the necessity for meeting the code stress

limits for the primary plus secondary stresses and

fatigue stresses that occur as a result of the pressure/

temperature cycles at that intersection.

All of the nozzles were designed to carry specified

pipe loads and the bimetallic or trimetallic safe end

attachment welds to the nozzle forging area were analyzed for differential thermal stresses, as well.

The core support structure in a BWR consists of an

annular plate welded at its outer periphery to the bottom of the cylindrical shell immediately above the shell

to the bottom head attachment weld. The inner periphery of this annular plate is supported by a series of

heavy Inconel posts or columns. The attachment of the

outside periphery of the annular plate and the bottom

of the Inconel columns (at the attachment to the bottom head) is analyzed in detail for all mechanical, thermal, and pressure conditions. Likewvise, the

attachments at the upper portion of the cylindrical

shell for the core spray spargers and the feedwater

spargers are thoroughly analyzed for specified

mechanical and thermal loads.

9.3 Degradation Sites

As mentioned earlier, three of the highest ranking potential degradation sites have been eliminated in the newer plants by design changes and

repaired in most of the older plants' These include:

austenitic stainless steel safe ends that have been

sensitized by postweld heat treatment, crevice areas

where thermal sleeves are attached to Inconel or

stainless steel safe ends, and austenitic cladding on

the inside corners and surrounding regions of the

nozzles (that experience fairly wide temperature

swings); Another possible area of concern that has

been eliminated in later design modifications is the

CRD penetration stub tube to vessel weld, heataffected zone; earlier plants that used stainless steel

stub tubes were susceptible to stress corrosion

cracking at this weld (leading to through-wall leakage). Failures did occur at two plants. The most

recent BWR-6 reactor pressure vessels do not have

stub tubes and the guide tube is welded directly to

the bottom head. This decreases the number of

potential degradation sites.

Regardless of whether the above design changes

have been made, the ranking of potential degradation sites would still vary from plant to plant,

depending on how they are operated. For example,

one of the more critical items would be the closure

head studs at the main closure flange. Because the

greatest range of stress amplitude in the studs is

produced during installation and removal of the

head (and the use of stud tensioners associated with

those operations), the number of times these occur

in a specific plant would have a great effect on

where the studs fall in the ranking of degradation

sites. Thus, real operating history rather than

design numbers and magnitudes of transients is

necessary in assessing critical degradation sites.

The threaded stud holes in the flanges without

bushings are subjected to wear because the flange

material is considerably softer than the studs. In

flanges with bushings, the threads for the closure

studs and flange bushings could be subjected to

fretting. However, the studs are removed from the

flange infrequently. In addition, the use of stud

tensioners for tightening makes fretting during this

operation unlikely.

The design of the flange assembly for most, if

not all, BWRs includes provision for sufficient preload to prevent slippage between the flange surfaces

during operation. Therefore, the flange surfaces

are unlikely to suffer physical damage except from

mishandling during removal and placement of the

top head. The O-rings are placed in the O-ring

grove between the flange surfaces to provide sealing. The O-rings have a soft metal layer on the outside to enhance sealing and reduce wear, and are

intended to be replaced when the head is removed.

Another important potential degradation site is

the outer end of the nozzle forgings, and the

attached safe end, for those nozzles that have water

with large temperature variations flowing through

them. The most severe nozzle of this type is the

feedwater nozzle. The specified flow temperatures

for this nozzle go from essentially the freezing

point of water to tfie operating temperature of the

vessel. The best operating example of problems

caused by flow through this nozzle is Millstone-I

where nozzle corner cracking was found. The

bypass of the feedwater around the thermal sleeve

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I

at this nozzle resulted in a temperature swing of

,\.20 0'F (I 10C) at approximately 1-Hz frequency

developing high-cycle fatigue cracking. Of course,

the upper end of that temperature swing was when

there was no bypass flow; therefore, this temperature range was larger than the variation in the line

temperature itself.

In a general sense, all locations of dissimilar

metal welds will be most sensitive to thermal

cycling. The effect of this thermal cycling will be

compounded by any mechanical loads occurring at,

the same location. Therefore, the most important

areas in this regard would be the nozzle safe end

attachments and the attachment welds to the pressure boundary for the core support structure.

The welds and the base metal in the reactor

beltline region are the main sites subjected to radiation embrittlement. It should be noted that repair

welds formed during construction and fabrication

are also present sometimes in the beltline region of

BWR vessels.

External attachments such as the support skirt and

the refueling bellows skirt are subjected to cyclic loading caused by differential radial thermal-induced

expansion during startup/shutdown. However, neither

of these locations are fatigue limited, i.e. fatigue usage

factors for these locations are far smaller than 1.0.

Rather, the limiting condition is primary plus secondary stresses being controlled by the 3-S. limit. There

are some corner weld joints in the refueling bellows

skirt assembly that may be fatigue limited for some

vendor's designs.

9.4 Degradation Mechanisms

As indicated in the previous discussion, the primary degradation mechanism will be fatigue

(induced through a combination of thermal and

mechanical cycling). This cycling also has an effect

on the main closure studs and is compounded by

the additional cycles of tensioning and untensioning during top head removal. The threads for closure studs, flange bushing, and stud holes in

flanges without bushing are subjected to wear, fretting, and corrosion, as discussed earlier. External

attachments such as the support skirt and the refueling bellows are subjected to cyclic loading, as discussed in the preceding section.

There should not be any locations in a BWR vessel except the earlier ones having stainless steel stub

tubes, where IGSCC should be a problem, because

most (if not all) of these problem areas have been

previously identified and modified. However, continued monitoring 6f these same areas should be

maintained to ensure'against :any reoccurrence.

Water chemistry control used to mitigate the potential for IGSCC in the recirculation piping also

would reduce any likelihood for problems in the

reactor pressure vessel.

' ' All of the welds in the ferritic steel components

have been postweld heat treated. Therefore, residual stress effects should not contribute to degradation. However, the welds joining Inconel, stainless

steel,' and the two materials to each other have not

been stress relieved. An exception to this last statement would be the build up of these materials that'

is made for the attachment of Inconel (or stainless

steel) to the ferritic material. These locations would

be at the outboard end of the nozzle forgings and

where the core support structure is attached. How-

, ever, even in the locations where some residual

stresses may exist, there is no need to be concerned

with excessive fatigue damage.

Radiation embrittlement is not as severe a problem with BWR vessels as it is with PWR vessels (see

Chapter 3). The fairly large diameters of the BWR

vessels, and the amount of water shielding the shell

from the reactor core, help alleviate this problem

for BWRs. However, radiation damage can eventually become a limiting factor, depending on material chemistries and actual accumulated fluence.

Most BWR vessels have an end-of-life fluence

(E>1 MeV) of about 5 x 1017 n/cm2 , although

some of the older vessels can be a bit higher; typical

PWR vessels have end-of-life fluences above

I x 1019 n/cm2 . However, irradiation embrittlement should not be totally discounted when assess:

ing life extension of BWR vessels. From an

operational standpoint, the effects of radiation

damage can be more of a nuisance. For example,

with the new Regulatory Guide 1.99, Revision 2

(discussed in detail in Chapter 3), hydrotesting of

BWR vessels after experiencing moderate radiation

exposure becomes more difficult because of the

higher temperatures required for the hydrotests.

However, the data used to develop the Regulatory

Guide 1.99, Revision 2 strongly dictate higher

transition temperature shifts than those previously

used for BWR vessels at these low fluence levels.

9.5 Potential Failure Modes

Because pressurized thermal shock and low-energy

ductile tearing are extremely unlikely in a BWR vessel

because of the low neutron irradiation conditions in

the beltline region, the only final failure mode is plastic

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I

overload leading to a leak in a region where a crack

exists (or failure of a stud). Thus, the conditions leading to this final failure mode are fatigue or corrosion

induced. All parts of the reactor vessel for which cyclic

loading has been specified have taken fatigue into

account and have fairly low usage factors (almost

always below 0.6 over the intended 40-year life). However, as indicated in Chapters 3 and 5, the calculation

of fatigue usage factors can vary and exact comparison

of numbers can be misleading. Consideration of

abnormal loading conditions that currently would be

considered as Levels C or D (emergency or faulted)

did not control any of the design features of a BWR

vessel. In the earlier BWR vessels with stainless steel

stub tubes, IGSCC-induced failures have resulted in

leakage through the CRD penetrations.

9.6 In-service Inspection and

Surveillance Methods

The PWR vessel in-service inspection (ISI) methods and requirements of Section XI of the ASME

Code were presented in Chapter 3 and are the same

for BWR vessels. However, many older BWRs have

very limited accessibility for external ISI of the vessel. Typically, 75 to 90%o of the vessel weld lengths

are exempted because of inaccessibility. The only

alternative is ISI methods of examination from the

inside surface. This is of particular importance at

the beltline welds. There were minor and major

repairs to shell plates during construction but some

of these cannot be inspected because of limited.

accessibility. This is also true for some of the pipeto-nozzle welds that were not configured to facilitate ISI. This was changed in later reactors after the

requirements of ASME Section XI were published.

Surveillance for irradiation embrittlement is dictated by federal law for BWRs as well as PWRs.

Therefore, monitoring of actual changes in Charpy

V-notch and tensile properties with regard to accumulated fluence for the most critical vessel materials is underway.

9.7 Summary and Conclusions

A summary of the degradation sites and mechanisms, stressors, potential failure modes, and ISI methods for BNVR vessels is presented in Table 9.1. Most of

the extensive conclusions listed in Chapter 3 also are

applicable here. The main difference is the much less

important effect of irradiation embrittlement for BNVR

vessels. However, this damage mechanism is present

and should never be ignored.

Other conclusions worth restating here are: nozzle fatigue usage should be monitored more closely

with regard to refined cycle counting and transient

severity so that better estimates of actual fatigue

usage factors can be determined, the closure studs

should be examined closely for replacement near

the normal 40-year end of life, and the largenozzle, CRD, and other small-nozzle penetration

welds should be carefully inspected.

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Table 9.1. Summary of degradation processes for BWR reactor pressure vessel.

Rank of

Degradation Site

Degradation

Site

- . ., : . I .~ .

Degradation Potential Failure

Stressors Mechanisms Modes

* ISI/Surveillance

Methods [

Nozzles (including

instrument and CRD

penetrations) plus

safe end welds

2 Closure studs,

flange bushings,

stud holes

3 Beltline region

4 External attachments

Mechanical

and thermal

stresses

Fatigue crack

initiation and

propagation,

IGSCC

Ductile overload I All large nozzle welds

l Ieading to a leak . inspected volumetrically

at each interval; visual,

external surface

inspections of small

- nozzles/penetrations

Mechanical Fatigue crack

and thermal . initiation and

stresses ' propagation,

fretting,

corrosion

Irradiation Neutron

embrittlement irradiation

(extent

depends on

vessel

materials)

Ductile overload

failure (can be

replaced)

Ductile high-energy

tearing leading to a

leak (not a serious

problem):

Ductile overload

failure

Volumetric and surface

inspections of all studs,

threads in flange stud

holes and bushings

100 volumetric

inspection; surveillance

as required by federal

law

Volumetric and surface

inspections

E

Mechanical I I

and thermal

stresses : .

Fatigue

F

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I

10. BOILING WATER REACTOR RECIRCULATION PIPING

H. Mantle

There has been an increasing incidence of intergranular stress corrosion cracking (IGSCC) failures

in boiling water reactor (BWR) recirculation piping

systems in the past 10 years: first in highly stressed

small-diameter piping and later in large-diameter

piping (12 in. and greater). While an actual pipe

severance has never occurred, the problem has

impacted plant availability and resulted in a substantial increase in radiation exposure to maintenance and inspection staff. Although IGSCC has

been the Achilles' heel of BWR plants, great strides

have been made in understanding the conditions

under which it occurs and in developing both shortand long-term solutions.' Other concerns that may

affect life extension have been raised (such as lowtemperature aging of duplex stainless steel castings,

crevice corrosion, and thermal fatigue), but these

have not yet been shown to be significant problems

in BWR recirculation piping systems.

10.1 Description

The recirculation system draws coolant from the

bottom of the reactor vessel through an outlet nozzle and downcomer spool into the recirculation

pump. The pump discharges this coolant into a

ring manifold that feeds the coolant by way of multiple risers to the associated inlet nozzles and jet

pumps, thus completing the loop. The reactor

recirculation system consists of two independent

loops. Primary piping in each recirculation loop in

a large BWR reactor (21100 NIWe) consists of a

28-in. pump suction line, a 28-in. pump discharge

line, a 22-in. pipe riser manifold, and five 12-in.

taps (ten total taps, five for each loop) to the ten jet

pumps inside the reactor. A simple schematic of the

basic arrangement is shown in Figure 10.1. Key

components include safe end nozzle welds, suction

and discharge valves, the recirculation pump, and

the ring manifold.

Historically, BWR recirculation piping systems

have been fabricated from wrought, AISI types 304

and 316 (piping) and cast types, CF8 and CF8M,

(pumps and valves) stainless steels. Replacement

piping has been made from Type 316 NG (nuclear

grade) that has been shown to be much more resistant to IGSCC than the conventional grades.

10.2 Stressors

Stressors include welding heat (sufficient to produce a sensitized microstructure), tensile stresses

(both residual and applied), oxygenated cooling

water, temperature and cyclic thermal stresses.

These factors, either in combination or separately,

can produce IGSCC, thermal fatigue, thermal

embrittlement, and crevice corrosion.

10.3 Degradation Sites

Degradation sites include weld heat-affected zones

and furnace sensitized components (safe ends), weld

metal, duplex stainless steel castings (elbows, valves,

T-joints), areas in the system with high thermally

induced stresses predicted by stress rule index analysisI

(nozzles, safe ends), and crevices at safe ends shielded

from the bulk fluid environment. Additionally, based

on the interpretation of the North American Electric

Reliability Council's 1984 Component Cause Code

Report, 2 the recirculation pump also is a primary degradation site (see Table 10.1). The report data indicate

that the recirculation pump and driver have a greater

outage frequency than the reactor coolant system

valves and piping. The cause for greater outage frequency is not reported in Reference 2 and is probably

not related to a loss in pressure boundary integrity.

10.4 Degradation Mechanisms

As discussed in previous surveys of aged nuclear

power plant facilities,3 degradation mechanisms

most responsible for current aging-related failures

can broadly be categorized as erosion, corrosion,

vibration (fatigue), and foreign material related.

Within the recirculation piping system of BWRs,

failures (leaks) have primarily been related to

IGSCC. Of secondary concern are crevice corrosion, low-temperature aging (embrittlement) of

duplex stainless steel, and thermal fatigue.

Degradation resulting from IGSCC has essentially been controlled by way of a multidirectional

effort, because IGSCC of austenitic stainless steels

requires the combined existence of the following

three conditions:

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'- Pump

discharge valve

74048

Figure 10.1. Schematic of a BWR recirculation piping system.

Table 10.1. Outage frequency: valves and piping versus pumps and drivers

(years 1975-1984, 29 units, 230 unit years)

Reactor Coolant

System Valves

and Piping

Forced 0.44

* Reactor Coolant

Recirculation Pump

I and Drivers

. I 0 5 .

0.56

1 .89

2.79 .

Unplanned derating

Forced maintenance,

planned and

equipment derating

0.26

* . 0.93

.i

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I

1. A sensitized microstructure

2. A local tensile stress that exceeds the bulk

alloy's yield stress

3. A corrosive (oxygenated) environment.

Even though these conditions theoretically exist

in the heat-affected zones of most BWR recirculation piping welds, only a very small percentage of

these welds have cracked. The primary factor that

determines whether a particular site will develop

IGSCC can therefore be inferred to be the existing

local tensile stress since it is the most variable factor. It is also the most difficult to quantify, because

it is made up of both applied (design loads) and

residual (weld shrinkage, forming, machining, and

grinding) stresses. Although IGSCC can be eliminated by controlling any one of the above three factors,4 the current practice is to attack two or (in

some cases) all three of the factors.

The first factor, sensitization of the base metal

and weld heat-affected zone, is affected by the

chemical composition of the base material (primarily carbon content) and the total thermal history

(high weld heat inputs increase sensitization). The

development of nuclear grade stainless steels with

maximum 0.02 wtqo carbon, nitrogen additions,

and tightly controlled levels of impurities has produced a material much less susceptible to sensitization and thus more resistant to IGSCC. Installation

contractors and nuclear steam system suppliers

also have attempted to control sensitization by rapidly cooling the pipe interior after deposition of the

root pass. This technique is called heat sink welding. This treatment reduces the degree of weld-zone

sensitization and produces a more favorable pattern of residual stress on the pipe inside surface.

The second factor, tensile stress, can be mitigated

by way of induction heat stress improvement

(IHSI) and last-pass, heat sink welding. The first of

these two techniques must be closely controlled to

achieve the desired effect (residual compressive

stress at the inside weld surface), particularly with

welds made by the gas tungsten arc process. 5 It also

has been suggested that IHSI may sensitize pump

and valve body castings adjacent to the weld zone.

The third factor, oxygenated cooling water, can

be attacked by careful control of water chemistry.

Under steady-state conditions, BWR coolant contains 200 to 400 ppb dissolved oxygen. This level

can rise to -.8000 ppb under start-up conditions,

thus further increasing the susceptibility to IGSCC.

Because oxygen is constantly being created from

the water coolant by way of radiolysis in the reactor

core, the only practical method to control the oxygen to levels below the threshold for IGSCC within

the recirculation loop is to inject hydrogen into the

cooling water. This produces an environment favoring the recombination of hydrogen and oxygen.

Ljungberg6 has shown that levels of 3 to 5 ppb

oxygen are achievable with this method. IGSCC is

completely inhibited at levels below 5 ppb oxygen.

Also of concern for life extension of BWR systems is

the effect that low-temperature aging has on duplex

(austenitic-ferritic) stainless steel castings (and possibly

weld metal) at operating temperatures and the potential

for thermal embrittlement in localized areas. Exposure

to temperatures in the 560 to 700'F (293 to 371IC)

range promotes the formation of secondary phases

(such as the G-phase, Type X, and alpha-prime phase)

and carbide growth in these materials. The embrittling

effect can be observed on materials with > 10°70 ferrite.

Materials with >20% ferrite show drastic reductions

in impact energy when aged at higher temperatures

[i.e., 750 to 850IF (399 to 4550C)]. Low-carbon

grades in general show greater resistance to this

embrittling effect, probably because of a reduction in

the formation of chromium carbides and other secondary carbide phases. There also is some evidence to suggest that careful initial heat treatment of cast duplex

stainless steel may produce a casting with superior

unaged impact energies 7 such that even after aging significant fracture toughness (measured by roomtemperature, Charpy V-notch energies) remains.

Typically, these duplex stainless steel components are

extensively weld repaired with attendant ferrite dilution

in the weld zone. Detailed records of welded repairs

and detailed ferrite surveys are not available. Ferrite

measurements and monitoring should be done on the

inside surfaces, which requires disassembly and extensive isolation provisions or draining the vessel after

unloading the core. Further study of the thermal

embrittling of the duplex stainless steel castings and

particularly the weld material is recommended. 8 Additional information on the thermal embrittling of the

duplex stainless steel components is discussed in Section 5.3.2, Chapter 5.

Crevice corrosion has been mentioned as a

potential degradation mechanism that could occur

over time at locations such as shaft sleeves and the

underside of socket welds if a corrosive environment and a difference in oxygen concentration

existed. To date, crevice corrosion has not been

reported as a problem in BWR recirculation piping.

This is probably a result of the tight controls maintained on water conductivity and oxygen levels in

this system.

Thermal fatigue and corrosion-assisted fatigue

also have been mentioned as potential degradation

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mechanisms. Frequent reactor scrams and

unscheduled outages produce transient thermal

stresses on both the cooldown and heatup cycles.

Highly stressed hot spots need to be defined and a

data base established to confirm whether this effect

has significance for BWR recirculation piping. The

weld overlays probably deserve specific attention.-

Fatigue considerations are currently being studied

for carbon steel systems, but stainless steel systems

(such as the BWR recirculation system) have not yet

been addressed.9

10.5 Failure Modes

Failure modes produced by the failure mechanisms above would be primarily through wall

defects resulting from cracks (IGSCC, embrittlement, or fatigue) or material wastage (crevice corrosion). So far, cracks in recirculation piping

materials have satisfied a leak-before-break criteria

and no guillotine breaks (complete separation)

have occurred. This latter type of failure is unlikely

to occur unless it can be shown that the duplex

stainless steel castings in actual components have

embrittled to such an extent that a brittle fracture

could occur at or near the operating temperature.

10.6 In-service Inspection

Methods (1S1)

The currently required in-service inspection

(ISIs) have not always been able to detect cracks

that have developed during service. Detection of

IGSCC using the currently available ISI techniques

is particularly difficult. Examination of cast stainless steel by ultrasonic testing is considered unreliable. The USNRC is currently sponsoring research

into improved methods for continuously monitoring pressure boundary components. 10 On-line

acoustic emission is being considered, but a question exists as to its ability to judge the severity of a

particular cracking event. Proposals have also been

made to install moisture-sensitive tape on the outside surface of critical piping to detect leaks. 1I

10.7 Summary, Conclusions, and

Recommendations

The degradation sites and mechanisms, stressors, potential failure modes, and ISI methods for

the BWR recirculation piping are summarized by

Table 10.2. IGSCC has been a major problem and

a variety of fixes and controls (discussed above) are

in place. Other degradation mechanisms are possible and should be evaluated, but to date they have

not been associated with actual outages or failures

in BWR recirculation piping systems.

As might be expected, the nuclear industry is

extremely optimistic about the possibility of

extending the life of nuclear power plants, at least

from the technical point of view.12'

1 3 The predominant view is that no insurmountable technical

obstacles currently exist that would preclude the

extended use of existing BWR plants, provided the

groundwork is laid now to support continued use.

The careful monitoring and recording of component performance from all operating plants is a key

factor in this effort. The actual degree'of thermal

embrittlement in the duplex stainless steel castings

should be monitored. The power industry, through

industry groups like the Electric Power Research

'Institute, and the nuclear steam supply system

manufacturers through programs like General

Electric's COMPASS computerized data base and

Westinghouse's Reliability Data Base, are attempting to develop useful data bases that will pinpoint

weak links and identify trends useful in the lifeextension effort. These data bases will be essential

in the effort to ensure and demonstrate the safety

and reliability of each plant being considered for

life extension.

Information on failure mechanisms and sites

should continue to be gathered from individual

utilities, nuclear steam supply system manufacturers (such as General Electric), and reviews of

USNRC reportable incident records to update the

data base of known weak links and material failures. Additional work needs to be performed to

analyze the existing data bases for trends in component failure mechanisms and locations. For

instance, failure locations may be one of the keys in

verifying the existence of so-called fatigue hot spots

in BWR recirculation systems. The existence of

such hot spots and their effect on overall system life

becomes of greater concern as the number and

severity of startup/shutdown cycles increase.

Additional work also should be done on duplex

stainless steel castings to determine whether (a) a BWR

environment will produce significant thermal aging

and (b) typical castings contain sufficient delta ferrite

to exhibit a significant embrittling effect.

The possible detrimental effects of hydrogen

additions to the BWR recirculation piping loop

should be studied to determine what reactions

occur between hydrogen and other system components such as feedwater piping.

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Table 10.2. Summary of degradation processes for BWR recirculation piping

Degradation Degradation Failure ISI

Rank Site Stressors Mechanisms Modes Method

WVeld heat-affected Tensile stress, oxygen

zone furnace environment, sensitized

sensitized safe heat-affected zone

ends

IGSCC Cracks, leaks Ultrasonic examination,

moisture-sensitive tape

Cracks, leaks Ultrasonic examination,

moisture-sensitive tape

2 High thermal stress Cyclic tensile stress, Thermal fatigue,

regions predicted corrosive environment corrosion fatigue

by stress rule

index analysis

3 Austenitic-ferritic High temperature, Thermal embrittle

stainless steel tensile stress, shock

castings with high loads

delta ferrite

levels

IT

!ment Cracks, leaks Ultrasonic examination,

moisture-sensitive

tape, impact test

specimens

4 Shaft sleeves,

other crevices

Corrosive environment, Corrosion

oxygen-starved areas

Material

wastage,

leaks

Moisture-sensitive tape

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10.8 References

1. P. W. Marriott, "Nuc!ear Plant Life Extension-A Comprehensive Approach to Plant Improvements," Nuclear Europe, 3, 1984, pp. 29-32.

2. North American Electric Reliability Council, 1984 Annual Report, Generating Ava'lability Data System (GADS), Component Cause Code Report, Princeton, New Jersey.

3. J. A. Rose, R. Steele, Jr., B. C. Cornwallj and K. D. DeWall, Survey ofAged PowerPlantFacilities,

NUREG/CR-3819, EGG-2317, June 1985.

4. J. C. Danko, "IGSCC-Recent Experience Reviewed," Nuclear Engineering International, 29, 356,

June 1984, pp. 27-30.

5. G. T. Krause, "Stress Improvement of GTAW-P-ME Welds in Austenitic Stainless Steel Pipe," Welding Journal, 65, 5, May 1986, pp. 21-29..-

6. L. G. Ljungberg, D. Cubicciotti, and M. Trolle, "Materials Behavior in Alternate (Hydrogen)

Water Chemistry in the Ringhals-l Boiling Water Reactor," Corrosion, 42, May 1986, pp. 263-271.

7. 0. K. Chopra and H. M. Chung, Long-term Embrittlement of Cast Duplex Stainless Steels in LWER

Systems: Annual Report October 1984-September 1985, NUREG/CR-4503, Volume 1, ANL-86-3, r

January 1986. -

8. 1. Spiewak, R. S. Livingston, and C. A. Negin, The Longevity of NuclearPobwer Systems, EPRI NP4208,

August 1985.

9. M. Vagins and J. Strosnider, Research Program Plan-Piping, NUREG-1 155, Volume 3, July 1985,

p. 4.

10. J. Muscara, Research Program Plan-Nondestructive Examination, NUREG-1155, Volume 4,

July 1985, p. 1.

11. P. Riddle, "Localized Leak Detection Using Moisture Sensitive Tape," NuclearEngineering International, June 1984, pp. 32-34.

12. D. R. Eggett and A. E. Meligi, "Extending the Life of Nuclear' Power Plants: Technical and Economic Issues," American Nuclear Society Transactions, 46, June 1984, pp. 583-584.

13. P. W. Marriott and J. P. Higgins, "The Impact of Power Plant Aging on Plant

Availability: I (Managing Nuclear Plant Aging)," American Nuclear Society Transactions, 46,

June 1984, pp. 579.

113
a

11. NONDESTRUCTIVE EXAMINATION METHODS

L. J. House

Safety and reliable performance typically are

achieved during the design life of nuclear reactors

by using an engineering approach based on overdesign and incorporation of generous safety factors.

Therefore, the demands on methods for detecting

and measuring flaws using nondestructive examination (NDE) are modest. Historically, many of the

American Society of Mechanical Engineers

(ASME) code requirements and standard NDE

methods were developed for detection and qualitative assessment of fabrication related defects generally associated with the quality of workmanship

and then adapted to in-service inspection requirements within the normal design life of the components. On the other hand, reactor plant-life

extension beyond the normal design life involves

engineering calculations for cases where safety factors may be reduced because of in-service degradation of critical reactor components. Thus,

life-extension inspection generally requires greater

detection reliability and a more quantitative assessment of defects and accumulated damage than traditional fabrication and in-service inspection.

Successful application of residual-life prediction

programs, therefore, depends on the availability of

adequate methodologies to detect and measure the

effect of accumulated damage on performancerelated properties of the key components.

The objective of this review is to address the efficacy

of standard NDE methods for detecting and evaluating

damage produced by degradation mechanisms occurring in reactor environments. It is well known that the

reliability and capability of NDE methodologies are

directly related to operator performance factors1 and

specific characteristics of the instrumentation, data

acquisition, and signal processing employed. 24 However, this review is intended to be of a general nature

covering the standard methodologies rather than

attempting to assess the many specific embodiments of

the techniques derived from those methods. Also, the

adequacy of NDE for any given application depends

on many other factors besides the performance of the

technique. These factors include cost-benefit analysis

(e.g., does inspection cost more than periodic component replacements?), availability of trained manpower,

accessibility, and inspection rates. Only the technicalperformance-related factors (i e., can the methodology

reliably provide the necessary measurement?) are

addressed in this review.

The methodologies generally practiced by the

nuclear and other industries and included in this review

are visual examination, penetrant and magnetic particle testing, x-ray radiography, eddy-current testing,

ultrasonic testing, and acoustic emission monitoring.

These standard NDE methods are discussed with

respect to their adequacy for providing the information

about accumulated damage in reactor components

necessary for residual-life estimation. The format for

this review consists of a background discussion of what

residual-life estimation requires of NDE, the type of

information standard NDE methods can provide, and

a summary of the general deficiencies in the standard

methods for use in residual-life estimation.

11.1 Background

In the transition from the task of providing a

simple quality control function to supplying quantitative assessment of accumulated damage for

fracture mechanics analysis and residual-life estimates, the reliability and adequacy of standard

NDE methodologies have rapidly diminished.

Many of the standard NDE methods were intended

to provide only an indication of the presence of a

flaw or developed specifically to detect a particular

type of flaw in a given component during fabrication and assembly. For residual-life estimates, those

same methods must be employed not only to detect

flaws but to determine the location, size, shape,

orientation, and type of flaw or damage in components with various service histories, complex geometries, and limited access. The inspection is often

performed in harsh environments with very severe

constraints on the time available to do the inspection. Consequently, a number of inadequacies in

standard practices of NDE have, and are continuing, to appear.

To estimate residual life using either a deterministic

or probabilistic approach, information about the effect

of various degradation mechanisms on performancerelated properties is needed. As illustrated in

Figure 11.1, the approach for each critical component

has been to identify the principal degradation sites,

establish the associated damage mechanism, determine

performance-related physical and material properties

affected by the damage mechanism, select the appropriate measurement technique to detect the relevant

property changes, and finally to assess the extent and

114
I

I I . . I . I

Il I I Accessibility

. . environment

Degradation site I . - configuration

I I material

I Degradation mechanism I'

I

Physical and material

properties of damage

produced by degradation

mechanism - . W

Damage measurement

.

I . methods H.

-{ .

Standard NDE methods

l

Miniature specimen testing I

l

Emerging measurement and

monitoring techniques F

7-3044

Figure 11.1. Typical approach to selecting a method for detection and measurement of accumulated damage in

critical reactor components. I

severity of the damage. 5- 12 The selection of the measurement methodology may include standard NDE,

miniature specimen testing, or one of the many emerging measurement and monitoring technologies. The

selection of the appropriate method depends on accessibility, environment, component configuration, and

the material at the potential degradation site.

The properties affected by degradation, particularly those detected and measured by standard

NDE methods, are generally only indirectly related

to the actual damage. For example, in the early

stages of plastic deformation, there is a marked

increase in dislocation density and associated

stacking faults. This causes a decrease in the electrical conductivity of the material. Eddy-current.

methods can measure changes in conductivity and,

thus, in principle can be used to measure

deformation-related properties, such as yield,:

strength. However, other factors, including variations in chemical composition, temperature, surface condition, and unrelated degradation

mechanisms, such as thermal aging, also can affect

conductivity. Also, changes in the concenitration of

interstitial solid solution hydrogen may distort the

crystal lattice of the metal alloy, causing a change in

its electrical conductivity. Therefore, the other

sources of conductivity changes preclude this

approach for most practical applications. Thus, if

the NDE method can not discriminate between the

types of damage or if a knowledge of the degradation mechanism causing the damage is unavailable,

then the nondestructive measurement-based on

empirical correlation of a physical property like

electrical conductivity with a particular type of

material degradation such as hydrogen damage or

plastic'deformation-can give unreliable results.

This issue becomes very complicated because of the

large number'of potential damage mechanisms that

exist in the'reactor service environment.'

The principal degradation mechanisms in key

reactor copronents include: embrittlement (neutron, hydrogen, and thermal), fatigue (corrosion,

thermal, low cycle, high cycle), corrosion (erosion,

wastage, pitting, denting,' and stress corrosion

cracking), wear and fretting, and aging (thermal

and environmenital). The issue for residual-life

assessment is whether' stauidard NDE methods can

detect the accumulated damage produced by these

various mechanisms and provide a measure'of the

severity and extent of the damage in a sufficiently

Tr

115
a

quantitative manner. For some of the degradation

mechanisms, like neutron embrittlement, it is

unclear what to measure using NDE methods in

order to detect and characterize the resulting damage. Ideally, however, NDE should be able to discriminate between or identify the type of damage,

measure the flaw size or extent of damage, measure

the flaw shape and orientation, and determine the

location of the flaw or damage. Although standard

NDE has many deficiencies in being able to provide

all of these necessary measurements for residuallife estimates, the information NDE can provide,

despite the attendant uncertainties and reliability

issues, is nevertheless essential.

11.2 Standard NDE Methods

Summarized in this section is the type of information provided by each of the standard NDE methods.

More specifically, the flaws and damage that can be

detected, the properties that can be measured, and the

deficiencies of each method with respect to residual-life

assessment are discussed.

11.2.1 Visual Examination.13 Inspectionbyvisual examination is the most widely used nondestructive method. Visual examination is important

for reasons other than its simplicity and low cost.

First, visual examination provides a qualitative

indication of surface conditions. In welds, visual

examination can reveal the presence of cracks, surface porosity, and other potential sources of

mechanical weakness, such as the extent of penetration, undercutting, sharp notches, or misalignment. Furthermore, visual examination may assist

in the determination of other nondestructive tests

that should be applied to a potential site of degradation in a reactor component and how they should

be applied. As with all nondestructive methods, the

usefulness of visual examination methods depend

on proper application and correct interpretation.

In visual examinations, more directly so than in the

other nondestructive methods, the results are subjective and dependent on the operator. The results

of visual examination are generally qualitative and

limited to assessing damage that manifests itself at

an accessible, surface of a component. Visual aids

such as optical microscopy to examine the microstructure of a prepared surface and a borescope to

examine the interior walls of tubing increase the

utility of this method. In tube bundles, borescopic

inspection provides information about the restriction of flow because of scale and deposit build-up

and wall thinning caused by wastage, pitting, or

uniform corrosion. The development of computerenhanced vision systems to automate visual examination and to permit remote visual examinations in

hostile or otherwise inaccessible environments have

been successful for viewing structures, but the

information provided is often qualitative in nature

and of limited value to residual-life estimation. Visual examination in most cases provides an indication that damage may have occurred but cannot

directly quantify the amount of material damage.

11.2.2 Penetrant and Magnetic Particle

Testing. Penetrant and magnetic. particle testing

are essentially an extension of visual examination in

that they are used as an enhancement technique to

improve the visibility of surface-connected flaws.

Magnetic particle testing also can indicate the presence of near-surface defects, such as cracks and

inclusions. Its principal utility as a NDE method is

to increase the inspection rate and probability of

detection of surface flaws, compared with unaided

visual examination. Liquid penetrants delineate

surface discontinuities making the inspection less

subject to the visual acuity of the operator. They

provide no direct information about the depth of

the crack although some very experienced inspectors have been able, in some cases, to make fairly

reliable estimates of the depth and size of defects.

The accuracy of these estimates depend on the

magnitude of the defect opening at the surface

being related to the depth of the defect. Penetrant

methods can be used to locate grinding cracks,

welding cracks, casting cracks, fatigue cracks,

shrinkage, blowholes, seams, laps, pores, cold

shuts, porosity, lack of bonding, pinholes in welds,

through cracks, forging laps, bursts, gouges, tool

marks, and die marks, provided the defects open to

the surface. In some work with radioactive penetrants, 14 residual activity of the penetrant has been

successfully correlated with crack size but complete

capillary diffusion of the penetrant into the flaw is

essential. In general practice, the correlation is not

a reliable indication of crack size. The deficiencies

of penetrant and magnetic particle testing are the

same as with visual examination, except that the

detection probability for surface-connected flaws is

improved over unaided visual examination.

11.2.3 Radiographic Examination. The standard

practice in radiographic examination is generally x-ray

film radiography. The method essentially provides a

recording of density variations that may occur because

r

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116
I

of a variety of causes: cracks, inclusions, porosity,

voids, lack of bonding, and dimensional changes. The

radiograph can'provide quantitative information about

flaws that affect the mass density or x-ray absorption

coefficients of the component; for example, it can indicate the length of cracks. A film'densitometer also can

give an estimate of the actual extent of damage given

suitable calibration and precise control of exposure

time and source intensity. However, two-sided access,'-

except in 'those few cases where autoradiography (a

technique using the intrinsic radioactivity of the component) is possible, the slow rate of examination, background radiation, and the time between the

examination and when the results are available after

film development and evaluation severely restrict the

application of this method. This is particularly true for

periodic in-service inspections of reactor components

where, to minimize costly down-time, the NDE activity

must not delay decisions to replace or return a compo-'

nent to service.

11.2.4 Eddy-Current Testing. The use of singleand multifrequency eddy-current testing is gener-'

ally limited to inspection of simple geometries for

near-surface defects. The only extensive app'lica-'

tion of this method in the'nuclear industry has been

to inspect pressurized water reactor steam generator tubing. Although the eddy-current signal is

affected by'the metallurgical condition of the material being tested and can provide an indication of

degradation or damage, the effect of changes in

phase and chemical composition, as well as other

metallurgical properties, on the eddy-current signal'

is neither quantitative nor predictable. A correlation between the measured impedance change of

the eddy-current coil and the desired structural or

serviceable characteristics must be established for

all of the various degradation mechanisms. Normal

metallurgical or chemical variations can mask the

effects produced by service-related damage in the

component structure and any condition affecting

the permeability or conductivity of the'material

being tested affects the measured eddy-current signal. Discriminating between changes in the mea-'

surement caused by normal material variations or

actual damage is often not possible. Single frequency eddy-current examination of steam genera--

tor tubing 'in 'accordance with the ASME -code

procedures has not always been reliable because, in

addition to the reasons cited above, other inspection variables, such as tube supports, tube sheets,

denting, and eddy-current'probe wobble, affect the'

results. Multifrequency methods remove the influence of many of these variables, but improvements

are still needed in the sensitivity and reliability of

flaw detection and characterization. One particular

deficiency is that standard eddy-current methods

and probes for tubes and methods are inadequate

for detecting and characterizing circumferential

type flaws, like intergranular cracks and general

intergranular attack, in tubes in the neighborhood

of and within the tube-sheet crevice. The design of

pancake coils for tube inspections enables the

-detection of circumferential flaws, but the use of

this type of probe is not presently a standard practice for tube inspections.

11.2.5 Ultrasonic Testing. After visual examination, the next most widely used volumetric nondestructive method for in-service inspection of components is

ultrasonic testing. The ASME code requires ultrasonic

inspection of welds and adjacent base material.

Although presently accepted as the most useful volumetric examination method, the reliability of standard

ultrasonic procedures and methods of flaw detection

has been increasingly questioned during the last several

years.3' 15 17 This is because'the ultrasonic examination methods, particularly procedures and practices

specified by the various ASME 'code requirements,

were developed primarily to assess overall quality by

detecting fabrication-related defects that correlated

- 'with workmanship. 16 However, the information

required for residual-life assessments and fracturemechanics-based integrity assessments requires detecting, locating, and sizing of more subtle flaws. Sizing of

flaws has always been a source of error because ASME

code sizing techniques are typically based on empirical

correlations of ultrasonic-echo-amplitude 'measurements with flaw size and not on basic physical principles.

The standard ultrasonic examination methods

generally use recording of pulse-echo amplitude

'and transducer position as the basis for flaw detection and sizing. The functional aspects of ultrasonic testing involve two operations. The first

operation is generally a rapid initial search mode to

detect the presence of a flaw in a component and a

second operation in which a more detailed inspection is performed to determine whether the flaw is

benign or represents a performance-limiting condition. Standard procedures, although satisfactory

for their original purpose, are not adequate and are

often unreliable for, providing the detection and

characterization of flaws needed to assess the residual life of reactor components. Some issues contributing to the inadequacy of standard ultrasonic

testing practices include'theoretical limits of detectability and resolution; external factors such as

r

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L.

117

17
a

operator training, equipment, and calibration deficiencies; acoustic coupling variabilities; and internal factors such as material variabilities, part

geometry, and flaw characteristics. An example of

internal factors that affect flaw-detection reliability

is the variability of ultrasonic echo strength with

mechanical load. A mechanical load of sufficient

magnitude can compress the two faces of a crack

tightly enough so that the ultrasonic reflection is

reduced or, in the extreme case, the crack becomes

transparent. Rough-clad weldments and components fabricated from cast stainless steels are virtually'uninspectable by standard ultrasonic methods.

Other deficiencies include unreliability of detection

of cracks under cladding, detection and sizing of

intergranular stress corrosion cracking (IGSCC),

and detection of thermal fatigue cracks and IGSCC

in wrought stainless steel welds.

It is clear that the limitations of standard ultrasonic test methods are in part caused by deficiencies and physical limitations of the available

technology. On the other hand, a previous evaluation of the reliability of Air Force field inspections

indicated that inspection results were poorer than

could be explained on the basis of the inherent

physical limitations of the technique.15,18 Poor

performance was attributed to human factors such

as operator boredom and the awkward positions in

which the operator was required to perform the

inspection, because of limited access.

A good summary of the capabilities of standard

ultrasonic NDE methods is given in Figure 11.2,3 in

which diagrams of the probability of defect detection (DD), correct rejection (CR), and correct

acceptance (CA) as a function of defect categories

are shown. The defect categories included-small

acceptable defects (a), small rejectable defects (b),

large continuous defects (c), and composite

(multiple-type) defects (d). The ultrasonic test procedures fell into three categories. Category A complied with the requirements of the Program for the

Inspection of Steel Components (PISC)-formerly

Plate Inspection Steering Committee-which

closely follow the requirements-of ASME XI. Category B involved minor modifications to the PISC

procedure, and Category C was markedly different

ultrasonic techniques providing'truly alternative

procedures. Only the large continuous defects are

properly detected and rejected by the Category B

and C ultrasonic test procedures. The Category C

procedures do result in better, but not perfect,

small-defect detection.

11.2.6 Acoustic Emission Monitoring. Acoustic emission measurement techniques have been

used extensively as an alternative to ultrasonic

examination for detecting flaws in pressure vessels

in the nuclear and petrochemical industry. Acoustic

emission is a passive technique that detects flaws

only during growth. One very useful application of

acoustic emission monitoring is leak detection in

pressure vessels. A major deficiency in acoustic

emission testing is the possibility of missing a large

flaw if it happens to not be growing during monitoring. Advocates of acoustic emission testing

counter with the argument that only flaws that are

growing can induce failure and continuous acoustic

emission should preclude missing the large flaws

during growth. A number of major evaluation programs have been performed on pressure vessels

with mixed results. Difficulties are associated with

discriminating between damage- or flaw-related

acoustic emission events and ambient background

noises, low-level emissions produced by flaw

growth in certain alloys, and accurate location of

the acoustic emission event in complex structures.

An intrinsic deficiency of standard acoustic emission monitoring is the absence of a reliable correlation between measurement of acoustic emission

events and the severity of the flaw. A present

advantage of most acoustic emission monitoring is

its use as a precursor to impending failure rather

than a method for quantitative measurements of

component damage for use in fracture-mechanics

estimates of flaw severity or residual-life assessment . Another advantage of acoustic emission

monitoring, unlike the previous methods discussed, is that a more global examination or largevolume interrogation is provided. A major

disadvantage of acoustic emission testing for

residual-life estimation at present is the absence of

a correlation between acoustic emission properties

and fracture-mechanics-related parameters. For

example, acoustic emission parameters can be used

for crack-growth detection or location, but the

measured parameters have not been reliably correlated with crack size.

11.3 Summary of General

Deficiencies of Standard

NDE Methods

NDE methods, as embodied in standard practice,

offer capabilities that are primarily limited to flaw

detection. Standard NDE methods are presently

_.

.,I I

t

118
a

DD

1

0.5

. 0

: ; . - I 11 .. ' ,_ .

1..CA, ,CR

B

PISC procedure

DO

1

0.5 :

'- A -'

I..,

I I

''a b c d a b c d

National standard or procedure in thespirit of ASME

Code (for preserviceor manufacturing inspection)

DD - CA CR

c- 1 1 r

0.5 cdc

a bc d a bc d

EL

Alternative techniques 7-3065

Figure 11.2. Average cumulative diagrams of defect detection (DD) and correct rejection or acceptance [CR(CA)] as

a function of the defect categories a, b. c, and d (from Reference 3).

measurements of physical properties that, based on

empirical correlation, are related to defects and

component damage. In other words, most NDE

measurements are only indirectly related to the flaw

size or damage state. These measurements are often

affected by geometry and material properties unrelated to damage and flaws. Changes in test conditions may produce entirely false indications and

some comparisons of damage state using different

inspection methods or simply different inspection

times may be invalid. Having to rely on flaw calibration standards and empirical correlations also

may make it prohibitively expensive to develop

NDE methods for all types of damage, materials

and heat treatments, component geometries, and

the various service histories of each component.

The high cost of down time and slow inspection

rates (area or volume of material inspected per unit

time) associated with the high-resolution techniques necessary for detecting small critical flaws

and providing flaw characterization represents

another deficiency of standard NDE methods.

Global inspection methods, even with low resolution, are needed to initially determine whether a

more detailed inspection is warranted and to minimize NDE bottlenecks in returning a system online. Acoustic emission is one of the few NDE

methods offering considerable potential for providing a global inspection approach.

Residual-life estimation and fracture-mechanics

analysis require more than flaw detection. Ensuring

structural reliability when extending components

119
I

beyond their original design life necessitates a quantitative determination of the effect of accumulated damnage on the performance related properties of various

components. When a NDE operator is asked to make

these types of measurements, an entirely different set

of problems arise than those encountered in simple

flaw detection. Standard NDE methods are presently

incapable of handling this new set of problems.

Eddy-current and ultrasonic methods are among the

most promising for making the type of damage-related

measurements needed for residual-life estimation, but

models for quantifying the effect of material properties

and associated degradation mechanisms are still being

developed. Presently, there is poor agreement between

models and experiments in actual structural components because models have been well developed only

for idealized defects in infinite media. There are deficiencies in the present understanding of the interaction

mechanisms between the various manifestations of

performance-related structural damage and nondestructive measurements. This requires considerable

empirical correlation to establish a reliable relationship

between the presence or extent of damage and a nondestructive measurement. The constraints of in-service

inspection limit most standard NDE to providing only

qualitative assessments. -

One approach for handling the uncertainties in the

reliability of flaw detection and sizing is a probabilistic

treatment of the NDE data. The probabilistic

approach estimates residual life with a specified level of

confidence given the uncertainty associated with the

detection and measurement of flaws. This does not

remedy the inadequacies of standard NDE methods

but provides an-approacli for assessing the risk associated with relying on the information that NDE can

presently provide.

The ultimate goal is to develop an NDE capability

adequate for a more quantitative evaluation of residual

life. Realization of this goal requires significant

improvements in flaw detection reliability and the ability to make quantitative measurements of the effect

accumulative, damage has on performance related

properties of a component. Discrimination between

types of flaws, accurate quantitative flaw sizing, and

quantitative measurement of performance-related

material properties is presently not adequately provided by standard NDE methods. Furthermore, a

more fundamental deficiency affecting the development of the necessary NDE methods for reliable estimation of residual life of components is related to the

limited understanding of the relation between the physical properties of the material, the degradation mechanisms, and their effect on component performance. A

better understanding of the effect of damage on these

physical properties is needed in order to identify the

physical measurements that need to be made.

T

ri

120

.IT
11.4 References

1. H. R. Chin Quan afd'l. G. Scott, "Operator Performance and 'Reliability in NDI," in Research

Techniques in Nondestructive Testing, Volume'3, New York: Academic Press, 1977.

L

2. S. H. Bush, Reliability of Nondestructive Examination, NUREG/CR-3 110, PNL-4584, Volumes 1-3,

October 1983.

3. 0. Buck and S. Wolf (eds.), "Nondestructive Evaluation: Microstructural Characterization and Reliability Strategies," Conference Proceedings of The Metallurgical Society of the American Institute of

Mechanical Engineers, October 5-9, 1980, p. 405.

4. Y. Segal et al., "A Systematic Evaluation of Nondestructive Testing Methods," in Research Techniques in Nondestructive Testing, Volume 3, New York: Academic Press, 1977.

5. R. B. Clough (ed.), Quantitative NDE in the Nuclear Industry, Metals Park, Ohio:- American Society for Metals, 1982.

6. J. T. Fong and 0. F. Heddon (eds.), Nondestructive Evaluation (NDE): Reliability and Human Factors, National Bureau of Standards Special Publication, Gaithersburg, Maryland.

7. M. J. Davidson et al., "Monitoring for Life Extension," Journal of Pressure Vessel Technology, 107, r

August 1985.

8. 3. Muscara, Research Program Plan-Nondestructive Examination, NUREG-1 155, Volume 4, July 1985.

9. F. A. Simonen, "The Impact of Nondestructive Examination Unreliability on Pressure Vessel Fracture Predictions," Journal of Pressure Vessel Technology, 107, February 1985.

10. G. D. Gupta et al., Remaining Life Evaluation and Extension of Boilers, FWC/FWDC/TR-84/14,

May 1984. r

L

11. G. H. Harth and T. P. Sherlock, "Monitoring the Service-Induced Damage in Utility Boiler Pressure

Vessels and Piping Systems," Journal of Pressure Vessel Technology, 107, August 1985.

12. G. D. Gupta et al., "A Practical Strategy for Life Evaluation and Extension of Fossil-Fuel-Fired

Boilers," Journal of Pressure Vessel Technology, 107, August 1985.

13. W. J. McGonnagle, Nondestructive Testing, second edition, New York: Gordon and Breach Science

Publishers, 1971. L

14. G. P. Hayward, Introduction to Nondestructive Testing, Milwaukee, Wisconsin: American Society of

Quality Control, 1978.

15. R. B. Thompson and D. 0. Thompson, "Ultrasonics in Nondestructive Evaluation," Proceedings

of the Institute of Electrical and Electronics Engineers, 73, 12, December 1985.

16. NondestructiveEvaluation in theNuclearlndustry-1980, Metals Park, Ohio: American Society for

Metals, 1981. fl

121
17. D. 0. Thompson and D. E. Chimenti (eds.), Review ofProgress in Quantitative Nondestructive Evaluation, Volumee2, New York: Plenum Press, 1983.

18. W. H. Lewis et al., Reliability and Nondestructive Inspections, Report SAALC/MME 76-6-38-1, San

Antonio Air Logistics Center, Kelly Air Force Base, Texas, 1978.

122

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12. ADEQUACY OF ASME CODE IN-SERVICE

INSPECTION METHODOLOGY

J. F Cook

Federal regulations1 require that the requirements of Section XI of the American Society of

Mechanical Engineers (ASME) Codea for inservice inspection (ISI) of nuclear power plant

Class 1, 2, and 3 components and systems be used..

The edition and addenda of Section XI that must

be followed also are specified in the regulations.,

The code edition/addenda to be used for the ISI is

determined by the date of the'start of plant operation and is normally updated whenever' a.new

addenda is issued. However, the updating of ISI

plans to later editions/addenda can be delayed by

extended shutdowns. Therefore, the code edition/

addenda followed varies among the operating

plants. This is further discussed in Reference 2.

Because there is no one set of current ISI requirements, this chapter discusses the general philoso-,

phy of the ISI specified"in' Section'XI, and the

status of improvements' in the ASME code requirements relative to selected components important to

plant safety and plant life extension. Advanced

NDE methodologies are also discussed.

12.1 ASME Section XI Code Rules

The acceptance criteria in ASME Section XI for

flaws found by nondestructive examination (NDE)

are based on fracture-mechanics analysis. This is in

contrast with the fabrication code (Section 111)

NDE acceptance criteria that are based more 'on

workmanship standards. However, the NDE.

requirements initially put in Section XI were based

on the same requirements used in the past for the

fabrication code NDE. This has resulted -in an

incompatibility between the Section XI acceptance

standards and the NDE capabilities. In many cases;

the NDE is not capable of detecting the Section XI

acceptance limit indications. For, example, inci--

dents' of through-wall intergranular stress corro--

sion cracking (IGSCC) have been found in piping,

systems by evidence of leakage that were not

detected by standard ASME code ultrasonic inspec-.

tion. Studies performed under the Program for the

Inspection of Steel Components3 (PISC) (formerly

the Plate Inspection Steering Committee) on

heavy-section specimens showed inadequacies in

standards, ASME code minimum flaw sizes, and

vessel ultrasonic examination procedures. Improvements have been made and work is in progress to

correct some of these limitations. Those improvements and work in progress are discussed below as

related to certain components identified as important to life extension.

The Section XI nondestructive examinations serve

as a complement to pressure tests specified in

Section XI that must be applied to all safety-class pressurized components. The pressure tests act as temporary stressors providing for: (a) some verification of

continuing structural integrity, and (b) identification

of flaws that cause leakage. Unlike NDE, which is generally applied only to welds and adjacent base metal,

the pressure tests check the complete pressure boundary. Also, the NDE is done on a sampling plan while all

components receive the pressure test. However, these

pressure tests only evaluate the integrity of the pressure

retaining boundaries; they do not check other components important to safety, such as supports, and they

do not'detect small defects that may later grow to a

critical size.

The NDE specified in Section XI is done on a sample of the welds. For example, 25°7 of the reactor coolant piping system welds must be examined during each

ISI interval. If flaws are found, then additional examinations are required.'Also components containing flaw

indications are required to be 'examined more frequently in the future (e'g., reexamined during the next

three successive inspection periods). The more frequent

examinations are designed to determine whether the

flaws are growing. Examples of these requirements are

contained -in Sections IWB-2420 and IWB-2430'of

Section XI.'

II

A code requirement that needs to be resolved is the

inspection schedule for times beyond the present 40-

year plant life. Two different' inspection programs are a. Wvhen used in this Chapter of~this report, unless otherwise

stated, references to the Code or Section Xl means Rulesfor In- presently allowed by the code. One program (Inspec- service -Inspection of Nuclear Power Plant Components, tion Program B) uses four 10-year intervals. Inspection

Section Xl, Division I ("Rules for Inspection and Testing of program A uses, from the start of operation, 3 years Components of Light-Water Cooled Plants") of the ASME

Boiler and Pressure Vessel Code. ' - , for the first interval, 7 years for the second interval,

123
I

13 years for the third interval, and 17 years for the

fourth interval. Most plants now use Program B. This

is fortunate because Program A reduces the ISI frequency as the plant operates longer and, thus, would

give less ISI information for assessing plant condition

relative to life extension.

The code has historically specified NDE requirements that included details on how to perform the

examinations (a cookbook approach). The philosophy of Section XI of the code is now changing to

specify NDE performance standards, leaving the

procedures up to the performing organizations.

The performance demonstration requirements are

intended to verify that the NDE system (equipment, procedure, and personnel) is capable of

detecting the minimum size flaws identified in the

Section XI acceptance standards. Additional performance demonstrations are being required for

flaw sizing. Examples of these requirements for

piping welds are contained in a draft of Code

Case N-409, Revision 1, now under consideration

by the Section XI ASME Code Committee. Implementation of the performance demonstrations

should result in more uniformity in the performance of NDE systems (procedures, personnel, and

equipment) for flaw detection and flaw length and

depth sizing.

12.2 NDE Improvements

Reference 4 contains extensive information on

the degradation of bolting in nuclear plants. Partly

as a result of the problems that are described in this

reference, the first code document that requires

NDE performance demonstrations (Section XI,

Appendix VI, "Ultrasonic Examination of Bolts

and Studs") was included in the Winter 1983

Addenda. The licensee is required to demonstrate

that the procedures and personnel can detect a

qualification, notch representing the maximum

acceptable size flaw. Data on field experience using

these requirements should be available following

USNRC reference of this code addenda in the regulations and subsequent update of plant ISI plans.

Piping-ultrasonic examination practices have

received great attention in the code because the

problems. with IGSCC in boiling water reactor

plants have increased. As a result of the IGSCC of

certain larger pipes,. first discovered at the Nine

Mile Point Plant, two USNRC Inspection and

Enforcement (I&E) bulletins have been issued. 5'

6

These. bulletins require qualification, of piping

examination procedures and, personnel on blind

specimens containing IGSCC. These practices were

subsequently covered by Code Case N-409. 7 Additional ASME code changes that are being developed may add performance demonstration

requirements for all piping weld ultrasonic examinations. This activity will significantly improve the

effectiveness of ultrasonic examination practices

for finding flaws in piping welds. However, there

are still problems in developing effective ultrasonic

procedures that apply to all piping. The geometry

of welds and adjacent fittings often prevent the

access required for conducting a complete ultrasonic examination. 8 In many cases, even special

ultrasonic techniques will not penetrate cast stainless steel piping and fittings. 9

Changes were published in the Winter 1983

Addenda to Section XI that simplified selection of,

and reduced the number of, Class 2 welds to be

included in examination samples. This change also

improved examination methodology in that all

surface-only examinations were eliminated. Data

collected from piping failures show that most problems (e.g., IGSCC and thermal fatigue cracks)

originate from the inside surface. Similar changes

should be made to Class I examination requirements. At present, Class I piping with less than a

4-in. nominal pipe size receives only a surface

examination. Confirmatory data on the effectiveness of ultrasonic examinations on small piping

would help in obtaining code committee approval

of changed Class I requirements.

Improvements in vessel ultrasonic examination

requirements are presently being acted on by the

ASME Boiler and Pressure Vessel Committee. It is

common practice now for most foreign and domestic plants to apply additional examinations beyond

the code minimum. As a result of international

cooperative studies on ultrasonic vessel examination practices known collectively as PISC and recommendations from the Pressure Vessel Research

Committee, requirements for an additional examination (70-degree examination angle) for nearsurface flaws (e.g. underclad cracks) are being

added to the code. In this same action, caution is

being added to the code language regarding using

amplitude/search unit movement techniques for

flaw sizing. The Section XI code action in progress

is scheduled to be published as Appendix I to Section XI. Another significant improvement in this

code action is a requirement for more sensitive

recording that will reduce the chance of missing

flaw indications.

Containment structure leakage rate testing is controlled by Appendix J of the USNRC regulations. As

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well as the leak testing, a general visual examination of

accessible internal and external surfaces is required.

However, minimum accessibility requirements are not

given. Thus, the extent of the examinations probably

varies among the plants. Subsection IWE of

Section XI of the code contains rules for the examination of metal containment structures. USNRC action

to require use of these rules is in progress. When imple-'

mented, these subsection IWE requirements, which are

concentrated on welds and other important parts of the

containment structure (e.g., bolting, seals, and gas-.,

kets), should result in more consistent metal containment examination practices. Most of the examinations

are visual and conducted on painted surfaces. Reference 10 indicates that visual examinations can appropriately detect degradation in painted metal structures.

Rules for concrete containment examinations are

now approved for Code' publication as

Subsection IWL of Section XI in the Winter 1986

Addenda. Additional tests that could enhance the

new subsection IWL rules relative to life assessment

for concrete containments are given in

Reference 10. These tests, controlled by British

standards, include checking carbonation depth,

ultrasonic pulse velocity, covermeter, and rebound

hardness measurements. These tests will 'be

reviewed and compared with the American Society

for Testing and Materials practices. Following this'

review, recommended changes will be presented to

the ASME for improving concrete containment

examination practices relative to plant-life extension. Previous work at the Idaho National Engineering Laboratory involved correlating ultrasonic

pulse velocity measurements with changes in concrete properties following exposure to radiation. 1

Reference 10 emphasizes the importance of

destructive tests for determining life forecasts of

concrete structures.

The Section XI steam generator tubing eddycurrent examination requirements are based on the

single-frequency methodology published in Regulatory

Guide 1.183.12 However, most examinations are now

performed using multiple-frequency methods, which

can enhance detection of flaws in areas such as those

adjacent to tube-support structures. The code has not

yet required that the eddy-current procedures be qualified for detection of cracks in tubing. Calibration using

flat bottom holes (simulating wall thinning) is specified. The eddy-current response to a flat-bottom hole is

usually significantly different than the eddy-current

response to a crack. Other steam generator examination practices that are now being used by industry such

as ultrasonic testing of tubing, profilometry measurements and radiography are scheduled to be covered by

the ASME; work has been started at the task group

level of the ASME code committee. Further discussion

of steam generator'examination practices is included in

Chapter 6 of this report.

The Section Xl NDE requirements are focused on

finding discrete flaws in components; detection of general degradation (e.g. radiation damage, thermal

aging) of material properties is not covered per se.

However, control specimens for destructive testing are

specified for inferring reactor vessel material properties

after periods of operation. As part of checking materials for aging relative to life extension, it may be desirable to have material property data on'other

components such as piping sections, supports, pumps,

etc. Samples of these components are not now available for destructive tests. However, miniature specimen

techniques (see Chapter 14 of this report) may prove

useful in this regard. Also, some of the required properties such as yield strength, tensile strength,: and

ductile-to-brittle transition temperature may be

inferred from NDE data using advanced techniques, as

reported in References 13 through 17. These references

indicate that measurements of ultrasonic attenuation

land velocity can be used to infer certain mechanical

properties because the mechanical and ultrasonic properties are all dependent, to a degree, on material microstructure. Measurement of changes in the mechanical

properties because of environmental factors, (such as

thermal aging and radiation damage) should be possible for a specific material once calibration data for the

material have been obtained. The results to date have

been obtained on laboratory test specimens; application of the techniques to actual components in the field

must await perfection of ultrasonic instrumentation

and data processing methods now being developed that

will allow measurements where only one material surface is accessible and will minimize surface and ultrasonic coupling variations. Further confirmatory tests

and field demonstrations will be required before any of

these techniques can be included in the ASME code.

Addition of material property NDE requirements to

Section XI would significantly improve the knowledge

of the actual in-service properties, and also might

reduce the need for testing surveillance specimens.

12.3 ASME Code ISI and Life

Extension

The current Section XI ISI requirements can

provide only part of the information needed for

making decisions regarding life extension. Several

weaknesses in NDE practices need to be resolved.

For example, in most cases, the code does not

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require any backup examinations to confirm results

from the primary examination. However, selected

utilities are applying backup examinations. If an

ultrasonic examination finds indications of cracks

in a pipe, it may be desirable to use an inspection

port to go inside the pipe with a remote visual or

surface examination to backup the ultrasonic

examination. Geometric conditions in components

sometimes produce indications that are confused

with indications from flaws. It is common practice

to always specify a backup examination for the

United States Air Force nondestructive inspection

programs. Backup examinations will increase the

confidence in NDE and provide better information

on the condition of components relative to plantlife extension.

Improvements are being made in the NDE methodology to enhance the reliability of flaw detection. After a flaw is found in a component, the flaw

must be characterized (sized) to determine the

effect of that flaw on the life of that component, or

else the component must be repaired or replaced.

Experience with the IGSCC problems in piping has

shown that qualifying equipment and personnel for

sizing is more difficult than qualifying for detection. Applying standard ASME code rules for sizing of flaws in vessels can in reality end up giving

the size of the ultrasonic soundbeam. Corrective

action is underway as shown in Supplement 12 of

Appendix I of the code. A significant problem in

developing flaw sizing requirements will be defining the qualification specimens. The economics of

fabricating an adequate number of vessel specimens will slow the code action on this problem.

However, life-extension considerations may

increase interest in solving this problem.

12.4 Summary, Conclusions, and

Recommendations

ISIs of nuclear power plants are controlled by the

USNRC regulations and Section XI of the ASME

code. Improvements in the ASME code NDE methodology relative to flaw detection are evolving and

are summarized in Table 12-1. These improved

code rules should be applied as part of evaluating

plant conditions relative to license renewal. Current

code methodology is not wholly adequate to assess

the residual life of the major light water reactor

components. More efforts are needed to develop

field-usable nondestructive examination techniques and equipment that can provide valid information on the sizing of defects and the material

properties of the major components.

126

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Table 12.1. Summary of code NDE improvements

Component NDE Methoda NDE Improvements Status

Bolting UT

Piping UT

Demonstration of procedure and

personnel qualification

Procedture and personnel

qualification using statistically

based specimen sets, containing

flaws

Near surface examination

requirement, correction of flaw

sizing techniques

Addition of Code examination

rules to supplement the

Appendix J leak testing

Published, W-83 Addenda

In committee, may be in

;W-87 Addenda

In committee, may be in

-W-86 Addenda .

NRC review in progress

Vessels UT r

L

Containments VT, LT

Steam generators ET, RT, UT

Profilometry

Code rules to cover all types of

NDE used by the industry

In committee

I a. . I .

a. UT, VT. LT, ET, and Kr refer respectively to ultrasonic, visual, leak testing, eddy-current, and radiograph nondestructive examination methods.

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127
12.6 References

1. Code of Federal Regulations, Title 10-Energy, Part 50-"Domestic Licensing of Production and

Utilization Facilities."

2. J. F. Cook and B. W. Brown, Evaluation of Nondestructive Examination Requirement Changes with

Updates of Section XI, Division I of the ASME Code, EGG-SD-6893, November 1985.

3. Commission of the European Communities, A Summary of the PISC-II Project, PISC-II Report No. I,

June 1985, General Directorate XII, Joint Research Centre, Ispra Establishment, S.A./l.07/Cl.85.13.

4. Proceedings of the Bolting Degradation and Failure in Nuclear Power Plants Seminar, Revision 1,

EPRI, December 1983.

5. U. S. Nuclear Regulatory Commission, "Stress Corrosion Cracking in Thick-Wall, Large Diameter,

Stainless Steel Recirculation System Piping at BWR Plants," USNRC Inspection and Enforcement

Bulletin 82-03, October 1982.

6. U. S. Nuclear Regulatory Commission, "Stress Corrosion Cracking in Large Diameter Stainless Steel

Recirculation System Piping at BWR Plants," NRC Inspection and Enforcement Bulletin 83-02,

March 1983.

7. American Society of Mechanical Engineers, "Procedure and Personnel Qualification for Ultrasonic Detec- L

tion and Sizing of Intergranular Stress Corrosion in Austenitic Piping Welds," Code Case N-409,

December 1984.

8. B. L. Baker, J. F. Cook, and H. A. Wilkin, Evaluation and Recommendationsfor Ultrasonic Examinations of ASME Class I and 2 Piping Melds, EGG-FM-5034, October 1979.

9. J. Muscara, Research Program Plan-Nondestructive Examination, NUREG-1 155, Volume 4, 1985.

10. G. H. Lewis, Degradation of Building Materials Over a Lifespan of30-100 Years, Commission of the [

European Communities, Final Report, Contract No. DE-A-001-UK, 1985.

11. A. D. Perkins, Analysis of Datafrom Irradiated Concrete Testing, EG&G Idaho, Inc., RE-M-80-002,

March 1980.

12. U. S. Nuclear Regulatory Commission, "Inservice Inspection of Pressurized Water Reactor Steam

Generator Tubes," USNRC Regulatory Guide 1.83, Revision 1, July 1975.

13. R. Klinman, G. R. Webster, F. J. Marsh, and E. T. Stephenson, "Ultrasonic Prediction of Grain L

Size, Strength, and Toughness in Plain Carbon Steel," Materials Evaluation, 38, 1980, pp. 26-32.

14. R. Klinman and E. T. Stephenson, "Ultrasonic Prediction of Grain Size and Mechanical Properties

in Plain Carbon Steel," Materials Evaluation, 39, 1981, pp. 1116-1120.

15. B. E. Droney, "Use of Ultrasonic Techniques to Assess the Mechanical Properties of Steels," in

Nondestructive Methodsfor Material Property Determination, C. 0. Ruud and R. E. Green (eds.),

New York: Plenum Press, 1984.

128
16. A. Vary, "Correlations Among Ultrasonic Propagation Factors and Fracture Toughness- Properties of

Metallic Materials," Materials Evaluation, 36, 1978, pp. 55-64.

17. A. Vary, "Correlation Between Ultrasonic and Fracture Toughness Factors in Metallic Materials," in

Fracture Mechanics, ASTM Special Technical Publication 677, 1978.

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I

13. CURRENT LIFE ASSESSMENT TECHNIQUES

C. E. Jaske

Both experimental and analytical techniques are

used to assess the remaining life of components that

are in-service. These include testing of surveillance

specimens, monitoring of operational parameters,

evaluation of samples removed from service-exposed

components, and prediction of damage accumulation

processes. In actual practice, one or some combination

of these techniques typically are used in making

residual-life assessments.

Surveillance programs are used to assess the

influence of neutron irradiation on the mechanical

properties of materials used in the beltline region of

reactor pressure vessels (RPVs). Owners of light

water reactors (LNVRs) are required by federal regulations to maintain such programs, and American

Society for Testing and Materials (ASTM) Designation: E 185 defines the "Standard Practice for

Conducting Surveillance Tests for Light-Water

Cooled Nuclear Power Reactor Vessels," which

provides guidelines for a minimum surveillance

program. Samples of base metal, weld metal, and

heat-affected-zone metal are enclosed in surveillance capsules. The capsules are placed within the

RPV at locations such that the amount of irradiation leads is greater than that at the inside surface

of the beltline region by a factor of I to 3. A schedule for capsule removal and sample testing is established so that the long-term effects of irradiation

can be assessed. Both tension and Charpy V-notch

tests are required. Other tests, such as hardness and

fracture toughness, are optional.

The status of surveillance programs in the United

States has been described in detail elsewhere. I The integrated surveillance program used for Babcock &

Wilcox plants has been described by Lowe.2 Data from

the surveillance programs provide indications of the

effects of radiation on tensile strength and Charpyimpact properties. Figure 13.1 shows the correlation of

brittle-to-ductile transition temperature shifts at

30 ft-lb (41 J) with tensile yield strength. A similar correlation of fractional decrease in upper-shelf energy is

shown in Figure 13.2. These two figures are from the

work of Odette,3 and they show that consistent trends

exist. The major shortcoming of such data is that their

relationship to more relevant properties for fracturemechanics analyses, such as fracture toughness and

crack-arrest resistance, is not well established.

Operational parameters, such as temperatures,

pressures, and amount of cycling, can be monitored.to provide an assessment of remaining life.

These parameters are not direct indicators of the

material's condition or state, so some type of damage accumulation algorithm is needed to estimate

their impact on material degradation. For example,

if thermal fluctuations were monitored at a critical

nozzle region, one would have to compute the corresponding thermal strain history in that region

and then use an appropriate fatigue-damage accumulation model to calculate estimates of expended

life and remaining life. A practical example of such

on-line monitoring of the condition of an oil-fired

1000-MW boiler is presented by Davidson.4 As

recently pointed out by Toman,5 equipment monitoring techniques should be (a) nondisruptive, (b)

noninvasive, (c) repeatable, (d) accurate, and (e)

cost-effective, in addition to yielding data that are

correlatable with the deterioration process. As an

example, he discusses the monitoring of the condition of an emergency diesel-engine generator system for a nuclear power plant.

Another approach to life assessment is to directly

measure the properties of samples that have been

removed from components. Typically, a boat, core,

or ring sample is removed from the component,

and the component is then repaired using procedures that satisfy the applicable regulations and

safety codes. As will be discussed later in

Chapter 14 of this report, there has been some

recent interest in removing miniature samples to

either eliminate or greatly reduce the amount of

required repairs. The main advantage of this

approach is that it provides a direct measure of the

mechanical properties of the material after service

exposure. However, archival data for that same

material are not usually available for comparison

and assessment of the degree of degradation. Using

typical properties for that comparison and assessment adds to the uncertainty of the analysis.

Another problem with this approach is that the

tests must be accelerated to obtain information in a

reasonable time period during an outage. If a timedependent degradation mechanism, such as corrosion or corrosion-fatigue behavior, is being

evaluated, the results of the accelerated testing

must be extrapolated to make remaining service-life

predictions, and the extrapolation introduces additional uncertainty.

As an example, the direct properties measurement approach was used by Battelle to provide data

for assessing the structural integrity of Type 316

11

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Li

0

i>

a Avy (MPa) Aay (MPa)

7-3077

Figure 13. 1. Plots of transition temperature' shifts indexed at 30 ft-lb'(Al J) versus static' yield stress changes for (a)

weld and (b) plate and forgings.

0.7

ILi

IC

a)

V)

co

0

0

ILL

USEO =

USEO =

Upper shelf energy for

unirradiated material:

Upper energy for

Irradiated material

[

EI.

0

0 :' hAUy (MPa) 400

7-3037

Figure 13.2. Fractional decreases in C, upper-shelf energy versus yield stress changes.

..:?/ - .1 ' , .. *.,

* L

131
I

stainless steel weld metal from the Maine Yankee

reactor coolant system.6 Fatigue test bars were

removed from boat samples of weld metal that contained microfissures. Low-cycle fatigue specimens

were machined from those bars and tested using

standard procedures comparable to the test procedures that are used to develop fatigue data for the

establishment of American Society of Mechanical

Engineers (ASME) code design fatigue curves.-The

fatigue data then were used to evaluate structural

integrity in terms of the expected fatigue usage factors and the safety factors incorporated in the

ASME code. As another example, tensile, fatigue,

fracture toughness, and Charpy specimens have

been taken from steam-turbine shafts to measure

properties for use in remaining life analysis.

Metallographic and fractographic examination

of samples from in-service components also is used

to help'assess remaining life. For example, if volumetric nondestructive evaluation (NDE) has indicated the presence and growth of a defect or both, a

sample of material containing the indicated defect

may be removed and sent to a laboratory for

detailed examination using conventional metallurgical techniques to ascertain the nature and extent

of the defect. Suspected regions of intergranular

stress corrosion cracking (IGSCC) in boiling water

reactor (BWR) piping systems have been examined

in this manner. In fossil plants, ring samples of

superheater and reheater tubes are often examined

metallographically to measure the degree of fireside and steam-side corrosion and the amount of

creep damage. 8 Similar samples of furnace tubes

from petrochemical plants have been examined

metallographically to measure the amount of creep

voids and fissures, where an empirical correlation

has been established between those factors and the

service-failure rate of the furnace tubes.8

The analytical prediction of damage accumulation is

an important technique used in remaining life assessment. In fact, some type of prediction of the accumulation of damage during anticipated future operations

is required to make any estimate of remaining life. The

models of material degradation processes discussed

previously are used to make the required calculations

of damage propagation starting from the current damage state of the material.

If no measure of the current damage, state is

available, one can estimate it by calculating the

damage accumulation during the past operating

history. However, knowledge of the past operating

history and the condition of the material that was

originally put into service is often incomplete, adding a great deal of uncertainty to this calculationof-current-condition approach. Thus, some

destructive or nondestructive measurement of the

current condition of the material is preferred for

remaining life analysis.

Analytical techniques for the prediction of

remaining service life have been pioneered by the

aircraft industry, where NDE and a fracturemechanics-based, damage-tolerant approach is

employ~ed. 9 .Complex computer programs have

been developed to compute fatigue-crack growth

and assess the potential for fracture in structural

components throughout the course of various

expected loading spectra. Because of the history

dependence inherent in niaterial degradation processes, the damage accumulation'calculations are

performed on event-by-event (cycle-by-cycle for

fatigue) basis throughout the course-of the simulated loading. For this reason, the computations

are usually time-consuming and demanding of

computer resources.

The type of analytical procedure described by

Jaske8 is often used to assess the remaining life of

steam-turbine rotors. A flow diagram showing the

technologies used in performing the life-assessment

analyses is presented in Figure 13.3. Examples of

the results are presented in Figure 13.4. Figure 13.4

(a) shows the ultrasonic indications that are

obtained from boresonic inspection of a rotor.

Cluster analysis is used to group the indications in

regions that are considered to be single flaws in the

flaw-growth analysis. The remaining parts of that

figure show the calculated stress distributions at

two flaw locations, the expected loading spectrum

for the rotor, and the predicted flaw-growth behavior during future service.

In performing plant-wide remaining-life assessments, it is useful- to have simplified, rule-based

approaches that can be applied rapidly and. easily to

compute conservative life estimates. Then the detailed

and extensive analyses are performed only for the cases

where the simplified rules show that a potential problem may exist. Simplified methods are usually application specific, so such procedures should be developed

for key reactor components, such as RPV welds, nozzle welds, piping Xwelds, coolant pumps, diesel generators, and internals, where the possibility of fatigue

damage is of concern.

Engle, 10 for example, has developed a simplified

approach for estimating fatigue-crack growth in

typical aircraft structural components under spectrum loading. Figure 13.5 shows a comparison of

predicted versus experimentally measured fatiguecrack growth for several design limit stress levels

that were made using his simplified approach.

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I

Analysis of

Inspection data

* G.E. inspection 1979

* Data plots/analysis

Stress and thermal

analysis

(NASTRAN/AXISOL)

* Finite element models/

analysis

* Isostress contours

* Flaw stresses and peak

stresses

. , i

.. , 1 r :

Identify NDI flaws

* Cluster analysis

.* Potential flaw

I link-up

s~~ ~~~~~~~~~~~~~~~~ I

Rotor steel material

fracture response data

* Literature/in-house

data daldN, Kic, KC

Residual life

I analysis

* NDI and assumed

flaws

* Service stress

spectrum-duty

cycle

Operational -

history data

* Cold/hot starts

* Cold/hot overspeeds

'- Structural Integrity

- evaluation

* Run/retire decision

* Reinspection Intervals

* Operation guides

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Figure 13.3. Life assessment analysis approach for steam-turbine components.

I

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150 -

140 -

130 _ 's 120 -

a 110 P

0,100 CD

a 90 _ 80 - 70 - 60 _ 50 2 42

I ._ i

Area of holding indic,

S

l 3

Flaw 3

* : 4.

X Individual indications0

a

43 44 45 46 47 48 49 50 51 52

Axial position (in.)

54 55

I (a) Boresonic data for rotor area showing flaw Indications

._

ci,

0

80 l l l

60

40-

Growth

20 _

0 1

0 1 2 3 4 5

r (In.)

Flaw 1 section

601

I-I

0

t

. 80

401

I I I I

-W Growth

I I I

1 2 3 4

r (In.)

Flaw 3 section

I 20

5

(b) Circumferential stress distributions

I

=1

0.

&V

Fracture Flaw 1 I

Flaw 3>Zz

Mao = 0.125 in.

/12co = 2.00 in. /

.- As~~~~~~~~~.1

I I I

. . . . . I.

0 2 Cycles 4 6 .8

Years of operation

10 12

(c) Yearly service spectrum for turbine rotor (d) Spectrum loading fatigue crack growth

behavior of bore flaws 1 and 3

7.3080

Figure 13.4. Example of results from a life assessment of a steam-turbine rotor.

134
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7000

.-Flights. 30

7-3060

Figure 13.5. Correlation of predicted with measured fatigue-crack growth rates under spectrum loading. 1.

Saxena et al.1I applied fracture-mechanics tech-.

nology to predict the residual life of a ships service

turbine generator (SSTG) casing using the general

approach illustrated in Figure 13.6. The region of

concern had a complex configuration as is shown in

Figure 13.7 and was subjected to cyclic thermal

stresses. They concluded that casings with cracks

<6.3 mm deep could be operated safely'-for

another 800 cycles, and they recommended intervals for future periodic inspections. This same type

of residual-life analysis could be applied to nuclear

reactor components with appropriate modifications to the materials properties, configurational

models, and failure criteria.

Broek9 has summarized the six basic types of

information that are needed to apply a fracturemechanics approach:

* The minimum detectable flaw size and.

type

* A prediction of the residual strength of the

structure with a flaw present and of the

critical flaw size associated with the failsafe load

* A description of the expected future loading history

* A determination of the crack growth response

from the minimum detectable size to the critical size associated with failure

A knowledge of the critical locations in the

structure where cracks are likely to

develop, such as geometric discontinuities

associated with fatigue-loaded welds

A reliable inspection of the various important regions of the structure or component, including consideration of the

accessibility to those regions.

With current NDE methods, defects can be detected

with relatively high confidence. However, as discussed

in chapters 11 and 12, !the sizing and classifying of

'those defects has much lower confidence and is an area

where major improvements are needed and where

much NDE research is concentrating.

Predictions of residual strength and critical flaw

size cani be made with reasonable confidence for

brittle fracture where linear elastic fracture

mechanics can be applied. For nuclear pressure

boundary comnponents, however, the possibility of

ductile fracture is often of concern. Inelastic fracture mechanics methods for dealing with ductile

fracture problems in piping and pressure vessels are

still evolving. Those methods need to model leakbefore-break and crack-arrest, so that fail-safe conditions occur at the critical crack size. For

applications to actual components, these methods

need to be developed and verified for dynamic as

well as static loading conditions.

I

I

135

I
Elastic finite element Cyclic stress-strain

analysis behavior

* Constrained p lasticity analysis

using Neuber approach Analytical

verification of

* Obtain elastic-plastic stress Neuber approach

distribution

* Calculation A J

* Convert to equivalent AK

= I

* Obtain A K versus crack length

Selection of failure criteria

th Crack growth analysis Crack size

r da f (AK) input from

ie dN nondestructive

N= lai f (^ K) ,inspection, ai

P red icted

remaining life

as a function of

initial crack

depth /

7-3074

Figure 13.6. Methodology for remaining life prediction of SSTG casings.

136

1
I

Plane CPlane B

I

Plane A,

2114% Cr. 1% MOLY

r

2-ft 3 114-In. dia.

1.0 in. = 2.54 cm

I

7.3059

Figure 13.7. Illustration or configuration of region where remaining life of SSTG casing was assessed. ' r

Descriptions of expected loading histories are The cyclic crack growth behavior of pressure vesobtained from past experience and from postulated ' sel steels in LWR environments has been well charworst-case scenarios that may' be encountered! acterized and that wo'rk is being xtended to piping

Because damage accumulation is strdngly history' materials. One area of concern is at very low cyclic

dependent, it is important that such loadinghisto- growth rates where there is so'me debate over the

ries be Well defined.- Unfortunately,`as pointed out existence of 'a threshold for 'cyclic crack growth.

by Yahr et al.,'12 this is noi the case in' the nuclear Additional work is needed to resolve this question.

industry, so they had to develop a unit histogram Asdiscussed in the earlier chapters of this report,

for the Zion-I plait for usein iheir fatiguedamage ',',the areas where cracks are likely to 'develop in

analyses. This type of work needs to be pursued'sot nuclear-plant 'components have 'been 'identified

that sets of reference histograms can be developed from operating experience and original design calfor routine use in remaining life calculations. ' culations. These include transition regions in the

L;

137

_. _1
I

RPV, closure studs, safe end welds, nozzle connections, elbows, control rod drive mechanism penetrations, and other welded regions in the pressure

vessel and piping system.'.'

Inspections of important regions of the nuclearplant components are. performed on a scheduled

basis. Improved quantitative descriptions of

detected flaws are required for use of a fracturemechanics approach.

All six of the items needed to apply a fracturemechanics approach are actually probabilistic in

nature. Ideally, they should be quantified and

employed in a probabilistic remaining life assessment methodology.

Recently, Sundararajan1 3 has critically reviewed

the state of the art in probabilistic assessment of the

reliability of pressure vessels and piping. Most of

the applications of that methodology have been in

the nuclear power industry. Figure 13.8 shows the

predicted failure rate of PWR RPVs for various

reactor shutdown pressure-temperature paths at the

end-of-life (EOL) fluence. For life extension, either

the core loading patterns would have to be modified to prolong the attainment of the EOL fluence

or the allowable pressure-temperature path would

have to be modified to provide an acceptable risk of

failure. In residual-life assessment, this type of

analysis should be made on a plant-specific basis to

allow for different materials properties and NDE

results.

As is illustrated in Figure 13.9, Sundararajan13

also has predicted the probability of protective barrier penetration from random fragment strike. If

the effective thickness or the penetration resistance

of the barrier is reduced during service exposure,

then the probability of penetration is increased. For

life extension, the barrier must be maintained and

refurbished as required to maintain an acceptably

low level of perforation probability.

Predictive methods for environmentally assisted

degradation and cracking are not as well established as

those for fatigue. A recent ASME symposium14 was

devoted to documenting the current state of the art in

this technical field. The typical approach to solving

corrosion'and stress-corrosion cracking problems has

been to adopt fixes that will eliminate or minimize the

occurrence of those problems in the future. 15,16 That

approach has been largely successful, so the predictive

methods may not be required. However, in order to

extend plant life, the components and regions of components that are'potentially susceptible to environmentally assisted degradation must be identified by

inspection so that the appropriate fixes can be implemented.

In the area of environmentally assisted fatigue

cracking, a great deal of work has been devoted to

developing predictive models14, 17 ,

18 and

Section XI of the ASME Code requires assessment

of environmental effects on fatigue-crack growth.

Bamford1 9 has reviewed the basis for reference

fatigue-crack-growth rate curves that are presently

included in Section XI. Based on detailed reviews

oftthe EPRI Database on Environmentally Assisted

Cracking (EDEAC), revisions of those reference

curves are being considered. 20 The EDEAC also

contains data from slow-strain-rate and stresscorrosion-cracking tests, as well as the data from

fatigue-crack-growth tests. Thus, it provides an

important resource of very relevant materials properties data that can be used in developing methods

for residual-life assessment.

Gilman21 '

22 used EDEAC to develop a useful

engineering model for predicting corrosion-fatigue

crack growth rates in LWR environments. The basis

for that model is illustrated in Figures 13.10 and

13.11. The lower bound of the cyclic crack growth

rate in BWR water, which coincides with the upper

bound of the cyclic crack growth rate in air, is

defined as the base rate. The actual time rate of

crack growth in LWR water then is correlated with

the base rate. Figure 13.10 shows that there is a

good correlation between the growth rates in BWR

water and the base rate times frequency, and

Figure 13.11 shows that there is a good correlation

between the upper bound of the growth rates in

pressurized water reactor (PWR) water and the

base rate times frequency. These correlations provide a useful tool to help estimate the remaining

service life in areas of the RPV where corrosionfatigue crack growth is a potential problem.

13.1 Summary

The major life-assessment techniques available

for application to LWR components are summarized in Table 13.1. For each technique, its characteristic features and its applications and/or

potential applications are briefly highlighted.

Testing of specimens from surveillance capsules

is used to assess the amount of radiation embrittlement that occurs in RPV steels during service

and is an important technique for ensuring the

structural integrity of the RPV at its EOL fluence.

Monitoring of operating parameters to establish

aging trends is not widely practiced, but with the

development of appropriate baseline data it can

provide a viable, economic tool to assist in

I

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I:

138

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I

^ff^n

2000 g

n 1500

a0 0

DE 1000

I.

104 failure perevent Act . // -

10-5 failure per event , //

10-6 failure per'event //

10 7 failure perevent' , j < / /

.0-~

, " t'0 Code limit curve

0.23% Cu and 23.0 F IRTNDT

EOL fluence

[

5001

l 1 I I.----

100 200 300

Temperature (OF)

400 500

743058

Figure 13.8. Constant failure rate and code-allowable pressure temperature paths for reactor shutdown at EOL

fluence.

I

Z

co

n

r0

Cl0.

0

a)-

0-

I

- -Barrier thickness (T).

7-3038

F Iur i 3 e P o r f o r a t i o n . b h

Figure 13.9. Probability of perforation versus barrier thickness.

139
I

Crack growth rates, A533B-1, 0.025% S in 288°C BWR water, R = 0.7, 0.8

10-3.

105- 4

10-5

2 10-6

", aK %S Key

p 1 o 6 / O / ~~~~~~~~~~~~~111 .2

010-716.5 025 v

S / 4 ~~~~~~~~~~~22 0.2

33 0

108 10-7 10-6 10-5 10-4 10-3

Base rate, 6.2 x 10-8 AK 2 -

5 x frequency (mm/s)

7-3081

Figure 13.10. Time rate of fatigue-crack growth in BWR water as a function of base rate.

Crack growth rates, A533B-1,in 2880C PWR water, R 0.5-0.8

10~-3

3.1 ~~~~~AK %S Key

20-6 6 0 1 1 0.025 A

S ~~~~~16.5 0.025 v

A ~~~ ~~22 0.025

O 10-7 33 0.025 0

A ~~~ ~~44 0.025 0

44 0.013

10-8 * ''it'1 I I II1,11 I Iufill

1 0.8 10-7 10.6 i0-5 10-4 1-

Base rate, 6.2 x i0-8 A K2-

5 x frequency (minis)

7.3057

Figure 13.1 1. Time rate of fatigue-crack growth in PWR water as a function of base rate.

140

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Table 13.1. Summary of current life assessment techniques

Technique Characteristics Applications

Surveillance programs

Monitoring of

operations

Testing of specimens

removed from

components

Metallographic and

fractographic

examination of samples

removed from

components

Analytical prediction

of damage accumulation

Uses specimens exposed to ' ' Measuring the radiation embrittlement

neutron irradiation in capsules of RPV steels. Limited amount of

within the RPV to measure test material is usually available.

tensile and impact properties. The relation of tensile and impact

properties to fracture toughness is not

well defined.

Measure and record tempera- Used to assess condition of

tures, pressures, and cycling mechanical equipment, such

for trending of performance. as pumps, bearings, engines, and

generators.'Need baseline data for

normal operating conditions.

Measure mechanical properties Used on piping welds in regions

of boat, core, or ring samples where repair is possible. Not widely

removed from the component. used because material removal is

destructive. Need to develop methods

for testing miniature specimens.

Measure metallurgical aging, Used on piping welds in regions where

void formation, and amount of repair is possible. Not favored because

microcracking. Identify removal of material is destructive.

mechanisms'of cracking. Need to develop methods that use

miniature samples or surface replicas.

Compute the amount of Widely used for areas where high

damage accumulation expected amounts of fatigue damage may

during future operations.. occur. Need to develop realistic

engineering approaches.

I

I

I

remaining life assessment. The destructive testing

and examination of samples removed from components provides useful data, but is not useful for

routine applications because repairs must be

made.

Analytical predictions of damage accumulations

are widely used in both design and remaining-life

assessments. Current approaches typically use very

simplified damage accumulation algorithms and

large safety factors to ensure conservative predictions. More realistic approaches that allow the

incorporation of history effects, fracture'mechanics, and probabilistic factors need to be developed,

especially to resolve differenice between the ASME

Code Section III and Section XI approaches to

fatigue damage analysis. r

EL.

141

I
I

13.2 References

1. American Society for Testing and Materials, Status of USA Nuclear Reactor Pressure Vessel Surveillancefor Radiation Effects, ASTM Special Technical Publication 784, Philadelphia, 1983.

2. A. L. Lowe et al., "Integrated Reactor Vessel Material Surveillance Program for Babcock & Wilcox

177-FA Plants," in Effects of Radiation on Materials, Twelfth International Symposium, American

Societyfor Testing and Materials, ASTM Special Technical Publication 870, 1985, pp. 931-950.

3. G. R. Odette, P. M. Lombrozo, and R. A. Wullaert, "Relationship Between Irradiation Hardening

and Embrittlement of Pressure Vessel Steels," in Effects of Radiation on Materials, Twelfth International Symposium, American Society for Testing and Materials, ASTM Special Technical Publication

870, 1985, pp. 840-860.

4. M. J. Davidson et al., "Monitoring for Life Extension," in Residual-Life Assessment, Nondestructive Examination, and Nuclear Heat Exchanger Materials, Proceedings of The American Society of Mechanical Engineers 1985 Pressue Vessels and Piping Conference, New Orleans, Louisiana, June 23-26, 1985, PVP-Volume

98-1, pp. 65-71.

5. G. J. Toman, "The Measurement of Equipment Degradation," Proceedings of the Thirteenth Water

Reactor Safety Research Information Meeting, NUREG/CP-0072, Volume 3, 1985, pp. 319-329.

6. R. D. Buchheit et al., Laboratory Studies of Type 316 Stainless Steel Weld Metal From the Maine

Yankee Reactor Coolant System, Final Report to Maine Yankee Atomic Power Company, 1971.

7. B. N. Leis et al., "Fatigue and Fracture Properties of Steam Turbine Shaft Materials," Fracture

Mechanics: Fourteenth Symposium- Volume II: Testing and Applications, ASTM Special Technical

Publication 791, 1983, pp. 11-101-II-I 19.

8. C. E. Jaske et al., "Life-Assessment of High-Temperature, Fossil-Plant Components," EPRI Conference on Life Extension and Assessment of Fossil Plants, Washington, D.C., June 1986.

9. D. Broek, Elementary Engineering Fracture Mechanics, Boston: Martinus Nijhoff Publishers, 1978.

10. R. M. Engle, Jr., "A Simplified Approach for Assessing Spectrum Loading Effects in Preliminary

Design," Advances in Life Prediction Methods, New York: The American Society of Mechanical

Engineers, 1983, pp. 223-227.

11. A. Saxena et al., Residual Life Prediction and Retirement for Cause Criteria for SSTG Casings, Part

II: FractureMechanicsAnalysis, Westinghouse Research Report 84-8D7-LARGE-P2, 1984.

12. G. T. Yahr et al., "Case Study of the Propagation of a Small Flaw Under PWR Loading Conditions

and Comparison With the ASME Code Design Life: Comparison of ASME Code Sections III and

XI," American Society of Mechanical Engineers 1986 Pressure Vessels and Piping Conference and

Exhibition, Chicago, Illinois, July 20-24, 1986.

13. C. Sundararajan, "Probabilistic Assessment of Pressure Vessel and Piping Reliability," Journal of

Pressure Vessel Technology, 108, 1986, pp. 1-13.

14. Predictive Capabilities in Environmentally Assisted Cracking, Proceedings of the WinterAnnual Meeting of the American Society of Mechanical Engineers, Miami Beach, Florida, November 17-22, 1985,

PVP-Volume 99.

142

I'
15. NV. E. Berry, "Stress-Corrosion Cracking in Light-Water Reactors," Thermal and Environmental

Effects in Fatigue: Research-Design Interface, PVP-Volume 71, ASME, New York, 1983, pp. 217-

230 .

16. B. M. Gordon and G. M. Gordon, "Materials Aspects of BWR Plant Life Extension," Proceedings

of the Thirteenth W'ater Reactor Safely Research Information Meeting, NUREG/CP-0072, Volume 3,

1985, pp. 293- 317.

17. Thermal and Environmental Effects in Fatigue: Research-Design Interface, PVP-Volume 71, ASME,

New York, 1983.

18. Proceedings of the Second International Atomic Energy Agency Specialists' Meeting on Subcritical

Crack Growth, Volumes I and2, NUREG/CP-0067, April 1986.

19. NV. H. Bamford, "Technical Basis for Revised Reference Crack Growth Rate Curves for Pressure Boundary

Steels in LWR Environment," Journal of Pressure Vessel Technology, 102, 1980, pp. 433-442. ..

20. R. Rungta, H. Mindlin, and J. D. Gilman, "Applications of EPRI Database on Environmentally -

Assisted Cracking in the Nuclear Industry," Paper No. 42, Corrosion '86, National Association of

Corrosion Engineers, Houston, March 17, 1986.

21. J. D. Gilman, "Application of a Model for Predicting Corrosion Fatigue Crack Growth in Reactor

Pressure Vessei Steels in LWR Environimen ts," Predictive Capabilities in Environmentally Assisted

Cracking, Proceedings of the Winter Annual Meeting of the American Society of Mechanical Engineers, MiamiBeach, Florida, November17-22, 1985, PVP-Volume99, pp. 1-16.

22. J. D. Gilman, "Application of a Model for Predicting Corrosion-Assisted Fatigue Crack Growth in

LWR Environments," in Proceedings of the Second internationalAtomic Energy Agency Specialists'

Meeting on Subcritical Crack Growth, Volume?2, NUREG/CP-0067, April 1986, pp. 365-384.

.43

.. . ..

' ' ''!, ' , .':;i 143

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I

14. NEW OR EMERGING METHODS FOR

INSPECTION AND LIFE ASSESSMENT

C. E. Jaske and L. J. House

As a direct result of the recent increased interest

in extending the operating life of all types of major

industrial plant equipment, including nuclear

power-plant equipment, new methods for inspection and life assessment of that equipment are

being developed. The inspection methods are

required to evaluate the current damage state of the

materials in such equipment without introducing a

significant amount of additional damage from the

measurement technique that is used. Current

remaining life-assessment methods, such as those

of Section XI of the American Society of Mechanical Engineers (ASME) Code, are based on inspection to find cracks or crack-like defecis' and using a

fracture-mechanics approach.

14.1 Nondestructive Examination

Techniques

The principal thrust of current research and development in nondestructive examination (NDE) is to

extend the state of the art from qualitative flaw detection and estimation of the size of large continuous

flaws to quantitative measurements of performancerelated material properties and more accurate quantitative flaw characterization. Emerging NDE

technologies consist essentially of variations and

improvements in existing measurement or NDE methods rather than new sensor technologies.

The fundamental basis of all NDE is predominantly

the measurement and analysis of either electromagnetic or mechanical radiation. Of the emerging technologies, ultrasonics continues to be the dominant

source of new inspection techniques and advances in

quantitative inspection technology.

The development of automated and expert systems for ultrasonic, eddy-current, and radiographic inspection also represents a major

contribution to advanced NDE. Improvements in

digital data acquisition of eddy-current and ultrasonic inspection data are now enabling real-time

analysis and archiving of inspection results and

permitting the application of adequately sophisticated models to provide more meaningful and

quantitative interpretation of the inspection data.

Advances in imaging systems also offer promise for

greater realization of the potential of radiographic

examination. Improvements in digital radiographic

image processing include qualitative image

enhancement and quantitative image analysis. The

development of video systems with improved resolution presents the possibility of practical real-time

radiography offering the advantages of immediate

presentation of a flaw image and potentially

enhanced flaw detection probability because many

angles can be examined. If combined with digital

processing, real-time radiography can provide a

powerful quantitative inspection tool.

A shift from the development of periodic inspection

methods to in situ monitoring techniques offers some

advantages. Because monitoring techniques rely on

detecting changes in material parameters rather than

making absolute measurements, a number of calibration problems that limit the usefulness and reliability of

periodic NDE techniques based on making absolute

measurements are avoided. With in situ monitoring,

the possibility arises for detecting microstructural or

metallurgical changes in a material, measuring changes

in component stresses, and measuring internal temperatutre. Of the techniques that provide large area monitoring capabilities, acoustic emission is one of the more

promising emerging technologies. Advances in analytical models for more quantitative analysis of acoustic

emission events, the development of small aperture

arrays, and greater sophistication in the hardware for

data acquisition offer considerable promise for acoustic emission to become a useful tool to monitor structural integrity.

Table 14.1 shows Broek's I summary of the seven

basic nondestructive inspection techniques that

were discussed in Chapter II of this report and are

used in conjunction with the fracture-mechanics

approach. The basic physical principles and the

typical application are summarized for each technique. The major challenge for quantitative use of

these techniques in a fracture-mechanics context is

to accurately determine the size, shape, location,

orientation, and type of crack or flaw. Furthermore, it would be useful to extend these techniques

to other types of material aging or damage.

As is illustrated in Figure 14.1, Masuyama 2 summarizes methods that could be used for the NDE of materials that have been in service for some period of time,

pointing out that both surface and internal cracks are

of concern. Most'of the traditional NDE methods can

be used to detect surface cracks. However, only

144
I

Table 14.1. Inspection methods (from Broek) I J.,

Method Principles

Visual

Penetrant

Naked eye, assisted by magnifying glass, low

power microscope lamps, mirrors.

Colored liquid (penetrant) is brushed on

material and allowed to penetrate into cracks.

Penetrate is washed off and quickly drying - suspension of chalk is applied (developer)..

Remnants of penetrant in crack are extracted by

developer and give colored line.

Part to be inspected is covered with a layer of a

fluorescent liquid containing iron powder. Part

is placed in strong magnetic field and observed

under ultraviolet light. At cracks, the magnetic

field lines are disturbed.

Application

Only at places easily accessible.

Detection of small cracks

requires much experience.

Only at places accessible.

Sensitivity of same order as

of visual inspection.

Magnetic particles Only applicable to magnetic

materials. Parts have to be

dismounted and inspected in'

special cabin. Also notches and

other irregularities give

indications.-Sensitive method.

X-ray X-rays emitted by (portable) x-ray tube pass

through structure and are caught on film.

Cracks, absorbing less x-rays than surrounding

materials, are delineated by black'line on film.

Ultrasonic

Eddy current

Acoustic emission

Probe (piezo-electric crystal) transmits highfrequency wave into material. The wave is

reflected at the ends and also at a crack. The

input-pulse and the reflections are displayed

on an oscilloscope. Distance between first

pulse and reflection indicates position of crack.

Interpretation: Reflections of cracks disappear

on change of direction of wave.

Coil induces eddy current in the metal. In turn,

this induces a current in the coil. Under the

presence of a crack the induction changes; the

current in the coil is a measure for the

surface condition.

Method with great versatility

and sensitivity. Interpretation

problems if cracks occur in

fillets or at the edge'of -

reinforcements. Small surface

flaws in thick'plates difficult to

detect.

Universal method because a

variety of probes and input

pulses can be selected.

Information about the size and

the nature of the defect (which

need not be a crack) are

difficult to obtain.'

Cheap method (no expensive

equipment) and easy to apply.

Coils can be made small

enough to fit into holes..

Sensitive method when applied

by skilled personnel. Little or

no information about nature

and size of defect.

Inspection while structure is

under load. Continuous

surveillance is possible."

Expensive equipment required.

Interpretation of signals is

difficult.

I" I' I I .

Measurement of the intensity of stress waves are

: emitted inside the material as a result of plastic

' deformation at crack-tip-, and as a result of

* crack growth.

I

145

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I

Damage mode Nondestructive detecting method

- For surface cracks _ _ Cracking * Penetrant test

( Surface cracks 0 Magnetic particle test

Internal cracks&* Ultrasonic test ® Internal cracks * Electric resistance method

- For internal cracks

* Radiographic test _ Jo * * Ultrasonic test

* Replication method \ } cracks Micro * Physical test (not established) I Ultrasonic

Electric

Magnetic

Radiographic,

* Replication method

Micro voids * Physical test (not established)

Ultrasonic

Electric

Magnetic

Radiographic

0 Replication method Structural * Extraction method

\* .8;-. ...... ^- degradation * Physical test

( Hardness

* "* X-raydiffraction

< P* : .. - nearlite Ele d* ctro-chemical tost

Virgin' Exposed

Ferritic steel

\/ *tN 5 'Examples i of precipitates analysis

a7 phase Fe Fe Cr Exposed . (precipitationJ I S MVg

Virgin Exposed Carbide a phase Intensity

Austentic steel depends on

exposed time

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Figure 14.1. Methods for nondestructive examination of materials after service exposure.

146

U
I

volumetric methods, such as radiographic and

ultrasonic testing, can be used to'detect internal

cracks. Metallographic replication can be used to

detect microstructural damage, such as-cracking

along grain boundaries or voids, at accessible surfaces, but physically based NDE methods for-the

detection of such microstructural daimiage have not

been developed for general use. "Time-dependent

changes in microstructural features, such as the ferrite phase in austenitic-ferritic stainless steels and

precipitate phases in stainless steels, can be detected

by replication, extraction, hardness testing, x-ray

diffraction, and electrochemical testing at accessi- --

ble locations.

Several general review articles on the current sta- !,

tus and future prognosis of NDE technology have

been published recently.3 -

5 Posakony3 pointed out

that the future thrusts in NDE development include

(a) defect characterization by means of advanced

systems, (b) continuous monitoring of equipment

by means of acoustic emission and vibration analyses, and (c) assessment of materials properties and

monitoring of material condition by various NDE

techniques for use in life extension efforts.

McClung4 indicated that improvements in NDE

are coming from (a) the increased use of computers

for acquisition, processing, and interpretation of -

large quantities of data in the field, while the

inspection is taking place, (b) the development of

visual display systems that provide enhanced

images for examination and interpretation, (c) a

better understanding of the physical phenomena

employed in NDE, and (d) the development of new

NDE techniques. Bernard5 described the basic

principles of time-of-flight-ultrasonics, and discussed its use in inspecting a pressure vessel with

steel walls that were almost one foot thick.

Muscara 6 has reviewed the current status and future

plans for the USNRC research program on NDE. That

report describes projects that are developing arid e&lu-''

ating methods both for in-service irispection (ISI) of

key reactor components usinig ultrasonics'and eddy-"

currents and for continuous monitoring of key reactor

components using acoustic emissions. Each of the foollowing subsections'discus'ses in more detail the emerg'-

ing NDE methods that could be applied to key reactor

components.' Relevant methods being develojed both

by the nuclear power iridustry and by other industries,'

such as fossil po&er, petrochemnical,4 and aerospace, are

included.

14.1.1 Advanced Processing of Data. Ultrasonic inspection is used for the reactor pressure vessel (RPV), piping, and nozzles. The synthetic

aperture focuiing'technique (SAFT) for ultrasonic

testing (UT)' has been developed to provide

enhanced visual images of flaws detected in components during inspection.6 The authors of Reference 7 reviewed their field experience with the

application of SAFT-UT to the inspection of welds

in the recirculation piping system at the Dresden-3

and the Vermont Yankee boiling water

*reactors (BWRs). The authors' initial results are

encouraging, but they indicated that destructive

examination of samples from the sites of the flaw

indications are required to fully qualify the results

of this new technique..

Figure 14.2 (from Reference 7) shows a conceptual diagram of a SAFT field system. It is evident

that the real-time processing involved requires a

major commitment of computer resources. However, with the rapid advances being made in digital

computer technology and parallel processing, such

resources will become readily available at reasonable costs within the next few years.

Bemard5 described a case history where SAFT-UT

was successfully used to inspect a refinery's internally

stsinless-steel-clad, ferritic-steel pressure vessel. The

vessel was inspected through the thickness of its

'-12-in. wall. SAFT was used to distinguish between

indications from reflectors and flaws.

The authors in Reference 8 reviewed emerging

computer-aided UT inspection systems for the

automated -examination of nuclear piping. That

review provides a'detailed summary of the hardware and software'that is used in six of those pipeinspection systems. The systems can be controlled

automatically and remotely, and the hardware is

assembled in reasonably small packages. An adaptable track-mounted, automated pipe scanner is

used with'four of ihe systems. All of the systems

provide fairly extensive data recording. The computerized iultrasibihic' data acquisition arid processing system and the ultrasohic'data' recording' and

processing system also provide data'analysis capabilities. As these systems'evolve,'it is expected that

expert software systems will be incorporated into

the data analysis to help the inspector in real-time

interpretation of the UT information'.

The authors in'Reference 9'developed a prototype syktern for 'detecting surface cracks' inp'iping

using ' the' dc-poienitial-drop m'ethod and have

applied it to detecting sirface cracks oni the inside

of a 12-in.'diameter austehitic' stainless' steel pipe.

Their equipment includes scannerslfor both the

inside and outside surfaces and a personal computer. They performed a large number of solutions

for the electric potential distributions around

147
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Pre-amplifier

and mechanical

Operator

console

Figure 14.2. Conceptual diagram of SAFT field system.

various sizes and shapes of cracks using the finiteelement method. During an inspection, they measured a distribution and compared it with their

library of solutions to estimate the crack configuration using the personal computer..

Pressurized water reactor (PWR) steam generator tubes are normally inspected by. means of

eddy-current testing. 6 The state-of-the-art eddycurrent techniques are being evaluated and statistically based reliability data are being developed in

an extensive evaluation of a steam generator that

has been retired from service. 6 To overcome problems associated with detecting circumferential

cracking and intergranular attack, a 16-element

pancake-coil eddy-current probe has been constructed and is being evaluated in conjunction

with that work.

Whaley and Flora1 0 discussed the application of

microcomputers to eddy-current testing. They have

developed a computerized system for detecting

flaws in steam generator tubes near the support

plates. They also have performed a preliminary

evaluation of the use of a deep-penetration, eddycurrent method for detecting flaws in the welds of

austenitic stainless steel pipe, with encouraging initial results. However, the computer-aided techniques for eddy-current testing are not as well

developed as those for UT.

Thomel I discussed the application of acoustic

(or ultrasonic) holography to defect imaging. To

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date, this technique has been'liiniate to laboratory

evaluation, where it provides excellent images of

defects. However, 'significant development of

angle-beam methods is required before this method

can be employed to inspect components.

14.1.2 On-Line Monitoring. The 'USNRC

research plan6 includes two major programs that

deal with on-line monitoring of reactor'components using acoustic emission (AE) technology.

One program is aimed at using AE to monitor crack

growth in pressure boundaries, 12 and the other

program is aimed at using AE for leak surveillance

in light water reactor (LWR) systems. 13

Methods for AE monitoring of crack growth

have been developed in laboratory studies, evaluated during a long-term fatigue test of a pressure

vessel, and demonstrated by field surveillance on

the Watts Bar-I reactor.12 That work has shown

that AE sensors and 'instrumentation can' be

deployed successfully in a reactor environment.

Fatigue-crack growth in pressure'vessels and, in

some cases, stress-corrosion crack growth in stain-:

less steel piping can be detected by AE monitoring,

even with interference from background reactor

flow noise. Thus, AE monitoring appears to be a

promising technique for on-line monitoring of critical areas in key reactor components.

Kupperman' 3 reviewed the current leak detection practices for the'cooling systems at 74 nuclear

plants. He concludes that the current methods may

not be sensitive and reliable enough for detecting

some types of possible leaks, especially with regard

to the Duane Arnold safe end cracking incident.

Work is underway to improve the existing leakdetection systems and to develop two new types of

systems-moisture-sensitive tape and AE monitors. Moisture-sensitive tape has been installed in a

few reactors for evaluation; it has a tendency to give

false alarms and, does not provide quantitative

information about the leak rate. AE leak detection

is being evaluated both through laboratory testing

and field demonstration. The initial results' are

encouraging, but more software development and,

field validation of the technique are needed before

AE leak detection can be used with confidence.,

Methods for on-line monitoring of pressure

boundary components also are being developed for

fossil plants. Davidson' 4 describes a prototype'system for on-line monitoring of boilers to record and

accumulate data during typical operating conditions, during startup, shutdown, and during

operating transients. That system is being evaluated

at Consolidated Edison's Ravenswood-30 Unit.

Similar systerfs'coiild be installed in a nuclear plant

to obtain important and valuable information on

key parameters during typical operating histories.

Fryl51 6 discusses the use of neutron-noise analysis for monitoring the conditions in a reactor core.

Ex-core neutron noise 'measurements have been

made at 13 LWRs, and a library of data for future

reference has been 'established. This 'approach

shows promise for monitoring vibrations in PWR

internals and possibly inadequate core cooling in

PWRs.' :Neutron noise could -be used to detect

abnormal vibration of instrument tubes and fuel

boxes in BWRs, but other vibrations in BWRs were

difficult to detect because of the large background

noise from boiling. This technique is not yet appropriate for on-line use, but work is continuing to

develop it further.

.GalpinI7 has outlined'a'practical, economic

approach to either manual or automatic monitoring of component performance to develop information on aging trends. He emphasizes that

standard procedures should be employed to monitor parameters, such as temperature, flow, pressure, vibration, and acoustic emission.' Repeatable,

accurate, and reliable data are needed to document

aging trends and use those trends to establish safe

operating limits.

14.1.3 Nondestructive Measurement of Materials Properties.; Most of the current and emerging

NDE techniques are limited to the detection and

characterization of flaws and cracks that are in the

form of mechanical or elastic discontinuities in

components. Recently there has been an increased

interest in the nondestructive measurement of

materials properties (NDMMP), such as tensile

strength, fracture toughness,' impact toughness,'

creep strength, grain size, and precipitate morphology, 'that 'heretofore have traditionally been measured by destructive 'testing and examination of

specimens. 'As illustrated by the papers in a recent

volume,18 the techniques being evaluated for

NDMMP include ultrasonic testing, eddy-current

testing, and x-ray diffraction analysis. Most of the

NDMMP work performed to date has been'

restricted to laboratory studies. Field applications'

will depend upon future development of apparatus

and techniques. i ' ' - --

Generaziol 9 investigated the use"of ultrasonic

attenuation for measuring the mean grain size of

copper and nickel. Based on the results of his study,

he has suggested a nondestructive method for

checking the heat treatment of metals. In cases

where aging causes microstructural changes to

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occur during service at LWR operating temperatures, this type of technique might be used to measure those changes. For example, in studies of

gas-turbine alloys, it was. found that ultrasonic

attenuation could be used to detect and quantify

microvoid creep damage.a It may be possible to

develop similar correlations for radiation-induced

damage in key reactor component materials.

Swanson 20 and Vary21 studied the use of ultrasonic attenuation measurements for ranking material fracture toughness. They have found

correlations between ultrasonic measurements and

the plane strain fracture toughness of various

grades of steel and certain titanium alloys. However, the measurements made by the ultrasonic

methods used in these studies required samples

with smooth plane parallel surfaces and a very

accurate measurement. of the sample thickness.

Other methods, such as those employing backscattering,2 2'

23 are being developed to eliminate this

restriction, but they require spatial- or directionalaveraging techniques to eliminate the effects on the

data produced by the coherent nature of grain scattering. The principal drawback of this approach is

the complex scanning requirements necessary for

spatial and directional averaging. Digital processing methods based on frequency averaging may

eventually provide an alternative to the complex

scanning requirements of spatial or directional

averaging necessary for eliminating the coherentgrain scattering effects.

Researchers 4 have successfully made ultrasonic

measurements of stress intensity factors caused by

surface cracks, but the study was limited to brittle

nonmetallic materials. The method was based on

the measurement of the reflection coefficient of a

Rayleigh wave incident to the crack. Clark2 5

extended the use of ultrasonic phase velocity measurements to the measurement of stress and stress

intensity factors around crack tips in aluminum. In

that work, the ultrasonic method used the birefringence of a horizontally polarized shear, wave that

occurs in a slightly orthotropic plate in a state of

plane stress.

Bussiere26 reviewed the techniques used for

NDMMP, with emphasis on electromagnetic and

ultrasonic methods for on-line measurement of the

properties of steel. He indicated that eight techniques are useful for NDMMP:

a. Unpublished results of studies by C. M. Jackson at Battelle

Columbus, Columbus, Ohio, 1982.

* Magnetic, including Barkhausen noisefor hardness,' grain size, precipitate size,

morphology, and tensile strength

* Ultrasonic-for formability, grain size,

tensile strength, and fracture toughness

* X-rays-for residual stress and stress

relaxation

* Resistivity/conductivity-for chemical

composition, phase transformations, and

state of cold work

* Thermal and thermal imaging-for grain

structure and extent of hardening

* Small-angle neutron scattering-for particle,

precipitate, and microvoid characterization

* Mossbauer spectroscopy-for microstructure and phase characterization

* Position annihilation-for damage caused

by plastic deformation, fatigue, and creep.

None of these techniques have been widely applied

to characterize material-aging processes for LWR

components, but with further development and the

establishment of reference standards some of them

may prove useful for that purpose in the future. For

example, it may be possible to use magnetic techniques to measure the volume fraction of ferrite

and mean ferrite spacing (Chopra and Chung27

have shown that both of these parameters are

important) in stainless steel castings to help assess

their potential for embrittlement by thermal aging

during lons-term service.

Yavelak2 8 used Barkhausen noise analysis

(BNA) to characterize the service-induced damage

in a secondary superheater outlet header that was

removed from a 240-MW boiler after 24 years of

service. The results indicate that this method has

potential for application as a tool to nondestructively measure damage. BNA has been used. to

inspect and locate damaged regions in headers and

steam lines in the field during outages.29

Eddy-current methods for measuring material

properties are being developed by Clark and

Metala.3 0 They have used eddy-current techniques

to measure applied and residual stress, tensile

strength, the amount of temper embrittlement, and

the degree of creep damage. Figure 14.3 shows their

measured correlations between the yield strength

and eddy-current response for a low-alloy steel

heat-treated to different strength levels. They point

out that a sufficient catalog of standard background data needs to be developed before the eddycurrent method can be employed as tool for

residual-life assessment.

e

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150

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8

6

5

4

8

7

Test frequency 2.0-i n. (5.08 cm a. kHz

probe 1. ~

Ground surface

ksi = 6.9 MPa

30 100 120 140 160 180 201

5.0 kHz

30 100 120 140 160 180 20

0

I

U,

C

0

CL

U,

C.ID

~0

w1

6

5

4

8

0

I *I l l I

1.0 kHz 7 _- 1.

6

5

4 L.

80

I I I I I

100 120 140 160 180 0

8

7

6

5

4

84

-

I I i I I

I0.5 kHz

_ I - .

. . I;

I

5 100 * 120. 140 160

Yield strenath (ksi)

. 180 - 200

- ''' . .. 7-3056

Figure 14.3. Eddy-current response versus yield strength (2-in.-diameter probe). 1.. . . p ., .

Electrochemical methods can be used to nofides-'

tructiely measure material properties: Because of

the 'problem of intergranular stress-corrosion

cracking (IGSCC) in BWR piping,- an extensive

effort was directed-toward developing a nondesItructive'electrochemical testing procedure for mneasuring the susceptibility of austenitic stainless steels

to IGSCC. This work resulted in the development

of the electrochemical potentiokinetic reactivation

(EPR) test, which has been described and evaluated

by Majidi and Streicher.31 In most cases, the EPR

test gives good, quantitative measures of the degree

of sensitization with a variety of operators and

instruments, and portable equipment is available

for performing it in the field. A double-loop EPR

test procedure has been developed to simplify the

making of field measurements.3 2

Shoji and Takahashi33 used an electrochemical

method similar to the EPR test to evaluate the shift

in Charpy transition temperature and the creep

damage in steam-turbine rotor and casing steels

that result from material degradation during

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long-term service. They found a correlation

between the measured electrochemical parameters

and the shift in Charpy transition temperature

caused by temper embrittlement and suggested that

their method provides a userul tool for nondestructively measuring the degree of temper embrittlement. It may be possible that a similar approach

could be employed to nondestructively evaluate the

degree of radiation embrittlement in RPV steels.

In heavy-section components, indentation hardness measurements may provide a nondestructive

evaluation of the degree of material aging. It may

be possible to correlate hardness readings with the

degree of radiation embrittlement in RPV steels34

and then make hardness measurements to assess

radiation embrittlement during extended service.

As an example, Cane35 outlined a procedure for

using hardness measurements to assess the remaining life of thick-section steel components in fossil

power plants.

14.1.4 Surface Replication. Making cellulose

acetate replicas of sample surfaces for detailed

examination at magnifications up to about

1O,OOOX using electron microscopy has been a typical laboratory procedure for many years. Recently,

there has been a great deal of interest in using replication as a tool for field NDE2 ,36-43 of components as part of remaining life-assessment

activities. That work has concentrated on headers

and steam lines in fossil plants rather than on

nuclear plant components, but the same techniques

could be adapted for use in LWR applications.

Figure 14.4 shows an illustration of the replication and extraction methods from the paper of

Masuyama. 2 The surface to be examined must be

accessible. The area to be replicated is first cleaned,

metallurgically polished, etched, and moistened

with an organic solvent. The cellulose-acetate tape

(plastic film) is then applied and allowed to sit in

place for a time while it softens and adapts to the

surface. The replica is then removed to provide a

negative image of the surface. For extraction, the

etching is deeper so the particles in the microstructure can be removed for examination. To produce

high-resolution images, the replicas are coated with

carbon or gold in the laboratory and examined

using optical or electron microscopy.

GPU Nuclear Corporation has developed a simplified replication technique where mechanical polishing time is minimized by the use of a final

electropolishing step and the replica is shadowed

with a felt-tip marking pen while it is 'drying.4 2

This approach permits replicas to be made more

rapidly (in \15 min) at the sacrifice of high resolution. These replicas are examined optically at low

magnifications using a portable microscope to

obtain preliminary information in the field and at

magnifications up to I,OOOX in the laboratory. A

potential danger of such a high-speed technique

(electropolishing) is that it may produce artifacts

that might be incorrectly interpreted as damage.

For many nuclear-plant applications, a remote

replication apparatus would be required. A prototype remote replication unit has been developed for

the metallographic examination of steam-turbine

rotor bores,43 and a more advanced rotor-bore

remote replication system is planned for development as part of a future Electric Power Research

Institute project.a Also, a replication manipulator

was developed for the remote examination of steam

generator tube bores.44 That equipment.was successfully used to inspect regions down to 6 feet in

the 0.57-in.-inner-diameter steam generator

tubes.45 In another study, replication was used to

detect and document creep-fatigue cracking near

welds in a test pressure vessel.4 The feasibility of

using remote replication apparatus has been demonstrated, but specific equipment needs to be developed for use in LWR applications.

Replication may be a valuable inspection tool for

areas near welds in LWR vessels, piping systems,

and internals. It also may prove useful for inspection of steam generator tubes. It can be used to

detect very small cracks or to assess changes in

microstructure because of aging. Replicas of cast

stainless steels could document the amount, distribution, and spacing of ferrite and might indicate

changes in the microstructure related to long-term

embrittlement of the ferrite phase. For these and

other possible applications, development and demonstration work will be required before replication

can be used as a routine ISI tool.

14.2 Life-Assessment Methods

As reviewed previously, the current lifeassessment techniques are both experimental and

analytical. The newer or emerging methods for life

assessment employ, miniature test samples that can

be removed from a component with negligible damage to the component, on-line procedures for the

calculation of damage, and improved models of

material degradation and damage accumulation,

a. The project. Steam Turbine Rotor Life Assessment and

Extension, EPRI RP2481-5, is scheduled to begin in 1987.

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152

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" 1.

Carbide Crack

\ / Surface of part

after polishIng ,77~k7~Iz and etching

Plastic film

| Softened film

adapts

-to surface

Carbide Crack

< Replica with

negative of

Rep n microstroctur

Replication method

Electron microscope

and analyzer

Precipitates

Surface of part

alter deep

Aetgtchlng

Plastic film~~~ Softened film

adapts

to surface

Precipitates

- I I. I ~ Replicawith

precipitates

Extraction method 7Tao I

Figure 14.4. Illustration of replication and extraction methods.

especially with the inclusion of probabilistic

parameters that are needed to assess 'safety on a

realistic basis. Each of these topics is'addressed in'

the remainder of this section of the report. -

14.2.1 Miniature Specimen Testing. Testing of

specimens removed from an actual component during service, as mentioned earlier, can provide a

direct measure of the degree of aging. If miniature'

samples are used, then the amount of material that

must be removed from the component will be small

so that little or no repairs will be required. Minia-'

ture specimens also can be valuable for surveillance

testing where only a limited amount of test material

is available and where space available for material

irradiation is restricted. For example, it may be pbs--

sible to make miniature specimens from the- ;'

untested grip portions of standard size specimens

to obtain additional data. Studies are in progress to

evaluate the applicability of- miniature specimen'

technology (MST) to residual-life evaluation.

As discussed by Jaske,4 7 there are two main con-:

straints to the application of MS1T

The specimen to be tested must be large

enough that it represents the material from

which it has been removed

* 'The specimen to be tested must be small

enough that its removal from the component being evaluated does not impair the

structural integrity of that component.

The first requirement (minimum size) expresses the

need for the miniature specimen to exhibit properties that are the same as those of the bulk material.

Experience has shown that the requirement can be

satisfied when the minimum specimen dimension is

five to ten times as large as the largest strengthcontrolling microstructural feature. For steels used

in LWR components, that feature usually, is the

grain size, which rarely is greater than 0.1 mm. The

minimum-size criterion is applicable, however, only

to yield strength, and maybe to ultimate strength.

A shear-punch test has shown promising results for

measuring stress-strain response. 48 That test uses small

disks of the same size as those used for transmission

electron microscopy (-3.0 by 0.25 mm). During testing, the load versus deflection response of ihe disk is

measured. Finite element stress analysis then is used to

develop stress-strain curves. For Type 316 stainless

steel, -this approach gave results comparable to those

obtained from conventional tensile tests. Similarly

promising results have been obtained using hardness

tests.4 9

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Fatigue cracking and fracture toughness, very

important properties in remaining life evaluation

and reliability analyses, are strongly influenced by

grosser microstructural features, such as inclusions

and banding.47 Both of these properties also are

notch sensitive, so predictions based on data from

unnotched specimens can be extremely misleading.

It also is well known that fracture toughness is size

dependent; the problem is that small specimens

provide nonconservative data compared with data

from large specimens. For these reasons, the carbon segregation in specimens from an American

Society for Testing and Materials (ASTM) A 508

pressure vessel steel limited the minimum specimen

size to 5 mm for obtaining successful measurements of plane-strain fracture toughness, JIc,

where the criterion for success was comparison of

the results with those obtained from larger specimens that satisfied the criteria of ASTM E 813.44

Recently, Misawa and Hamaguchi 50 evaluated

two types of miniature specimens for measuring the

fracture resistance of ferritic steels. They used a

24-mm-diameter, disk-shaped, compact-type specimen with thicknesses from 1.8 to 5.0 mm to measure the tearing modulus. Also, they used

small-punch specimens (0.25xlOx1O mm) taken

from the ends of broken Charpy specimens to measure the ductile-to-brittle transition temperature

and obtained a linear correlation with transition

temperature measured using Charpy specimens.

The fracture-toughness data obtained using MST

show promise. However, one must be careful in

selecting a specimen size for the measurement of

fracture toughness and fatigue cracking. Several

task groups of the ASTM currently are working on

guidelines for the application of MST to fracturetoughness testing.

Battelle's Columbus Division is currently undertaking a major in-house study of MST. 51 The

objective of that laboratory study is to evaluate the

ranges of parameters over which miniature specimens can be successfully employed. Applications

being evaluated include characterization of deformation' behavior, fracture toughness, and crackgrowth resistance.

Miniature specimens also are being evaluated for use

in assessing the remaining life of fossil-plant components. Askins 52 described the development and validation of a miniature-specimen testing technique for

evaluating the tensile creep and stress-rupture properties of steels. As is illustrated in Figure 14.5, their specimen design employs a 3-mm-diameter by 13-mm-long

piece taken from a component in service. Extension

pieces are electron-beam welded to each end of that

piece to provide material for gripping in the test fixtures. The final specimen has a configuration with a

2-mm-diameter by 10-mm-long gauge section. For

accelerated testing at temperatures near 10000F

(600'C), an argon gas environment is used to avoid

erroneous results from excessive oxidation of the test

specimen. For a lower temperature of 10000F (5380C),

Saxena 53 tested similarly sized creep specimens from a

retired l-l/4Cr-I/2Mo steel header in air and obtained

good agreement with comparable data from tests of

standard-size specimens, as is shown in Figures 14.6

and 14.7. Thus, miniature tensile creep testing has been

demonstrated to be a viable tool for residual-life assessment.

Although validation work would be required, it

is expected that the same type of miniature tensile

specimen configuration used for creep testing

would provide valid results for the measurement of

tensile strength and deformation response of samples removed from reactor components. For example, a study by VanEcho and Williamsa showed

that miniature tensile specimens are useful for measuring the through-thickness properties of thin

('.12 mm or less) steel plate. Miniature specimens

(including the disk configuration discussed earlier)

would be particularly useful for characterizing the

properties near welds because samples could be

taken from the weld metal and heat-affected zone

in cases where standard-size specimens would be

too large to sample the properties in the desired

region.

142.2 Reconstituted Specimen Testing. Another

procedure developed to fully use the limited amount of

irradiated material that usually is available for evaluation is to reconstitute a Charpy V-notch impact specimen from one half of a previously tested

specimen.54'

55 Figure 14.8 illustrates the sequence of

steps involved in reconstitution. 54 The fracture face is

cut off one half of the tested specimen, and two end

tabs are stud welded to it. Final machining and notching of the specimen then is completed. Properly reconstituted specimens yield results that agree with those

obtained from the original specimens, as is shown for a

typical set of data in Figure 14.9.54 Thus, the use of

reconstituted specimens can triple the amount of

Charpy data obtained from the same volume of material when only original specimens are used.

a. Unpublished research results, Battelle Columbus, Columbus, Ohio, July 1981.

154

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;

. Ex-service

sample: ' EIB

weld

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* J~-13mm -71 N Extension piece

for grip

;Collarto provide

run-in material for this

weld to prevent undercut

Figure 14.5(a). Schematic diagram of an electron beam welded blank and test piece for tensile testing.

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Am 1 ,/ ~~~~~~~~~~6mm

Parallel gauge 7.3062

Figure 14.5(b). Schematic diagram of a test specimen machined from the assembly, shown in Figure 14.5(a).

_~~~~~ ! - -

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MPa

10 20 40 60 80 100 10-2

10 i- 4

0.

S!

0)

C.,

0)

0)

1 2 4 6 8 10 20

Applied stress, a (Ksi) 7.3198

I

Figure 14.6. Minimum creep rate as a function of stress for hot- and cold-end materials.

156

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N.t

CO

104

7.3199

F r , i b. .

Figure 14.7. "Comparison between creep strain 'versu's te'st time data from standard and miniature specimens.

I ,; ,, i ; , ,- I -. , ,,,, ; , . I - I --: . -. . .1 . . , I , '' I :1 I

157

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Original specimen

As-tested specimen

* Half of broken

specimen Aluminum ball

:' C': )

End tabs, machined

half of broken specimen,

and arc shields

(exploded view)

I.

Fli

As-welded specimen

iished specimen

7-3061

Figure 14.8. Sequence of stages in preparing a reconstituted Charpy specimen. I

14.2.3 On-Line Damage Calculation. On-line

damage and remaining life calculations are being

evaluated as a possible method for residual life

assessment.14 This approach uses a digital computer and thermal, stress, creep, and damage analysis software to compute the damage accumulation

during service in conjunction with input data from

an on-line temperature monitoring system. All of

the equipment is installed at the plant and must be

able to survive and perform satisfactorily in an

industrial environment. If the calculations are performed in full detail, this approach is quite

demanding on computer resources. Thus, in practice, appropriate simplifications must be made to

the analyses. If these analyses are performed properly, then the approach appears to be promising

and can provide useful information about equipment operation. The major shortcoming, at

present, is the lack of validated simple analytical

models of the damage accumulation processes.

It would be possible to apply a similar on-line damage accumulation approach to nuclear systems where

key operating parameters are being monitored on-line.

For example, if local temperatures near a nozzle were

being monitored during operation, then on-line damage analysis could be employed to compute the thermal

stresses and strains and the resulting fatigue damage

accumulation during startups, shutdowns, and major

operating transients. Such a system would require

development and field demonstration and validation

before it could be implemented with confidence.

14.2.4 Effective Engineering Models. Successful implementation of life extension and life assessment strategies requires engineering models that

incorporate reasonable simplifying assumptions,

158

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Temperature (CC)

a,Z

C

w

a)

C

w

Original specimen

Reconstituted specimen

-250 -150 -50 50 - ; 150 ! 250

Temperature (°F)

350 450

7-3055

Figure 14.9. Correlation between Charpy data for original and reconstituted specimens.'

but still predict the major, important features of

material damage evolution (aging). In other words,

the models must provide a balance between sophistication and practical utility so that good assessments can be made economically. Where

simplifying assumptions a-re necessary; they should

be made on the conservative side to ensure a high

probability of safe future operation.

Balkey and Furchi 56 outlined an engineering

approach for the life extension of PWR vessels, as

is shown in Figure 14.10. First, onemakes an initial

assessment of the RPV using the current guidelines

(Regulatory Guide 1.99). If the required remaining

life can be achieved with the. current operating

parameters, then operations can be continued in

the present manner. If not, then fuel management

may be evaluated using traditionral or-new

approaches. Other modifications to reduce the>

neutron flux, such as shielding,-may beconsidered.

The appropriate actions are then taken to obtain

the desired remaining safe life. Thistype of assess-'

ment is carried out in a probabilistic fashion to help.

ensure that safe and reliable decisions are made.

Sidey57 used a model for combined corrosion

and stress-rupture damage of fossil-plant superheater and reheater tubes to develop a simplified

engineering procedure for estimating the remaining

life of reheater tubes. When a reheater tube failed

after 78,000 h of service, he used that result to calibrate the model and develop a simple plot of wall

thickness versus remaining life, as is shown in Figure 14.11. Now, when ultrasonic testing is used to

measure tube-wall thicknesses during an outage,

that relationship is used to estimate remaining life

of other reheater tubes. Using detailed models to

predict remaining life trends based on past operating experience provides a valuable,' straightforward engineering approach.

Probabilistic factors need to be incorporated in

residual-life estimations. Clark58 u'sed probabilistic relationships to illustrate the role of NDE in failure risk assessments, as is shown in Figure 14.12.

The probability of failure increases as the inspection interval and as the probability of crack initiation, Pi, increase. Increased probability of defect

detection,'PF, reduces the probability of failure.

This type of approach provides a rational and realistic method for evaluating and minimizing the risk

of failure.

I

159

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Implementation

phase

7.3073

Figure 14.10. A cost-effective approach to address reactor vessel life extension.

T4 -

0)

C

xY

0)

0.

21

Material - 1.25 Cr 0.5 Mo

Outside diameter - 54 mm

Uncorroded wall thickness - 4.18 mm

I I I I

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II

III

I

A\

0 20 40 60 80 100 I

Useful remaining life (h) x

7.3040

Figure 14.11. Remnaining life graph for TI I reheater tubing after 78,000 hours in service.

160

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o- o1 

1 -2 -

a.-

.0

0

2 -

k10-5 I

10-6

I

10o- 71L.

0

I I I

5 10 15 20

Inspection interval, years

25

7-3039

.. .. Figure 14.12. Probability of failure versus inspection interval for example cases.

14.3 Summary

The major emerging methods for nondestructive

inspection and life assessment of key primary system pressure boundary components are summarized in Table :14.2.. For each method, both its

characteristics and its applications are briefly highlighted. The applications indicate areas where the

method is used and areas where the method has

potential use if further advancements'and developments are achieved.

Advances -in traditional flaw-detection

approaches are being achieved through automation

and computerized processing of data. With state-

* of-the-art systemsand well-trained, knowledgeable

operators, flaws can be located with high probability and sized with moderately good probability.

Clearly, there is a need to develop better methods of

quantifying the size, nature, and distribution of

flaws for use in fracture-mechanics analyses and

'development 'work is underway in this area.

Acoustic emission monitoring of crack growth

and pressure boundary leaks needs to be further

developed for routine field application. On-line

neutron-noise and material damage analysis offer

long-range potential, but a great deal of development work is required before they can be routifiely

used with a high degree of confidence.'

The nondestructive measurement of materials properties shows a great deal of promise for use in residuallife'assessmerits, but a large amount of research and

development is needed before the laboratory techniques can be made commercially viable. Surface replication and miniature specimen testing also offer some

potential uses. For real-world use, effective engineering

models of material-aging processes are essential. These

models must provide a balance between ease of use and

complex representations of physical phenomena.

I

161 '

I
I

Table 14.2. Summary of emerging methods for inspection and life assessment

Method Characteristics Applications

SAFT-UT Computerized processing of

ultrasonic data to provide enhanced

significant visual images of flaws in

real time.

Welds in the RPV, in piping,

and at nozzles. Requires

computer resources resources

resources and has not been fully

developed and verified for field

applications. V

Computer-aided UT pipe

inspection systems

DC potential drop pipe

inspection systems

Computerized eddycurrent inspection

systems

Acoustic emission

monitoring

On-line damage analysis

Automated ultrasonic inspection of

piping by means of state-of-the-art

hardware and software.

Automated inspection and

computerized analysis of DC

potential distributions around cracks

on both inner and outer surfaces of

piping.

Advanced eddy-current coil designs

coupled with real-time computer

analysis of data.

Uses computerized processing

of data to monitor noise from

crack growth or water leakage.

Monitors temperature and pressure

during operation and computes

damage accumulation on-line in

real time.

Piping and IGSCC cracks near

welds in piping. Need more

real-time analysis and

interpretation of data to aid

operators in decision making.

Promising new method for

inspection of pipes. Prototype

system needs to be developed,

demonstrated, and verified for

routine commercial applications.

More reliable inspection of

PWR steam generator tubes. The

advanced methods need to be

fully evaluated and

demonstrated. Deep penetration

methods show some potential for

pipe inspection, but they need

considerable development.

Monitoring of fatigue-crack

growth in RPVs, possibly

stress-corrosion crack growth

in piping, and water leaks in the

primary pressure boundary.

Requires advanced techniques to

remove background noise.

Results of initial field studies are

encouraging, but further

demonstration and validation are

needed.

Creep-fatigue damage in fossil

boiler headers and fatigue

usage factors in nuclear

plant nozzles. Only prototype

systems are currently available.

Much uncertainty with damage

accumulation models that are

used.

fI

162

I1
I

Table 14.2. (continued)

Method Characteristics Applications

Neutron-noise analysis

Nondestructive

measurement of

materials properties

Surface replication

Measures the ex-core neutron noise

and analyzes the recorded data to

diagnose potential problems within

the RPV.

Uses ultrasonic, eddy-current,

x-ray, and electrochemical testing

methods to measure properties, such

as tensile strength, fracture

demonstrated, toughness, and impact

strength.

Produces plastic film replica of

material surface that can be . c

examined at high magnifications

using both optical and electron

microscopy to detect small cracks,

voids, and microstructural changes.

Shows promise for monitoring

vibration of PWR internals

and BWR instrument tubes

and fuel boxes. Need reference

data for normal operations and

'hardware and software for

on-line use.

Most applications have been

limited to laboratory studies.

Offers great potential when

methods are developed,

and verified by means of future

basic and applied research. Need

to develop a library of

well-characterized correlations

between nondestructive

measurements and properties.

Widely used to inspect for creep

damage in fossil-plant headers

and steam piping. Has potential

use for the detection of small

cracks and microstructural

changes caused by aging in key

nuclear components. Could be

used to measure ferrite amount

and spacing in cast stainless

steels. Needs to be developed for

nuclear-plant applications.

Laboratory studies are being

performed to define the range of

use for this technique. Shows

promise for the measurement of

deformation, such as creep and

stress-strain response. May have

applications for measurement of

fracture toughness and crack

growth. Further verification of

the technique is required for field

application.

Trending the aging of key

components from a knowledge , of the operating history. Most

current approaches seem either

too detailed and cumbersome to

use or too simplistic to provide

realistic predictions.

1.

I

Miniature specimen

testing

Uses a small specimen removed from'

a component to measure materials

properties.

. I

Effective engineering

models

Use of simplified engineering models

that provide reasonable descriptions

of material aging as a -,

function of operating history.

I

163
i

14.4 References

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12. P. H. Hutton and R. J. Kurtz, "Current Capabilities for Continuous AE Monitoring to Detect Flaws

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1985 Pressure Vessels and Piping Conference, New Orleans, Louisiana, June 23-26, 1985, PVP-Volume

98-1, pp. 217-222.

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. I'~-
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Mechanical Engineers 1985 Pressure Vessels and Piping Conference, New Orleans, Louisiana, June 23-

26, 1985, PVP-Volume 98-1, pp. 65-71.

15. D. N. Fry et al., "Analysis of Neutron-Density Oscillations Resulting From Core Barrel Motion in a

PWR Nuclear Power Plant," Annals of Nuclear Energy, 2, 1975, pp. 341-351.

16. D. N. Fry et al., Use of Neutron NoiseforDiagnosis of In- VesselAnomalies in Light-WMater Reactors,

NUREG/CR-3303, 1984.

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18. C. 0. Ruud and R. E. Green, Jr., (eds.), Nondestructive Methods for Material Property Determination, New York: Plenum Press, 1984.

19. E. R. Generazio, "Scaling Attenuation Data Characterizes Changes in Material Microstructure,"

Materials Evaluation, 44, 1986, pp. 198-202.

20. G. D. Swanson, Development of Fracture Toughness Measurement Capability by Ultrasound, BDX613-2625, June 1981.

21. A. Vary, "The Feasibility of Ranking Material Fracture Toughness by Ultrasonic Attenuation Measurements," Journal of Testing and Evaluation, 4,14, 1976, pp. 251-256.

22. K. Coebbels and P. Holler, "Quantitative Determination of Grain Size and Detection of Inhomogeneities in Steel by Ultrasonic Backscattering Measurements,"' Proceedings of the First International

Symposium on Ultrasonic Materials Characterization; Gaithersburg, Maryland, June 7-9, 1978, NBS

Special Publication 596.

23. A. Hecht et al., "Nondestructive Determination of Grain Size in Austenitic Sheet by Ultrasonic

Backscattering," Materials Evaluation, 39, 1981, pp. 934-938.

24. M. T. Research et al., "The Acoustic Measurement of Stress Intensity Factors," Applied Physics

Letters, 34, 3, 1979, pp. 182-184.

25. A; V. Clark et al., "Acoustoelastic Measurement of Stress and Stress Intensity Factors Around Crack

Tips," Ultrasonics, 21, 1983, pp. 57-64.

26. J. F. Bussiere, "On-Line Measurement of the Microstructure and Mechanical Properties of Steel,"

Materials Evaluation, 44, 1986, pp. 560-567.

27. 0. K. Chopra and H. M. Chung, "Investigations of the Mechanisms of Thermal Aging of Cast

Stainless Steels," Meeting of the Metal Component Subcommittee of the Advisory Committee on

.Reactor Safeguards, Columbus, Ohio, July 1-2, 1986. -

28. J. J. Yavelak, "Correlation of Diametral Measurements and Cracking with Barkhausen Noise Patterns in a Thick-Wall Croloy 1-1/4 (1-1/4Cr-1/2Mo) Header," Residual-Life Assessment, NondestructiveExamination, and NuclearHeatExchangerMaterials, Proceedings of TheAmerican Society of

Mechanical Engineers 1985 Pressure Vessels and Piping Conference, New Orleans, Louisiana, June 23-

26, 1985, PVP-Volume 98-1, pp. 73-78.

29. T. P. Sherlock et al., "Advanced Techniques for Assessing the Remaining Life of Boiler Components," EPRI Conference on LifeExtension and Assessment of Fossil Plants, Washington, D.C., June

1986.

165

1
30. W. G. Clark and M. J. Metala, Jr., "Eddy Current Determination of Material Properties," EPRI

Conference on Life Extension and Assessment of Fossil Plants, Washington, D.C., June 1986.

31. A. P. Majidi and M. A. Streicher, "Potentiodynamic Reactivation Method for Detecting Sensitization in AISI 304 and 304L Stainless Steels," Corrosion, 40, 8, 1984, pp. 393-408.

32. A. P. Majidi and M. A. Streicher, "The Double Loop Reactivation Method for Detecting Sensitization in AISI 304 Stainless Steels," Corrosion, 40, 11, 1984, pp. 584-593.

33. T. Shoji and H. Takahashi, "Non-Destructive Evaluation of Materials Degradation During Service

Operation by Means of Electro-Chemical Method," EPRI Conference on Life Extension and Assessment of Fossil Plants, Washington, D.C., June 1986.

34. G. E. Lucas et al., "Effects of Composition, Microstructure, and Temperature on Irradiation Hardening of Pressure Vessel Steels," Effects of Irradiation on Materials: Twelfth International Symposium, ASTM Special Technical Publication 870, Philadelphia, 1985, pp. 900-930.

35. B. J. Cane et al., "A Mechanistic Approach to Remanant Creep Life Assessment of Low Alloy

Ferritic Components Based on Hardness Measurements," Journal of Pressure Vessel Technology, 107,

1985, pp. 295-300.

36. B. Neubauer and U. Wedel, "Restlife Estimation of Creeping Components by Means of Replicas,"

Advances in Life Prediction Methods, New York: The American Society of Mechanical Engineers,

1983, pp. 307-313.

37. P. Auerkari, "Remanant Creep Life Estimation of Old Power Plant Steam Piping Systems,"

Advances in Life Prediction Methods, New York: The American Society of Mechanical Engineers,

1983, pp. 353-356.

38. R. Viswanathan, "Dissimilar Metal Weld and Boiler Creep Damage Evaluation for Plant Life Extension," Residual-Life Assessment, Nondestructive Examination, and Nuclear Heat Exchanger Materials, Proceedings of The American Society of Mechanical Engineers 1985 Pressure Vessels and Piping

Conference, New Orleans, Louisiana, June 23-26, 1985, PVP-Volume 98-1, pp. 9-18.

39. G. H. Harth and T. P. Sherlock, "Monitoring the Service Induced Damage in Utility Boiler Pressure

Vessels and Piping Systems," Residual-Life Assessment, Nondestructive Examination, and Nuclear

Heat Exchanger Materials, Proceedings of The American Society of Mechanical Engineers 1985 Pressure Vessels and Piping Conference, New Orleans, Louisiana, June 23-26, 1985, PVP-Volume 98- 1, pp.

19-23.

40. J. F. Henry and F. V. Ellis, Plastic Replication ProceduresforDamageAssessment, progress report on

EPRI Project RP2253-1, September 1983.

41. S. 0. Hilton and M. F. Steakley, "Steam Line Inspection Program for Tennessee Valley Authority's

High Temperature Steam Piping Systems," EPRI Conference on Life Extension and Assessment of

Fossil Plants, Washington, D.C., June 1986.

42. J. A. Janiszewski et al., "A Production Approach to On-Site Metallographic Analysis," EPRI Conference on Life Extension and Assessment of Fossil Plants, Washington, D.C., June 1986.

43. T. Takigawa et al., "Life Assessment and Extension of Steam Turbine Rotor-Overview of MHI's

Activities, " EPRI Conference on Life Extension and Assessment of Fossil Plants, Washington, D. C.,

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166
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44. T. R. Padden and C. A. Daniels, "Replication Manipulator and Procedure for Tube Bore Surfaces,"

Annual Meeting of Scanning Electron Microscopy Inc., Dallas, Texas, April 1981.

45. T. R: Padden and C. A. Daniels, In Situ Examination'of the Small Scale Steam Generator Model,

WARD-1i10134-1, Westinghouse Electric Corporation, October 1976.

46. T. R. Padden, Metallurgical Characterization of the Creep Ratcheting Test Article Failure, WAESDHT-94000-28, Westinghouse Electric Corporation, June 1983.

47. C. E. Jaske et al., "Life-Assessment of High-Temperature, Fossil-Plant Components,' EPRI Conference on Life Extension and Assessment of Fossil Plants, Washington, D. C., June 1986.

48. M. P. Manahan et al., "The Development of a -Miniaturized Disc Bend Test for Determination of

Post Irradiation Mechanical Properties," Journal of NuclearMaterials, 103-104, 1981, pp. 1545-1550.

49. F. M. Haggag et al., "The Use of Miniaturized Tests to Predict Flow Properties and Estimate Fracture Toughness in Deformed Steel Plates," International Conference and Exhibition on Fatigue, Corrosion Cracking, Fracture Mechanics, and Failure Analysis, Salt Lake City, Utah, December'1985.

50. T. Misawa and Y. Hamaguchi, "Miniature Specimen Tests for Evaluating Fracture Resistance With

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(ed.), Metals Park, Ohio: American Society for Metals, 1986, pp. 521-527.

51. Battelle Columbus Division, Nuclear Tech Notes, 3, Spring 1984. -

52. M. C. Askins et al., "Estimating the Remanant Life of Boiler Pressure Parts," Meeting of the EPRI

Advisory Group on Remaining Life of Boiler Pressure Parts, Palo Alto, California, June 4, 1985.

53. A. Saxena et al., "Evaluation of Remaining Life of High Temperature Headers: A Case History,"

EPRI Conference on Life Extension and Assessment of Fossil Plants, Washingion, D.C., June 1986.

54. -J. S.-Perrin et al., "Preparation of Reconstituted Charpy V-Notch Impact Specimens for Generatingl

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582-593.

55. R. P. Shogan, S. E. Yanichko, and W. S. Galloway, "The Use of Reconstited Charpy Specimens to

.Extend R. E. Ginna Reactor Pressure Surveillance Data," Nuclear Technology 72, 3, March 1986.

56. - K. R. Balkey and E. L. Furchi, "A Cost Effective Approach for Extending the Life of Pressurized

water Reactor Vessels," Residual-Life Assessment, Nondestructive Examination, and Nucleai Heat

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176. .-

57. D. Sidey, "Utility Plant Life Assessment as a Continuous Procedure," EPRI Conference on Life

Extension and Assessment of Fossil Plants, Washington, D.C., June 1986.

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- :Prediction Methods, New York: The American Society of Mechanical Engineers, 1983, pp. 293-299.

*, ,6* ' ,- ...

167
I

15. SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS

The first step in the residual life assessment task was

to identify the major light water reactor (LWR) components that are critical to nuclear plant safety during

normal operation, off-normal situations, design-basis

accidents, or severe accidents. This report has identified and prioritized these components according to

their role in preventing the release of fission products to

the public. Table 15.1 presents the list of selected pressurized water reactor (PWR) components for the residual life assessment task and reasons for their ranking.

This table also identifies the most likely and other

potential degradation sites and aging mechanisms for

each component. Similarly, Tible 15.2 presents the list

of selected boiling water reactor (BWR) components.

The next step in the residual life assessment task was

to perform a qualitative analysis of the degradation

processes active in each of these components and carefully identify corresponding degradation sites, stressors, degradation mechanisms, and potential failure

modes during normal operation and accident conditions. This report presented the results of the analyses

for seven major components: four PWR components

(vessel, containment, primarycoolant piping, and

steam generator), two BWR Components (vessel and

recirculation piping), and supports for both PWR and

BWR vessels. The currently employed in-service

inspection (ISI), surveillance, monitoring, and lifeassessment methods were also evaluated. Similar studies will be completed in, 1987. of the, degradation

processes associated with the remaining major components: four PWR components (reactor coolant pump

bodies, pressurizers, control rod drive mechanisms,

and reactor internals), four BWR components (containments, recirculation pump bodies, control rod

drive mechanisms, and reactor internals), cables and

connectors, and emergency diesel generators in PWR

and BWR plants.

This review of potential or, in many cases, actual

degradation sites, mechanisms, and potential failure modes associated with the major LWR components and structures has raised a number of

questions or issues deserving further attention. A

summary of the results, presented in this report,

conclusions based on these results, and the recommendations for further work are grouped by component and listed below. (these recommendations

are indicative of general needs and not necessarily

tasks that might be accomplished at the Idaho

-National Engineering Laboratory as part of the

Nuclear Plant Aging Research program). Finally, a

summary of the results, corresponding conclusions, and recommendations related to the current

and emerging nondestructive examination (NDE)

and life-assessment methods is presented.

15.1 PWR Reactor Pressure

Vessels

Table 15.3 summarizes the important degradation

sites, stressors, degradation mechanisms, potential failure modes, and current ISI and surveillance methods

associated with PWR pressure vessels. Note that the

ranking of sites is based on the safety impact of that

site and not on the possibility of that site having severe

degradation. The major conclusions and recommendations are as follows:

1. Radiation embrittlement is the primary

safety concern that: may limit vessel life

extension. Therefore, continued close

monitoring of the pressure vessel material

condition is required. The current correlations used to estimate the shifts in RTNDT

and upper-shelf-energy changes, are

uncertain because of the limited accuracy

and range of the surveillance data base.

The data base does not include the effects

of high-fluence and low-flux conditions,

and treats the effects of spectrum and irradiation temperature in an approximate

manner. A comprehensive plan to obtain

reactor pressure vessel (RPV) materials

data from environments (i.e. irradiation,

thermal, and chemical) that reasonably

represent the realities of long-term inservice behavior shouid be developed.

2. Use of the temperature shift at 30 ft-lb

(41 J) level, ARTNDT, from the Charpy Vnotch curve for certain RPV materials may

underpredict the actual temperature shift

at the measured transition-temperature

fracture toughness of 100 ksi %/W Therefore, the validity of using Charpy test

results for estimating temperature shifts to

determine transition-temperature fracture

toughness needs further assessment. However, the conservatisms in the overall reference toughness curves and in the shifting

of these curves may outweigh this difference between Charpy V-notch and fracture

toughness tests.

.

168
I

Table 15.1. Key PWR components for residual life assessment

Rank Component

I Reactor

pressure

vessel

Reasons for Ranking

Severe safety impact

of catastrophic failure

of an embrittled

vessel

Degradation Sites

(most likely, others)

.Weldments in the belt

line region, hot-leg and

cold-leg nozzles, weldment

at the lower shell to

bottom head junction,

vessel flanges and studs,

instrument penetrations in

lower head

Degradation Mechanisms

(most likely, others)

Neutron embrittlement,

corrosion fatigue, thermal

-fatigue

2 Containment Public protection

and basemat during an accident

Posttensioning system

(tendons and anchors),

concrete, metal liner,

reinforcing steel

Hydrogen embrittlement and

stress corrosion cracking of

anchor heads, corrosion and

relaxation of tendon material,

environmental degradation of

concrete, corrosion of

reinforcing steel and liner

3 Reactor

coolant

piping and

safe ends

4 Steam

generator

Severe safety impact

of a large break

Tube rupture will

provide a passage from

the primary system

directly to the

environment for the

primary coolant, -

relatively poor

operating experience

Weldments at the safe

ends, branch nozzles, cast

stainless steel components

(elbows and

t-connections), complete

* primary loop piping in

new Westinghouse plants

Inside tube surfaces

at U-bends and tube

sheet, outside surfaces

at tube-to-tube sheet

crevices, divider plate,

tube support plate,

feedwater nozzle,

girth weld

Fatigue, thermal embrittlement

I.

Wastage, denting,

intergranular stress

corrosion cracking, pitting,

fretting, intergranular attack,

thermal fatigue, corrosion

fatigue

I

5 Reactor

coolant

pump casing

6 Pressurizer

7 Control rod

drive

mechanism

Primary protection

from any internal

failure of pump

elements (impeller

parts, shafts), primary

coolant pressure

boundary

Primary coolant

pressure retaining

boundary

Failure may lead to a

reactivity initiated

accident, an

anticipated transient

without scram, or a

loss-of-coolant

accident

Casing wall, flanges,

seal housing bolts

Surge and spray nozzles,

spray head, upper shell

barrel, seismic lugs

Drive rod assembly,

control rod pressure

vessel

Corrosion fatigue, thermal

embrittlement

Thermal fatigue, erosion,

thermal embrittlement

Wear and thermal

embrittlement

!

I

169

-I
I

Table 15.1. (continued)

Rank Component

8 Safetyrelated

cables and

connectors

9 Emergency

diesel

generator

10 Reactor

internals

II RPV

supports

Reasons for Ranking

Active during normal

operation in mitigating

operational transients

and accidents

Needed to operate

critical safety

equipment in the event

of a loss-of-offsite

power

Failure may lead to

disbursement of fuel

into the coolant, or a

reactivity-initiated

accident

Failure will challenge

the integrity of the

primary coolant

piping and the small

lines connected to the

RPV; the neutron

shield tanks maintain

acceptable levels of

radioactivity in the

containment

Degradation Sites

(most likely, others)

Cable insulation, inserts

in connectors

Governor in the

instrumentation and

control system, cooling

system pumps and piping,

fuel injector pumps,

turbocharger

Lower core plate,

baffle-former assembly,

upper support column

bolts, control rod bolts,

control rod guide tube

sheath and support pins,

thermal shield bolts,

in-core instrument

nozzles, flux thimble

tubes bolts, in-core

instrument nozzles, flux

thimble tubes

Inside surface of the

support structure (neutron

tank or column support)

at the core horizontal

midplane level, lubricant

in sliding foot assembly

of neutron tank support

Degradation Mechanisms

(most likely, others)

Thermal aging and creep of

insulation, thermal

embrittlement and

corrosion of connectors

Fatigue and vibrations

Neutron embrittlement, wear,

high-cycle fatigue, SCC,

relaxation

Neutron embrittlement,

corrosion, degradation of

lubricant

I

i

I

170

1
Table 15.2. Key BWR components for residual life assessment

Degradation Sites

(most likely, others)

Degradation Mechanisms

Rank Component Reasons for Ranking (most likely, others)

I Containment Public protection during an

and basemat accident

Concrete, metal liner, and

reinforcing steel in

suppression pool and

primary containment wall,

'base metal, welds and vent

pipes in metal containment

Nozzles, closure studs,

beltline region, welds at

the stub tubes

Safe ends, austenitic

stainless steel fittings

Environmental degradation

of concrete, fatigue and

corrosion of metal

containments

2 Reactor

pressure

vessel

3 Recirculation

piping, safe

ends, and

safety system

piping

Primary pressure boundary, poor

operating experience with reactor

pressure vessel nozzles

Primary pressure boundary,

relatively poor operating

experience

Thermal fatigue, neutron

embrittlement, intergranular

stress corrosion cracking

-Intergranular stress

corrosion cracking,

thermal aging

4 Recirculation Primary protection from any

pump body internal failure of pump elements,

primary coolant pressure boundary

Heat affected zones near weldments in the wall

Thermal aging, corrosion

fatigue, crevice corrosion

5 Control rod

drive

mechanism

6 Safety-related

cables and

connectors

7 Emergency

diesel

generator

8 Reactor

internals

9 - Reactor

pedestal

10 Biological

shield

Failure may lead to a reactivity

initiated accident, an anticipated

transient without scram (ATWS), or a small-break, loss-of-coolant

accident

Active during normal operation

in mitigating operational transients -and accidents

Needed to operate critical

safety equipment in the event

of a loss-of-offsite power

Failure may cause fuel

failure or problems in

scraming the reactor

Failure will challenge integrity

of reactor pressure vessel

Maintains acceptable level of

radioactivity in containment

Drive rod assembly,

control rod pressure

vessel

Cable insulation, inserts

in connectors

Governor in the

instrumentation and

control system, cooling and

lubrication system pumps

and piping, fuel injector

pumps, turbocharger,

generator windings

Wear and thermal

embrittlement

Thermal aging of insulation,

thermal embrittlement and

corrosion of connectors

Fatigue and vibrations

Irradiation-assisted

stress corrosion

cracking, intergranular

stress corrosion

cracking, fatigue

Core shroud, top guide

plate, core plate,

holddown beam in a

jet pump, feedwater

and core spargers .

Concr

steel

Inside

biolog

core hi

level

ete and reinforcing ---Thermal cycling of concrete,

corrosion and fatigue of

reinforcing steel

surface of the Neutron irradiation,

ical shield at the gamma-ray heating;.

orizontal midplane 'environmental

; degradation of concrete,

corrosion of reinforcing steels

i, -

! ; 'i-

: . : I

171 . ;
Table 15.3. Summary of degradation processes for PWR reactor pressure vessels

Rank of

Degradation

Site

Degradation

Site Stressors Degradation Mechanisms Potential Failure Modes ISI Surveillance Methods

I Beitline region Neutron

irradiation,

mechanical

and thermal

stresses

Irradiation embrittlement

(degree is dependent on

individual vessel

materials and flux

spectrum history)

Ductile high-energy

tearing leading to

leakage (net section

over-load)

Brittle fracture (i.e.,

pressurized thermal

shock)

100% volumetric during

first inspection; one

weld for subsequent

inspection

Surveillance program for

assessing irradiation

damage is required by law V

Ductile low-energy

tearing (low uppershelf toughness)

Environmental

fatigue

Ductile overload leading

to a leak, possible brittle

fracture if PTS occurs

2 Outlet/inlet Mechanical

nozzlesa and thermal

stresses

3 Instrumentation

nozzles

(penetrations)

and control rod

drive mechanism

housing nozzles

Mechanical

and thermal

stresses

Fatigue crack initiation

and propagation

Fatigue crack initiation

and propagation

Fatigue crack initiation

and propagation (possibly

corrosion-assisted)

Ductile overload leading

to a leak; possible

brittle fracture if

pressurized thermal

shock occurs with

some irradiation

embrittlement

Ductile overload leading

to a leak

Ductile overload failure

(can be replaced)

All nozzle welds inspected

volumetrically at each

interval

Visual, inspection of

external surface; 253% of

nozzles inspected at first

interval; remaining 750e

spread out over next three

intervals

Volumetric and surface

inspection of all studs

and threads in flange stud

holes at each interval

4 Flange closure Mechanical

studs and thermal

stresses

a. Welds at the vessel to safe end are covered in Chapter 5 on primary system piping; the comments here are also applicable except

that volumetric and surface examinations are required.

II

3. Current RPV materials surveillance programs have limited space and limited numbers of specimens of critical material.

Therefore, miniature and reconstituted

specimen testing should be performed, so

that additional mechanical property data

can be extracted from the given limited

amount of critical material. Also, the current surveillance programs should be modified as necessary to ensure that specimens

are available for surveillance testing during

the renewed license period.

4. Thermal fatigue of the RPV nozzles and penetrations is an important safety concern associated with license renewal. Techniques for

better monitoring of fatigue transients for

nozzles and penetrations should be developed

so that the actual usage factors may be determined. To improve the life-prediction capability of the American Society of Mechanical

Engineers (ASME) Section III and XI

fatigue design analyses, a modified procedure

for crack initiation should be developed,

based on the local-strain approach that

I

172

I
I

accounts for the effect of both stress and

strain concentrations at notches and of the

sequence of variable amplitude cycles. In

addition, an updated version of the damagetolerant procedures that incorporate an

improved appreciation of environmental

factors, low-growth-rate behavior, and the

role of small flaws should be developed.

5. Inspection and transient logging records as

well as all material properties (unirradiated and surveillance results), and design

data need to be maintained. Collection of

these records is needed to develop lifeextension plans.

6. Thermal annealing of reactor pressure ves-_

sels should be pursued as a long-term

option for license renewal. Annealing

mechanisms and kinetics need to be better

understood and corresponding models,

need to be developed.

15.2 PWR Containments and

Basemats

Tables 15.4, 15.5, and 15.6 (for prestressed concrete, reinforced concrete, and steel PWR containments and basemats, respectively) summarize the

important degradation sites, stressors, degradation

mechanisms, potential failure modes, and the current

ISI, surveillance, and testing methods. The major conclusions and recommendations are as follows:

1. Hydrogen embrittlement, pitting, and

microbiological-inrduced corrosion are the

major degradation mechanisms associated

with the posttensioning system anchors,

tendon wires, and grease in the prestressed

concrete containments. Monitoring methods that will detect the degradation of the

anchors and decomposition of the tendon

grease are needed.

2. Aggressive environments and internal

chemical reactions are the major stressors

damaging the concrete in concrete containments. Corrosion is the major degradation

mechanism associated with the reinforcing

bars and steel liners in concrete containments. Additional information about the

long-term degradation of reinforced and

prestressed concrete containments is

needed. Data should be collected from the

older LWR containments and facilities

that have been shut down after extended

service. Accelerated aging techniques

could also be investigated and, if found

appropriate, used to obtain additional

data. The results from these studies should

support the development of quantitative

residual life models.

3. A comprehensive and standardized ISI

program can and should be developed to

identify and quantify degradation in reinforced and 'prestressed concrete containments. State-of-the-art ISI methods used

- - in Europe,-Japan, and the United'States

need to be evaluated and included in the

program where appropriate.

4. The impact of the concrete-liner interaction

on the integrity, of the aged reinforcedconcrete and prestressed-concrete containments during severe accidents should be

evaluated. The relevant research programs

sponsored by the Electric Power Research

Institute are expected to provide useful information.

15.3 PWR Coolant Piping and

Safe Ends

Table 15.7 summarizes important degradation sites,

stressors, degradation mechanisms, potential failure

modes, and current ISI methods associated with the

PWR coolant piping and safe ends. The major conclusions and recommendations are as follows:

1. Low-cycle fatigue is the main degradation

mechanism for the main coolant pipe nozzles

and the dissimilar metal welds at terminal

ends. However, the number, type, and severity

of transients are generally not recorded.

Therefore, an in-depth study of past operations in selected nuclear power plants would

provide some insight into the rate at which the

design life is being consumed because of

fatigue.

2. Cast austenitic-ferritic (duplex) stainless

steel components, i.e., main coolant pipe,

elbows, T-connections, and valve bodies,

are subjected to thermal embrittlement.

The actual degree of thermal-aging

embrittlement of these cast stainless components should be monitored.

173 -

I'
Table 15.4. Summary of degradation processes for prestressed concrete containment

vessel

Degradation

Rank Site

I Posttensioning

system anchorage

2 Posttensioning

tendon wire or

strand

Degradation

Mechanisms

Potential

Stressor Failure Modes

Material properties

and trapped water

Moisture, trapped

water, or breakdown

of grease material

ISI Method

Hydrogen

embrittlement

Loss of stress Tendon surveillance

program

Pitting.

microbiologicalinduced corrosion

Loss of stress Tendon surveillance

program

3 Steel liner dome

and wall

Moisture, acidic

environment, and

stress

Corrosion and

cracking

Liner-concrete

interaction,

leakage of

radioactive gases

Leakage of

radioactive

material

Leakage testing

(10 CFR 50,

Appendix J)

4 Steel liner over

base slab

Moisture, acidic

environment, and

stress

Corrosion Leakage testing

(10 CFR 50,

Appendix J)

5 Dome, wall, and

base slab

reinforcing

steel

Aggressive

environment

Corrosion Loss of structural

integrity

Visual

6 Concrete Aggressive

environment, and

internal chemical

reactions

Cracking and

spalling

Loss of integrity,

corrosion of

reinforcing steel

Visual, rebound

methods, core

samples if required

174

I'
I

Table 15.5. Summary of degradation processes for reinforced concrete containment

vessel

Degradation

Rank Site

Degradation

Mechanisms

Potential

Stressor Failure Modes ISI Method

I Dome and wall

reinforcing

steel

2 Base slab

' reinforcing

steel

Aggressive

environment

Aggressive

environment

Corrosion

Corrosion

*Loss of

structural

integrity

Loss of

structural

integrity

Visual

Visual

3 Steel liner

over dome and

wall

Moisture, acidicenvironment, and

stress

Corrosion Liner-concrete

interaction,

leakage of

radioactive

gases

Leakage testing

(10 CFR 50,

Appendix J)

4 Steel liner

over base slab

5 Dome and wall

concrete

6 Base slab

concrete

Moisture, acidic

environment, and

stress

Aggressive

environment,

internal chemical

reaction

Aggressive

environment,

internal chemical

reactions

Corrosion

Cracks and

spalling

Cracks and

spalling

Leakage of

radioactive

material

Loss of

integrity,

corrosion of

reinforcing

steel

Loss of

integrity,

corrosion of

reinforcing

steel

Leakage testing

(10 CFR 50,

Appendix J)

Visual, rebound

methods, core

samples if

required

Visual, rebound

methods, core

samples if

required

175

I
Table 15.6. Summary of degradation processes for steel cylinderlsteel sphere

containment vessel

Rank

Degradation

Site

I Shell welds and

base metal

2 Interface

between shell

and concrete

slab at base

of shell

Stressor

Stresses, vibration,

cyclic loading,

aggressive

environment

Stresses, vibration,

cyclic loading,

aggressive

environment

Degradation

Mechanisms

Corrosion

Corrosion

Potential

Failure Modes

Loss of structural

integrity, leakage

of radioactive

gases

Loss of structural

integrity, leakage

of radioactive

gases

ISI Method

Visual, leakage

testing

Visual, leakage

testing

3 Discontinuities

in the shell

such as hatches

and penetrations

4 Steel bottom

of shell

embedded in

concrete

5 Base slab concrete

Stresses, vibration,

cyclic loading,

aggressive

environment

Corrosion Leakage of

radioactive

gases

Leakage testing

(IO CFR 50,

Appendix J)

Aggressive

environment

Corrosion Leakage of

radioactive

material

Leakage testing

(10 CFR 50,

Appendix J)

Aggressive

environment,

internal chemical

reactions

Cracks and

spalling

Corrosion of

reinforcing

steel, corrosion

of steel bottom

of containment

shell

Visual, rebound

methods, core

samples if

required

Table 15.7. Summary of degradation processes for PWR primary system piping

Rank of

Dezradation Site

Degradation

Site

I Main coolant

pipe nozzlesa

2 Terminal end

dissimilar

metal weldb

Stressors

System

operating

transients

System

operating

transients

System

operating

transients

Degradation

Mechanisms

Fatigue crack

initiation and

propagation

Fatigue crack

initiation and

propagation

Thermal

embrittlement

leakage

Potential

Failure Modes

Through wall

leakage

Through wall

leakage

Through wall

leakage

ISI Methods

Volumetric

inspection

for diameter a 4-in.

Surface

inspection

for diameter

<4-in.

3 Cast stainless Volumetric

steel

a. The nozzles in the primary coolant piping are the highest ranking degradation sites. Additional specificity is avoided because the

most severely loaded nozzles are difficult to determine. Reported results are heavily dependent on type of analysis performed. Actual

damage will completely depend on the actual transient usage that occurs in the plant. This is the basis for the recommendation.

b. In Westinghouse plants, the dissimilar metal welds are at the reactor vessel and steam generator nozzles.

176
I

15.4 Steam Generators:

Table 15.8 summarizes the important degradation sites, stressors, mechanisms, potential failure

modes, and current ISI methods associated with

the recirculating and once-through steam genera-,

. . tors. The major conclusions and recommendations

' are as follows: .. . .

1. Intergranular stress corrosion cracking

* (IGSCC) and intergranular attack (IGA) are

- the major degradation mechanisms for the

Table 15.8. Summary of degradation processes for steam generator tubes

Degradation Degradation u

Ranka Site Stressors i Mechanisms Failure Modes ISI Method

I Inside surface of

U-bends and rolltransition regions

Tube rolling stresses,

Corion phenomenon,

denting

IGSCC Cracking Eddy-current testing

2 Outside surface of*

hot-leg tubes in

the tube-to-tube

*sheet crevice region

3 Cold-leg side in

sludge pile or where

scale containing

.copper deposits is

found

Alkaline environment, IGA, IGSCC

presence of S0 4 and

-CO3 anions

Brackish water, air, Pitting

and copper

May eventually result in

cracking ,

Local attack and tube

thinning may eventually

lead to a hole

Eddy-current testing

Eddy-current testing,

optical scanner

system, sonic leak

detector system

4 Outside surface of

tubing above tube

sheet l

Phosphate chemistry,

chloride concentration,

resin leakage from

condensate polisher bed

Wastage (thinning) Uniformn attack, tube

thinning may eventually

wear out the material

Eddy-current testing

5 Tubes in the tubesupport regionsOxygen, copper oxide, Denting

chloride, temperature,

pH, crevice conditions

Flow blockage in tubes

caused by plastic

deformation

Helium leak and

sonic leak testing,

optical probes,

hydrogen evaluation

monitoring, pulse -

echo ultrasound

method

6 Contact between tube

and anti-vibration

bar

Flow-induced

vibrations

Fretting Wear out of material

caused by rubbing and/or

; fatigue -

'Eddy-current testing

7b Once-through steam

generator tubes -

Velocities, sizes, __ '. Erosion-corrosion'

shapes, impact angle, :',frorn impingement

and hardness of . - of particles

particles "

Wear out of material Eddy-current testing

8b Once-through steam - Chemicals, (localized -Fatigue

generator tubes in ,,corrosion) vibrations

the upper iube sheet .

region

a. Based on operating experience for steam generator defects.

Primary to secondary

leaks

Eddy-current testing

_ . .' . . .

b. Denotes once-through steam generator (Items 7 and 8 do not reflect rank order). The first six items are for recirculating steam

generators.

177

I'
recirculating steam generator tubes. Pitting is

another important degradation mechanism

which progresses at an accelerated rate once

initiated, and failures often occur suddenly.

Therefore, methods to monitor the degradation of steam generator tubes caused by

IGSCC, IGA, and pitting are needed. Also, a

better understanding and better models of

these corrosion mechanisms are needed.

2. Several changes in the nuclear power plant

water chemistry, and the design and materials of construction of the recirculating

steam generators (SG) and/or other plant

components have been effective in mitigating the degradation damage. The water

chemistry changes include use of an allvolatile treatment for the secondary-side

water chemistry rather than a phosphate

treatment, addition of boric acid to the

secondary coolant, use of full-flow condensate polishers, and use of tube sheet

crevice flushing. The changes related to the

plant design and materials include a

change from copper to titanium or stainless steel as condenser tube material, use of

stainless steel for the feedwater lines, use

of a quatrefoil design and Type 405 stainless steel for the SG tube support plates,

SG tube rolling to the full depth of the tube

sheet, shot peening of the tubing inside

surfaces in the U-bend and roll-transition

regions of the SG tubes, and use of thermally treated Inconel 690 in place of

Inconel 600 for the SG tubes. The gap size

between the antivibration bars and the SG

tubes has also been reduced in some

designs to avoid fretting. The effects of

these mitigating techniques on the failure

rates of the SG tubes and tube support

plates should be monitored.

3. Erosion-corrosion and environmental fatigue

are the major degradation mechanisms associated with the once-through steam generators. The damage caused by environmental

fatigue has been reduced by blockage of the

untubed lane with a lane blocker.

4. Bobbin-type eddy-current coils have not been

successful in all cases in detecting IGA

defects. However, successful detection has

been achieved by the use of pancake coils that

are arrayed with their axes parallel to the tube

radius and multiplexed to permit use of available electronic packages.

15.5 Reactor Pressure Vessel

Supports

Table 15.9 summarizes the important degradation sites, stressors, degradation mechanisms,

potential failure modes, and current ISI methods

associated with the five different types of RPV supports: neutron shield tank, column, cantilever,

bracket, and skirt-type supports. The first four

types of supports are used with PWR RPVs, while

the fifth type of support is used on BWR RPVs.

The major conclusions and recommendations are

as follows:

1. The neutron shield tanks, columns, and

cantilever-type supports may be susceptible to catastrophic brittle failure caused by

irradiation embrittlement. The neutron

spectra have a very large component of

low-energy neutrons (E < 1 MeV), and the

support temperatures are less than 4500F

(2320C). Both the fracture toughness and

strength of the RPV support steels at high

fluence and low irradiation temperatures

need to be determined. Also, a correlation

of fracture toughness with Charpy V-notch

properties at temperatures less than 4500F

(2320C), and as a function of displacements per atom is needed.

2. A better determination of the radiation environment (neutron spectra and flux levels) is

required for the neutron shield tank, cantilever, and column-type support structures.

3. The dry lubricant in the sliding foot assembly of the neutron tank support degrades

because of neutron irradiation. The effects

of the radiation levels on the lubricants

that are used in the RPV nozzle supports

should be investigated.

15.6 BWR Reactor Pressure

Vessels

Table 15.10 summarizes the important degradation

sites, stressors, degradation mechanisms, potential failure modes, and current ISI and surveillance methods

for the BWR pressure vessel. The major conclusions

and recommendations are as follows:

1. Low-cycle fatigue of the RPV nozzles is

the major degradation mechanism of concern. Refined cycle counting and transient

178

I,
I

Table 15.9. Summary of degradation piocesses for PWR a'nd BWR RPV supports

Rank Degradation Sitea

Degradation Potential

Stressor - Mechanism Failure Modes ISI Methodb

I Neutron shield tank at

the core horizontal

midplane elevation

2 Column support at

the core horizontal

midplane elevation

3 Cantilever support in

the active height of

the core

4 Threaded parts in

sliding foot assembly

Neutron irradiation,

tensile stresses, operati

temperature, water

chemistry

Neutron irradiation,

tensile stresses, operati

temperature

Neutron irradiation,

tensile stresses, operati

temperature

Tensile stresses,

operating temperature

ng

Neutron embrittlement, Catastrophic

corrosion ' : brittle failure

I ' Monitoringc

Monitoring and sampling

Monitoring and sampling

Neutron embrittlement

ng

Neutron embrittlement

ng

Stress corrosion

cracking

Degradation caused by

; neutron irradiation

Catastrophic

brittle failure

Catastrophic

brittle failure

5 Dry lubricant in Neutron irradiation,

sliding foot assembly operating temperatures

Binding that causes

possibly excessive

stresses in the primary

system during heatup

and cooldown

Binding that causes

possibly excessive

stresses in the primary

system during heatup

and cooldown

Ductile overload

failure

6 Skirt support Mechanical and thermal Fatigue

stresses

a. The bracket and skirt RPV supports will experience no neutron embrittlement. x ,

b. There are no national standard ISI methods per se to determinie state of degradation.

c. Monitoring of neutron field near RPV support.

severity estimates should be used to monitor nozzle fatigue usage so that actual

fatigue usage factors can be determined.

2. Closure studs, 'flange bushings,' and stud

holes in flanges without bushing are subjected to wear, fretting, and corrosion.

These components should be examined

and replaced 'or repaired as appropriate

near the 40-year end of life.

3. Brittle failire of a BWR vessel caused by

irradiation embrittlement is extremely

unlikely because of the relatively low neutron

flux and fluence in the beltline region. How-

- ever, the effects of irradiation embrittlement

of BWR vessels should be monitored. Current RPV materials surveillance programs

should be modified for license renewal consideration.

4. Many of the older BWRs have limited

accessibility for external ISI of the welds in

the beltline region. ISI methods for inspection of these welds from the vessel inside

surfacerneed to be developed.'

15.7 BWR Recirculation Piping

Table 15.11 summarizes the important degradation sites, stressors, mechanisms, potential failure

modes, and current ISI methods for the BWR'recirculation piping. The'major conclusions and recommendations are as follows:

1. Aging degradation of the recirculation

piping caused by IGSCC has been a significant problem at a number of BNVRs. The

179

I'
Table 15.10. Summary of degradation processes for BWR reactor pressure vessel

Rank of Degradation Degradation Potential ISI/Surveillance

Degradation Site Site Stressors Mechanisms Failure Modes Methods

I Nozzles (including Mechanical Fatigue crack initiation Ductile overload All large nozzle welds

instrument and CRD and thermal and propagation, IGSCC leading to a leak inspected volumetrically

penetrations) plus stresses at each interval;

safe end welds visual, external surface

inspections of small

nozzles/penetrations

2 Closure studs,

flange bushings,

stud holes

3 Beltline region

Mechanical Fatigue crack initiation Ductile overload Volumetric and surface

and thermal and propagation, failure (can be inspections of all

stresses fretting, corrosion replaced) studs, threads in flange

stud holes and bushings

Irradiation Neutron irradiation Ductile high- 100%6 volumetric

embrittlement (extent depends on vessel energy tearing inspection; surveillance

materials) leading to a leak as required by federal

(not a serious law

problem)

4 External attachments Mechanical Fatigue

and thermal

stresses

Ductile overload Volumetric and surface

failure inspections

Table 15.11. Summary of degradation processes for- BWR recirculation piping

Degradation Degradation

Rank Site Stressors Mechanisms Failure Modes SI Method

I Weld heat-affected Tensile stress, oxygen

zones furnace environment, sensitized

sensitized safe ends heat-affected zones

IGSCC Cracks, leaks Ultrasonic examination,

moisture-sensitive tape

2 L High thermal stress . Cyclic tensile stress,

regions predicted by corrosive environment

stress rule index

analysis

Thermal fatigue,

corrosion fatigue

Cracks, leaks Ultrasonic examination,

moisture-sensitive tape

3 Austenitic-ferritic

stainless steel

castings with high

delta ferrite levels

High-temperature,

tensile stress, shock

loads

Thermal embrittlement Cracks, leaks Ultrasonic examination,

moisture-sensitive tape,

impact test specimens

4 Shaft sleeves, other Corrosive environment, Corrosion

crevices oxygen-starved areas

Material wastage, Moisture-sensitive tape

leaks

180
heat-affected zones of the recirculation

-piping welds have been particularly susceptible. Three conditions are required for

IGSCC of austenitic stainless steel to exist:

a sensitized microstructure, high tensile

stresses, and an oxygenated environment.

Several techniques have effectively mitigated IGSCC damage. These techniques

include solution heat treatment of the

welds, heat sink welding, induction heating stress improvements of the welds, lastpass heat-sink welding, mechanical stress

improvement processes, weld overlays,

and use of hydrogen water chemistry.

Nuclear-grade stainless steels are much less

susceptible to sensitization and more

resistant to IGSCC, and therefore are used

as alternate materials for the recirculation

piping material.

2. The possible detrimental effects, if any, of

hydrogen additions to the BWR recirculation piping loop need to be evaluated.

Addition of hydrogen will reduce the oxygen level in the BWR coolant and may lead

to degradation of the feedwater lines

because of a erosion-corrosion mechanism

similar to that which caused the recent

Surry pipe break.

3. A program should be developed to monitor the actual degree of thermal-aging

embrittlement in the cast stainless steel

components in the recirculation piping

loop.

4. The long-term fatigue behavior of the weld

overlays needs to be better understood.

15.8 Nondestructive Examination

Methods

The NDE methods used in standard practice are visual examinations, penetrant and magnetic particle testing, radiographic examinations, eddy-current testing,

ultrasonic testing, and acoustic-emission monitoring.

These standard NDE methods, which are employed to

satisfy the present ASME code ISI requirements, are

not entirely adequate for residual life assessments.

Inspection methods are needed to accurately determine

the size, shape, location, orientation, and type of both

surface and internal flaws, so that fracture mechanics

approaches may be used for life assessment. Several

emerging methods being developed to satisfy these

needs are summarized in Table 15.12. Several emerging

inspection methods that can be used to detect timedependent changes in microstructural features (such as

the ferrite phase in duplex stainless steels) are also summarized in Table 15.12. Recommendations for further

work on emerging inspection methods are as follows:

1. Development of the synthetic aperture focusing technique (SAFI) should be continued.

SAFT ultrasonic testing (UT) can provide

enhanced visual images of flaws detected during inspection. The initial results of the application of SAFT-UT in the nuclear industry

are encouraging.

2. The dc potential drop method and computeraided eddy-current testing techniques need to

be evaluated and their ability to detect flaws in

pipe needs to be determined.

3. The use of acoustic emission technology

for on-line monitoring of crack growth in

pressure boundaries, and leak surveillance

in LWR systems should be evaluated.

4. NDE methods for measuring the material

properties changes caused by aging of the

major LWR components are needed. Some

of the promising methods are eddy-current

methods to measure residual stresses, electrochemical methods to measure the susceptibility of austenitic stainless steel to

IGSCC, and indentation hardness measurements to assess radiation embrittlement of RPV steel.

15.9 Emerging Life-Assessment

Methods

Emerging life-assessment methods include miniature specimen testing (MST), reconstituted specimen testing, on-line damage monitoring, 'and

effective engineering models applications.

Table 15.12 summarizes two of these lifeassessment methods. These methods may provide

valuable information concerning several important

degradation mechanisms. For example, MST can

be used for surveillance testing where only a limited

amount of test material is available. On-line damage monitoring can be applied to calculate fatigue

usage factors for RPV nozzles during startups,

shutdowns, and major operating transients, where

nozzle temperatures are monitored during operation. The engineering models can then be usedalong with the information from the improved ISI

methods, data from MST activities, and on-line

181
monitoring data-to properly estimate the residual

life of various components. These estimates will

form the basis for LWR license renewals. Recommendations for further work on the emerging lifeassessment methods are as follows:

1. The usefulness for RPV surveillance testing of the miniature specimen and reconstituted specimen testing methods needs to

be evaluated. The validity of the tensile

property, fracture toughness, and fatiguecrack initiation and growth data obtained

with these two methods also needs evaluation.

2. Development, field demonstration, and

validation of the on-line damage monitoring methods are required before they can

be implemented with confidence.

3. Simplified engineering models should be

developed to provide reasonable estimates

of material aging as a function of operating history.

;

182
Table 15.12. Summary of emerging methods for inspection and life assessment

Method Characteristics Applications

SAFT-UT Computerized processing of ;

ultrasonic data to provide enhanced

significant visual images of flaws in

real time.

.Welds in the RPV, in piping,

and at nozzles. Requires

computer resources resources

resources and has not been fully

developed and verified for field

applications.

Compute

inspectio

r-aided UT pipe Automated ultrasonic inspection of

n systems piping by means of state-of-the-art

hardware and software.

itial drop pipe Automated inspection and

n systems computerized analysis of DC - potential distributions around cracks

on both inner and outer surfaces of

piping.

DC potei

inspectio

Piping and IGSCC cracks near

welds in piping. Need more --real-time analysis and

interpretation of data to aid

operators in decision making.

Promising new method for

inspection of pipes. Prototype

system needs to be developed,

demonstrated, and verified for

routine commercial applications.

More reliable inspection of

: PWR steam generator tubes. The

advanced methods need to be - fully evaluated and

demonstrated. Deep penetration

methods show some potential for

pipe inspection, but they need

considerable development.

I

Computerized eddycurrent inspection

systems

Advanced eddy-current coil designs

coupled with real-time computer

analysis of data. -

.. I ....

Acoustic emission

monitoring

On-line damage analysis~~~~

Uses computerized processing

of data to monitor noise from

-crack growth or-water leakage. :

Monitoring of fatigue-crack

growth in RPVs, possibly

stress-corrosion crack growth -

in piping, and water leaks in the

primary pressure boundary.

Requires advanced techniques to

remove background noise.

Results of initial field studies are

encouraging, but further

demonstration and validation are

needed.

, Monitors temperature ani

during operation and con

damage accumulation on

real time. ;

d pressure Creep-fatigue damage in fossil

2putes boiler headers and fatigue

-line in ; usage factors in nuclear

i- plant nozzles. Only prototype - ; systems are currently available.

Much uncertainty with damage

accumulation models that are

used.

183
Table 15.12. (continued)

Method Characteristics Applications

Neutron-noise analysis

Nondestructive

measurement of

materials properties

Surface replication

Miniature specimen

testing

Effective engineering

models

Measures the ex-core neutron noise

and analyzes the recorded data to

diagnose potential problems within

the RPV.

Uses ultrasonic, eddy-current,

x-ray, and electrochemical testing

methods to measure properties, such

as tensile strength, fracture

demonstrated, toughness, and impact

strength.

Produces plastic film replica of

material surface that can be

examined at high magnifications

using both optical and electron

microscopy to detect small cracks,

voids, and microstructural changes.

Uses a small specimen removed from

a component to measure materials

properties.

Use of simplified engineering models

that provide reasonable descriptions

of material aging as a

function of operating history.

Shows promise for monitoring

vibration of PWR internals

and BWR instrument tubes

and fuel boxes. Need reference

data for normal operations and

hardware and software for

on-line use.

Most applications have been

limited to laboratory studies.

Offers great potential when

methods are developed,

and verified by means of future

basic and applied research. Need

to develop a library of

well-characterized correlations

between nondestructive

measurements and properties.

Widely used to inspect for creep

damage in fossil-plant headers

and steam piping. Has potential

use for the detection of small

cracks and microstructural

changes caused by aging in key

nuclear components. Could be

used to measure ferrite amount

and spacing in cast stainless

steels. Needs to be developed for

nuclear-plant applications.

Laboratory studies are being

performed to define the range of

use for this technique. Shows

promise for the measurement of

deformation, such as creep and

stress-strain response. May have

applications for measurement of

fracture toughness and crack

growth. Further verification of

the technique is required for field

application.

Trending the aging of key

components from a knowledge

of the operating history. Most

current approaches seem either

too detailed and cumbersome to

use or too simplistic to provide

realistic predictions.

III

I

184

I
NRC FORM 335 U.S. NUCLEAR REOULATORY COMMISSION I REPORT NUMBER IAslgs,.ob, TIDC add V2 ho. Ac."l

C BIBLIOGRAPHICDAAS NUREG/CR-4731 ,,0,, 3702 BIBLIOGRAPHIC DATA SHEET EGG-2469 Volume 1

SEE NSTIRUCTIONSON THE REVERSE

J TITLE AND SUBTITLE J LEAVE BLANK

Residual Life Assessment of Major Light Water Reactor

Components--Overview Volume 1 A DATE REPORT COMPLETED

MONTH ',EAR 6 I 87

S AUTHORISI

V. N Shah and P. E. MacDonald e DATE REPORT ISSUED

MONTH YEAR

6 1. 87

AGE oGANf12ION ,4R NAME AND MAILING ADDRESS I'Iciud. ZMI Coad) I B PROJECTITASK.WORK UN?7 NUMBER calm 9 ano, lnc.

P. 0. Box 1625 9 FIN OR GRANT NUMBER

Idaho Falls, Idaho 83415 A6389

10 SPONSORING ORGANIZATION NAME AND MAILING ADDRESS Ilcluoe Za Cod.l II TYPE OF REPOF°Final, Technical

U.S.Nuclear Regulatory Commission

Washington, D. C. 20555 b PERIOD COVERED liu~ doal/

17 SUPPLEMENTARY NOTES

13 ABSTRACT (ZO.1d. e,-.t.

This report presents an assessment of the aging (time-dependent degradation) of

selected major light water reactor components and structures. The stressors, possible

degradation sites and mechanisms, potential failure modes and currently used nondestructive examinations, in-service inspection (ISI) and life assessment methods are

discussed for seven major light water reactor components: pressurized water reactor

(PWR) and boiling water reactor (BWR) pressure vessels, PWR containment structures, PWR reactor coolant piping, PWR steam generators, BWR recirculation piping, and reactor pressure vessel supports. Unresolved technical issues related to life

extension of these components, including requirements for advanced ISI and life

assessment methods, are also discussed.

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