ML20112C246

Forwards Info Re Purge & Vent Valve Operability Qualification,Per Draft SER Open Item 26j & in Response to 840906 Request

ML20112C246
Person / Time
Site:	Vogtle
Issue date:	03/18/1985
From:	Bailey J GEORGIA POWER CO.
To:	Adensam E Office of Nuclear Reactor Regulation
References
GN-560, NUDOCS 8503210362
Download: ML20112C246 (30)

Text

Georg!a Power Company .

Project Management e, Rout 3 2, Box 299A Wayrtesboro, Georgia 30830 Telephone 404 724 8114 404 554 9961 h

Vogtle Project a

March 18, 1985 Director of Nuclear Reactor Regulation File: X7BC35 Attention: Ms. Elinor G. Adensam, Chief Log: GN-560 l

Licensing Branch 54 Division of Licensing

- U. S. Nuclear Pz;gulatory Commission Washington, D.h. 20555 REF: EISENHUT TO FOSTER, DATED 9/6/84, ACCEPTANCE REVIEW OF APPLICATION FOR OL NRC DOCKET NUMBERS 50-424 AND 50-425 CONSTRUCTION PERMIT NUMBERS CPPR-108 AND CPPR-109 VOGTLE ELECTRIC GENERATING PLANT - UNITS 1 AND 2 DSER OPEN ITEM 26j - DEPENDABILITY OF CONTAINMENT ISOLATION

Dear Mr. Denton:

In Enclosure 7, Attachment 4 of the referenced letter, a request was made for information on purge and vent valve optrability qualification. This request was carried forward in the DSER as open item 26j. Attached for your staff's review are five copies of the requested information.

If your staff requires any additional information, please do not hesitate to contact me.

Sincerely, J. A. Bailey Project Licensing Manager JAB /sm Attachment xc: D. O. Poster L. Fowler R. A. Thomas M. A. Miller G. F. Trowbridge, Esquire L. T. Gucwa J. E. Joiner, Esquire G. Bockhold, Jr. I g C. A. Stangler 0074m 24 B5032 PDR 0362 BSDOCK DR 05 E

t Operability. Qualification of\ Attachment Open Item 26j Purge and Vent Valves Demonstration of operability of the containment purge and vent valves and the ability of these valves to close during a design basis accident is necessary to assure containment isolation.

This demonstration of operability is required by NUREG-0737,

" Clarification of TMI Action Plan Requirements," II.E.4.2 for containment purge and vent valves which are not sealed closed <

during operational conditions 1, 2, 3 and 4.

1. For each purge and vent valve covered in the scope of this review, the following documentation demonstrating compliance with the " Guidelines for Demonstration of Operability of Purge and Vent Valves" (Attachment 2) is to be submitted for staff review:

A. Dynamic Torque Coefficient Test Reports (Butterfly valves only) - including a description of the test setup.

I Response Testing has been performed by the butterfly valve supplier i to obtain dynamic torque coefficients. See attachment 2.

l B. Operability Demonstration or In-situ Test Reports (when used)

P.esponse Not Applicable C. Stress Reports

Response

l The stress analysis for these valves are provided as part of the attachments.

D. Seismic Reports for Valve Assembly (valve and operator) and associated r3rts.

Response

i The seismic reports for the valve assemblies are provided as part of the attachments.

\

-E. Sketch or description of each valve installation showing the following (Butterfly valves only):

1. direction of flow

2. disc closure direction

3. curved side of disc, upstream or downstream (asymmetric discs)

4. orientation and distance of elbows, tees, bends, etc. within 20 pipe diameters of valve S. shaft orientation

6. distance between valves

Response

See Table 1.

F. Demonstration that the maximum combined torque developed by the valve is below the actuator rating.

Response

REQUIRED BETTIS ACTUATOR YOKE ARM (DEGREES) ACTUATOR SPRING TORQUE (in-lb.) TORQUE (in-lb.)

0 3714.05 18828 10 1834.61 16628 20 2057.18 15602 30 2244.81 14268 40 .

3182.96 14345 50 3946.42 14952 60 6197.98 16088 70 8157.09 18845 80 11807.46 23362 90 11807.46 29640 See Table 2 and 3

2. The applicant should respond to the " Specific Valve Type Questions" (Attachment 1) which relate to his valve.

Response

See attachment 1.

\ '

3. Analysis, if used, should be supported by tests which establish torque coefficients of the valve at various angles. As torque coefficients in butterfly valves are dependent on disc shape, aspect ratio, angle of closure flow direction and approach flow, these things should be accurately represented during tests. Specifically, piping installations (upstream and downstream of the valve) during the test should be representative of actual field installa-tions. For example, non-symmetric approach flow from an elbow upstream of a valve can result in fluid dynamic torques of double the magnitude of those found for a valve with straight piping upstream and downstream.

Response

Although testing was not performed to establish torque coefficients at various valve angles, analytical techniques used by the valve supplier are based on model testing experience and scaling developed over many years. More information is presented in attachment 2.

4. In-sitt tests, when performed on a representative valve, should ce performed on a valve of each size / type which is determined to represent the worst case load. Worst case flow direction, for example, should be considered.

Response

Not applicable. In-situ tests are not performed.

5. For two valves in series where the second valve is a butterfly valve, the effect of non-symmetric flow from the first valve should be considered if the valves are within 15 pipe diameters of each other.

Response

The valves are located 22'-5" and 22'-9 9/16" from each other (HV-2626B to HV-2627B and HV-2628B to HV-2629B respec-tively) or approximately a distance of 19 pipe diameters, which is not within the 15 pipe diameters limitation.

6. If the applicant takes credit for closure time vs. the buildup of containment pressure, he must demonstrate that i

the method is conservative with respect to the actual valve closure rate. Actual valve closure rate is to be determined under both loaded and unloaded conditions (if valves close faster at all angles of opening under loaded conditions, no load closure time may be used as conservative) and periodic inspection under tech. spec. requirements should be performed to assure closure rate does not ine sase with time or use.

Response

No credit was taken for closure time vs. the buildup of containment pressure. See attachment 2.

I

.4_

. .. . . .- i ,

TABLE 1

! Valve Tag Number HV-2626B -

HV2627B HV2628B HV2629B

! 1. Direction Inlet-butt weld end Inlet-raised face Inlet-raised face Inlet-butt weld end of Flow outlet-raised face outlet-butt weld end outlet-butt weld end outlet-raised face

2. Disc closure Clockwise Clockwise Clockwise Clockwise Direction

3. Curved side Upstream Downstream Downstream Upstream of disc l 4. Orientation and Tee-19'5" upstream, Tee-3'-0" down- Tee-18'-9-9/16 up- Tee-4'-0" downstream, i distance of elbows, horizontal line stream horizontal stream, horizontal line. horizontal line.

tee, bend within 90' elbow-11'-6-1/4" line. 90* elbow-11'-9-3/4" 90' elbow-10'11-13/16 20 pipe diameter upstream horizontal 90* elbow-10'-10-3/4" upsream, horizontal line. downstream horizontal-of valve. line. downstream, horizontal Tee-3'-9-3/4" upstream, line.

Tee-6'-9-11/16" upstream, line. vertical line. Tee-16'-11'-3/16 down-vertical line. Tee-15'-7-5/16" down- 90* elbow-l'-8-3/4" stream vertical line.

90' elbow-2'8-11/16 stream, vertical line. upstream, vertical line. 90* elbow-21-0-3/16" uphr. ream, vertical line. 90' elbow-19'-8-5/16" Flow orifice-25'-4-9/16" downstream, vertical Flow orifice-23'-6-5/8 downstream vertical upstream, horizontal line, upstream, vertical line. line. line. Flow orifice-2'-7" Flow orifice-l'-10-1/4" riow orifice-l'-1-5/8" Flow orifice-1'10-1/8" upstream, horizontal downstream, horizontal upstream, vertical downstream, horizontal line.

line. line. line. Flow orifice-Flow orifice- Flow orifice- 24'-7-10/16" down-24'-3-1/4" do'instream, l'-10-1/8" downstream, stream, horizontal horizontal line. horizontal line. line.

90

5. Shaft orientation. Vertical Vertical Vertical Vertical

6. Distance between 22'-5" (between - V-2626B & 2627B) 22'-9-9/16" (between HV-2628B & 2629B) j valves. .

I e'

TABLE 2 CCTUITOR TORQUE CALCULATIONS KEV. 8 s

9 TEMP. 320 DEG F REF. NO. CASE 4 VALVE TYPE 9280 VALVE SIZE 14 INCH SHAFT F AT. SA-564 H1075 31-39 BUSH. HAT. BRONZE / GRAPHITE DISC-SHAFT CON. PINNED DRIVE-SHAFT CON. M YED Flow DIR. TOWARD HUB OATE 02-11-85 BY JON C. WHITESELL INPUT DATA-

)!

0.0 10.000 20.000 30.000 40.000 50.000 60.000 70.000 80.000 90.000 ANG 13 244 13 244 13 244

• 13 244 13 244 13.244 13 244 13.244

~ DOISC 13.244 13 244 50.000 50.000 DELTP 60.000 50 000 50 000 50.000 50 000 50 000 50.000 50.000 0.906 0.906 0 906 0.906 0.906 0.906 0.906 0.906

, DLO 0.906 0 906 0.0 00 00 0.0 00 0.0 00 0.0 T5 2n37.000 00 0.0 0.0 0.0-TI 0.0 0.0 00 0.0 00 00 0.0 0.0 0.0 0.0 00 0.0 0.0 0.0 0.0 DM 0.0 0.0 1145.000 15.400 32.000 50.000 108.000 194 000 368.000 742.000 1145.000 Dir 0.0 50.000 50.000 DPIN 60.000 50.000 50 000 50.000 50.000 50 000 50.000 50.000 0 400 0 300 0.250 0 250 0 200 0.200 0.140 0.140 0.140 DELTPF 60.000 98 000 98.000 98.000 98.000 STSH 98.000 98 000 98 000 98.000 98 000 98 000 85 000 e5 000 85.000 85.000 85 000 85.000 85.000 85.000 85.000

• SbOSH 85.000 0.500 0.500 0.500 0.500 Cl 0.500 0 500 0 500 0.500 0.500 0.500 0.750 0.750 0.750 0.750 0.750 0.750 0.750 0.750

C2 0.750 0.750 2 2 2 2 2 2 2 2 2 2 BSHTYP 1.748 1.748 1.748 DSHFT 1.748 1 748 1 748 1 748 1 748 1 748 1.748 GENERATED WARIABLES 51450.00 51450.00 51450 00 51450.00 51450.00 51450.00 51450.00 51450.00 51450.00 ST 51450.00 25725.00 25725.00 25725.00 25725.00 SS 25725.00 25725.00 25725.00 25725 00 25725.00 25725.00 8500.00 8500 00 8500.00 8500.00 8500.00 8500 00 8500.00 8500.00 SH M500.00 8500.00 OUTPUT ACT.10Ru 3714.0564 1834.6165 2057.1843 2244.8140 3182.9634 3946.4241 6197.9805 8157.0937 11807.4687 11807.4687 STRSI 8087.2656 6147.9375 6225.5703 6300.2148 6779.6562 7269.5391 9003.0859 10684.2070 13981 1289 13981 1289 STRS2 4509.1367 3166.1660 3243.7979 3318.4426 3797.8850 4287.7695 6021.3125 7702.4336 10999.3555 19999.3555 STRS3 5044.8281 2982.0837 3194.6082- 3373.7705 4269.5R20 4998.5898 7148.5391 9019.2461 12504.8906 12504.8906 5T954 5490 2148 2130.3176 2555.5364 2914.0049 4706.3477 6164.9492 10466.5742 14209.4844 21183.5664 21183 5664 STRSS 4730.4961 2336.7039 2620.1831 2859.1621 4054.0593 5026.4570 7894.2109 10389.4844 15038.8750 15038.8750 1128.0229 1128.0229 1128.0229 1128.0229 1128.0229 1128 0229 1128.0229 STGS6 1353.6267 1128.0229 1128.0229 iL (1) e

TABLE 2 VARIABLE NAME VARIABLE DESCRIPTION INPUT DATA ANG ANGLE IN DEGREE DDISC DISC DIAMETER (UNITS: IN)

DELTP PRESSURE DROP IF ITYP = 1 (UNITS: PSI)

DSHTT SHAFT DIAMETER IF ITYP = 2 (UNITS: IN)

DSHFT SHAFT DIA'ETER IF ITYP = 3 (UNITS: IN) *o DLO HUB TO BUSH DISTANCE (UNITS: IN) FROM DRAWING g TS SEATING TORQUE (UNITS: IN-LES)(FOR 0 DEGREE ANGLE 0;;LY)

REFER TO TEST REPORT #92.3-44 SIZING EQUATION FOR 9200 (NEW EQUATION)

TI INERTIAL TORQUE (UNITS: IN-LES)(FOR FAST STROKING SPEED ONLY 3 LL5s THAN gst AMOUNT OF CAM (UNITS: IN)

CENTERLINE OF DI; DM [

FROM DRAWING . DM SHAFT DTF DYNAMIC TOROUE FACTOR FROM SALES HANDBOOK: CFC 40B-10, C VALUES FOR 9200 SERIES, FLOW AGAINST WB USE 1.5 C DPIN INLET PRESSURE (UNITS: PSIC)

TROM REQUISITION IF ITYP = 3 LEAVE IT BLANK

! DELTPF EFFECTIVE PRESSURE DROP FACTOR i FROM SALES BANDBOOK: CFG'40B-10. TABLE 1 (D0 NOT INCLUDE P , ONLY COEFFICIE T)

, 3 STSH SHAFT STRENG111 FACTOR 1 FROM SALES BANDBOOK: CPC 20D-10 SBUSH BUSH. STRENGTR FACTOR 1 FROM SALES BANDBOOK: CFG 20D-10

C1 DISC TO SHATT CONCETRATION FACTOR FOR FIN = 0.5, EEY = 0.75 C2 DRIVE TO $!!ATT CONC 511LATION FACTOR FOR FIN = 0.5, KEY = 0.75 (2)

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

TABLE 2

~. -

VARIABLE NAME VARIABLE DESCRIPTION ESHTYP BUSH. TYPE.NO.

FROM SALES HANDBOOK: CFG 20D-10 NOTE: 1. ALL ARE REAL NUMBERS EXCEPT BSHTYP (BSHTYP-INTEGER)

2. REPEAT CARDS 2 THROUGH 3 FOR NLMSER OF SHATT CALCULATIONS

3. REPEAT CARDS 3A AND 3B FOR NUMBER OF ANGLES FOR EACH SHATT CALCULATION .

G

. g e I

(3) -.

m. . - , _ , . _ . _ . .

TABLE 3 s n i

7 PRESSURE PRESSURE PRESSURE YOKE ARM SPRING PRESSURE ANGLE TORQUE TOROUE TOROUE TORQUE TORQUE (degrees) (in Ib) C 50) psi ( 60) psi ( 70) psi ( 80) psi 0 18828 19187 26587 33986 41386 15 15602 12435 17949 23463 28977 30 14268 9444 14147 18851 23554 i

45 14394 8053 12542 17031 21521

'l 21837 60 16088 7532 12300 17069 75 20224 7572 13256 18941 24625 90 29640 7778 15600 23422 31245, 1

58088 . _ _ _ _ , _ _ _ ,

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E E E E E E E E E E

- N TT) v W 843 h CX3 C73 M ARN ANGl.E 4

i ATTACHMENT 1 SPECIFIC VALVE TYPE QUESTIONS _

AND RESPONSES FOR VEGP PURGE VALVES

Operability Qualification of Purge and Vent Valves Specific valve Type Questions The following questions apply to specific valve types only and

. need to be answered only where applicable. If not applicable, state so.

A. Torque Due to Containment Backpressure Effect (TCB)

For those air operated valves located inside containment, is the operator design of a type that can be affected1.e., by the containment pressure rise (backpressure effect) where.the containment pressure acts to reduce the operator torque capability due to TCB. Discuss the operator design ,

with respect to the air vent and bleeds. Show how TCB was calculated (if applicable).

Response

a. The valves are equipped with BETTIS NT316B-SR2-M3 spring-return actuators. The spring side of the cylinder actuator is vented to the local ambient conditions; if the pressure side is vented (through-the solenoid) to the same local ambient conditions, no pressure differential will exist across the cylinder as a result of-surrounding local pressure rise. The spring will still drive the actuator to the safety-mode (closed) position and maintain that position as long as the solenoid remains de-energized and as long as no subsequent re-opening signal is received.

In the event of a delay in solenoid de-energization, a local ambient' pressure rise will reduce the AP across the cylinder, which will initially partially close the valve. When the solenoid subsequently de-energizes and vents, the spring would complete the closure stroke.

! b. In the event the external ambient pressure is maintained at an elevated level for a prolonged period with the solenoid still energized, and providing the regulator l

is vented to the same ambient level with a sufficiently -

high supply, the regulator would eventually adjust the air supply to the cylinder actuator to re-establish the initial full-open position.

c. Adequate spring-driven torque output is available from the actuator to control the valve from any open or closed position (regardless of external ambient pressure) ,

providing the cylinder casing is vented (locally) .

The torque available is well within the capabilities

of the NT316B-SR2-M3.

l i

E

+- , ---w, ,w,--,---._m,,,---y-y- ,,mww.,,,,,,w __,,-,._,,,,,,.-.,,,,,,v-. ---,ww m-- m --,-,--, w w.- w,,pc ..----,-m--w-,.w t

B. Where air operated valve assemblies use accumulators as the fail safe feature, describe the accumulator air system configuration and its operation. Discuss active electrical ,

components in the accumulator system, and the basis used to l determine their qualification for the environmental conditions I experienced. Is this system seismically designed? How is i the allowable leakage from the accumulators determined and j monitored?

Response - Not Applicable.

C. For valve assemblies requiring a seal pressurization system (inflatable main seal), describe the air pressurization system configuration and operation including means used to determine their qualification for the environmental condition experienced. Is this system seismically designed?

Response - Not Applicable.

D. Where electric motor operators are used to close the valve

~has the minimum available voltage to the electric operator under both normal or emergency modes been determined and specified to the operator manufacturer to assure the adequacy of the operator to stroke the valve at accident conditions with these lower limit voltages available? Does this reduce voltage operation result in any significant change in stroke timing? Describe the emergency mode power source used.

Response - Not Applicable.

E. Where electric motor and air operator units are equipped with handwheels, does their design provide for automatic re-engagement of the motor operator following the handwheel mode of operation? If not, what steps are taken to preclude the possibility of the valve being left in the handwheel mode following some maintenance, test, etc. type operation?

Response - Not Applicable.

F. For electric motor operated valves have the torques developed during operation been found to be less than the torque limiting settings?

Response - Not Applicable.

2

ATTACHMENT 2 VEGP DEMONSTRATION OF OPERABILITY OF PURGE AND VENT VALVES

E . .

i is Guidelines for Demonstration i of Operability of Purge and '

Vent Valves operability 4

In order to establish operability, it must be shown'that the valve actuator's torque capability has sufficient margin to overcome or resist the torques and/or forces (i.e., fluid dynamic, bearing, seating, friction) that resist closure when stroking

- from the initial open position to full seated (bubble tight) in the time limit specified. This should be predicted on the pressure (s) established in the containment following a design basis LOCA. Considerations which should be addressed in assuring valve design adequacy include:

1. Valve closure rate versus time - i.e., constant rate or other.

Response

In the case of a cylinder, spring-return actuator (such as the BETTIS NT316B-SR2-M3, Push Down to Open), full travel is achieved at a cylinder pressure of approximately 60 psig. Cylinder pressurization up to 60 psig is done to provide the maximum torque capability (hold-open torque) at 90 travel. Therefore the spring will not start to move the valve shaft linkage until some cylinder pressure bleed-down through the solenoid occurs (when the solenoid is de-energized).

When calculations were done to determine tl.. allowable P at open angles, the assumption was made that the peak containment pressure (50 psig) was present immediately at the onset of the l LOCA accident condition. This is a conservative approach since-the buildup from ambient to 50 psig has been specified to occur i in 5 seconds. No time-history study was made, since maximum j conditions were assumed at the onset.

! The Type NT316B-SR2-M3 actuator cylinder pressurized volume has to be expelled through a 3/4" NPT connection (Cv=7.5) connected to a NP831674E solenoid, in order for closure action to occur. i There is minimal restriction in the solenoid (Cv=5.5), so that

the pressure line to the actuator is immediately vented when the

normally closed solenoid is de-energized upon receipt of the closure (safety-mode) signal .

Production valve stroking times were demonstrated as required (for each valve assembly) prior to shipment. At this point, the best stroking time data could be obtained by timing the installed valve assemblies during a field stroking test, conducted at the plant site.

When stroking time calculations or testing is done, lag times

! due to cylinder overpressure venting are considered and included r

4

.,-- ._m.er - 4~-%.,-,_--,e .,y,m,,.,,-yn%, , _m-e-,---.. .%.,,_.,._.-,__,.-m_,.

when supplying stroking data, i.e. the closure / opening times are noted from the time the solenoid signal is received, not from the time the actuator starts to move.

2. Flow direction through valve /k across valve.

Response

Peak containment (LOCA) pressure (50 psig, @ 320*F) was used in determining the fluid conditions across the valve at all open angles of rotation (10' to 90 ). P across the valve was considered equal to peak containment pressure (PSIG).

Material properties were evaluated at peak containment temperatures.

The effect of compressible flow in sizing Fisher butterfly valves is best explained by the following:

AIR VS WATER SERVICE Whenever a Fisher butterfly valve is in a gas flow application, the effects due to compressible flow are taken into considera-tion while determining the dynamic torque effects for each individual valve selection. This consideration is built into our valve selection procedures and requires a conscious liquid or gas decision in calculating the effective pressure drop of wich the dynamic torque is a function.

Fisher's philosophy concerning the effects of compressible flow on butterfly valves is presented in ISA Transactions, Vol. 8, No. 4 entitled, "Effect of Fluid Compressibility on Torque in Butterfly Valves", written by Floyd P. Harthun (Manager, Product Evaluation, Fisher Controls Co. (Attachment 4).

3. Single valve closure (inside containment or outside contain-ment valve) or simultaneous closure. Establish worst case.

Response 1

No credit is taken for reduced pressure due to valves in series.

Vogtle valves are designed for the maximum in-containment pressure.

4. Containment back pressure effect on closing torque margins of air operated valve which vent pilot air inside containment.

Response

See response to Item A in attachment 1.

5. Adequacy of accumulator (when used) sizing and initial charge for valve closure requirements.

Response - Not Applicable.

6. For valve operators using torque limiting devices - are the settings of the devices compatible with the torques required to operate the valve during the design basis condition.

Response - Not Applicable.

7. The effect of the piping system (turns, branches) upstream and downstream of all valve installations.

Response

All Fisher sizing data is based on dynamic torque determination tests which were performed with uniform flow profiles and on valve discs with representative geometries. The effects of a non-uniform flow profile, due to piping elbows, "T"-connections, etc., upstream, are discussed below.

The concern over geometrical piping system effects is relevant, since Fisher typically sizes butterfly valves assuming a uniform could produce a non-uniform flow as illustrated by Figure A of Attachment 3.

a. Valve / Flow Orientation, Figure A The plant layout is such that the valve is oriented to the flow as depicted in Fig. A (Attachment 3), the non-uniform fluid profile will not produce an additional torque on the valve disc since both " wings" of the disc (as split by the shaf t) will be subjected to the same flow with respect to time.

8. The effect of butterfly valve disc and shaft orientation to the fluid mixture egressing from the containment.

Response

See Item No. 7 response above.

i

I Demonstration i Demonstration of the various aspects of operability of purge and vent valves may be by analysis, bench testing, in-situ testing or a combination of these means.

Purge and vent valve structural elements (valve / actuator assembly) 1 must be evaluated to have sufficient stress margins to withstand loads imposed while valve closes during a design basis accident.

Torsional shear, shear, bending, tension and compression loads /

stresses should be considered. Seismic loading should be addressed.

Once valve closure and structural integrity are assured by t analysis, testing or a suitable combination, a determination of the sealing integrity after closure and long term exposure to the containment environment should be evaluated. Emphasis should be directed at the effect of radiation and of the containment spray chemical solutions on seal material. Other aspects such as the effect on sealing from outside ambient temperatures and debris should be considered.

The following considerations apply when testing is chosen as a means for demonstrating valve operability:

Bench Testing A. Bench testing can be used to demonstrate suitability of the in-service valve by reason of its traceability in design to a test valve. The following factors should be considered when qualifying valves through bench testing.

1. Whether a valve was qualified by testing of an identical valve assembly or by extrapolation of data from a similarly designed valve.

Response

In determining allowable pressure drops across a particular butterfly valve at various angles of the disc, Fisher Controls uses classical " mechanics of materials" type equations to calculate stress levels at various worst-case locations in the valve assembly (specifically, various locations along the valve shaft). The approach to the analysis, the equations used, and the combination of the calculated stresses all make up a portion of Fisher's design philosophy for butterfly valves. This analysis approach addresses all of the different states of shea:: and stress which are applicable to the loading conditions defined.

Establishing the loads that actually exist makes up the remaining portion of our design philosophy for butterfly valves. These loads range from easily calculated loads, such as bending due to pressure differential across the disc, to loads such as packing and dynamic torques which

s require a certain amount of testing combined with scaling in order to analyze al1~ valve sizes. It is the factor of ,

dynamic torque that produces different stresses at different I disc rotations and disc geometries. Through testing and )

scaling, Fisher has produced dynamic torque factors for incremental disc rotations. i I

The model tests used to establish the dynamic torque values used in sizing were conducted using 4" and 6" test valves with various aspect ratios ranging from 2:1 to 14:1 (such as 3:1, 4 :1, 5:1, 8 :1, 11:1, and 14 :1) . The dimensionless aspect ratio (defined as the ratio of the disc diameter to the hub diameter) was judged to be a significant parameter for evaluation of dynamic torques at various opening angles.

The tests were conducted using the Fluid Controls Institute

~

(FCI) specifications for test arrangement and conduct, per FCI paper 58-2.

No published data is.available describing the precise scaling procedure used in establishing the sizing tables.

However, the general approach is described in some detail in.ISA Transactions, Vol. 8, No. 4, " Effects of Fluid Compressibility on Torque in Butterfly Valves" (Attachment 4).

All of the testing, scaling and analytical results referenced above have been compiled and tabulated for all available valve sizes, classes, and materials, making it relatively easy to determine the allowable pressure drops, torque requirements, and capacities for specific constructions.

2. Whether measures were taken to assure that piping upstream and downstream and valve orientation are simulated.

Response

No credit was taken for downstream back pressure on the design of these valves. All design constraints for upstream piping provided by the valve supplier are factored into the system design.

4

3. Whether the following load and environmental factors were considered

a. Simulation of LOCA

b. Seismic loading

c. Temperature soak

d. Radiation exposure

e. Chemical exposure

f. Debris i

o 1

2

Response

Simulation of LOCA, seismic loading, temperature soak, radiation exposure, and chemical exposure are delineated in Fisher Qualification Report FQP-llA. Debris was not considered because debris screens have been provided upstream of the valves to prevent debris from interfering with valve closure. Seismic loading is supplemented by seismic analysis, resonant frequency search testing, and seismic static loading of a Vogtle production valve.

B. Bench testing of installed valves to demonstrate the suita-bility of the specific valve to perform its required function during the postulated design basis accident is acceptable.

1. The factors listed in Items A.2 and A.3 should be considered when taking this approach.

Response - Not Applicable.

In-Situ Testing In-situ testing of purge and vent valves may be performed to confirm the suitability of the valve under actual conditions.

When performing such tests, the conditions (loading, environment) to which the valve (s) will be subjected during the test should simulate the design basis accident.

NOTE: Post test valve examination should be performed to establish structural integrity of the key valve /

actuator components.

Response - Not Applicable.

3

Attachment 3: Piping System Sketch, Figure A i

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l. FLOYD P. MARTHUNt k Fisher Goternar Company i Marshalltown,lona

> A technique is presented by which the shaft torque resulting from fluid fiow through butterfly vahes can be determined with reasonable accuracy for both compressib!e and incompressibfe flow. First, the generat torque relationship for incompressib'e flow #5

, estabbshed. Then, an effective pressure differentialis defined to extend this relationtHp

• to includp the effect of fluid compressibility. The application of this techn'que shc 9ed very good agreement with experimental test results.

INTRODUCTION DEVELOPMENT OF GENERAL TORQUE Tite Arrt.trATios of butterfly valves in various automatic RELATIONSHIP control systems requires proper actuator sizing for The total shaft torque required to operate butterfly efficient control. Thus, a thorough knowledge of the vahes can be separated into two major components:

fluid reaction forces acting on the valve disc is required.

Extensive experimental work"' has been performed in 1. Dynamic torque-that portion of the total the past to establish a relationship to determine these operating torque attributable to the fluid reaction forces and thus determine the resultant shaft torque. The force of the flowing medium acting on the vahe general form of this relationship has been established and disc.

conGrmed. However, by using the classical fluid momen. 2. Friction torque-that portion of the total tum approach, a similar relationship can be obtained in operating torque attributable to friction in the which the torque is shown to be directly proportional to packing and bushings.

the measuted valve pressure differential for a F iven disc .

position. This relationship along with most of the Since each of these components is independent of the previously publishec' torque information is adequate for othc., a separate evaluation of cach component affords incompressible Dow Although the effect of fluid com- the best approach to this problem. This investigation n pressibility on torque has been recognized, no useful limited to an evaluation of the dynamic torque com-relationship has been developed. The primary ob. ponent. If the friction on the valve shaft is assumed to be jective of this insestigation is to extend the established {ndependent of direction of rotation,it can be readily torque relationship to include the en'eci of fluid com. Isolated. The imque required to totaic the valve disc is pressibihty. ".icasured in a clockwise and a counterclocl.ane direc-tion through full trasel. Since friction always opposes

. * 'Proente.t at the tws isA Annual conferenec:reuicd August. lHa. motion the difference between these values will be Iwiec tResearch I.nnneer. the actual shaft friction.

2G1 IS.4 Transactions . l'ol. 3. No. 4 L

n ,

The dynamic torque for butterfly valves is a function Combining Equations (7). (8), and (0) of the fluid reaction forces acting on the valve disc. It would be diflicult to determme these forces by purely To = B,2B2 BsB,nD'AP 00) r 4 A unal)tical techniques. Esperimental determination ofIhe ~

pressures and velocity profiles in the immediate area of or the dise w ould al3o be quite dillicult. Ilowever.if a cont rol 3 volume is selected so the bounitaries are points of known To = K D AP i (10-A) pressure and selocity,an analysis of these forces can be where made from the change in fluid momentum through this control volume.- y,2 B 2B, Ben To A,i =

4

=

p33p 00@

INCOMPRESSIBLE FLOW Equation 00-11)is defined as the dimensionless torque An expression for dynamic torque is developed cocflicient which can be determined experimentally from assuming incompiessible flow. This torque is a function tests conducted with incompressible flow, of the fluid reaction force.T,and a moment arm. D,which is a characteristic dimension of the vahe disc.

COMPRESSIBLE FLOW To = f(T,D) 0)

Using the fluid momentum upproach, the force, T is The dynamic torque for butterfly valvesis proportional to the mass flo.v rate and velocity change through a g:ven by:

selected control volume for both compressible and f = AfAl' (2) i.ncompressible flow (i.e., To x AfAl'). Therefore, the approach used to obtain an expression for this torque wl ere assuming incompressible flow can be extended to T = sum of external forces acting on fluid compressible flow by re-defining these two sariab!es.

Af = mass flow rate First, assume that the ulocity at the valve disc. IL, Al'= fluid selocity change through the control is proportional to the velocity change through the control solume voh'me. Then, the dynamic torque can be expressed as The mass flow rate, Af.is gisen by 7 , x Aft; g 1)

' Af = PA l' (3) The velocity at the valve disc is given by

~

' By using a proportionality constant, B i,'the mass flow Af 02) rate can also be defmed as Ik = psA Af = B A(PAP)n (4) 11y combining Equations U1) and 02) the dynamic Equations (3) and (41 are combm.ed to obtain the follow- torque is shown to vary directly as the square of the mass ing expression for fluid selocit): flow rate and imersely with the fluid density at the vahe V = B i(AP/p)n: (5) disc.

The velocity change through the control volume, Al', Af*

To x 03) in Equation (2)can be expressed in terms of the velocity at the valve disc by use of a proportionality constant, B 2 Determining the flow rate of a compressible fluid F = B 2Af t' (6) through a control vahe by unal 3tical techniques is quite difficult because of valve geometry.The major problem By substituting the expressions for mass flow rate is to establish the pressure differential between the vahe Equation (4) and fluid velocity Equation (5)into Equa, inlet and the vena contracta. However, b) definmg the lion (6) the force on the-valtc' disc is physical system in which the valve is installed to conbrm F=B,8BAAP2 (7) with specifications given by the Fluid Controls Institute (FCI),'2' empirical relationships developed specitically For a given valve sire, the flow area. A, for any angle of for determining flow rate for control valves can be disc rotation,0,can be written as considered. Several such empirical relationships have frp2 been developed: however. only one, the Universal Gas A = Bq (8) Siring Equation. has been shown to accurately define the flow rate for any valve configuration. This equation The force. T.aets upon a moment arm u hich is a function is gisen by of the dise diameter. D. Now, the dy namic torque can be utitten as [EU ~ 5M4 A.l7 e f- (14)

I *

/-

0 = \ 67.P C C C, sin f C \2 Ps.es i 2 Tn = B 3FD . (9) 2B2 ISA Tnmwctiwn . l'ol 8. Nm 4

Equation (14) can be rewritten to cbtain an equivalent incompressible flow. ,

expression for mass flow rate. -g'g' 2 r _

'To = K D'P, sin: 0 . (24) 5 -09 64 r Jf = 1.06/p,P C,C C, sin ~9.M[AP~

(15) For convenience the form of Equation (24)is simplified, To = K D 8AP, i (25)

The sine function in Equations (14) and (15) is used t where derme the transition beiween incompressible flow occur-

-CC 8 2-a sina g g ring at low pressure ratios (AP/Ps) and critical flow. Ap, p, Let _59.64,

~59.M

~

Equation (26) is defined as the pressure differential 0= (16) contributing to the dynamic torque on butterfly valves 7' "' with conditions of compressible flow.

Rewriting Equation (15)in the following manner:

EXPERIMENTAL RESULTS M = 1.06/piPC C C,F i (17) 1he first step in the experimental evaluation was to The factor, T is bounded by the following: establish the dimensionless torque coefficient. K , as a F = sin 0 for 0 < n/2 function of valve disc rotation as defined by Equation (18) (10-B). A test was conducted on a 4-in. vahe under the F = 1.0 for 0 2 n/2 following controlled conditions:

By substituting Equation (17) for the mass flow rate in 1. The valve was installed in a 4-in. test line with a Equation (13). the dynamic torque for a gisen valve is minimum of 12 pipe diameters of straight pipe given by upstream.

pi P(C C, sin 0): 2. The pressure taps were located according to FCI To oc (19) specifications and attached to the test line

1. d according to specifications in the ASME Power The only parameter in Equation (10) that cannot be and Tc5r Code."'

readily obtained is the density at the valve disc, p,. 3. Water at ambient temperature was used as the Assuming that the change in the ratio of fluid density at .

flowing medium.

A the sahe inlei to fluid density at the vafve dise with 4. The inlet pressure and outlet pressure were held increasing pressure ratio is small relative to the total constant.

change in mass flow rate, the torque expression can be 5. The test was conducted at a low pressure ratio simplified in the following manner: (AP/P, = 0.088) to ensure incompressible flow.

To cc Pi(C i C sin 0)2 (20)

Therefore, for compressible flow:

32 To - K P (C i iC sin 0)2 (21)

For small values of pressure ratio (AP/P ) Equation (21) 28 reduces to the incompressible torque relationship given by Equation (10-A). 24 As AP/P - 0 i 20 sin 0 = 0(radians)

% [\

7, = K (59.64) AP (22)

J h12 The expression in Equation (22) is equivalent to the 3 expression in Equation (10 A): a 8 K:(59.64)'AP = K D'AP x

~

)

K" (23) "

9' 2 0 O" 10 20 30 40 50 60 70 80 90 By substituting the expression in Equation (23) for the ANGLE OF ROTATION.0EGREES coefhetent K in Equation (21). a general expression for dynamic torque for compressible flow is obtained using rigo,. i. oimon.; ni... ,,, . ...,ricient. 4 in. butterfly

- the dimensionless torque coeflicient established for vaivo incompressibt. f sovv: Pe n 100 poio. P m A10 psi.

283 ISA Transactions Vol. B. No. 4

l 180 720

( '

14 0 o-IEST DATA 560 6 iTg;fsp g

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a o K, AP l 4 120 H g 480 o-TEST DATA - AIR 10 0 E 400 80 320

/ \

60 0 240 u

{ 40 J  ;

5 160 f \

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0^ 1 0 10 20 30 40 50 60 70 80 90 0 10 20 30 40 50 60 70 80 90 ANGLE OF ROTATION. DEGREES ANGLE OF ROTATION. DEGREES Figure 2. Dynamic torqun vs. angte of cline rotation 4.in. Figure 4. Dynamic torque vs. angle of disc rotation,8.in.

butterfly valve, compariron of experimental results with butterfly valvn, comparison of experimental results with calculated torque, incompressible flow: P, m 100 psig, calculated torque, incompressible flow: P, a 100 psig, AP = G psi. AP = 5 psi.

Torque measurements were made at selected incre- ment betw een measured torque and the torque calculated ments of disc rotation (0-90'). A transducer. consisting using this coeflicient.

of a steel bar with strain gayes attached. was fixed to the The next step was to verify that the torque coefficient valve shaft and used in conjunction with an oscillograph is indeed appbcable to other valve sizes provided geo-

' to m:asure and record the shaft torque.The data from metric similarity is reasonably well maintained. The

= this test w ere used to determine the dimensionless torque results on Firures 3 and 4 again show very good agree-coefficient plotted as a function ofdne rotation on Figure ment between measured torque and calculated torque

1. The etirves plotted on Figure 2 show escellent agree- for two 8-in. valves.

720 360 640 320 o ko -K[D'oF 4- FEST JATA e [

560 280 INCOMPRESh!BLE' o-g o $ N FLOW

- 480 r Kf, oP -

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I a-1 EST DATA .

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160 80

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/ 10 20 30 40 50 60 70 80 90 0 10 20 30 40 50 60 70 80 90 0 ANGLE OF ROTATION. DEGREES  ! wVE PRESSURE DROP AP . PSI Figure 3. Dynamic torque vs. nnette of disc rotation, S.in. Figure 5. Dynamic torque vs. valve pressure drop, 4 in.

' butterfly valve, comparison of esperimental results with butterfly walve, GO' disc rotatien, comparison of esperi.

, calculated torque, incompressible flow: P, = 100 psig, mental results with calculated torque, compressible flow:

( P, as 214.4 psia, flowing medium a air.

AP = 5 psi.

ISA Toansanium l'ol. 5. No. .! 284

l 5 l e  !

1 90 400 I 80 360 ht 0-TEST DAT o-TEST DATA j

{ h p a-T=K, g -

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ANGLE OF ROTATION. DEGREES o

' l 0 10 20 30 40 50 60 70 60 90 f Fipure G. Dynamic torque vs. engle of disc rotation,4.in.

butterfly valve, cometarison of caperimental results with ANGLE OF ROTATION. DEGREES calculated tort co, compressibic flow: P, a: 114.4 psia, AP er 5 poi ( AP/P, er 0.044G), flowing medium = eir. pg ,g ,g g l

butterfly valvo, comparison of itst results with calculated It should be noted that discs in the two 8 in. valvcs torrue, com stet hible flow: P, = 64.4 psia, AP = 30 psi were of substantially different geometric shape. Using ( AP/P, = c.4ss), (critical flowl flowing mea.um = .ir.

Ihe ratio of disc diameter to hub diameter as an indicalor, these ratios wue 4.56 I cud 3.55.1 for the salves used to similar to the disc in the 4 in. test sahe used to estabbsh

, obtain the data for Figures 3 and 4. respectively. The the torque coemeient. Ki .

cr' difTerence in torque magnitude for thvse vahes with a The estemion of the dynamic torque relation > hip to

~

f ,,= 5 psi pressure dilTerential shown in Figures 3 and 4 is include the effect of nuid compressibihty is accomplished the result of this difference in geometry. The disc in the by defining an effecthe pressure differential as shown in S.in. vahe used for the test in Figure 3 was geometricall'v Equation (25) The cuts es on Figurc $ 3how the tran<ition from incompressibic flow to critical flow with incre.ning pressure ratio for a 4 in. vahe set at 60' dise rotation.

18 0 Here again there is Scry good agreement between the t rque calculated using Equa"on (24) and the esperi-16 0 < mental resuhs. The incompressib!c torque curve is also TfK D'oP shown on Figure 5 to emphasize the clieet of fluid 14 0 a- i "dKtD'o P, n compressibihty.

d120 I t /\ The curves on Figures 6 through 8 are presented to o- TEST'0ATj compare esperimental results with torque calculated Z

100 i[ T i using Equation (24) for full 90* disc rotation. At low pressure ratios, the torque using air as the flowing 80

/" medium is essentially equal to the torque for incom-pressible flow (Fi;;urc 61. As the pressure ratiois incicased.

• - -

• the effect of fluid compressibihty becomes more pro-V 60 nounced as shown in Figure 7. Once critical flow hes been attained, no further increase in torque is reali7ed 40 20 J [  !

by increasing the valve pressure differential as show n on Figure 8.

O' 0

' f 10 20 30 40 50 60 70 80 90

\

CONCLUSIONS ANGLE OF ROTATION. DEGREES , A technique is pr esented w hich ean be used io determine Figure 7. Dynamic torque vs. angte of disc rotation,4.in.

H uM,s gAes w% mn.&

trutterfly valve, comparison of esperimental results with accuracy. The basie torque rel.itionship developed for calcuisted torque, compressible view: P, s- 64.4 psia. incompressible flow i< estended io inclu,le the c0cet of

• . AP sa to psi ( AP/P, as o.1151, t'owing medium == ir. fluid compressibilit). The method presented is des eloped 285 IS.4 Tromsactiom . D'o!. 3, No. 4

m i

e AP, = Pressure dArrmi.il affeetmg d>n. inia torque ming the Univerul Gas Sizm;: Equation to define an effective pressure differential foi the transition frcan O. - f in* rate inmineres iore flu = sni C, = rio. raie mnpeewihte rivid. .cfh t

r incompressible flow to critical flose. Application of this 7 = Absoluit tensiwrature.*K method shows cuellent agreement with experimental 7, i33 ,,,n,e,,,,,p,,,;,.i, ,

. l' . riu.d seloest), en la felt results. ri = l'luid densii3 at upstream preuurc tar. lb in *

1. 8 r, = Fluid densit) at sabe dix.Iblin NOTATION 8

A = Flow area.in Bs.Br. REFERENCES 3, 8, = Constant. of proportionalit3 C, = C,rr, 1. Kctler. I. C, and Salimann. I. F. January 194. Acrod n.smic 3 C, = Correction f.icter for variation in spenfic heat ratin M.* del Tests on liutterfly Vafscs " DrI,.r.frps Ar.,i.9 C, = Gas sirint coi fh.icnt r C, = Ylom coelho:na 3. Reteennaensial S*olunusurv Soamlardsfor JIrasure ment Pro.~r.l.o efor Dcarrmmeng Control s'ori ri.m C.rpent.r.1956. Fluid Contrels C = Nominal sahe dianietcr.in Institute.Inc paper FCI 55 2 T = force.15 G = Specific truity 3. Buresh.) F and Schuder.C. B. October 1964 "The Doctorment K = Dim:nsionlew torque coefficient of a Uniserul Gas Samp Equaison for Control Valsen " lS A f r.mi.

Af = MassI;ow rate.lb s 3:322-328 P, = Inl:t irturt. r'i

• 4 E hm' \ fen'"'em.*nt: In'r!***r"rS ond <ffr.uru' Surris' menu ro che AS Alf; hmer Tc.it Ca.s/cs. AS\tE seport l'TC 19.5.4 1939 of' = % abc pressure d.Jerentral. pse

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ISA Toamartions l'ol. 8. No. 4 2CG