ML20197B036

Final Deficiency Rept (55(e)-86-20) Re MSIV Leakage Rates. Initially Reported on 860909.Actions Being Taken to Restore Valves to Acceptable Condition & Reduce Wearing Stress to Acceptable Level

ML20197B036
Person / Time
Site:	Nine Mile Point
Issue date:	10/20/1986
From:	Mangan C NIAGARA MOHAWK POWER CORP.
To:	Kane W NRC OFFICE OF INSPECTION & ENFORCEMENT (IE REGION I)
References
(55(E)-86-20), (NMP21-0918), (NMP21-918), NUDOCS 8610270451
Download: ML20197B036 (108)

Text

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, , M V NIAGARA RuMOHAWK' NIAGARA RAOHAWK POWER CORPORATION /300 ERIE BOULEVARo WEST, SYRACUSE N.Y.13202/ TELEPHONE (315) 474-1511 October 20, 1986 (NMP2L 0918)

Mr. W. Kane, Director U.S. Nuclear Regulatory Commission Region I Division of Reactor Projects 631 Park Avenue King of Prussia, PA 19406 Re: Nine Mile Point - Unit 2 Docket No. 50-410

Dear Mr. Kane:

Enclosed is a final report, in accordance with 10CFR50.55(e), for the problem concerning the Main Steam Isolation Valve leakage rates. This problem (55(e)-86-20) was reported via telecon to G. Meyer of your staff on September 9,1986 and an interim report was submitted on October 8,1986.

Very truly yours, N L C. Y. Mangan Senior Vice President CVM/ GAG /cla (1889H) xc: J. M. Taylor, Director of Inspection and Enforcement U.S. Nuclear Regulatory Commission Washington, DC 20555 W. Hz.ughey, NRC Project Manager W. Cook, NRC Resident Inspector Project File II'IlIII I2130SBS 4

e6 0270$Edg $$o8!hi N0103U-03Al3333 PDR

i e UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION In the Matter of )

Niagara Mohawk Power Corporation ) Docket No. 50-410 (Nine Mlle Point Unit 2) )

AFFIDAVIT C. V. Mangan , being duly sworn, states that he is Senior Vice President of Niagara Mohawk Power Corporation; that he is authorized on the part of said Corporation to sign and file with the Nuclear Regulatory Commission the documents attached hereto; and that all such documents are true and correct to the best of his knowledge, information and belief.

/o_/WAA M A

~

Sub cribed and sworn to before me, a Notar gblicinandfortheStateofNew York and County of OnRy/Ana , this day of Octobt't , 1986.

O PAiiatne <luitn l Notary Public in and for

! d/UrdMa. County, New York 6

My Comission expires:

DeusTWEAusTM d

1

% E FINAL REPORT 10CFR50.55 (e)

MSIV LEAKAGE NINE MILE POINT UNIT #2 NIAGARA MOHAWK POWER CORPORATION OCTOBER 1986

TABLE OF CONTENTS

1.0 INTRODUCTION

1.1 Objectives 1.2 Executive Summary 1.3 Safety Evaluation

2.0 BACKGROUND

2.1 valve Description 2.2 Valve Selection 2.3 Initial Qualification 2.4 Other Operating Experience and Corrective Action 3.0 PROBLEM DESCRIPTION 3.1 Problem Discovery 3.1.1 Test Method 3.1.2 Test Data 3.2 Description of Failure 4.0 EVALUATION AND ANALYSIS 4.1 Root Cause Analysis 4.1.1 Mechanical Analysis of Seat and Ball 2

' =

4.1.2 Stress Analysis 4.1.3 Metallurgical Evaluation 4.1.4 Root Cause 4.2 Description of Inplant Tests and Results 4.2.1 Purpose 4.2.2 Test Parameters 4.2.3 Test Combinations 4.2.4 Test Results 4.3 Other Testing 4.3.1 Laboratory Testing 4.3.2 Site Frictional Test 5.0 CORRECTIVE ACTION A'AD TECHNICAL JUSTIFICATION 5.1 Corrective Action 5.1.1 Recoated Balls 5.1.2 Modified Springs 5.2 Mechanical Analysis of Re-Coated Ball and Modified Springs 5.3 Stress Analysis 5.4 Evaluation and Analysis 5.4.1 Operating Conditions 5.4.2 Steam Effects 5.4.3 Temperature Effects 5.5 Evaluation of Test with Blended Balls 6.0 ADDITIONAL CONFIRMATORY TESTING 6.1 Leakage Testing 3

' O 6.1.1 Initial Type C" Testing 6.1.2 Confirmatory Type "C" Testing 6.1.3 Mid Cycle Test 6.2 Developmental Testing 6.2.1 Test Objectives 6.2.2 Test Organization i 6.2.3 Testing Format and Schedule 6.2.4 Initial Prototype Test Program 7.0 CONTINGENCY PLANNING 7.1 Leakage Control System 7.2 Y-Pattern Globs Valves

8.0 CONCLUSION

S 9.0 APPENDICES 9.1 Description of Materials of Ball Construction 9.2 MSIV Chronology 9.3 MSIV LLRT Procedure Summary 9.4 Ball Positioning (Thrust Washer) 9.5 Union Carbide Testing 9.6 Site Frictional Testing 4

'l.cl* INTRODUCTION During March 1986 the eight NMP2 Main Steam Isolacion Valves (MSIV) passed their formal " Type C" Local Leak Rate Tests (LLRT) conducted in accordance with the requirements of Appendix J of 10CFR50.

On September 2, 1986 NMPC conducted additional LLRT's on the MSIV's.

During August 1986 the valves were operated in excess of 100 times.

This was an attempt to correct actuator problems.

The September leak testing was initiated to provide data supporting the acceptability of the LLRT method used at NMP2 for this type of ball valve. During this testing it was discovered that all eight MSIV's had exceeded allowable leakage. Immediately following this discovery, an extensive series of inspections, analyses and tests were performed.

1.1 objectives This report presents the original bases for selecting these valves for this application, discusses the details of the leakage problem, presents a root cause analysis, and describes the corrective actions that have been taken including those tests performed to justify our resolution of the leakage problem. The justification for the corrective action is based on analytical and test results, together with additional planned confirmatory testing and contingency plans. This document is the final report required by 10CFR50.55(e) 5

addressing the MSIV leakage problem. This report supplements information contained within the Application for Schedular Exemption Related to Further Analysis of and Possible Modification to the Main Steam Isolation Valves submi t ted to the NRC on October 2, 19RA.

1.2 Executive Summary Niagara Mohawk believes that the problem resolution plan presented in this report includes sufficient testing and analyses to demonstrate that the MSIV's will remain leak-tight through the first operating-cycle. Furthermore, Niagara Mohawk is confident that the mechanism which caused the excessive valve leakage is understood. Specifically, it has been determined to be wearing stress (excessive contact stress combined with f riction) which causes localized delamination of the tungsten carbide ball coating. The removed material scratches the stellite seats, causing excessive leakage. Actions are being taken to restore the valves to an acceptable condition and to reduce the wearing stress to an acceptable level. On site cycle testing and analysis of operating conditions provide confidence that the coating delamination will not occur during the first onerating cycle.

i Analyses have been completed which demonstrate that normal and emergency operating conditions, including effects of temperature i

and valve closure under steam flow, do not add significantly to l

the condition causing delamination on the ball. Further, comparison 1

6

_.. - .- _ ___ , ._ , ~ ._____ _

• of calculated maximum bearing stress and as-tested tungsten carbide bearing strength shows that in situ forces should remain below those necessary to cause delamination of the coating.

Niagara Mohawk is committed to continued testing and contingency programs. The long-range testing programs provide further assurance that the root cause of previous failures are thoroughly confirmed and documented. If modifications are required to assure reliable service of the valves throughout plant operating life they will be completely developed and thoroughly tested for implementation during the first refueling outage. Also, design and procurement activities for the addition of a MSIV leakage control system will continue to proceed on an expedited basis should future testing or analysis show the need for such a measure. Finally, design and procurement planning for Y-pattern globe valves and their associated leakage control system remain in ptogress.

4 In conclusion, the current MSIV ball valves are safe for a minimum of one cycle of unit operation. Ongoing analyses and testing are being expedited to assure early identification of concerns so that any long term modifications can be implemented. Contingency programs are also being expedited to ensure they can be implemented if necessary.

1.3 Safety Evaluatior.

The quick-closing MSIVs function to isolate the reactor and containment 7

systems in the event of a break in a steam line outside the primary containment, a design basis loss of coolant accident (LOCA), or other events requiring main steam line or containment isolation.

In the case of a main steam line break, the isolation valves would terminate the blowdown of reactor coolant in 3 to 5 seconds thereby

. preventing an uncontrolled release of radioactivity from the reactor vessel to the environment.

Fesults of NRC staff standard plant analyses, which use conservative assumptions for considering the offsite consequences of a postulated design-basis LOCA coupled with uncontrolled leakage of the main steam isolation valves above technical specification limits, have i nd icated that the calculated deses would be in excess of 10 CFR 100 guidelines.

Even though the valves did not meet leakage specifications, they would close and would have termir.ated blowdown of the reactor vessel.

l Had the. leak rate through the main steam isolation valves not been corrected, a design-basis LOCA could result in offsite doses i in excess of 10 CFR 100 guidelines; or a main steam line break i could result in excessive doses in plant occupied areas.

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. . - - y . -. ._____ _ _ -._-___7_ , _ , . -m_ . - - _ . , _ _ . . . _ . , . - _ . _ _ . . - . -

' 2.0* BACKGROUND 2.1 Valve Description NMP2 is provided with two MSIV's on each of four main steam lines (see Figure 2-1). The valves are designed to provide isolation within 3 to 5 seconds during an emergency.

The MSIV's also provide redundant isolation between the reactor vessel and the main turbine generator, thereby permitting normal operation and maintenance of the steam plant systems while the reactor is at operating temperature and pressure. Valve operation under normal c o ndi't io n s takes approximately five (5) minutes to open and twenty (20) seconds to close. A more detailed discussion of normal closure is contained in Section 5.4.1.

The MSIV's are ball type valves, welded into a horizontal pipe run of each of the four main steam lines; one valve is close to the inside of the primary containment and the other is located just outside the containment.

I

! A complete description of the valves is detailed in Gulf & Western I

l Topical Report No. G&W FSD 2538 submitted to the NRC on January i

24, 1979. FSAR Figure 5.4-7 (attached to this section) shows a

! cutaway view of an MSIV. Each 24-inch reduced-port ( 21- i n ch) ball type valve has a full-ported ball with an integrally cast top and bottom trunnion. At rated steam flow through each valve I

l-

the pressure drop through a valve is calculated to be 1.2 psi.

The valve internals are top loaded into the valve which allows disassembly without removing the valve from the piping system.

The ball, when rotated within the body, is aligned in two roller-bearing assemblies (upper and lower) . Valve seal assemblies are fit into the valve body and held in contact with the ball by the force of multiple springs. The seat-to-body interface is sealed by a multiple ring packing seal that is compressed by the seal springs. The seal is produced by the spring-loaded force of the seat against the ball's surface. In a closed position the upstream seat's sealing force is aided by system pressure acting against the projected annular area of the seat. A multiple ring packed stuffing box seals each ball trunnion against the body. The bonnet closure is sealed by a metallic, pressure-type seal ring. The ball is designed with a vent hole between the flow hole of the ball and the bottom of the ball. During operation with the valve open, this equalizes pressure between the body cavity and the steam flow area, thereby reducing seat differential f pressure loading. Details of the materials of ball construction are contained in Appendix 9.1.

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I I

10 l

2.2 Valve Selection This ball valve was selected as the NMP2 MSIV based on the significant benefits compared with the standard Y-pattern globe valve, and based on the extensive testing and analysis of smaller valves of the same design that demonstrated their suitability for the proposed application. At the time of purchase these benefits were determired to be long range enhancements to the unit's overall availability since Y-pattern valves were considered a maintenance problem. An MSIV chronology is contained in Appendix 9.2 The main advantages of the ball valve are summarized below:

o Each valve contains two sealing surfaces.

o operating characteristics which result in less wear and stress on valve internals and thus less maintenance.

o Ease of disassembly and maintenance will result in less plant down time and radiation exposure, o Low pressure drop providing optimum steam conditions i

for power generation.

o Simplified piping, support and restraint configurations.

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o Low operating (rotational) velocity even during emergency i

11

closure (5 rpm or 1 ft/sec maximum rotational velocity).

o Sealing components are not used in the deceleration and stopping of the valve in the fully closed position.

o Seating surface maintenance can be accomplished in remote areas thereby minimizing personnel exposure.

2.3 Initial Qualification Energy Products Group (EPG) Fluid Systems Division of Gulf and Western Manufacturing Company was the original supplier of the MSIV's and publisher of the Topical Report. Most of the initial qualification testing was performed by this organization, and is described in their Topical Report. A summary of this testing is provided in the following paragraphs.

Dynamic qualification of the valve assembly was demonstrated through a combination of testing and analysis which evaluated all parameters affecting its function. These parameters included flow, temperature and combined seismic and hydrodynamic loads, in both normal and accident conditions.

The ball valve and actuator design was subjected to extensive testing at the EPG facility in Warwick, Rhode Island. In August 1976 a prototype 8 inch valve and actuator were tested with 1300 psig saturated steam. The valve was in the open position with 12

a downstream block valve initially closed. The block valve was then opened to allow steam flow through the 8 inch. ball valve while venting to the atmosphere. The ball valve was then closed against the saturated steam flow of approximately 1.8 x ig6 LB/HR.

Testing was repeated several times against varying flow rates.

During the test, the valve body was subjected to torsional, axial, and bending loads that approximated plastic deformation limits of a 10 inch pipe. Data were taken regarding the stresses incurred in the body and actuatot as well as determining the closure torque requirements.

Subsequent to the flow test at Wyle Laboratories in Huntsville, Alabama, the valve was shipped back to the manufacturer for leak rate testing. The leakage was found to be excessive. Upon disassembly it was noted that the ball had sustained galling on its upstream face. This was determine.1 to be caused by debris being carried in the steam and impacting the ball during the closing cycle.

Substantiation for this conclusion was the considerable mill scale and oxidation (rust) found in the body cavity. For this test, both the 8 inch ball and seats were overlayed with Stellite which further contributed to the galling effect once the coating was 1 damaged on the ball face.

The ball and seats were reworked and the reassembled valve exhibited a very low leakage rate.

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I This testing resulted in several key conclusions:

o Flow has little effect on required closure torque, o The ball required a more durable coating and properly matched seat material.

o Mathematical models could be developed for valve body and actuator stress analysis.

In addition to the Wyle Laboratories tests, a quench test was conducted on a 3 inch ball at EPG. This demonstrated that the tungsten carbide coating on the ball is suitable for the thermal transients anticipated during all modes of plant operation. The test was performed using a ball overlayed with stellite and coated with tungsten carbide. The ball was heated slowly to 4750 F and then quenched in water. The ball was visually inspected with -

no relevant indications found.

2.4 Other Operating Experience and Corrective Actions Beaver Valley has three ball valves originally manufactured by EPG and which are identical to those at NMP2. This unit has not operated their valves enough at this time to have useful information about leakage related problems.

The Swiss utility Kern Kraftwerk Leibstadt (KKL) has four of these 14

• valves. These volves are the semo size, material and provided by the same vendor as those at NMP2. These valves are normally open and used only for low pressure containment isolation. Each valve is the third isolation valve in the main steam line; the first two valves are Y-pattern globe type valves.

KKL identified corrosion (rusting) of the carbon valve body under the spool seat packing. A noncorrosive material was overlaid in the packing area. KKL also uses a low salt (chloride) Titan packing to avoid spool bore corrosion. Based on the Swiss experience, NMPC developed and implemented a repair program to correct this problem. Incoloy 625 was applied to the valve body adjacent to the spool packing se'al. This enhancement was completed in 1985.

MFL has also identified wearing of the tungsten carbide ball coating and measured leakage rates greater than their allowable amount.

The FKL valves have experienced hot steam conditions in the open position; they have not been closed against design steam flow. KKL has Jixty-115 pound springs equally spaced on their seal ring.

The NMP7 valves will have a modified spring arrangement.

Uoon inspection of their valves, KKL observed radial cracks on one of the spool seats. Following discussions with KKL personnel it was determined that this was an isolated incident relating to a manufacturing defect. These radial cracks were repaired and the spool seat was returned to service.

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• Following rework to remove seat scratches, the NMP2 seal rings will be thoroughly inspected by appropriate NDE methods.

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3.G ' PROBLEM DESCRIPTION 1.1 Problem Discovery 3.1.1 Test Method The NMP2 MSIV's are allowed a maximum leakage of 6 standard cubic feet of air per hour (SCFH) per valve using a test differential pressure of 40 psi. This leakage can be measured in either of two ways; in the normal leakage flow direction through the valve (inboard to outboard) or between the valve's upstream and downstream seats by p r e s s u r i z i ng the MSIV body cavity. Appendix 9.3 provides additional details on MSIV Local Leak Rate Test (LLRT).

Recent leakage test results have confirmed that the between seat method is more conservative in determining valve leakage conditions. This method directly tests both seats at the required pressure. Through the valve testing allows a substantial pressure drop through the upstream seat, lowering the driving pressure for downstream seat leakage, thereby resulting in lower values for total measured leakage. Figure 3-1 demonstrates this condition.

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3.1.2 Test Data Between seat " Type C" tests were performed for informational purposes after valve assembly in April 1985. These tests were conducted to verify proper installation of essential sealing components. Formal, between seats,

" Type C", leak tests were conducted in March 1986.

An estimated 15-20 cycles of the valves occured between April 1985 and March 1986. Most of these cycles used manual cycling which takes approximately 5' minutes to stroke the valve fully open or closed. The last 2-3 cycles prior to formal testing were with the actuator.

The valve closure speed during these strokes was within the required 3-5 seconds. All data demonstrated leakage well below the 6 SCFH allowable value. Table 3-1 (attached) provides a sum ary of these two tests.

On August 28, 1986 the NRC questioned the test method since it did not simulate the in situ flow direction.

The NMPC position as previously stated was that the between seat test is conservative. On September 2, i 1986 in an effort to demonstrate this position, NMPC conducted a leakage test between two valves. This test indicated excessive seat leakage. Since our previous Type "C" tests were between the seats, we duplicated our testing technique and recorded excessive leakage i

j between the seats of the valve. All valves failed the 9

20

, , 6 SCFH ccceptanca critoria. Tablo 3-2 (attached) indicates the actual leakage values.

From March 1986 to September 1986 all eight valves were stroked during actuator testing. It is estimated that each valve was cycled over 100 times as part of this testing. The tests' results dictated that visual inspection of the ball and seats was required. The first ball was removed for inspection on September 8, 1986.

3.2 Description of Failure All MSIV's were disassembled and inspected. Visual inspection revealed that the eight valves exhibited similar conditions, markings and defects. Figures 3-2 and 3-3 show typical damage locations.

o There were scarred areas on the hard surfaced ball where patches of the tungsten carbide were missing. The removed areas were near the ball open position to the right of the ball vertical centerline at the top and bottom of the flow hole. Approximately 8 to 10 mils of material had been removed.

There was also some galling in the scarred areas.

o There was some evidence of galling on each ball's machined bottom surface which bears on a 304 Stainless Steel thrust f

washer. Galling of this surface is addressed in Appendix 9.4.

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4 There were some scratches in the direction of ball rotation o

and there was material on the ball in areas where the seat would seal against the ball. Laboratory analysis of the material on the ball and in the valve body established that this was common rust.

o The Stellite seal ring showed signs of scratching due to the cutting by tungsten carbide particles.

The major portion of damage was evident in areas near the valve open position. The closed sealing surfaces were in relatively

'However, the stellite seal ring scratching was

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good condition.

sufficient to degrade the valves' sealing capabilities.

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INFORMATIONAL 1

FORMAL i TYPE "C" TYPE "C"

APRIL, 1985 M AR C H, 1986 VALVE LEAKAGE (SCFH) LEAKAGE (SCFH) i 6A 0.89 1.09 l

i 68 1.37 0.54 1

1

] 6C 0.321 0.158 I

i 6D 0.99 0.215 1

j 7A 0.34 0.084 l

! 78 2.778 1.183 7C 0.798 0.199 i 7D 0.306 0.088 i

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INITI AL l TYPE "C"

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awusivuo48 T ABLE 3-1

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! SEPTEMBER 2,1986 i'

VALVE LEAKAGE (SCFH) 6A 22 68 40.2 ,

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! 6D > 42 l . 7A 30.3 I

! l l 78 > 42 7C 23.6 7D 16.7 SEPTEMBER, 1986 l TYPE "C" l TEST i

NuusivuG49 TABLE 3-2

4.0 EVALUATION AND ANALYSIS 4.1 Root cause Analysis As discussed in Section 3.2, all eight balls showed consistent damage oatterns of tungsten carbide delamination at the top and bottom of the flow holes. The location of the damage matched the position of the seal ring seat on the ball surface near the full open position. This damage occurred f rom contact pressure, friction and a considerable number of operating cycles.

An analysis of the mechanical interaction of the seat and ball follows, with a st'ress analysis, and metallurgical evaluation of the damage.

4 .1.1, Mechanical Analysis of the Seat and Ball Figures A-1 and 4-2 show the details of the seal assembly and ball with the valve in the full closed position.

The seat is held against the ball with sixty-115 pound springs, placed uniformly around the circumference, which exert a total force of 6900 pounds on the ball.

The seal assembly is free to move in the spool bore with a diametrical clearance of approximately 80 mils.

Because of the relationship between the seat outer diameter and the flow hole inner diameter, when the valve is 28

near the full open position, the seat approaches a point of instability. As shown in Figure 4-3, at this point, almost half of the spring loaded seat is not in contact with the ball. This load combined with rotation of the seal ring due to f rictional moment causes the seat to pivot near the edge of the flow hole and " rock" the seat area off the ball. When " rocking" occurs, the entire 6999 pound spring load is concentrated at the pivot points near the edge of the ball flow hole, which results in a high contact stress on the ball.

An analysis of the mechanical forces on the seat and ball while opening and closing the valve follows.

valve closing When the valve operates from the open to closed position, rotation of the ball is clockwise as viewed from the top of the valve. Referring to Figure 4-4, there is a friction load on the slightly larger than one half area of the seat in contact with the ball. This friction l load causes the seat ring to tend to pivot in the spool bore around the seat packing in a counter clockwise direction as shown in Figure 4-4. The counter clockwise rotation of the seat ring due to friction assists il

! keeping the seat in contact with the ball while closing i

j 29 l

1 l

L

the valve. Therefore, forces do not exist to cause instability or " rocking" of the seat area off the ball while closing the valve.

Valve Opening When the valve operates from the closed to open position, rotation of the ball is counter clockwise as viewed from the top, and the friction forces are reversed. As shown in Figure 4-5, the friction load on the seat causes the seat ring to tend to rotate in a clockwise direction.

This rotstion of the seat "into the hole" could result in instability or " rocking" of the seat area off the ball while opening the valve.

Mathematical Model Figure 4-6 shows a simplified mathematical model of the seat ring loaded uniformly by the seat springs.

A calculation has been performed which determines the coefficient of friction required to cause incipient

" rocking" of the seat while the valve is opening to be 0.46.

Appendix 9.5 provides a summary of test data for the coefficient of friction between the tungsten carbide i .

1 30 i

(

i l

ball and the stellite seat. This testing was performed l l

by Union Carbide and showed that the coefficient of friction between the stellite and tungsten carbide to be between 9.3 to 0.5. This substantiates the calculated value of 9.46.

4.1.2 Stress Analysis From the preceding section, it was shown that the seat area can " rock" off the ball during valve opening.

This section provides an estimate of the resulting contact stress on the ball.

When " rocking" occurs, the full spring load of 6999 pounds is concentrated at two relatively small areas near the edge of the ball hole. Estimating the contact area between the seat and ball of approximately 9.955 square inches, the contact stress is estimated as shown below:

Stress = 6999 lbs. x 0.74 = 46,499 psi 2x (9.955) sq. in.

This concentrated stress is illustrated in Figure 4-7.

31

• 4.1. 3 Metallurgical Evaluation Evaluation of tests performed both at Union Carbide and at the NMP2 Site further substantiates the tungsten carbide failure mode. From this testing it has been concluded that the tungsten carbide coating failed due to high localized wearing stress and not due to improper coating application.

4.1.4 Root Cause The root cause of the tungsten carbide coating failure and resulting valve leakage has been determined to occur in the following sequence. This sequence is predicated on the original spring arrangement.

o Friction between the ball and the seat while the ball rotates from the closed to open position causes an overturning moment on the portion of the seat remaining in contact with the ball. This moment tends to rock the seat.

o The point at which the rocking starts to occur causes the area of the seat in contact with the ball to become very small, yielding high localized stresses on the tungsten carbide ball coating.

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o The high localized stress combined with friction caused delamination and removal of the tungsten carbide coating, o The removed coating becomes trapped between the seat and ball causing the removal of additional coating and scratching of the seats.

o The scratched seats do not seal properly when the valve is in the closed position thus failing the leakage testing.

A discussion of coating failures that occurred during the testing program is included in the following section -

(Section 5) .

33

a

'4.7 *DSCcription of Inglant Tootn and Docultn 4.7.1 Purpose Mased upon preliminary evaluations of the valve failure, a test plan was devised. The testing plan included a reasonable number of valve cycles to envelop one plant operating cycle, and included sufficient leakage testing to correlate between seat and through valve Type "C" leakage tests.

1 Three tests using different ball and seal spring combinations were run. Variables in each of the tests were carefully controlled. The number of cycles, the type of leakage testing i

and the hydraulic actuator configuration were held as constants.

4.7.2 Test Parameters i o A preliminary root cause analysis had determined that the seat area rocked off the ball. A mathematical model indicated that the seal ring spring forces had to be <

revised such that counter balancing would not allow-i the seal ring to rock on the ball. Figure 4-8 shows the modified spring arrangement.

o The number of valve cycles was established to be 75 (75 openings and 75 closings strokes). This is the number of cycles conservatively estimated to the first i

34

._- ,________...-..m_ - - _ . _ - _ . _ - _ _ _ _ _ , _ _ _ _ . _ - _ _ _ . . ~ _ .

m ___________,._m. _m,.c, , - _ . _ . . , . , . . . , . _ _ _ , ,

. . rofuoling of tho plant, approximatoly 30 months from initial fuel loading. Table 4-1 provide 1 a basis to this number of cycles, o The valves were cycled a specified number of times followed by a Type "C" test. Table 4-2 (attached) shows the stroke and leakage test requirements for the test plan.

These requirements ensure suf ficient through valve testing would be performed to establish a correlation between the two test methods.

o For the purposes of this testing, the MSIV actuator was tempo ~rarily configured to trip the valve closed in 3-5 seconds by venting the hydraulic cylinder to the hydraulic system tank. This venting was done through two solenoid operated valves. This is the configuration that will be used on the MSIV's when the final hydraulic actuator modifications are complete, thereby simulating operating conditions. The actuator modifications will be addressed in a separate report.

d.7.1 Test Combinations 4.2.1.1 The first combination consisted of:

o New Ball (this was a spare ball that had never been installed) 35

o Modified Springs section 5.1.2 discusses the modified spring arrangement, also see Figure 4-8 4.7.3.2 The second combination of ball and seat spring configuration consisted of:

o Blended Ball - (see Figure 4-9)

This is a ball on which the damaged area has been uniformly removed and edges blended into the surrounding areas. The blended areas were approximately 8-10 mils lower than the remaining tungsten carbide coating. Blending was? selected as an option to quickly make available test balls with good closed seal s u r f a.ce s . As discussed in Section 3.2, the damaged area of the ball was located in the open seating surface. In both blended balls tested the mating seal rings were lapped to the ball in the closed position to assure initial leak integrity.

o original Spring Configuration 36

i 4.2.3.3 The third combination consisted of:

o Blended Ball o Modified Spring Configuration 4.2.4 Test Results o Leakage Table 4-3 (attached) presents the leakage test summary for the three test configurations.

In summary, test results show the blended balls behave in a similiar fashion with either an original or modified spring configuration. The modified spring configuration when used in conjunction with a new ball results in no damage to the ball or seats.

o Visual Examination After testing was completed the three test valves were disassembled and inspected. Technical Consultants from the valve manufacturer (Crosby) and tungsten carbide coating applicator (Union Carbide) were part of the inspection team.

37

New Ball, Modified Spring Configuration -

There was no evidence of ball coating degradation due to seat rocking. The only indications were slight signs of polishing from the interaction of ball and seat.

These indications appeared as lines on the ball. The consultants concluded and site testing confirmed that this ball would have been capable of acceptable performance with more than 75 cycles. The leakage data also confirms the sealing integrity and continued capability.

The surface of the ball has discolored rings in both closed s' eating areas due to the bluing used to check contact area. Polished bands are located above and below these discolored areas, due to lapping of the ball to the seat. A slight difference in surface texture can be felt across these areas due to the polishing effect. The seating areas around the centerbore show some evidence of wear. The area contains gray smears and black streaks. This type of wear became evident in both the NMP2 and Union Carbide tests. This wear was present at the lower loadings during the tests and did not propagate into coating failure.

A visual inspection of the seats did not reveal any damage due to the 75 strokes. The surface finish meets vendor surface requirements. Light scratches appear 38

at different locations around the seat and also some polishing is evident due to lapping operation. The scratches do not appear to be of significant depth since they did not effect the valve leak rate.

Blended Ball, original Spring configuration -

The tungsten carbide coating adjoining the blended area had been removed. Except for location and extent, the damage appeared consistent with that observed on the original eight balls. The seats had trapped some of the removed material and the damaged area extended near the closed seating area of the ball.

Blended Ball, Modified Spring Configuration -

The damage was slightly less than on the other blended ball. The revised spring configuration apparently helped to stabilize the seat, otherwise the tungsten carbide removed area was to the right (facing ball) of the original removal area.

l I

1 i

39

4.3 Other Testing l l

t 4.3.1 Laboratory Testing In an effort to determine material interface characteristics Union Carbide performed wear tests between the tungsten carbide and Stellite #6 materials. Thermal cracking tests were also performed. Appendix 9.5 summarizes this testing. A copy of the Union Carbide formal test report will be submitted to the NRC under separate cover by October 24, 1986.

4.3.2 Site Frictional Tests i

Testing was performed on site to duplicate the mode of tungsten carbide failure experienced. The testing was performed on the blended ball (with modified spring pack) that had been -

I removed. The testing was able to demonstrate a failure of j the tung:: ten carbide at a high contact stress with friction.

Appendix 9.6 provides test results. A copy of the on-Site i

, formal test report will be submitted to the NRC under separate cover by October 24, 1986.

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I VALVE ASSEMBLY STROKES 5 PRE OPERATIONAL TESTING 5 POWER ASCENSION TESTING 5 SURVEILLANCE TESTING 30 MAINTENANCE CLOSURES 5 i

] CONTINGENCY 25

.l 1 TOTAL 75 l

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ACCUMULATED LEAKAGE RATE, SCFH LEAKAGE RATE, SCFH LEAKAGE RATE, SCFH-

. MSIV OPENING NEW BALL BLENDED BALL BLENDED BALL CYCLES MODIFIED SPRINGS ORIGINAL SPRINGS MODIFIED SPRINGS 4

BS

• TS** BS

• TS** BS

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15 4.9 2.9 51 5.5 59 -

25 4.7 -

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0 j 45 4.2 - - - -

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l NOTES:

• BS = BETWEEN SEATS

• • TS = THROUGH SEATS TEST LEAKAGE

• *

• THROUGH SE AT TESTING NOT COMPLETED DUE TO POOR INITIAL TYPE "C" BETWEEN SEAT TESTING.

SUMMARY

wuusivuos2 T ABLE 4-3

5.9* CORRECTIVE ACTION AND TECHNICAL JUSTIFICATION 5.1 Corrective Action 5.1.1 Recoated Balls Analytical and test results confirmed that blending of damaged areas on the balls would not solve the problem.

Recoating is required to provide balls in the same condition as the ball used in the test configuration. The recoating process involves the following:

o Removal by hand chipping of the old tungsten carbide coating, o Machine grinding of the ball to provide a uniform new surface for recoating.

o Local weld repair of Haynes 25 and/or base material and regrinding as required.

o Application of the tungsten carbide.

o Machine grinding to proper tolerance and finish requirements.

o Lap mating seats to new coating.

56

The above steps assure that each ball is returned to originally specified requirements identical to the ball used in testing. A metallurgical comparison of a new ball to a recoated ball is shown in Table 5-1.

5.1.2 Modified Springs The modification for reducing the concentrated load to the ball surface consists of removing four 115 pound springs on the left side of the seal ring (f acing ball) and adding eight 57.5 pound springs on the right side.

The net sealing forces remain virtually constant.

5.2 Mechanical Analysis of Re-Coated Ball and Modified Springs The modified spring pack, as described in Section 5.1.2, will prevent the seat from " rocking" off the ball and will, therefore, eliminate the high contact stress which damaged the tungsten carbide coating.

As shown in Figure 5-1, the modified spring pack provides a counter clockwise moment on the seat ring which counteracts the clock-wise moment produced by friction when the valve is opening. A mechanical analysis similar to the analysis dicussed in Section 4.1.1 was performed for the modified spring pack. The analysis demonstrated that the coefficient of friction that causes the seal ring to rock on the ball is 0.73. Since the maximum coefficient of friction at expected loads is 0.3 to 0.4, the seal ring will not rock.

57

Figure 5-2 shows a simplified mathematical model of the seat ring loaded by the modified spring pack.

5.3 Stress Analysis Sinca the modified springs will prevent the seat from rocking, the contact area of the seat on the ball is considerably larger.

Calculations show a reduction of an order of magnitude compared to the original spring pack. The calculated bearing stress is approximately 2700 psi.

Conclusion o The modified spring pack will prevent the seat from

" rocking" off the seat area of the ball.

o The seat spring loading will be distributed over a relatively large contact area on the ball.

o The estimated contact stress is below the allowable value.

o Damage to the tungsten carbide coating will be precluded.

o Site testing has demonstrated that the tungsten carbide coating and seal ring are not damaged.

58

5.4 EVALUATION AND ANALYSIS 5.4.1 Operating Conditions 5.4.1.1 Normal Opening and Closure During normal Plant Start-up, the MSIV's are opened with no differential pressure across the valves. valve opening time is approximately 5 minutes. It is permissible to open the MSIV's at rated temperature and pressure.

The pressure will equalize across the valves prior to reaching the critical valve open position.

During normal plant shutdown, the MSIV's remain open until reactor pressure reaches approximately 150 psi at which point they are closed and the plant put in shutdown cooling mode. Valve closure time is approximately

~

20 seconds in this mode.

During monthly surveillance testing the valves are partially stroked closed. This test is performed with no dif ferential t

I pressure.

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5.4.1.2 Abnormal Conditions o Upset and Emergency Closure of the MSIV's is assumed to occur with the design flowrate through the valve at 100% (3.57 x ig6 lb/hr) and Reactor Pressure Vessel (RPV) pressure and temperature equal to 1965 psia and 5520 F. Valve closure occurs within 3-5 seconds, o Faulted The ' faulted operating condition for the MSIV's is defined by a rupture of the main steam line downstream of the MSIV's. Under this condition, the design flowrate through the valves is limited to 200% of maximum normal flow (7.14 x 106 lb/hr) by flow restrictors in the main steam line. The valves will close against this maximum flowrate with the RPV pressure and temperature equal to 1965 psia and 5520 F. Valve closure occurs within 3-5 seconds.

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60

5.4.2 Stecm EffGcts Figures 5-3, 5-4, and 5-5 show three positions in the sequence of valve closure or opening against steam:

1. Valve fully open

2. Valve partially closed (or open)

3. Valve fully closed i Provided below is a summary of an evaluation of the effects i

of drag and pressure forces on the seats due to steam and resultant contact stresses on the ball.

As seen in Figures 5-3, 5-4, and 5-5, there are two seats, an upstream and a downstream seat. These seats are affected differently by steam flow and pressure.

The following paragraphs discuss the effect of steam flow and pressure on these seats during valve opening and closing.

Steam Flow The flow on the upstream seat produces an insigificant drag load (approximately 155 pounds) due to skin friction since the seat is not protruding into the flow stream. The flow on the downstream seat, on the other hand, produces a direct impingement load on the downstream seat when the valve is 61

partially closed. It should be noted that the impingement force of 20.6 kips on the downstream seat tends to rotate the seal ring in a manner that would keep the seat on the ball and would assist in preventing the seat " rocking" that was previously discussed. This results in a contact stress of 600 psi.

Steam Pressure When the valve is normally open, the differential pressure across both the upstream and downstream seats is zero because of a vent in the ball which equalizes the cavity pressure, as shown in figure 5-6. When the valve is closed against steam flow and pressure, the differential pressure across both seats will increase. The pressure on the upstream seat will cause the seat contact stress against the ball to increase.

However, the differential pressure across the downstream seat will reduce the contact stress against the ball and the seat will actually move away f rom the ball at a dif ferential pressure of just over 130 psi. When the valve is opening or closing there is a non-uniform differential pressure across the upstream seat which will tend to rotate the seat ring in a manner that would assist in preventing the seat from " rocking".

62

Valve Closing It was concluded in Section 4.1.1, that seat " rocking" will not occur during valve closure. This conclusion is also applicable for valve closure against steam flow and pressure.

Seat " rocking" is caused by a moment on the seal ring due to friction, which tends to rotate the seat "into the hole",

that is, in a counter clockwise direction when viewed from the top. Drag and pressure forces on the seat due to steam will change the magnitude of seat forces on the ball and, consequently, the magnitude of the friction on the seat ring.

However, the direction of the frictional force and resulting moment on the seal ring is governed by the relative motion between the seat and ball. While closing the valve, with or without steam, the ball is rotated in a clockwise direction, as viewed from the top, producing a counter clockwise rotation of the seat. Therefore, the seat will not " rock".

Valve Opening As discussed in Section 5.4.1, the valve may be opened against steam pressure. For this case, the pressure will equalize l quickly af ter the ball begins to rotate open, and the dif ferential pressure across the seat will be zero before the seat reaches the postion required to initiate " rocking". In addition, as discussed previously, there will be a non-uniform dif ferential pressure across the seats which assists in preventing " rocking".

l l

! 63

, - - ~

• Therefore, the analysis and' testing of the modified spring pack envelop conditions which will exist while opening the valve during operation.

Conclusion:

Steam Effects o During valve opening and closure, the differential seat pressure assists in preventing seat " rocking" o Drag loads due to steam flow are small on upstream seats compared to seat spring loads o Drag loads oh~the downstream seat assist in preventing seat

" rocking" o The analysis and 75 cycle test with the modified spring pack are representative and applicable to full steam conditions, including drag and differential pressure across the seats.

64

5.[.3 Temperature Effects The following discussion will address the suitability of the internals for the thermal environment and will focus on the effects of the difference in coefficients of thermal expansion between the ball and its coating and the effect of differential expansion between valve componets.

o Ball o Kern Kraftwerk Leibstadt The exposure of these valves at reactor operating J

temperatures has not revealed evidence of cracking of the coating due to any temperature related phenom-ena. These valves are coated by the same process as the NMP2 balls.

o 3-inch Ball Test 1989 i

A quench test was performed on a 3 inch diameter ball in 1989 by Gulf and Western Fluid Systems i Division. This test was performed to show the acceptability of the tungsten carbide coating application method. The method tested was the detonation gun process. This method employs high particle velocity to provide a coating with low i

- 65

o O porosity compared to other application methods.

Coating temperatures remain cooler, generally under 3000F, than by other methods.

l o Thermal Testing of Coating on MSIV Ball To provide additional assurance that the coating of the balls will not crack at operating temperature a test was conducted on the ball from Valve No. 7D which had been removed for recoating due to damage described in Section 3.2.

~

Tes t! ' De scr iption : The ball was encased in thermal blankets and heated on the inside of the 21 inch diameter hole of the valve. It was heated at a rate which did not exceed 1300F per hour, up to a temperature of 5000F, while performing a series of tests to check the effect of temperature on ball coating. The ball was then cooled at a controlled rate which did not exceed 1000F per hour. Temperature was monitored at four locations both on the bore and by a calibrated surf ace pyrometer on the surface.

Inspection of the ball surf ace revealed no indication of coating failure due to the heating process.

l 66

o Other Valve Components o Seat Springs The seat springs are fabricated from Hastelloy C276 a trademarked alloy of the Cabot Corp. This alloy was chosen for its good high temperature creep properties and no degradation of the spring force is expected.

i l

67

5.$ EVALUATION OF TESTS WITH BLENDED BALL As discussed in Section 4.2.4, two tests were performed on balls with a 6 inch portion of the tungsten carbide coating " blended" that is, ground in the area of the observed damage. One test was run with the original seat springs and the other with the modified spring pack.

Although both tests failed, due to further damage to the balls and subsequent excessive leakage rates, the tests further strengthen the arguments presented in the root cause evaluation and corrective action. Also, both tests showed subtle differences which are expected considerihg the failure mechanism.

A blended ball actually causes the seat ring to be unstable as the ball is rotated either open or closed. The instability is strictly due to the fact that more than one half of the spring loaded seat ring is not supported on the seat area of the ball,

regardless of the friction effect discussed in Section 4.1. The instability exists for both the test with the original spring pack and the test run with the modified spring pack.

The very rapid deterioration of the coating for both tests was due to the instability occuring during both the opening and closing of the valve. From observations of the two balls, and as discussed in Section 4.2.4, the ball with the original spring pack sustained more damage than the ball with the modified spring pack. In ef fect, 68

. , th3 modificd seat springs attempted to counteract the instability, and although the damage was less severe than the ball with the uniform springs, the countering moment was not large enough to overcome the instability.

69

BALL I

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ATTRIBUTE NEW BALL RECOATED BALL '

i BAi L MATERIAL 316 SS C ASTING ORIGINAL CASTING USED EXTENSIVE WELD REPAIRS MINOR ADDITIONAL WELD REPAIRS SOL'N ANNE AL AFTER REPAIRS NO SOL'N ANNEAL AFTER REPAIRS 4

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HARDFACING HAYNES 25 BY SAW ORIGINAL DEPOSIT USED

! MATERIAL WELD REPAIRS BY GT AW WELD REPAIRS BY GT AW NO HEAT TREATMENT NO HEAT TREATMENT 1  ; .080 TO .100 FINAL THICKNESS .076 TO .097 FINAL THICKNESS FINAL SURFACE GROUND FINAL SURFACE GROUND l

FINAL PT FINAL PT j TUNGSTEN CARBIDE APPLIED BY UNION C ARBlDE APPLIED BY UNION C ARBIDE l

COATING MATERIAL SURFACE PREP BY BL ASTING SURFACE PREP BY BL ASTING APPLIED BY D-GUN APPLIED BY D-GUN O.010 MIN THICKNESS 0.010 MIN THlCKNESS FINAL SURF ACE GROUND FINAL SURF ACE GROUND LAPPED TO SE AT IN SHOP LAPPED TO SEAT IN SHOP l

LAPPED TO SE AT IN FIELD LAPPED TO SEAT IN FIELD PSMMSiv MP8 6 TABLE 5-1

~

6.d ADDITIONAL CONFIRMATORY TESTING 6.1 Leakage Testing 6.1.1 Initial Type "C" Testing Post assembly verification of the valve requires a limited number of cycles (1 to 5 cycles) in order to properly align the ball and seats. Local Leak Rate Tests (LLRT)

(between-the-seats method) will be performed subsequent to each cycle as a proof test for proper alignment.

The final LLRT will serve as the valve's formal LLRT (Type "C E )' in accordance with 10CFR50 Appendix J.

In preparation for fuel loading, the above procedure will be performed on one valve in each of the four steam lines. However, the valve actuators used to stroke tnese valves will not yet be modified with the revised hydraulic system. Therefore when these valves' actuators are modified and the valves are then stroked for timing and limit switch adjustment, another LLRT will be performed to ensure the required leak tightness.

l l

The remaining four valves will then be assembled, but may not undergo initial stroking and LLRT until after their actuators have been modified. In this case, these valves will be cycled for timing and limit adjustment I

l l

t 77

in conjunction with the alignment stroking and LLRT. Again, the final LLRT with acceptable leak rate results will serve as the LLRT for plant operation.

6.1.2 Confirmatory Type "C" Testing In addition to leak testing the MSIVs after they are repaired and reassembled (and whenever they are required to be by Technical Specifications), Niagara Mohawk will test the valves following the MSIV full isolation test.

This leak test will be conducted no later than the first outage following (but within 30 days of) the 100 hour0.00116 days <br />0.0278 hours <br />1.653439e-4 weeks <br />3.805e-5 months <br /> warranty run whichever is earlier.

6.1.3 Mid Cycle Test The next leak rate test will be performed during tne mid-cycle outage tentatively scheduled for the Fall

'87, approximately 12 months af ter initial power operation.

All MSIV cycles will be logged in order to establish a total number of cycles on each valve prior to this I

mid-cycle testing. This outage has been scheduled in l order to perform surveillance testing as required by l

the Technical Specification (e.g., snubber inspections).

It will also serve as the time period to perform all 18 month required surveillances (of which type "C" testing is a part) prior to continued power operation until l

78 l

ee

the first refueling outage tentatively scheduled for the Spring of 1989.

It should be noted that leak rate testing of the MSIV's will be accomplished with the between-the-seats testing as the preferred method. This method has been demonstrated to be a conservative test compared with normal through-the-valve test method. However, should this test method produce unacceptable test results (due to its highly conservative nature), NMPC may employ the through-the-valve test method in order to satisfy Technical Specification requirements.

6.2 DEVELOPMENTAL TESTING 6.2.1 Test Objectives Niagara Mohawk is initiating a developmental program for both the valve and actuator of the MSIV's. This program includes plans to completely review the existing valve and actuator design, perform additional analyses of specific design features, investigate alternate materials and test a full scale prototype of the valve and actuator.

The prototype testing will include operation of the valve under steam flow conditions. A principal objective of this program is to confirm during the initial prototype testing the acceptability of these valves for the first l

79

plant operating cycle. Another principal objective of the developmental program is to demonstrate the ability of the MSIV's to perform their intended function beyond the first operating cycle. The program will also identify any changes in design or materials which will improve the long term reliability of the valves.

6.2.2 Test Organization A task force, under the direction of Niagara Mohawk, has been organized to manage the developmental program.

The task force includes representatives from Crosby Valve & Gage Company, General Electric, MPR Associates, Stone & Webster Engineering Corporation, and Westinghouse Electric Corporation. These representatives will insure access to a wide range of experience and technical expertise on valves, materials, and mechanica1 design. This experience and expertise will be used to the extent necessary in defining and implementing the developmental program.

6.2.3 Testing Format and Schedule The developmental program is being structured into 3 general phases. Phase 1 will include a complete review of the existing valve design, materials and operational experience; a thorough re-analysis of problems which have occured; and identification of any additional areas I

80

. . of conc @rn. Phcsa I alsa includes an initial prototype test program which is discussed in more detail in Section 6.2.4 of this report. Phase 2 will involve identification of alternate design f eatures and materials, evaluation and testing of proposed changes, and the selection of specific changes for implementation. Phase 3 will be the detailed design, procurement, f abrication, qualification and installation of any identified valve and/or operator modifications. Organization and definition of the developmental program is in progress at the present time. Preliminary reviews of the valve and operator relative to long term operation have started. Completion of the Phase 1 effort is scheduled for April 1987.

Completion of the entire Phase 2 effort is scheduled by December 1987. The Phase 3 design work should begin October 1987 with materials for modifications ordered by March 1988. Actual installation of modifications is scheduled for the first refueling outage which is currently scheduled for Spring 1989.

6.2.4 Initial Prototype Test Program 6.2.4.1 Valve Configuration to be Tested The valve configuration to be tested will duplicate to the maximum extent possible the valve and actuator configuration at NMP2. Specifically, the test configuration l

l 81

i will include a full scale MSIV and a hydraulic actuator which is functionally identical to the installed actuator.

Features to be included are the modified force spring pack; a new or recoated ball (equivalent to the balls to be installed); and reference design stellite seat rings, packing, thrust washer, bearings and other internals.

Actuator components will be essentially identical to the installed actuator.

6.2.4.2 Test Conditions Test conditions will duplicate to the extent practical normal p l ant operating and test conditions. Specifically, prototype test conditions will include:

o Ambient pressure and temperature conditions o Normal operating pressure and temperature conditions ,

o Steam flow rates consistent with test facili.ty limitations o Technical Specification leak test conditions 82

l 6.2.4.3 Test objectives The primary objective of the prototype test program is to verify the operability for at least one operating cycle under anticipated normal operating and test conditions, of the MSIV and actuator design and materials installed in NMP2. This will include:

o verification of the mechanical integrity of the valve and actuator for the expected operating and test cycles.

^

o Demonstration of valve leak tightness for the expected valve duty cycles.

o Demonstration of the ability to close the valve within Technical Specification limits under normal operating pressure and temperature steam conditions.

o Verification of the conservatism of the between-the-seat leak test method as an alternative to across-the-valve seat leakage tests.

A second objective of the prototype test program is to provide baseline data for evaluation of (1) the long term suitability of the valve and (2) potential design and material improvements.

83

. - 6.2.4.4 Tost Scop 7 .

Details of specific tests to be performed are under development by NMpC. It is expected that the prototype tests will include the following types of tests:

o Cyclic operation tests which simulate the expected valve duty cycle, e.g., ambient condition cyclic tests, valve closure time tests, partial closure surveillance tests under normal steam conditions and full closure tests at high steam flow rate at operating pressure and temperature.

o Valve leak tests duplicating in-plant Type "C" leak tests. Tests would be performed periodically during the valve cyclic test to monitor the effect of wear / degradation on valve leak rate. Tests will include across-the-valve pressurization and between-the-seats pressurization. Leakage of other

(' valve seals will also be monitored.

o Periodic disassembly and examination of critical components of the valve and actuator during the test program.

The test valve, actuator and f acility will be instrumented to allow monitoring of valve and actuator performance.

84 t

. - _ - - _ - . , - - - . -- - - - - - - - - - - - - - - - - - - - - - - - - ' - - - - - - - ~ - - - ~ ~ ~ ' - ' ~ ~

Additionally, test instrum?ntation may be installed to investigate specific phenomena as considered appropriate.

Special instrumentation requirements will be included in the test plan.

l 6.2.4.5 Schedule The schedule for initial prototype testing is based on availability of the test valve and a c tua to r and a suitable test facility. At present the schedule is as follows:

o Procurement of test valve -

In progress. Delivery expected by January 1, 1987.

o Procurement of test valve actuator - Arrangements being finalized to obtain an actuator from a cancelled power plant. Expect delivery by January 1, 1987.

o Identification and preparation of test facility In progress. Expect to complete f,acility in time to support tests starting in February, 1987.

o Testing -

To be performed in February-March, 1987 time frame with the objective of completing initial prototype tests by April 1, 1987.

85

i 6.2.4.6 Reporting Reporting milestones include:

o Test plan prior to initiation of test o Final report of tests by May 15, 1987.

7.9 CONTINGENCY PLANNING In the unlikely event, that the additional testing or developmental testing programs described earlier indicate that leakage or reliability characteristics of the ball valves are not satisffctory, NMPC is actively developing two contingency plans.

7.1 Leakage Control System One plan would provide for the installation of a leakage control system to be used in conjunction with the MSIV ball valves to ensure that any leakage past the seats is collected and discharged in a monitored, controlled manner to the environ-ment through the Standby Gas Treatment System and main stack.

This Leakage Control System is being designed using the guidance contained in Regulatory Guide 1.96. Preliminary schedule information indicates that the leakage control system could be installed and operable by late January 1987 at the earliest.

However, this schedule assumes all activities are expedited 86

_ -. .,.m

and has no contingency for unknowns.

I 7.2 Y-Pattern Globe Valves The second contingency plan involves the procurement of fully qualified Y-pattern globe valves for installation in place of the ball valves. The globe valves would also be provided with a leakage control system meeting the same requirements as the system described above for use with the ball valves.

Preliminary schedule information indicates that the globe valves and their requisite leakage control system could be installed and operable to support fuel load by late January 1987 at the (a'rliest. Again, the schedule assumes expedited activities and has no contingency for unknowns.

8.0 CONCLUSION

S NMPC believes this report demonstrates that the current MSIV

design coupled with the modified seal spring arrangement is reliable for a minimum of one plant operating cycle.

NMPC believes this report demonstrates that the problem resolution 1

presented herein assures the MSIV's will remain leak-tight through the first plant operating cycle. Niagara Mohawk is committed to continued testing and contingency programs to provide further assurance of reliable service throughout I

l plant operating life.

i l

l 87

9.9 APPENDECIES

9.1 DESCRIPTION

OF MATERIALS OF BALL CONSTRUCTION 9.2 MSIV CHRONOLOGY 9.3 MSIV LLRT PROCEDURE

SUMMARY

9.4 BALL POSITIONING (THRUST WASHER) 9.5 UNION CARBIDE TESTING 9.6 SITE FRICTION TESTING 88

. . APPENDIX 9.1 Description of Materials of Ball Construction o The ball core consists of a type 316L stainless steel casting made to SA351, GR. CF8M.

o A nominal 0.090 inch layer of Haynes 25 is deposited by welding. Haynes Alloy No. 25 is a cobalt-based superalloy with a nominal chemical composition of approximately 50%

cobalt, 10% nickel, 20% chromium and 15% tungsten. It has excellent" corrosion resistance. It was deposited on the 316L MSIV balls using the submerged arc welding process.

The hardness of this as-deposited alloy is approximately 80 to 90 on the Rockwell B scale.

o The tungsten carbide coating (0. 008 " to 0. 010") is deposited on top of the ground Haynes 25 layer using the detonation gun process. Detonation gun process is used because it imparts minimal heat to the workpiece (the ball in this

[ case) and therefore presents the least risk of distortion.

It is a process specifically developed for the deposition i

j of hard, wear-resistant materials such as tungsten carbide.

Typical applications include jet-engine seals and aircraft compressor and turbine blades. The tungsten carbide coating has a Union Carbide coating designation of LW-5 and a nominal composition of 68 W, 22 Cr, 5 Ni, S C.

89

o As used on the MSIV seats, Stellite 6 is a cocalt-based weld deposited hardfacing alloy with a nominal chemical composition of 1.1% carbon, 28% chromium, 4 % tungsten and balance cobalt. Its microstructure consists of a network of chromium and tungsten carbides in a cobalt-chronium matrix. It was deposited on the 316 stainless steel spool seat using the gas tungsten arc welding process. The average hardness of this as-deposited alloy is approximately 35 on the Rockwell C scale. However, the average hardness may not be completely indicative of the alloy's behavior in wear situations since the carbide microconstituents are extremely'hard.

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90 l

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APPENDIX 9.2 MAIN STEAM ISOLATION VALVES CHRONOLOGY August 1976 8" Prototype Ball Valve Testing October 1977 Purchase order for MSIV's was placed with the Fluid Systems Division (FSD) of Gulf and Western (G&W). The Energy Product Group (EPG) of FSD, successor to the EFCO Ball Valve Company (EBV) who actually manufactured the valve.

January 1979 Topical Report submitted to NRC.

February 1981 Valve bodies arrived at the jobsite.

March 1983 Hydraulic actuators arrived at the jobsite.

November 1984 Crosby Valve and Gage Company (Division of Geosource Incorported) has announced the 4

acquisition of the FSD of G&W.

December 1984 Applied corrosion resistant cladding to spool bore area.

April 1985 All MS1V's passed " informational" Type "C" leak tests.

91

- . , _ , . . - - _ , _ , . - . . . - . . , - , ,. ..n-

.. - __ - - _ _ . .-- - -= _

Au' gust 1985 Main Steam Line & RPV hydrostatic test and system flushing.

f December 1985 Installed actuators on valve bodies, i

February 1986 MSIV Pre Operational Testing began on site.

March 1986 All MSIV's passed " formal" Type "C" leak tests.

l March-August 1986 All MSIV's stroked for actuator testing.

September 1986 All MSIV's failed Typa "C" leak tests.

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92

e . APPENDIX 9.3 MSIV Local Leak Rate Test Procedure Sumunary o Pressurize line or valve body to 49 psig with air.

o Vent opposing test boundaries, o Allow sufficient time for air in-leakage readings to stabilize.

o Record iri-leakage value in SCFH.

The leakage test schematic and valve line-ups are shown on the following pages.

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i 93

g . APPENDIX 9.3 (Cont'd)

TYPE "C" VALVE LINE-UPS o Through Valve Test - Inboard

1. Main steam Line Plug in Place

2. Vent "A" Closed

3. Inboard MSIV Closed

4. Inboard Body Cavity Drain Closed

5. Vent "B" Open o Through Valve Test - Outboard

1. Main Steam Line Plug in Place

2. Vent "A" Closed

3. Inboard MSIV Open

4. Inboard Body Cavity Drain Closed

5. Vent "B" Closed

6. Outboard Body Cavity Drain Closed

7. Vent "C" Open o Between Valve Seat Test - Inboard or Outboard *

1. MSIV Closed

2. Vent "A" & "B" or "B" & "C" Open as Applicable 94

4

APPERIDIX 9.4 Ball Positioning (Thrust Washer)

Galling of the thrust washer during valve movement is due to rubbing of the 394 SS thrust washer with the 316L SS ball. On some valves, the galling was observed 3600 around the thrust washer while on others it was evident only on a portion of the circumference.

The galling was not considered to be excessive considering the normal characteristics of a loaded stainless steel to stainless steel interface. Since the thrust washer is at the bottom of the valve assembly"'it is also possible that tungsten carbide chips worked their way into the thrust washer area and initiated galling.

In order to eliminate this galling, a new thrust washer material has been selected.

The material selected is a bronze bearing material. This material is normally used in this type of load bearing application and is not susceptible to galling. The bronze material also will reduce frictional forces on ball movement.

Another critical attribute of the thrust washer is its thickness.

The washer must be sized to allow thermal growth of the ball without interfering with proper seat and ball interaction. In the cold condition, the thrust washer is sized to be twenty to thirty mils smaller than would be required to align the ball bore to the valve 95

i

..; . body boro (sco Figuro 9.4-1) . This undersize allows for thermal growth and permits the seat to self align without interference with the body bore. If sized thinner, the seat could bottom out j on the body bore, if sized thicker the seat could top out on the body bore in the hot condition. In either case seat or ball damage could occur.

+

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96

. . _ - _ . _ _ _ _ , - - _ _ _ . - . _ _ . _ _ . . . - - . . , . , _ _ _ _ . _--__,___-,..m__,. - _ _ _ _ . _ _ - , -

T =

(1/2 A+B) -

(1/2 C+D) -

(0.020" TO 0.030")

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l j o.020" I TO 0.030" l A -U I

_ C i $ ODY JL BALL l ORE l BORE l V l n I u

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lhLVE THRUST i BODY WASHER THRUST

! WASHER

SEAT l

l THRUST WASHER SIZING suusivuo4i FIGURE 9.4-1

L

• APPENDIX 9.5 UNION CARBIDE TESTING Dynamic Rub Test The dynamic rub test is performed by fixturing three 1/4 inch diameter Stellite 96 pins into a chucking device and mounting it in a vertical press. Mounted below the pins is a tungsten carbide coated SS with Haynes 25 overlay and seal plate, simulating the MSIV ball configuration. The pins are lowered to the plate and a known load is applied. The chuck containing the pins is then rotated to approximately 900 then returned. The applied stress, velocity of rotation and average coefficient of friction are recorded and plotted for each cycle. Figure 9.5-1 shows this arrangement.

A typical rub test would subject a specific material combination to a specified number of cycles on the test appuratus or until coating failure. The plate speciman is then visually examined, liquid penetrant examined, if required, and physical measurement of removed tungsten carbide are taken.

Rub Test Results During tne rub test, the unit loading on the Stellite $6 pins varied f rom 3,000 psi to 5,000 psi in five steps. Each load increment 98

was cycled 25 times and values for coefficient of friction were established. In all cases the value for coefficient of friction ranged from 0.3 to 0.5, with 0.3 to 0.4 being the values at the 3,999 psi value. Table 9.5-1 summarizes this data. This table also shows the depth of scars remaining on the tungsten carbide after testing. This depth is the depth of the area over which the Stellite #6 pin transversed during cycling.

Thermal Cracking Test .

An additional test was run to determine the effects of temperature on the tungsten carbide coating. The test heated a plate overlayed with Haynes 25 and tungsten carbide to 6200 F at a rate of approximately 1990 F per hour and acoustically monitored the sample to record the number of acoustic events. Only a single acoustic emission was recorded at 6200 F during the test. The speciman was then inspected under a 10x magnification. No signs of cracks were present. Should temperature increments cause cracking, far more emissions would have occured.

l I

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l 99 i ---

1 4 1 O

l LOAD APPLIED THROUGH O O HYDRAULIC ACTUATION l

A-A  :

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ROTATES 90*

( j 50 FPM I

f A

VERTIC AL MOVEMENT A ik J LA FIXED PL ATE OF TEST MATERIAL l

UNION C ARBIDE WE AR TEST 100 M ACHINE NMMSIVMOSS FIGURE 9.5-1 1 _.-__ ___ _. _ _ _ _ _ _ . _ . - . .. .. - - . _ .a . . - _ _ _ _ .

WEAR TEST RESULTS ,

UCAR L W.- 5 CN HAYNES 25-WELD OVERLAY WITH STELLITE 6 PIN S-UNIT LOAD COEF.OF DEPTH PSI CYCLES FRICTION OF SCAR

, 3,000 1-25 .3 .4 NONE l 26-50 .3 .4 51-75 . 4' .25 10,000 1-25 .4 .5 0.1 T O O.15 26-50 .5 .4 51-75 .4 .3 l 20,000 1-25 .5 .4 0.1

! 26-50 .4 51-75 .4 30,000 1-25 .5 .4 0.3 26-50 .4 .5 51-75 .4 .5 40,000 1-25 .4 0.3 T O O.4 26-50 HIGH 51-75 .4 AND HIGH 50,000 1-25 HIGH 0.5 26-50 HIGH Muusiv uP4 2 T ARE 9.b1

O 4 APPE! DIX 9.6 SITE FRICTION TESTING Mock Up Friction Test In an effort to simulate the wearing condition that was evident on the eight original balls, a test was devised. The test would use one of the MSIV balls with damaged coating and a simulated portion of seat made from Stellite 46. Figure 9.6-1 shows the test rig used.

The ball was str'o'ked across the seat which had a contact area of 0.06 square inches. As the stroking occured a hydraulic ram was used to load the seat on the ball to a given pressure.

Mock Up Test Results Tne testing was performed hot (4500F) at 56,000 psi seat load and cold (650F) at 3,500 psi. The pressures were established by estimates of seat to ball pressure. Calculations later provided numbers that were within a reasonable percent dif ference of estimates.

Taoles 9.6-1 and 9.6-2 show test results. The low pressure (3,500 psi) condition envelops the expected maximum seat pressure with the modified spring pack while the higher pressure 56,000 psi exceeds the calculated rocking pressure without the modified spring pack.

102 l

a b It is clear from the results that coating failure at nigh contact pressure is likely to occur in a relatively few strokes (approximately

50) while at the lower contact pressure failure could not be predicted.

103 a

N 1 O 6 J I N 9 N

U PN E R

R T UO U I TG U

K TSF I

C CE I

T OR MF 1

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'1 6 E E T E EC AL E P

T TA A I

P I P L L LF P L

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TA C S.

B SH C S.

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- RE OBM PI PT U 4 SEL 5 D

M LD V LA I S

AR M BC 5" M N

56,000 psi BALL PRESSURE ,

SEAT AREA -

.06 IN SO.

TEMPERATURE -

450*F I

NUMBER OF STROKES MSIV BALL SURFACE SIMULATED SEAT 10 DARKENED CONTACT AREA, POLISHING

.: SLIGHT WE AR 20 VISIBLE WE AR, SOME VISIBLE WE AR l TR ANSFER OF M ATERIAL

! a j 30 SME ARING, POCKMARKS POCKMARKS GR AINY SURF ACE APPE AR ANCE SCRATCHES l

l 40 MORE WE AR, SIGNS OF MORE WEAR FRACTURE CRACKS MATERIAL REMOVED ACROSS WEAR SURFACE, LAYERED TEXTURE TO SURFACE 50 WC F AILURE - CO ATING MORE WEAR FLAKED OFF MOCK UP FRICTION TEST Nuusiv up3 3 T ABLE 9 6-1

_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ __ . _ _ _ _ , _ _ _ =

3,500 PSI BALL PRESSURE SEAT AREA -

.06 IN SO.

TEMPERATURE -

65*F NUMBER OF STROKES MSIV BALL SURFACE SIMULATED SEAT 3 15 NO WEAR NO WEAR 1 BL ACK LINE INDIC ATING CONTACT AREA CONTACT AREA POLISHED l 45 NO CHANGE NO CHANGE

- 75 NO CHANGE VISIBLE WE AR 8 SCR ATCHES & PIT TING t

l 12 5 ROUGHER SURFACE WE AR HAS INCRE ASED l

l i 17 5 SHINY ARE AS POSSIBLE WE AR ON SE AT

TRANSFER OF SEAT HAS SLOWED DOWN M AT ERI AL TO THE B Al.L 225 WE AR RATE HAS SLOWED SAME AS ABOVE DOWN - AMOUNT OF WEAR WOULD REOUIRE REL APPING

, OF THE WC COATING MOCK UP F RIC TIO N TEST RMMSivuP32 TASLE 9 6-2