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EGG-SSRE-1 0039

May 1993

An Evaluation of the Effects

of-Valve Body Erosion on

Motor-Operated Valve Operability

Thomas H. Hunt

Michael E. Nitzel
EGG-SSRE-1 0039

An Evaluation of the Effects of

Valve Body Erosion on

Motor-Operated Valve Operability

Thomas H. Hunt

Michael E. Nitzel

NRC Technical Monitor

G. H. Weidenhamer

Published May 1993

Idaho National Engineering Laboratory

EG&G Idaho, Inc.

Idaho Falls, Idaho 83415

Prepared for the

Division of Engineering

Office of Nuclear Reactor Research

U.S. Nuclear Regulatory Commission

Washington, D.C. 20555

Under DOE Idaho Operations Office

Contract DE-AC07-761D01570

FIN A6389
ABSTRACT

Engineers at the Idaho National Engineering Laboratory evaluated the

effects of erosion-induced valve wall thinning on motor-operated valve

operability. The authors reviewed reports that identified the extent and

location of erosion damage in nuclear plant valves and chose a globe valve

with severe erosion damage to assess the potential for loss of operability.

They developed a finite element model of the selected valve and performed

a structural analysis with valve closing forces to analyze the effects of the

erosion on structural integrity. The results indicate that sufficient margin to

yield stress remained. Therefore, erosion-related wall thinning is not likely to

create an operability problem for MOVs.

A6389-Aging-Components and Systems IV

iii I
EXECUTIVE SUMMARY

Wall thinning from erosion has'causeddamage in valves typical of those installed ini a variety

of systems in nuclear power plants. The available damage reports'indicate that 'among'safety-

'related systems,' residual heat removal (RHR) systems appear to have the'highest in cidence of

damage from erosion. Service water systems seem to incur a high incidence of 'erosion-related '

damage in nonsafety systems. '

When extensive erosio does occur, excessive localized or distributed wall thinning is

possible, which could create undue stresses and deformations in the valve'body. These

deformations could lead to relative displacements, which may in turn induce disc or stem binding.

Binding would prevent the valve from performing its safety or control function.

-Because of these'concerns, the NRC commissioned the Idaho National Engineering

Laborato y' (INEL)' to evaluate the effects of erosion-induced valve wall thinning on motoroperated valve (MOV) ope'rability. This investigation was sponsored as part'of the NRC's

Nuclear Plant Aging Research (NPAR) program.; The evaluation consisted 6f the following

elements:

* Reviewing reports that identify the extent and location of erosion damage in nuclear

plant valves

* Choosing a specific valve where severe damage was occurring

* Developing a finite-element model of the valve

* Performing a structural analysis of the valve with valve closing forces after reducing the

valve's wall thicknesses to simulate the eroded conditions

* Analyzing the results of the structural analyses.

This report only evaluates the effect of erosion-related wall thinning on MOV operability.

This limits the investigation to damage occurring to the valve itself. It excludes any damage that

may occur downstream in the connected piping as a result of flow perturbations occurring in the

valve. While this damage would affect the operability of the system, it does not directly affect the

operability of the valve and therefore is not part of this study.

As the first step in our study, we reviewed existing reports and failure events to identify

where erosion damage is occurring in valves, and the extent and severity of the damage.

Appendix A summarizes a representative selection of these reports. This review revealed that

valves in the RHR system in boiling water reactors (BWRs) are among those most likely to

experience erosion. It also indicated that a globe valve that is frequently throttled would be most

affected. Therefore, we chose a 16-in. carbon steel globe valve used in the RHR suppression

pool test line as a typical valve. The chosen valve had experienced the most severe damage

reported for the RHR system.

v
Our next step was to perform a finite element analysis of this valve. The model developed

represents the complete valve body, seat, and bonnet regions of the valve, including the areas

where erosion had been observed. Since the selected valve is normally closed or nearly closed,

the model included full system pressure (225 psig maximum pump dead head pressure) upstream

of the valve seat. Because of the normal operating configuration of the valve, no internal

pressure was modeled downstream of the valve seat. Normal operating temperature of 150TF was

accounted for in the model's material properties. Conservative estimates of the valve stem force

required to seat the valve were also modeled. The maximum moments on the valve caused by the

attached piping were also included as loads on the valve.

As part of the analysis, we made two "scoping" analysis computer runs simulating degraded

valve body conditions. These computer runs simulated valve erosion damage worse than that

which would be expected to occur. The first run was a conservative model of the erosion that

was observed on the Brunswick valve (see Appendix B). The second run simulated through-wall

erosion that would leak and be detected. The results of both analysis runs were evaluated and

stresses below yield were indicated in all cases. The results indicate that sufficient margin to yield

stress remained so that erosion-related wall thinning was judged not likely to create an operability

problem for MOVs.

vi
CONTENTS

ABSTRACT. ! ....................... iii

EXECUTIVE SUMMARY ........................... v

INTRODUCTION .................... 1 . . . . .~ . ~ . ~ ~ ~~ ~~~. 

EROSION PROBLEMS IN THE NUCLEAR INDUSTRY . 2

FINITE ELEMENT MODEL . 3

STRUCTURAL ANALYSIS. 7

CONCLUSIONS. 11

REFERENCES. .......................................................... 12

APPENDIX A-SUMMARIES OF EROSION DAMAGE REPORTS .A-1

APPENDIX B-BRUNSWICK EVENT DESCRIPTION .B-1

APPENDIX C-ANALYSIS OF A 16-INCH GLOBE VALVE WITH ERODED WALLSDESCRIPTION OF THE FINITE ELEMENT MODEL DEVELOPMENT .C-1

APPENDIX D-ANALYSIS OF A 16-INCH GLOBE VALVE WITH ERODED WALLSDESCRIPTION OF THE ANALYSIS RESULTS .D-1

FIGURES

1. Cross-sectional and overhead views of Anchor-Darling 16-in. globe valve. 5

2. Wire-frame drawing of model of 16-in. globe valve. 6

3. Stress contour plot of the bonnet cylinder region of 16-in. globe valve. 8

C-1. Simplified view of 16-in. globe valve and erosion areas .C-4

C-2. I-DEAS finite element model of 16-in. globe valve with eroded walls .C-6

C-3. Bonnet region elements with the areas of maximum erosion for 16-in. globe valve

model .C-7

D-1. Representative stress contour plot of eroded bonnet region .D-6

vii
TABLES

1. Percent of throttleable valves larger than 4 in. in RHR systems of BWRs and PWRs ...... 4

2. Maximum stress results with tmin = 0.10 in ...................................... 9

3. Maximum stress results with tmn = 0.010 in .................................... 10

D-1. Maximum stress results with tmin = 0.10 in ................................... D-7

D-2. Maximum stress results with ton = 0.010 in .................................. D-8

viii
An Evaluation of the Effects of

Valve Body Erosion onMotor-Operated Valve Operability

INTRODUCTION:

Since 1983, the U.S. Nuclear Regulatory Commission (NRC) has issued information notices,

bulletins, technical reports,'and a generic letter addressing nuclear plant problems related to'

erosion damage to pumps, piping, and valves (see the Reference section on page 12).. Erosion in

nuclear plant systems continues to be a concern to the NRC.

Wall thinning from erosion has caused damage in valves typical of those installed in a variety

of systems in nuclear power plants. The available damage reports indicate that among safetyrelated systems, residual heat removal (RHR) systems appear to have the highest incidence of

damage from erosion. Service-water systems seem to incur a high incidence of erosion-related.

damage in nonsafety systems. -

When extensive erosion does occur, excessive localized or distributed wall thinning is

possible, which could create undue stresses and deformations in the valve body. These

deformations could lead to relative displacements, which may in turn induce disc or stem binding.

Binding would prevent the valve from performing its safety ,or control function.

Because of these concerns, the NRC commissioned the Idaho National Engineering

Laboratory (INEL) to evaluate the effects of erosion-induced valve wall thinning on motoroperated valve (MOV) operability., This investigation was sponsored as part of the NRC's

Nuclear Plant Aging Research (NPAR) program._.The evaluation consisted of the following

elements: ; .

Reviewing reports that identify the extent and location of erosion damage in nuclear

plant valves . - i , .

Choosing a specific valve where severe damage was occurring

, * Developing a finite-element model of the valve

Performing a structural analysis of the valve with valvefclosing forces after reducing the

valve's wall thicknesses to simulate the eroded conditions

Analyzing the results of the structural analyses.- .

This reportonlyevaluates the effect of erosion-related wall thinning on:MOV operability.

This limits the investigation to damage occurring to the valve itself. It excludes any damage that

may occur downstream in the connected piping as a result of flow perturbations occurring in the

I
valve. While this damage would affect the operability of the system, it does not directly affect the

operability of the valve and therefore is not part of this study.

As the first step in our study, we reviewed existing reports and failure events to identify

where erosion damage is occurring in valves, and the extent and severity of the damage.1-11,a

Appendix A summarizes a representative selection of these reports. This review revealed that

valves in the RHR systems of boiling water reactors (BWRs) are among those most likely to

experience erosion. It also indicated that a globe valve that is frequently throttled would be most

affected. Therefore, we chose a 16-in. carbon steel globe valve used in the RHR suppression

pool test line as a typical valve. The chosen valve had experienced the most severe damage

reported for the RHR system. Details of the finite element and structural analyses are given in

subsequent sections of this report.

EROSION PROBLEMS IN THE NUCLEAR INDUSTRY

Most of the erosion-related problems in the nuclear industry have occurred in piping and

valves. In the case of valves, the erosion damage usually results in internal leakage past the

seating surfaces. All of the catastrophic failures have occurred in nonsafety system piping,

primarily the main turbine extraction steam and the main feedwater systems. Service water

systems also have experienced a large number of erosion-related failures, including through-wall

leakage.

In the first-line safety systems, primarily emergency coolant injection and containment

isolation systems, few occurrences of through-wall erosion damage are reported. However, there

have been examples of valve body damage caused by erosion. One of the more significant

instances occurred at the Brunswick Nuclear Plant. In 1989, evidence of severe erosion damage

was discovered during normal maintenance activities on the Unit I RHR inboard low-pressure

coolanrt injection (LPCI) valve. An investigation on the redundant LPCI train valve indicated that

it, too, had suffered severe damage. The investigation determined the cause to be cavitationinduced erosion because of throttled flow through the valve. An expanded investigation revealed

four more valves in the RHR systems and one valve in the high-pressure coolant injection (HPCI)

system also had suffered cavitation-induced erosion damage. Additional details of this incident

are provide in Appendix B.

The cavitation-induced erosion damage in valves prompted the NRC to initiate a study by

Oak Ridge National Laboratory (ORNL). This study reviewed industry experience with erosion

damage and determined that the service *water and condensate/feedwater systems accounted for

78% of the reported through-wall valve failures. Their report did not identify any through-wall

failures in first-line safety systems, a fact they attributed to the lower operating hours for these

systems. To perform a more detailed study on erosion, ORNL studied a single system in depth.

ORNL chose the RHR systems at both BWRs and pressurized water reactors (PWRs) for three

reasons: (a) they are critical safety systems, (b) the previously mentioned Brunswick event

occurred in the RHR system, and (c) they are operated more often than other first-line standby

safety systems.

a. Private communication with D. Casada, Oak Ridge National Laboratory, Oak Ridge, Tennessee, 1991.

2
Throttleable valves greater than 4 in. nominal size were considered to be of primary

importance. The review of RHR systems identified approximately 900 throttleable valves larger

than 4 in. in the RHR systems of BWRs and PWRs; nearly 500 of these were located in carbon

steel piping systems in BWRs. Table 1 shows the distribution of the valves'identified. ORNL's

review showed that BWRs account for the majority of the valves large'r than 4 in. used in RHR

systems, and the number of globe valves in BWR RHR systems is larger than the total number of

valves used in all of the PWR RHR systems.

Our primary purpose in reviewing the ORNL study and other operating history was to

determine an appropriate valve to be modeled in this study. Although erosion appears to be most

prevalent in the service water and feedwater systems, neither of these systems tends to have

valves that are subjected to high stresses during accident conditions. In concurrence with ORNL's

reasons for detailed study of RHR systems, we elected to model an RHR valve. Also, because

the ORNL study showed that more valves exist in BWR RHR systems, and most of these are

globe valves, we chose to model a carbon steel globe" valve typical of one 'used in BWRs. The

Brunswick event provided detailed descriptions of actual erosion damage and the locations of the

damage. Consequently, we chose to model the Anchor-Darlingb 16-in. globe valve used in the

suppression pool test line at Brunswick. Figure 1 shows simple cross-sectional and overhead views

of this valve.

The locations where the erosion damage occurred are illustrated by'the'shaded areas in

Figure 1. These areas ranged from about 1 to 5 in2. The original thickness of these areas was

about 1.25 in. thick. The actual damage removed nearly 1 in. of material thickness in some areas.

The damage was severe enough to reduce the wall thickness to less than the required code

thickness for a 300-lb valve, and an engineering analysis was required to justify continued

operation followin g the outage without repair or replacement of the valve.

FINITE ELEMENT MODEL-.

A primary concern for this study was that erosion damage could occur without detection

until the valve failed in service. Unless the bridge structure (the part of the valve body that

encloses the seat ring of the valve) or valve body experienced through-wall erosion, the damage

would not be detected during normal operation. This occurred at' the Brunswick plant, where the

damage was found when the valve was disassembled for routinemaintenance.: -:No prior indication

of the damage had been apparent. If the fluid cavitation had continued to erode the valve body,

the valve may not hive retained adequate structural integrity to operate when required.

To evaluate this problem, we modeled the valve using the I-DEAS modeling and analysis

software. A wire frame drawing of the model is shown in Figure 2. The intent of the model is to

bound the stresses that the valve would experience and to determine whether the valve material

yield stresses would be approached. To do this, we decided it was necessary only to perform an

elastic analysis.-- If -the results showed elastic stresses were not exceeded, the strain would be in

b. Mention of specific products and/or manufacturers in this document implies neither endorsement or

preference, nor disapproval by the U.S. Government, any of its agencies, or EG&G Idaho, Inc. of the use of

a specific product for any purpose.

3
Table 1. Percent of throttleable valves larger than 4 in. in RHR systems

of BWRs and PWRs.

Valve type

Babcock & Wilcox Plants

Stainless steel butterfly

Stainless steel globe

Total

Combustion Engineering Plants

Stainless steel butterfly

Stainless steel globe

Total

General Electric Plants

Stainless steel globe

Carbon steel globe

Carbon steel butterfly

Total

Westinghouse Plants

Stainless steel globe

Stainless steel butterfly

Carbon steel globe

Carbon steel butterfly

Total

Percent of

valve population'

0.5

2.6

3.1

5.4

11.2

16.6

5.2

47.4

3.1

55.7

0.2

0.5

18.8

4.8

24.3

a. Total valve population is approximately 925 valves. Percentages are the

fraction of all valves.

the elastic range, displacements would be small, and valve operability most likely would not be

inhibited. If the results showed elastic stress values were exceeded, additional analysis would be

needed to study the effect of valve wall thinning on MOV operability.

4
Bonnet area

Plug guide

indication

I ,. I , - I I

I -

I I _ . I :

Valve : :I . .

inlet

Valve

I outlet - I I

_771.

L93 0405

_ I

Erosion damage area

M Plug guide

Figure 1. Cross-sectional and overhead views of Anchor-Darling 16-in. globe valve.

5
Inlet IEros ion Area

:~~~~~ I

Seat RingOl

'"Oule

Figure 2. Wirc-framc drawing of model of 16-in. globe valve.
Using the Brunswick event as .a template .for modeling erosion damage, we duplicated the

seven areas of damage in our model. The damage areas with the greatest damage occurred

downstream of the seating surfaces in the valve body on each side of the three plug guides'. The

seventh location was in the bonnet cylinder area. This last area had only minor uniform damage,

which covered almost the entire circumference of the upper region of the bonnet cylinder.

-The damaged areas around the guides in the Brunswick valve varied from about 1 to 5 in.2

with losses of about 0.375-0.75 in. of wail thickness. The eroded areas tapered down in thickness

from the nominal wall thickness (1.25 in.) to the minimum value that remained (0.5-0.875 in.).

In the finite element model, the wall thickness was decreased to 0.10 in. for the first run and

0.010 in. for the second run. Since the original thickness of the walls was greater than 1.25 in.,

the minimum values chosen for the analysis were conservative representations of the maximum

thinning observed. This allowed us to reasonably simulate conditions for a through-wall leak

while maintaining the numerical stability of the model.' Also, Brunswick personnel performed an

engineering analysis to determine the minimum allowable wall thickness for the valve at operating

pressure (about 195 psig), which is 0.492 in.. Since the minimum wall thickness values chosen for

the finite element analysis were both well below 0.492 in., we believe that the results of the

analysis provide a conservative indication'of structural deficiencies under the loading conditions

considered. -

'For both runs, the valve was modeled in the closed position, with internal pressure only

upstream of the' seat ring. The other loads included in each run were the piping forces and

moments on the'valve ends, and plug/seat forces and their reaction forces on the upper bonnet.

*sThis loading provided the maximum load across the seat for a conservative structural analysis.

- Because there is no pressure load downstream of the seat,' the locally thinned elements did not

have direct pressure stress. -This is a local effect,lhowever, and would not'affect the'overall ability

of the valve body to resist the deformations required to produce binding of the plug or stem and

failure of the valve. Since the eroded areas in the upper bonnet cylinder had only minor uniform

damage, we duplicated this area in the model without any modification or increase in' its severity.

Appendix C describes the model and modeling assumptions in detail.-

STRUCTURAL ANALYSIS

The I-DEAS software was used to develop the globe valve model and perform the stress

analysis. -Details of the model and the load combinations considered in the model are identified

in Appendix D. -After completing the stress calculations, we made stress contour plots. Figure 3

shows a representative midwall stress contour plot resulting from the analysis of, the valve with a

minimum eroded wall thickness of 0.010 in. Using the color scale in Figure 3, one can see that

the highest stresses are occurring approximately in the middle of the figure. These high stresses

are located in the outlet regions of the valve near the plug guides. For the run where the

minimum wall thickness was 0.10 in., the stress contour plot would appear similar to that shown in

'Figure 3 but with lower stresses. The maximum stresses for both cases are listed in Tables 2 and

3. A detailed description of the analysis results is given in Appendix D.

The analytical results show that the stresses are well below the yield stress for both cases.

--Tables 2 and 3 summarize the results of these calculations. As shown in the tables, the yield

stress for the valve material 34.4 ksi at an operating temperature of 150'F. The maximum stress

7
Figure 3. Stress contour plot of the bonnet cylinder region of 16-in. globc valve.
Table 2. Maximum stress results with tmin = 0.10 in.

Calculated Von Mises stresses3

Model area

Seat ring

-

,: I Element

;surface .. .. sufc . .. 

Top

Middle

Bottom

- Load

combination lb

9.33

-*-, 9.83

1 1.; 10.5

* I... Load

combination 2 ;,

9.53 . . 10.

10.2

,.,, 11.1

Yield

stress

'(ksi)

34.4

34.4

34.4

. Bridges Top

Middle

Bottom

11.7

5.26

-- , 8.08

- 12.1

,. - 5.41

I * . ; 8.34

* 34.4

34.4

34.4

* Upper Bonnet

(Area 7)

Top

Middle

Bottom

i'1⁄4 r1 4.87

3.02

4.09

4.95

." % 3.05

- 4.06

-, 34.4

34.4

34.4

Bonnet Cylinder .;

(Areas 1-6)

Top

Middle

Bottom

A 16.6

11.7

:'- 17.5

17.1 , .

11.5

- 18.0

34.4

; 34.4

34.4

a. Von Mises stresses were used because of their accuracy in representing material behavior under

static loading. , . , ;

b. Load combination 1 is the combination of operating pressure, valve seating force, and pipe end

moments. . * ,;

c. Load combination 2 is the combination of pump dead head pressure, valve seating force, and pipe

end moments.

., i I

I . ! , . - ' -, "
Table 3. Maximum stress results with tin = 0.010 in.

Calculated Von Mises stressesa

Yield

Element Load Load stress

Model area surface combination lb combination 2c (ksi)

Seat ring Top 9.50 9.77 34.4

Middle 10.4 10.8 34.4

Bottom 11.3 11.9 34.4

Bridges Top 12.8 13.4 34.4

Middle 4.82 5.00 34.4

Bottom 9.87 10.2 34.4

Upper bonnet Top 5.38 5.44 34.4

(Area 7) Middle 3.25 3.20 34.4

Bottom 3.84 3.81 34.4

Bonnet cylinder Top 24.5 24.2 34.4

(Areas 1-6)

Middle 25.0 24.8 34.4

Bottom 29.1 29.0 34.4

a. Von Mises stresses were used because of their accuracy in representing material behavior under

static loading.

b. Load combination I is the combination of operating pressure, valve seating force, and pipe end

moments.

c. Load combination 2 is the combination of pump dead head pressure, valve seating force, and pipe

end moments.

indicated in Table 2 is approximately 52% of the yield stress, while the maximum stress shown in

Table 3 is approximately 85% of the yield value.

10
One of our initial concerns was that the erosion damage observed at Brunswick that

occurred on the bridge area could lead to'loss of support for the seat ring. Although this area

was modeled with much greater than observed damage, the bridge has an even greater margin to

yield than the outlet areas of the'valve. '

CONCLUSIONS

The analysis described above is not a complete analytical treatment of the vaive. 'The

objective of this analysis was to conduct a scoping study that would indicate the structural margin

that may be available when wall thinning is encountered similar to what,.was observed at

Brunswick. This information could then be used to predict how the erosion might affect MOV

operability. '

The results of our analysis show that even with conservative modeling of the erosion, the

maximum stresses remain below'the yield stresses for the valve material. Therefore, we conclude

that erosion-related valve wall thinning is not likely to create an operability problem for MOVs.

.- .....

11
REFERENCES

1. NRC, Service Water System Problems Affecting Safety-Related Equipment, Generic Letter

89-13, Supplement 1, April 4, 1990.

2. NRC, Service Water System Problems Affecting Safety-Related Equipment, Generic Letter

89-13, July 18, 1989.

3. NRC, Excessive Valve Body Erosion at Brunswick, AEOD/E908, September 29, 1989.

4. NRC, Summary of January 11, 1989 Meeting Held with BWR Owner's Group Regarding

Valve Degradation By Erosion, January 31, 1989.

5. NRC, Valve Body Erosion, NRC Information Notice 89-01, January 4, 1989.

6. NRC, Pump Damage Caused by Low Low-flow Operation, IE Information Notice 89-08,

January 26, 1989.

7. NRC, Summary of January 10, 1989 Meeting Held With CP&L Regarding Valve Degradation

by Erosion, January 31, 1989.

8. NRC, Potential Safety-Related Pump Loss, NRC Bulletin 88-04, May 5, 1988.

9. NRC, Thinning of Pipe Walls in Nuclear Power Plants, NRC Bulletin 87-01, July 9, 1987.

10. NRC, Erosion in Nuclear Power Plants, AEOD/E416, June 11, 1984.

11. NRC, Misapplication of Valves by Throttling Beyond Design Range, IE Information Notice

83-55, August 22, 1983.

12
Appendix A

Summaries of Erosion Damage Reports

A-1
.Appendix A

Summaries'of Erosion Damage Reports

Office of Analysis and Evaluation of Operational Data (AEOD) '

The AEOD sponsored several reports reviewing and investigating instances of erosionrelated problems occurring in the commercial nuclear industry. These reports investigated trends

in erosion-related failures and evaluated industry responses and the programs intended to resolve

these problems.

NRC, Erosion in Nuclear Power Plants, AEOD/E416, June 11, 1984.

This report documents problems with erosion-related failures in both steam and water piping

systems. It identifies areas having potential safety impact and suggests review of six areas: service

water systems, J-tubes in steam generators, emergency feedwater systems, erosion-related leakage

of main steam isolation valves, potential for rupturing steam lines, and personnel safety.

NRC, Ecessive Valve Body Erosion at Bnunswick, AEOD/E908, September 29, 1989.

This report discusses the valve erosion damage at Brunswick Unit 1. Several groups of

erosion-related failures are identified that may have a potential impact on nuclear safety. These

include failures from cavitation-induced damage, vibration damage to valve assemblies and piping

supports, piping erosion downstream of throttled valves, severe degradation of pumps, low-flow

damage to pumps, and erosion fouling of heat exchangers.

Office of Nuclear Reactor Regulation (NRR)

NRR issued two bulletins in 1987 and 1988 that addressed concerns about erosion-related

degradation of piping systems and pumps.

NRC, Thinning of Pipe Walls in Nuclear Power Plants, Bulletin No. 87-01, July 9, 1987.

This bulletin addresses the problems associated with erosion-corrosion failures in carbon

steel piping systems. Large-diameter piping systems previously were not routinely monitored for

erosion-corrosion damage because of the limited number of failures that have occurred in carbon

steel piping systems. The bulletin requires all licensees to provide information on their programs

to monitor piping wall thicknesses in carbon steel piping systems.

A-3

! I.
NRC, Potendial Safety-Related Pump Loss, Bulletin No. 88-04, May 5, 1988.

This bulletin discusses the problems associated with inadequate minimum flow protection for

safety-related centrifugal pumps. As flow is decreased for an operating pump, cavitation and

other problems may occur that could damage the pump. This low-flow condition allows the fluid

in the pump to begin to cavitate, causing vibration and damage to the impeller and pump casing.

A-4
Appendix B

Brunswick Event Description

B-I
Appendix B

Brunswick Event Description

In December of 1988, Brunswick Unit 1 technicians conducted maintenance on the stem and

backseat of the residual heat removal (RHR) system's low-pressure coolant injection (LPCI)

valve, 1-E1l-FO17B (a 16-in. carbon steel angle body globe valve). When the valve was

disassembled, localized pockets of erosion damage were noted in the walls of the valve body just

above the seat ring on the outlet volute of the valve. Followup inspection of these pockets

determined that more than half (2.3 in. of the original 3.5-4 in.) of the wall thickness had eroded

away. This reduced the wall thickness below the minimum required by the appropriate design

code [American National Standards Institute (ANSI) 16.5]. The redundant RHR subsystem LPCI

valve, 1-E1-FOl7A, was then inspected and similar indications were noted in this valve, although

the minimum wall thickness requirements were met.

During this same period, the RHR suppression pool cooling inboard isolation valve,

1-E11-F024B, failed a local leak rate test, and was disassembled for repair and inspection. The

inspection noted areas of localized erosion damage in the walls of the valve body and on the

bridge supporting the seat. These indications were above the seat ring and adjacent to the plug

guides. In one location, the wall thickness did not meet the minimum required. The redundant

RHR subsystem suppression pool cooling inboard isolation valve, 1-El1-F024A, was then

disassembled for inspection. Similar indications were noted, but the wall thicknesses did not fall

below the minimum values.

Following the discovery of the problems on the Unit 1 RHR valves, 29 valves in the RHR,

core spray, reactor core isolation cooling, and high-pressure coolant injection (HPCI) system were

inspected. Only one other valve had indications of erosion damage, the HPCI full flow test

throttle valve (1-E41-F008). The damage to 1-E41-F008 was limited to eroded plug guides, with

only minor damage to the valve walls.

The Unit 2 RHR LPCI injection and suppression pool cooling inboard isolation valves

(2-Ell-F017A/B and 2-Ell-FO24A/B) also were inspected. These inspections revealed similar

areas damaged to a lesser extent than the Unit I valves.

The root cause of the erosion damage was throttling flow through the valves. This throttling

caused cavitation and resulted in the localized erosion of the valve. The investigation also showed

that the cavitation occurred at all flow rates for the suppression pool cooling inboard isolation

valves and became more prevalent at higher flow rates. The LPCI injection valves also cavitated

at all flow rates, but the cavitation increased at the lower flow rates.

B-3
Appendix C

Analysis of a 16-Inch Globe Valve

with Eroded WallsDescription of the Finite Element

Model Development

C-1
Appendix C

Analysis of a 16-Inch Globe Valve

with Eroded WallsDescription of the Finite Element

Model Development

MODEL DESCRIPTION,

Valve Description

The intent of this analysis is to perform a "scoping analysis" togauge the effect of erosion on

the valve and to determine whether more detailed analysis is warranted. It is not intended to be a

detailed stress analysis that considers every possible load case.

The analyzed valve is a 16-in. nominal size globe valve with an American National Standards

Institute (ANSI) 300-lb class rating manufactured'from A216, grade WCB, cast carbon steel. The

component was manufactured-by Anchor-Darling Industries' and is representative of several that

are installed in the residual heat removal (RHR) piping systems of both nuclear power generating

units at the Brunswick plant. A simplified view of the valve and a plan view of some of the

erosion areas are shown in Figure C-1. The valve was originally qualified for operation under the

ANSI B31.1 Power Piping'code,-1973 Edition (hereinafter referred'to as' "the Code") and the

ANSI B16.5 Standard, "Steel Pipe Flanges and Flanged Fittings." '

Modeling Assumptions

No engineering or casting drawings were available for use in constructing the model of the

16-in. globe valve. -Consequently,-we used field measurements-and sketches obtained from

Brunswick plant personnel and Anchor-Darling'personnel during the model development. In

some cases, assumed values were used where no other information was available. Also, since this

is a scoping analysis, certain simplifying assum'ptions were appropriate... Major model features and

simplifying assumptions 'Emiployed in this analysis are discussed in the paragraphs below.

The model was developed using the I-DEAS software package written by the Structural

Dynamics Research Corporation. This software includes drafting, solid modeling, analysis, and

post processing capabilities. Interfaces to other widely used analysis codes (e.g., ABAQUS) are

also included within I-DEAS. -.The solid modeling feature of the code was used to construct a

representation of the complete valve, which contained 5,026 nodes and 4,992 quadrilateral plate

a. Mention of specific products and/or manufacturers in this document implies neither endorsement or

preference, nor disapproval by the U.S. Government, any of its agencies, or EG&G Idaho, Inc. of the use of

a specific product for any purpose. , , , ; ; - : - . .

C-3
-

16.5 in.

Inlet

Bridge

wall

Indication #4 -

Wall thickness = 0.85 in.

- Gu

Indication #2

Wall thickness = 0.85 in.

Indication #3

Wall thickness = 0.70 in.

Outlet

L93 0406

Indication #1

Wall thickness = 0.78 in.

Figure C-1. Simplified view of 16-in. globe valve and erosion areas.

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elements. The finite element model is shown in its entirety in Figure C-2. For clarity, a view of

only the bonnet region elements, including the areas of maximum erosion (noted as indications

1-6), is shown in Figure C-3.

The inlet, outlet, and bonnet sections of the valve were modeled with circular cross sections.

The actual physical dimensions of these sections changed somewhat at different locations on the

valve, but exact dimensions were not available at all locations. Based on analytical experience,

this simplifying assumption was deemed appropriate to the overall goal of the analysis effort.

A second simplification was the omission of the valve plug guides. These guides are located

on the bonnet cylinder side walls and function to guide the plug in its vertical travel during

operation and provide lateral restraint for.the plug. The guides would have a stiffening effect on

the valve walls in their vicinity. Because of the curved surfaces blending the guides to the walls,

no significant additional stress would be expected in these areas. Thus, omission of the guides will

tend to make the'results error on the conservative side because the deflections, and thus the

stresses, that are calculated will be somewhat Iarger -than what would actually occur. Appendix D

discusses the analysis procedure and results in greater detail.

Operating Conditions and Imposed Loads

The following temperature and pressure conditions were used for the 16-in. globe valve:

TD = design pressure 375 psig' -

-POP = normal operating pressure = 160 psigb

TD = design temperature '' = 2250F -

Top = normal operating temperature '-150TF.

In actual operation, the'valve'would be subjected to a variety of thermal conditions, internal

pressures, and piping system structural loads. For the purposes of this scoping analysis, however,

we wanted to limit the number of loads that were considered: For this reason, the load cases

considered were limited to the following combinations:

P0 p + Fw t + M -

Ppdh + Fsmat + Mpe

.where

Font = valve plug seating force (lb)

Mpipc = piping system moments at the valve weld ends (in.-lb)

b. 160 psig is the normal operating discharge pressure; the rated discharge pressure is 195 psig.

, C-S
IrAlelt

0n

Figure C-2. I-DEAS finite element model of 16-in. globe valve with eroded walls.
Indication /- - I / t N

- ''l~civ- /\I k~-i~ndication 6

Indication 2 .; .C1\tn --

rdeication I

Figure C-3. Bonnet, region elements with the areas of ihaximum 'erosion" for 16 in. globe vialve model.
Pop +,F,., + Mpie

Ppdh + Fst + Mpipe

where

Fw, = valve plug seating force (lb)

Mpipr = piping system moments at the valve weld ends (in.-lb)

Ppdh = pump dead head pressure (psig).

Brunswick plant personnel provided the necessary pressures and seat force information and a

table of piping system moments for various thermal, weight, and seismic conditions. The piping

system moments were examined and a worst case combination of weight, thermal, and seismic

moments was constructed to conservatively represent the structural loads applied by the piping

system. The resulting loads were

Pop = 160 psig

Ppdh = 225 psig

Fseat = 92,464 lb

Mpipe = -58,824 in.-lb (about the X global axis)

= -48,504 in.-lb (about the Y global axis)

= 229,392 in.-lb (about the Z global axis).

Material Properties

The valve was manufactured using A216, grade WCB, cast carbon steel. This material has a

maximum carbon content of 0.30% and an ultimate tensile strength of 70.0 ksi. Table A-1 in the

ANSI B31.1 Code lists a maximum allowable stress of 14.0 ksi in the operating temperature range

of -20 to 650'F. Interpolation of data contained in Table C-1 in the Code indicated a Young's

modulus of 27.78 x 106 psi at the operating temperature of 150'F. Similarly, interpolation of data

from Table I-2.1 in Section III of the American Socicty of Mechanical Engineers (ASME) Code

gives a yield stress of 34.4 ksi at the operating temperature of 150'F.

Load Application

Throughout our analyses, the valve was assumed to be in the closed position. To represent

this situation, internal pressure (either normal operating or pump dead head) was applied normal

to the plate surfaces of the elements representing only that portion of the valve upstream of the

seat. Since the combination of piping bending moments, plug seating force applied to the seat

ring, and the differential pressure across the bridge structure elements should produce the

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maximum stress condition, the closed valve assumption is considered to be a conservative

approach to satisfying the analysis goals.'

The seating force applied by the plug to' the seat ring was applied as a line load 'around the

inner circumference of the seat ring. Appropriate force components were used to simulate the

actuial direction of the seat force as it 'is 'applied to the ring. In' te a6tual valve, the seat load

reaction path is through the bonnetf cylinder' walls and bonnet flange. -This load reaction was

simulated by forces applied around the tofp of the bonnet cylinder.'

The piping system moment loading was obtained from' Brunswick 'plant personnel as

described earlier. These moments were converted to the mro'del global coordinate system-and

applied -to the valve by using the'c`onstraint element feature found in the I-DEAS code. The'

moments were applied to the outlet (downstream) end of the valve while sufficient boundary'

conditions were applied at the upstream end to properly restrain the valve.

Boundary Conditions

The model boundary conditions consisted of restraints on the upstream end of the valve

node. The nodes in the plane of the weld end (upstream end of the valve) were restrained so

that rigid body rotation could not occur and so that all applied loads would be reacted properly.

Modeling of Eroded Areas

Information received from Brunswick plant personnel indicated that seven general areas of

erosion had been observed. In general, erosion occurred on the bonnet cylinder walls on both

side of the three plug guides. These eroded areas varied in size from approximately 1 to 5 in2.

The vertical locations varied from 4 to 6 in. above the upper surface of the seat ring. The

maximum erosion depths ranged from 0.375 to 0.75 in. deep. The original wall thickness in these

areas was not available; however, it was assumed to be at least equal to the 1.25-in. thickness

measured elsewhere. These conditions occurred in six of the erosion areas (termed Indication

1-6 in the Brunswick data). The eroded areas tapered down in thickness from the nominal wall

thickness (=1.25 in.) to the minimum value that remained (0.5-0.875 in.). This was simulated in

the finite model by beginning at the outside edge of an erosion area and "stepping down" the

element plate thicknesses to the minimum value selected for the analysis. Thus, each erosion area

was modeled with several groups of elements with different thicknesses to model the transition

into the minimum value of the eroded area. This technique not only provided reasonable

representation of the actual erosion patterns but also helped to avoid the possibility of numerical

instabilities that large relative differences in adjacent element properties might cause during

solution of the finite element model.

For the first of the two computer runs, a wall thickness value of 0.10 in. was used as the

minimum value for all six of the eroded areas described above. This value was used instead of the

actual minimum thickness remaining because we wanted to investigate the effect that a minimum

wall thickness condition would have on the stress levels. The Code calculation for minimum

thickness is discussed in further detail in Appendix D. The second computer run used a wall

thickness value of 0.010 in. as the minimum thickness to simulate thinning that might result in a

C-9
through-wall leak. Also, the erosion areas were enlarged, thus enabling the thicknesses to be

stepped down in additional increments. However, the erosion areas containing the elements with

the minimum thickness values remained the same for both computer runs.

The seventh erosion area occurred in the upper region of the bonnet cylinder, beginning

about 2 in. below the bonnet flange and extending vertically down near the tops of the plug

guides. This erosion also extended around most of the circumference of the upper bonnet

cylinder. The information received from the Brunswick plant personnel indicated that the erosion

in this region was fairly uniform and had removed about 0.125 in. of the wall material thickness.

This configuration was represented in the finite element model by simply reducing the thickness

of the appropriate elements in the bonnet cylinder by 0.125 in. The erosion in this area of the

model was held constant for both computer runs.

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Appendix D

Analysis of a 16-Inch Globe Valve

with Eroded WallsDescription of the Analysis Results

D-1
,.-Appendix D.

Analysis of a 16-lnch Globe Valve

with Eroded Walls

Description of the Analysis Results .... . , ..-. ,i

ANALYSIS

,

Analytical Approach

The intent of this analysis is to perform a "scoping analysis" to gage the effect of erosion on

the valve and to determine whether more detailed analysis is warranted. It is not intended to be a

detailed stress analysis that considers every possible load case. -,

As discussed in Appendix C, we developed the model using the I-DEAS software package

written by the Structural Dynamics Research Corporation.' This software includes drafting, solid

modeling, finite element solution, and post processing capabilities. Interfaces to other widely used

analysis* codes (e.g., ABAQUS) are also included within I-DEAS. We used the solid modeling

feature of the code to construct a representation of the complete valve that contained 5,026

nodes and 4,992 quadrilateral plate elements. The finite element model is shown in its entirety in

Figure C-2. For clarity, a view of only the seat ring and bridge structure elements is shown in

Figure C-3.

;, .; ' C fis ~4.4 .. * 4 . .. -, ,,;.1!',*! ;.t'''

This analysis used Version VI of the I-DEAS code -as installed on the Idaho National

Engineering Laboratory Engineering Department's computer aided engineering computer system.

This version of I-DEAS was subjected to the benchmark tests and quality control requirements

required by the Structural Dynamics Research Corporation before release of the code.. This

version of the code also is currently undergoing benchmark testing by Applied Mechanics Group

personnel. Benchmark runs using plate elements have been successfully completed and the

res~ults have been-verified fororrectness. An Applied Mechanics Group Computer Program -

' 'Configuration `Documentation form'is being completed for this version of I-DEAS and will be

available for reference if desired.. ;, ,. . . 4 . :- .4-

r . . . ,4 . .. . , , . .j, Ffjj ; .; t :' i,,,., ' } ' ; .

'Me" globe valve normally would be subject to dynamic loads during some phases of actual

operation; however, for this analysis,'we selected linear elastic static analysis as sufficient to satisfy

the requirements of a scoping analysis.. Therefore, we used the linear static solution option within

I-DEAS to calculate results discussed below. ' -

4~~~~~ 4 . . 4 ~ 4 ~ 4 ,.- .... . 4 4..

4 .4. 4, , . -, , : . . .

a. Mention of.specific products and/or manufacturers in this document implies neither endorsement or

preference, nor disapproval by the U.S. Government, anyof its agencies, or EG&G Idaho, Inc. of the use of

a specific product for any purpose.

D-3
Minimum Wall Thickness Considerations

The 16-in. globe valve originally qualified for operation under the 1973 edition of the ANSI

B31.1 Power Piping Code (hereinafter referred to as "the Code") and the ANSI B16.5 Standard,

'Steel Pipe Flanges and Flanged Fittings."

Equation (3) in Subparagraph 104.1.2 of the code can be used to calculate the recommended

minimum wall thickness for components subject to internal pressure. This equation is

= °PD +, A (D-1)

m 2(SE + Py)

where

P = design pressure (psig)

Do = outside diameter (in.)

SE = maximum allowable stress due to internal pressure and joint efficiency (psi)

y = coefficient from Table 104.1.2.A of the Code (0.40 for this case)

A = additional thickness as defined in Subparagraph 104.1.2 of the Code (in.).

When A is assumed to equal zero, a minimum wall thickness of 0.212 in. is required by the

Code to contain the design pressure of 375 psig. This calculation, based on Equation (D-1) (the

equation for pipe components), was used as an approximation to choose a conservative value for

the model. The minimum value, reported by Brunswick, was developed from valve design

equations.

Based on the calculations described above, the finite element model in the first computer

run used a value of 0.10 in. for those elements representing the areas of maximum erosion. This

value allowed us to investigate the stress levels when the walls had eroded almost to the minimum

thicknesses allowed by the Code. Since the object of this analysis was to determine whether

severe erosion would be likely to impair the valve's operability, we considered this thickness value

for the most highly eroded areas to be a reasonable starting point for the scoping analysis. The

second computer run used a value of 0.010 in. in the finite element model for those elements

representing the areas of maximum erosion. This value was chosen to enable the investigation of

the stress levels when the walls had suffered severe erosion that might lead to a through-wall leak.

Stress Analysis Results

We used the linear static solver option within the I-DEAS computer code to solve the finite

element model when subjected to the combined load sets discussed in Appendix C. The load

combinations are repeated below for clarity.

D-4
Pop ±~.Fseat +.Mpip

Ppdh +Fst + Mppewhere

P~p = normal operating pressure' (psig)

Feat = valve plug seating force (lb)

Mpipe = piping system moments at the valve weld ends (in.-lb)

Ppdh = pump dead head pressure (psig).

When the calculations were completed, we used the post' processing functions within

I-DEAS to review'the results.. Figure D-1 shows a representative color stress contour plot of the

eroded bonnet area of the valve. We reviewed plots such as this for a variety of stress

components and locations on the valve.

The post processing options within the I-DEAS computer code offer the capability to

calculate and display a variety of stresses such as principle stresses, maximum shear stress, and

Von Mises stress. Depending on the element type being used, results can also be plotted at

different surface locations through the element thicknesses. We chose the Von Mises stresses as

the values for comparison because research has shown that these stress results will provide a

highly accurate'indication of material behavior under static loading. We then compared the Von

Mises stresses to the material yield strength to predict the onset of nonlinear response.

We originally anticipated using several computer runs to investigate the behavior using

different values of wall thickness in the most highly eroded areas. However, the results using the

0.010 and 0.10-in. values provided sufficient information'to satisfy the goals of the scoping

analysis. The analytical results for the two load cases considered are included in Tables D-1 and

D-2.
Figure D-1. Representative stress contour plot of eroded bonnet region.
Table-D-1. Maximum stress results with tin = 0.10 in. J.'.

- ~ - ~-~~ -- ~- --~ - -~~~ ~-- ~~--Calculated Von Mises stressesa - .- - -

Yield

Element Load Load stress

Model area surface combination lb combination 2c (ksi)

Seat ring -Top 9.33 - 9.53 34.4

Middle 9.83 10.2 34.4

Bottom 10.5 11.1 34.4

Bridges Top 11.7 12.1 34.4

Middle 5.26 5.41 34.4

Bottom 8.08 8.34 34.4

Upper bonnet Top 4.87 4.95 34.4

(Area 7) Middle 3.02 3.05 , _.: 34.4

Bottom 4.09 4.06 34.4

Bonnet cylinder Top 16.6 17.1 34.4

(Areas 1-6) Middle 11.7 11.5 . 34.4

Bottom 17.5 18.0 34.4

a. Von Mises stresses were used because of their accuracy in representing material behavior under

static loading. :: ; i . . ..

b. Load combination 1 is the combination of operating pressure, valve seating force, and pipe end

moments. -- - :

c. Load combination 2 is the combination of pump dead head pressure, valve seating force, and pipe

end moments. i- * ' ;' . j 

D-7
Table D-2. Maximum stress results with tmin = 0.010 in.

Calculated Von Mises stressesa

Yield

Element Load Load stress

Model area surface combination lb combination 2c (ksi)

Seat ring Top 9.50 9.77 34.4

Middle 10.4 10.8 34.4

Bottom 11.3 11.9 34.4

Bridges Top 12.8 13.4 34.4

Middle 4.82 5.00 34.4

Bottom 9.87 10.2 34.4

Upper bonnet Top 5.38 5.44 34.4

(Area 7) Middle 3.25 3.20 34.4

Bottom 3.84 3.81 34.4

Bonnet cylinder Top 24.5 24.2 34.4

(Areas 16) Middle 25.0 24.8 34.4

Bottom 29.1 29.0 34.4

a. Von Mises stresses were used because of their accuracy in representing material behavior under

static loading.

b. Load combination I is the combination of operating pressure, valve seating force, and pipe end

moments.

c. Load combination 2 is the combination of pump dead head pressure, valve seating force, and pipe

end moments.

D-8