ML20024C158
| ML20024C158 | |
| Person / Time | |
|---|---|
| Site: | Limerick  |
| Issue date: | 06/30/1983 |
| From: | Krueger J, Sansone R BURMAH TECHNICAL SERVICES, INC. (FORMERLY CLOW CORP.) |
| To: | |
| Shared Package | |
| ML20024C157 | List: |
| References | |
| 6-06-83, 6-6-83, NUDOCS 8307120333 | |
| Download: ML20024C158 (158) | |
Text
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CLOW l Clow Corporation do Chestnut Avenue 312 789 8930 Engineered Products Division Westmont, IL 60559 l
l PURGE AND VENT VALVE OPERABILITY QUALIFICATION ANALYSIS Report No. 6-06-83 PREPARED FOR PHILADELPHIA ELECTRIC C0.
LIMERICK GENERATING STATION UNIT 1 by James E. Krueger
/
Robert C. Sansone June 1983 Work performed under Bechtel Purchase Order Number 8031-P-144 AC, Rev.1 items 1.1, 1.3, 1.5, 1.7, 1.9, 1.11, 1.13 Clow Job Numbers: 82-2053-01, 02, 03, 04, 05, 06, and 07(N)
This report covers Valve Mark Nos: M0-57-109, 112, 115, 135, 147, 161, 162, 163, 164, and A0-57-104, 114, 121, 123, 124, 131.
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1 8307120333 830707 -
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i CERTIFICATION This is to certify that all valves (Tag Nos. MO-57-109, M0-57-112, MO-57-115, M0-57-135, M0-57-147, M0-57-161, M0-57-162, M0-57-163, M0-57-164, A0-57-104, A0-57-114, A0-57-121, A0-57-123, .
A0-57-124,A0-57-131) have been evaluated for operability under the installed conditions indicated in Bechtel Material Requisition 8031-P-144 Rev. I and accompanying specifications as amended by Clow exceptions and processed Bechtel SDDR's. The information contained in this report is the result of complete and carefully conducted analyses and to the best of our knowledge is true and correct in all respects. The information presented in combination with the supporting documents referenced, represents a demonstrated qualification of the subject valves to the best of our knowledge for the required service application.
Paper written and analyses by bw I buuA (dmes E. Krueger '
M/f/J 3 Design Eng. Mgr., Nuclear Clow Corp.
w w Robert C. Sansone bff93 Design Engineer Clow corp.
Paper reviewed and approved M'
' Theodore E. Thygse[/ ~
Professional Engineer Registration No. 062-034780 State of Illinois
ii TABLE OF CONTENTS Page LIST OF TABLES y LIST OF FIGURES vii
1.0 INTRODUCTION
1 ,
1.1 Testing Performed 2 1.2 Qualification Method 4 2.0 DESIGN OF VALVE AND ACTUATOR ASSEMBLY 9 2.1 Valve Design 9 2.1.1 Geometry 9 2.1.2 Materials 12 2.1.3 Operation 21 2.2 Actuator Design (Pneumatic / Spring Return) 25 2.2.1 Geometry 25 2.2.2 Actuator Design Materials 30 2.2.3 Actuator and Valve Operation 31 (Pneumatic Spring Return) 2.2.3.1 Actuators and Accessories Supplied 31 2.2.3.2 Pneumatic Actuator Output Torques 34 2.2.3.3 Operating Time 41 2.3 ActuatorDesign(Electric) 42 2.3.1 Basic Electric Actuator Description 42 (Limitorque Design) -
2.3.2 Nuclear Electric Actuator Materials 53 of Construction 2.3.3 Actuator And Valve Operation 53 2.3.3.1 Description 53 )
1
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I iii TABLE OF CONTENTS (con't)
Page 2.3.3.2 Actuator Output Torques And 58 Operating Times Design Require-ments 2.3.3.3 Operating Time 60 ,
3.0 VALVE OPERATING AND INSTALLATION REQUIREMENTS 61 3.1 Valve Operating Conditions 61 3.2 Valve Installation Configurations 63 4.0 VALVE STRUCTURAL INTEGRITY UNDER SEISMIC AND 69 OPERATIONAL LOADINGS 4.1 Valve Frequency And Stress Analysis 69 4.2 Bettis Actuator Resonant Frequency Test 70 4.3 Asco Solenoid Valve Resonant Frequency Test 70 4.4 Static Load Test During Simulated LOCA Flow 71 I
4.5 Bettis Actuator Seismic / Hydrodynamic 72 J Operability Test 5.0 VALVE AERODYNAMIC TORQUES 78 5.1 Model Tests 79 5.1.2 Tests With An Upstream Elbow 85 5.1.3 Tests With Two Valves In Series 86 .
1 5.1.4 Downstream Piping Effects 93 l 5.2 Model Data Verification 94 5.3 Application of Model Aerodynamic Test to 95 Full Size Valve Operability 5.3.1 Valve Operating Times Expected in 95 Service 5.3.2 Aerodynamic Torques For Valves As 96 Installed
iv TABLE OF CONTENTS (con't)
Page 5.3.3 Conclusions Concerning Valve 112 Operability 115 ,
6.0 VALVE SEALING CHARACTERISTICS 6.1 Normal Sealing 115 6.2 Long Term Sealing 117 6.3 Debris Effects On Sealing 118 6.4 Sealing Under Temperature Variations 119 120
7.0 REFERENCES
APPENDIX A APPENDIX B APPENDIX C
\
v LIST OF TABLES TABLE TITLE PAGE 1 ACTUATOR ACCESSORIES 33 2 GUARANTEED TORQUE RATIOS 34 3 VALVE BENCH TEST OPERATING TIMES (AIR OPERATED) 41 4 EFFICIENCY BREAK-AWAY VALUES 48 5 ELECTRIC ACTUATOR TORQUE CHARACTERISTICS 59 6 VALVE BENCH TEST OPERATING TIMES (MOTOR OPERATED 60 SEISMIC LOADINGS FOR ALL VALVES 61 7
8 PRESSURE DIFFERENTIALS APPLIED TO VALVES 62 62 9 ALLOWED SEAT LEAKAGE RATES LOWEST VALVE RESONANT FREQUENCIES 73 10 CONDITION APPLIED FOR STRESS ANALYSIS 73 11 12 ALLOWED STRESS 73
SUMMARY
OF ALLOWABLE STRESSES - 4" VALVE 74 13 14
SUMMARY
OF ALLOWABLE STRESSES - 6" VALVE 75 15
SUMMARY
OF ALLOWABLE STRESSES - 18" VALVE 76 16
SUMMARY
OF ALLOWABLE STRESSES - 24" VALVE 77 17 TEST VALVE SCALED SIZES (CRITICAL ELEMENTS) 83 COMPARISON OF PRODUCTION VALVES TO VALVE MODEL SIZES 84 18 (CRITICAL ELEMENTS) 97 19 NORMAL FLOW CALCULATIONS 4" VALVE NORMAL FLOW CALCULATIONS 6" VALVE 98 l 20 99 21 NORMAL FLOW CALCULATIONS 18" VALVE NORMAL FLOW CALCULATIONS 24" VALVE 100 22
l vi LIST OF TABLES (con't)
TABLE TITLE PAGE 23 EMERGENCY FLOW CALCULATIONS 4" VALVE 101 24 EMERGENCY FLOW CALCULATIONS 6" VALVE 102 25 EMERGENCY FLOW CALCULATIONS 18" VALVE 103 ,
26 EMERGENCY FLOW CALCULATIONS 24" VALVE 104 27 VALVE NO. MO-57-161 (4") PREDICTED TORQ'JE 107 28 VALVE NO. M0-57-163 (4") PREDICTED TORQUE 107 29 VALVE NO. M0-57-109 & A0-57-121 (6") PREDICTED TORQUE 108 30 VALVE NO. M0-57-162, M0-57-164, & A0-57-131 (6") 108 PREDICTED TORQUE 31 VALVE NO. A0-57-104 (18") PREDICTED TORQUE 109 32 VALVE NO. M0-57-112 (18") PREDICTED TORQUE 109 33 VALVE NO. A0-57-114 (24") PREDICTED TORQUE 110 34 VALVE NO. A0-57-123 & A0-57-124 (24") PREDICTED TORQUE 110 35 VALVE NO. M0-57-115, M0-57-135, & M0-57-147 (24") 111 PREDICTED TORQUE 36 FNEUMATIC ACTUATED VALVE TORQUES 113 37 ELECTRIC ACTUATED VALVE TORQUES 114 38 VALVE SEALING CHARACTERISTICS 116
,m - * = . m se e
vii LIST OF FIGURES FIGURE TITLE PAGE 1 TRICENTRIC VALVE OFFSETS 10 2 4" MOTOR OPERATED VALVE ASSEMBLY AND HATERIALS 14 3 6" MOTOR OPERATED VALVE ASSEMBLY AND MATERIALS 15 4 6" AIR OPERATED VALVE ASSEMBLY AND MATERIALS 16 5 18" MOTOR OPERATED VALVE ASSEMBLY AND MATERIALS 17 6 18" AIR OPERATER VALVE ASSEMBLY AND MATERIALS 18 7 24" MOTOR OPERATED VALVE ASSEMBLY AND MATERIALS 19 8 24" AIR OPERATED VALVE ASSEMBLY AND MATERIALS 20 9 DISC WITH CLOSING FORCES APPLIED 23 10 ACTUATOR SCOTCH YOKE DESIGN 26 11 TYPICAL TORQUE OUTPUT FOR DOUBLE ACTING SCOTCH YOKE 28 ACTUATOR 12 FAIL SAFE, SPRING RETURN ACTUATOR DESIGN 28 13 TYPICAL TORQUE OUTPUT CURVES FOR A SPRING RETURN 29 ACTUATOR 14A CALCULATED TORQUE DATA T312 SR5 35 14B CALCULATED TORQUE PLOT T312 SR5 36 15A CALCULATED TORQUE DATA T820 SR5 37 38 ISB CALCULATED TORQUE PLOT T820 SR5 16A CALCULATED TORQUE DATA T820 SR4 39 16B CALCULATED TORQUE PLOT T820 SR4 40 A TYPICAL HBC GEAR OPERATOR 43 17 45 18 HOBC THRU H3BC COMPONENT LISTING H4BC THRU H7BC COMPONENT LISTING 46 19
viii LIST OF FIGURES (con't)
FIGURE TITLE PAGE 20 SMB MAJOR COMPONENTS 49
- - 21 HELICAL GEAR SET 51 1
22 WORM GEAR SET 51 23 SMB - INTERNAL ASSEMBLY SHOWING POSITION OF LIMIT 52 SWITCHES 24 TORQUE SWITCH ACTIVATION 56 25 INSTALLED ORIENTATION OF 4" VALVE MO-57-161 64 26 INSTALLED ORIENTATION OF 4" VALVE M0-57-163 64 27 INSTALLED ORIENTATION OF 24' VALVE A0-57-131 65 28 INSTALLED ORIENTATION OF 6" VALVE MO-57-109 65 29 INSTALLED ORIENTATION OF 24" VALVE A0-57-121 65 30 INSTALLED ORIENTATION OF 24" VALVE M0-57-164 66 31 INSTALLED ORIENTATION OF 18" VALVE M0-57-162 66 32 INSTALLED ORIENTATION OF 24" VALVE A0-57-114 AND 67 MO-57-115 33 INSTALLED ORIENTATION OF 18" VALVE A0-57-104 AND 67 M0-57-112 34 gINSTALLED ORIENTATION OF 24" VALVE A0-57-124 AND 68 M0-57-147 35 INSTALLED ORIENTATION OF 24" VALVE A0-57-123 AND 68 M0-57-135 36 POSSIBLE ORIENTATION OF TWO CLOW VALVES INSTALLED 87 IN SERIES 37 WATER TABLE STUDY OF CHOKED FLOW PATTERN WITH DISC 88 0
FULL OPEN (90 )
38 WATER TABLE STUDY OF CH0KED FLOW PATTERN WITH DISC 89 '
PARTIALLY OPEN (60 )
l 1
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ix LIST OF FIGURES (con't)-
P'GE FIGURE TITLE 90 39 WATER TABLE STUDY0 OF CHOKED FLOW PATTERN WITH DISC PARTIALLY OPEN (40 )
91 40A TEE WITH FLOW FROM TWO SIDES 91 40B TEE WITH FLOW FROM ONE SIDE 92 41 VALVE ORIENTATIONS RELATIVE TO UPSTREAM ELB0W
Page 1
- 1. INTROD'JCTION The Nuclear Regulatory Commission has, since 1979, been highly concerned about the operability of purge and vent valves during certain postulated occurrences. Their study in this area
- has shown that many valves were designed only to operate under normal flow requirements. For a postulated loss of coolant accident, such valves may fail to close in the time required to prevent discharge of radioactive gases to the outside environment.
Such a failure could exceed 10 CFR guidelines and present a significant hazard to the health of persons in the area.
NRC Branch Technical Position CSB 6-4 gives some backgroud on operations of purge and vent systems and basic requirements for their design. For the valves used in such systems, further guide-lines are provided in " Guidelines for Demonstration of Operability of Purge and Vent Valves", which was provided to nuclear plant operators by an NRC letter in September 1979. This set of guide-lines covers twenty-one points (less two) which are to be addressed by the plant operator. This paper addresses those items which may be answered by the valve manufacturer based on the conditions provided by the plant operator for the postulated loss of coolant accident.
This paper describes the design of Clow's Tricentric butterfly valve, the Bettis pneumatic actuator, and Limitorque electric l l
actuator used to operate the valve. In addition descriptions of I various tests performed to determine flow and torque characteristics l
l Page 2 l
l and application of this test data to the installed condition of the subject valves are presented. Information as to the structural integrity of the valve and operator assembly under seismic and other inplant loadings are also presented. This information, in ,
combination with the supporting detailed technical reports (see 7.0 references), represents a demonstrated qualification of the subject valves to the best of our knowledge for the required service application.
1.1 Testing Performed Clow became involved with design of butterfly valves specifically for purge and vent containment isolation early in 1981. A test program was initiated to determine the mass flow and aerodynamic torque characteristics of the Tricentric butterfly valve design. Tests were performed for 12", 24", 48", and 96" scale model valves (scaled to 3" pipe size) in a straight pipe run for both unchoked and choked flow regimes. Pressure ratios for choking, flow coefficients for mass flow, and aerodynamic torque coefficients were determined in these experiments. The experi-mental set ups met the ISA test requirements for compressible flow measurement. All measurements were automatically read, digitized, and recorded on magnetic tape. The obtained data was then evaluated by other computer programs. Subsequently, a computer program, CVAP was developed using the measured data base to predict flow and torque values for full size valves in a straight run. l
Page 3 In the Spring of 1981. Clow personnel met with representatives of the NRC to review the test program to that point and to obtain recommendations for additional testing. As a result, Clow and it's fluid dynamic consultant set up two additional programs to determine ,
how the aerodynamic torque characteristics of the Tricentric valve varied with installed piping conditions. For such effects of both upstream and downstream piping elements (elbows, tees, reducers, etc.) were considered. From results of backpressure tests performed in the first set of experiments and water table studies previously l
done by Clow, it was determined that upstream piping elements would present a worst case condition. Further, due to the numerous types l of upstream elements (upstream elbow (mitered, 90 , other angles, short radius, long radius), tees, reducers), a worst case had to I be selected for evaluation. A 90 mitered elbow was selected due to the fact that this element presented the worst separated flow region at the inner corner and biased a major portion of the flow 1
to the outer corner. A second set of tests was developed to i obtain information about the effect on each other of two valves in series (the common plant installed practive). Due to the fact that each experiment required an increasing amount of test combin-ations, the experiments were done in a phased approach.
The upstream elbow tests were performed first for a scale model of a 12" valve in 3 orientations relative to the elbow and and a 3 spacings (2, 4, & 8 diameters) from the' elbow. From the results a worst case was determined to occur at 2 diameters.
Page 4 Thus the scale models of the 24" and 48" valves were tested only at 2 diameters. Upstream elbow effects diminished significantly at 4 diameters and were barely detectable at 8 diameters.
From these results, the two valves in series tests were restricted to spacings of 2 and 4 diameters. As in the elbow experiments, the worst case occurred at 2 diameters and at 4 diameters the results approached those for the single valve experiments.
To substantiate the model tests and show the validity of scaling the model data to full size valves Clow performed a ,
i choked flow operational test of a full size 12" valve with a pneumatic spring return actuator at Vought Corp., Dallas, Texas, in November of 1981 (see the appendix for a basic description). .
The test showed that the valve would operate under the choked l
flow tast conditions, that mass flows were as predicted, and that l
use of the CVAP program to predict torques was a conservative method (peak measured torque was approximately 65% of that pre- ,
l dicted). The test also incorporated a static 11.0 g load to the I actuator simulating a severe seismic / hydrodynamic induced loading.
t 1
It further validated the directional effects of aerodynamic torque (in the test all torques tended to close the valve) as measured in model tests.
1.2 Qualification Method Clow provides certific: tion of operability 'of valves produced for purge and vent containment isolation service by a combination of tests and analysis. The following items are considered and 1--__ . - - - -
Page 5 covered in this and supplemental reports.
A. Environmental All portions of the Clow Tricentric is of completely metallic construction other than stem packings and the '
asbestos seal laminations. The valve seals by metal to metal contact between the seat and seal. The asbestos seal laminations used to separate the SST laminations do contain a SBR binder which may degrade under radiation but the asbestos is uneffected. Further, the asbestos laminations are shielded by the SST laminations and disc components. Although the asbestos may become embrittled on the periphery, the valve will still per-form its sealing function (see Radiation Sensitivity Analysis Report Wyle 17629-01). The packings will per-form their function under the required environment as long as they are replaced at recommended intervals.
Actuators used on the valves are qualified for the environment by the actuator manufacturer to codes, standards, or test procedures accepted by the valve buyer.
B. Structural (For Seismic and Other Loadings)
Clow provides for each valve design, a finite element analysis of the valve structure and hand calculations of selected components. These analyses show the valve to be constructed within ASME Section III requirements and that elements not covered by the code are designed with adequate safety margin. Analyses can be found in this l
e Page 6 I
Qualification Report, the code required Design Report, and the Structural Analysis Report. The elements considered by these reports include:
- 1. Valve body
- 2. Valve disc
- 3. Valve disc shaft
- 4. Valve disc shaft connection
- a. Dist ear
- b. Drive keys
- c. Dowel pin (retains shaft from hydrostatic end load only)
- 5. Actuator mounting structure
- a. Adaptor flange
- b. Bolting Actuators are qualified separately by the .
manufacturer by generic test results.
C. Operability Under Flow Operability under maximum flow conditions is based on a combination of a bench test of each unit (timed test with no flow) and an analysis of the torque characteristics of the subject valve. The bench test shows the closing cycle time when no aerodynamic torque is imposed. This data, combined with conservative (see assumptions below) calculations of the aerodynamic torque, is used to show the valve will close in the required time. Bench tests of actuators and valve assemblies include operation during
Page 7 worst case conditions (minimum voltage, air supply, or maximum backpressure for pneumatic actuators if applicable).
The following method is used to show operability:
- 1. Determine no flow worst case operating time from ,
- 2. Using Clow program CVAP calculate aerodynamic torques for straight pipe conditions.
- 3. Determine a torque modification factor based on the installed (from buyer prints) or a worst case upstream piping condition using the mitered elbow or two valves in series test data.
- 4. Determine predicted torque values for all disc angles based on 2 and 3 above.
- 5. Provide tabulation or plot of actuator output torque for all actuator angles
- 6. Show that actuator output provides sufficient margin to evercome aerodynamic and other torques (bearing, packing, disc wt.) to close the valve.
- 7. From the above data, actuator type, and Vought full size test valve data, project a closing rate under the conditions analyzed above.
In the above calculations, the following assumptions are employed:
- a. Containment pressure is at a m'aximum value and I
full flow is developed before the valve starts to close. !
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Page 8 ,
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- b. The pressure downstream of the valve is atmo- )
spheric. In the elbow experiment it was noted that downstream elbows may choke before the valve for certain disc angles, producing a higher back- ,
pressure and lower torques,
- c. Upstream piping components may produce a less severe torque condition than the experimental element (mitered elbow worse than radius elbow) used as a basis for the analysis,
- d. Torque coefficients used in the CVAP program are worse case values. In the experiments a band of coefficients was observed with some dependence on pressure ratio. The high end of the band was used in the CVAP program,
- e. Scaling of torques to larger size valves by the 03method may be largely conservative as was shown by the Vought Test.
The net result of all such calculations and tests to date, continue to show that the design and sizing of all components use in the valve or the actuator exceed the aerodynamic closure requirements as a result of designing for suitable torques to seat and seal the valve.
I
Page 9 2.0 DESIGN OF VALVE AND ACTUATOR ASSEMBLY 2.1 Valve Design 2.1.1 Geometry The Tricentric valve uses a geometry that is unique not only to purge valves but to butterfly valves in general. This feature gives the Tricentric functional characteristics which are-desirable in purge valve applications. Thru use of a conical sealing surface with, the cone axis offset from the pipe axis and a rotation point selected so that it is offset from both the pipe axis and the seal plane, a metal to metal seal can be obtained. (Fig. 1) The sealing is a result of normal forces ,
acting between the sealing surfaces rather than sealing due to surface interference typical of other butterfly valves with elastomeric seals.
One of the major advantages of the conical seal design is that it provides a non-jamming action. This characteristic results from controlling the cone angle so the angle of friction of the material is exceeded. This has been proven in actual l 1
tests similar to the test described here: 1 A 20 inch Tricentric wafer valve was closed by applying 20,000 in.lbs. of seating torque. Then the unseating torque was measured. This was repeated 3 times to determine an average value for the unseating
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Page 11 torque. The test was repeated with the seating torque increased by 10,000 in.lbs. increments until a maximum seating torque of 100,000 in.lbs. had been achieved. !
During the entire test, the seat seal interface was dry (highest angle of friction) and no pressure was applied to the valve. The smallest value of torque that could be accurately measured was 1000 in.lbs. and at no time was more than 1000 in.lbs. required to unseat the valve regardless of the seating torque applied.
Since the shaft is offset in 2 directions, one from the pipe axis and one from the seal plane, 2 performance advantages result. ,
The first is the sealing surface is continuous thru 360 degrees with no interruptions from the shaft penetration. This eliminates the leakage and wear associated with the shaft penetration areas.
The second advantage comes from the shaft being offset (eccentric) from the pipe axis. This eccentricity produces unequal areas about the rotation point, so when the valve is closed and pressure is applied to the shaft side of the disc (normal direction), a closing moment results. This will result in increased sealing forces between the seat-seal interface as pressure increases.
This force, in combination with the mechanical torque produced by the actuator, results in the tight sealing capability achieved with the Tricentric. A definite relationship between these b
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Page 12 2 offsets is required to provide a valve that has no binding or interference problems as the seal is rotated out of the seat.
This relationship is detemined analytically to provide the best performance without overdesigning the valve components. ,
All of these features have been incorporated into the lugged wafer body that results in a very rugged and sturdy valve design capable of meeting or exceeding all the requirements set forth in the specification.
2.1.2 Materials A complete list of valve component materials used on Bechtel s
Purchase Order 8031-P-144-AC may be found on the General Arrange-ment Drawings (D-0699 thru D-0705) which follow this section.
Since purge and vent valves must perform safety related functions not only during normal conditions but also during and after upset, emergency end faulted conditions, the material selections were based on a worst case event. Because the valves are required to prevent discharge of radioactive gases to the outside environment during a LOCA, the seat and seal materials are critical to the operation of the valves. During normal operation the valves are exposed to the air in the containment and outside air, but during a LOCA the media may be made up of AMd steam ai; M which may be radioactive and at elevated temperatures. The seat material selected for this application was SA479 316L SST. The 316 grade was selected due
Page 13 to its corrosion resistance and ability to withstand all of the possible medias that may come in contact with the seat. The L grade of 316 SST was further specified because the seat is welded to the body (SA516 GR70) and the L grade has a lower carbon content that will reduce the carbide precipitation in the heat affected zone of the seat. The seal is a laminate of 316 SST and asbestos.
The 316 SST was chosen in the " straight" grade since no welding is done on the seal. The asbestos used is made of John Manville style 60 or Klinger K-61 material . The laminated type seal was selected for its ability to seal with less torque than would be required for a solid seal. The laminate allows each SST member to act independently and to conform to the contour of the machined seat as seating torque is applied. The asbestos member not only allows each SST member to act independently but also reduces the seal area in contact with the seal and therefore, results in application of higher normal stresses to the seal for any given seating torque.
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Page 29 Since the output of the unit is a function of the thrust applied, a new torque output curve must be used because the air cylinder not only moves the " Scotch Yoke" but must now also compress the spring. A typical torque output graph is shown here for both the pressure stroke and the spring return stroke.
A description of actual output torque values will be presented in the Operation Section.
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FIGURE 13 - Typical torque output curves for a spring return actuator 1
Page 30 2.2.2 Actuator Design Materials The Bettis actuators used for this job are the T series actuators. These were further specified to be the N version for nuclear service and qualified per IEEE 323-1974, ,
IEEE 344-1975, and IEEE 382. These actuators incorporate use of special materials for nuclear service as listed below.
Special Material:
Grease - Mobil 28 Seals - Ethylene Propylene (certified to 1.4 x 10 8 rads)
Internal cylinder coating - Molybdenum disulfide Yoke pin and rollers - Ryton coated It should also be noted that since these units are of the fail safe type, the spring is a critical safety component.
All springs supplied on this order were 100% magnaflux inspected to insure the spring quality.
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Page 31 2.2.3 Actuator and Valve Operation (Pneumatic / Spring Return) 2.2.3.1 Actuators and Accessories Supplied A complete list of all accessories used on each pneumatically actuated valve can be found in Table 1 and each is further described here.
An Asco solenoid valve (s) is used on each actuator to control l l
the air supply to the actuator and, to " dump" the air in the cylinder
)
which allows the valve to open or close as required. The solenoid valves are 3 way, internal piloted diaphragm valves. The solenoid valves are controlled by a 120 VAC coil. When the coil is de-energized by intentional or faulted conditions, the cylinder port is allowed to l discharge through the exhaust port and thereby allow the spring return actuator to perform its required function. When the coil is energized, the supply pressure is directed into the cylinder and rotates the valve in a direction opposite to spring induced rotation. Two solenoid valve models are used, one is a NP831664E. This valve is designated for use in nuclear power applications which consists of providing IEEE compliance and a waterproof solenoid enclosure.
It is also a high flow valve which has 1/2 in. NPT ports and a 5/8 in. orifice. All elastomeric materials of construction are Ethylene Propylene material. The other solenoid valve used is a NP8316E34E which is identical to the NPB31664E except the port size is 1 in. NPT and the orifice is 1 in.
Limit switches are also provided, mounted on the actuator to indicate full open or closed position. One of each model no.
switch is provided, one set for the open position and the other set for the . closed position. The switch model Nos. are Namco
.. ___~________-________~______________.
. - - ~'
Page 32 EA 180-31302 and EA 180-32302 which are DPDT switcht; with 2 NO and 2 NC contacts and are quick make-quick break type.
The switches meet NEMA 1, 4, and 13 and also all applicable IEEE requirements. The switches are of the spring return type .
With one model being CW operation and the other CCW operation.
Both switches use the same lever arm which is a Nam:0 model EL-060-56500.
L _
w t
TABLE 1 PNEUMATIC ACTUATED UNITS 1
Actuator Accessories l Fail-sa fe Asco Namco limit switches i and lever arm ;
Bettis Rotation Fail- Solenoid Valv2 Actuator (viewed sa fe Valve Model Nos. l 512] Mark Clow Model from top Valve Model (2 closed position switches) Other Job No. No. of unit) Position No. (2 open position switches) Accessories (in.) Nos.
6 A0-57-121 82-2053-03(N) NT312- CW Close NP831664E EA 180-31302 L.S. Rosedale Filter SR5 (Oty. 1) Y6-3/4-25-B A0-57-131 EA 180-32302 L.S. Fisher Regulator 95H-41 EL 060-56500 L.A. Hoffman Enclosure A-1008 CHNF A0-57-104 82-2053-05(N) NT820- CW Close NP8316E34E EA 180-31302 L.S. Rosedale Filter 18 SR5 (Qty. 2) Y15-40-80 EA 180-32302 L.S. Fisher Regulator 95H-53 EL 060-56500 L.A. Hoffman Enclosure A-1008 CHNF 24 A0-57-114 82-2053-07(N) NT820- CW Close NP8316E34E EA 180-31302 L.S. Rosedale Filter SR4 (Qty. 2) Y15-40-80 EA 180-32302 L.S. Fisher Regulator A0-57-123 95H-53 EL 010-53337 L.A. Hoffman Enclosure ?
A0-57-124 A-1008 CHNF %
U\
I
Page 34 2.2.3.2 Pneumatic Actuator Outpct Torques The torque plots provided in this section represent the
, calculated output torque of the actuators for the spring and various supply pressures shown. The only listed guaranteed .
output torque that Bettis provides is for the yoke arm at 0 degrees and the spring fully extended. The ratio of guaranteed torque to calculated torque is shown below for the three actuator sizes used.
Table 2 Actuator Guaranteed Torque / Calculated Model Torque %
NT312-SR5 5,510/5,810 .95 NT820-SR5 63,300/62,160 -
1.02
~
NT820-SR4 95,500/93,098 .1.03 The graphs which follow show how the torque output varies for the pressure stroke as a function of supply pressure. It can also be seen that the spring output torque is not a function of supply pressure. The graphs also demonstrate that the output torque (pressure on spring stroke) is a function of yoke position.
The graphs provided are based on the numerical data provided from the actuator manufacturer.
s a
FIGURE 14A Page 35 l i
~312 SR5 DA"A:PU" CYLINDER DIAMETER Cin)= 12.33 CENTER OR TIE 8AR DIAMETER (in)= 0.875 PIST0H ROD DIAMETER (an)= 1.375 HUMBER OF FIST 0HS = 1 ,
MOMENT ARM (in)= 2.612 SPRING LOAD A (1bs)= 1396 SPRIPG LOAD C (Ibs)= 3409 8REAK EFFICIENCY (1)= 70 RUHHING EFFICIENCY (1) = B5 ENDlHG EFFICIEHCY (1) = 74 PRESSURES (psi) = 40 50 60 70 ACTUATOR TYPE,C8=1,HD=2,T=3,04P=4, = 3 YOKE ARM SPRING PRESSURE PRESSURE PRESSURE PRESSURE EFFICIEHCY ANGLE TOROUE TCROUE TOROUE TOROUE TOROUE SPR. PRES.
(degrees) Can 1b) ( 40) psi ( 50)pst ( 60)psa ( 70) psi 1 1 0 5810 12884 17479 22073 26668 74 70 5 5719 10910 15004 19098 23191 77 73 10 5608 9453 13166 16880 20594 79 76 15 5508 8353 11777 15201 18625 81 78 20 5437 7509 10713 13917 17120 82 80 25 5403 6855 9894 12933 15973 83 82 33 5411 6346 9266 12187 15107 84 83 35 5466 5949 8791 11632 14474 85 84 40 5573 5644 8442 11240 14038 85 85 45 5738 5413 8200 10988 13776 85 85 50 5971 5245 8055 10866 13676 85 85 55 6281 5131 7998 10865 13733 84 85 60 6687 5065 8026 10987 13948 83 84 65 7210 5041 8138 11235 14332 82 83 70 7883 5053 8336 11619 19901 80 82 75 8753 5096 8626 12155 15685 78 81 l
80 9887 5157 9013 12868 16723 76 79 l 85 11389 5219 9504 13788 18073 73 77 90 13421 5242 10100 14957 19815 70 74
a T312 SR5
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FIGURE ISA Page 37
~82Z S 5 cus ;s : x >;-
CYLINDER DIAMETEk (in)= 19.58
. CENTER OR TIE BAR DIAMETER (an)= 1.000 PISTON ROD DIAMETER Can)= 1.750 HUMBER OF PIST0HS = 1 MOMENT ARM (in)= 8.000 SPRING LOAD A (1bs)= 5250 SFRING LOAD 8 (Ibs)= 7834 BREAK EFFICIEHCY (1)= 70 RUNHING EFFICIENCY (1) =
85 ENDING EFFICIEHCY (1) =
74 PRESSURES (psi) =
40 50 60 70 ACTUATOR TYPE.C8=1,HD=2,T=3,RsP=4, =
3 YOKE ARM SPRING PRESSURE PRESSURE PRESSURE ANGLE TORQUE PRESSURE EFFICIEHdY TOROUE TORQUE TORQUE TORQUE SPR. PRES.
(degrees) (in Ib) ( 40) psi ( 50) psi ( 60)psa ( 70)psa
___...............___................... 1 1 0 62160 74313 107591 140869 174148 74 70 5 57004 64133 93781 123429 153078 77 73 to 52975 56555 83451 110348 137244 79 76 15 49868 50819 75617 100415 125213 81 78 20 47525 46431 69635 92839 116043 82 80 25 45834 43067 65078 87090 109102 83 82 30 44719 40502- 61655 82808 103961 84 83 35 44129 38586 59166 79746 100326 85 84 40 44042 37211 57475 77740 98004 85 85 45 44454 36307 56498 76688 96878 SS 85 50 45388 35828 56183 76538 D6893 85 85 55 46890 35750 56516 77282 98048 84 85 60 49038 36065 57511 78956 200401 83 84 65 51948 36784 59213 81642 104072 82 83 70 55794 37931 61706 85480 199254 80 82 75 60825 39552 65116 90680 116244 78 81 80 67408 41709 69631 97552 125474 76 79 85 76096 44482 75514 106547 137500 73 77 90 87741 47965 83145 118324 153504 70 74
Te20 sRS Page 38 FIGURE ISB IN.___,___,______,______,___.______,
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! Page 39 FIGURE 16A sy
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CYLlHDER DIAMETER Cin)= 19.58 j
. _! CENTER OR TIE BAR DIAMETER (in)= 1.000 i PISTON ROD DIAMETER (in)= 1,750 -
MUMBER OF PISTONS = 1 MOMENT ARM (in)= 8.000 SPRING LOAD A (1bs)= 7863 SPRING LDAD 8 (Ibs)= 12902 BREAK EFFICIEHCY (1)= 70
. RUHHING EFFICIENCY (1) = 85 ENDIHG EFFICIEHCY (1) = 74 PRESSURES (pst) = 60 70 80 90 ACTUATOR TYPE,C8=1,HD=2,T=3.RsP=4, = 3 YOKE ARM SPRING PRESSURE PRESSURE PRESSURE PRESSURE EFFICIENCY AHGLE TORQUE TORQUE TOROUE TORQUE TOROUE SPR. PRES.
(degrees) (in Ib) C 60)psa ( 70)ps: ( BO)ps: C 90)ps 1 1 0 93098 111604 144882 178160 211438 74 70 5 86357 95385 125034 154682 184330 77 73 10 80989 83363 110260 137156 164052 79 76 15 76813 74277 99075 123873 148671 81 78 20 73674 67317 90521 113725 136928 82 80 25 71452 61949 83961 105973 127984 83 82 30 70064 57809 78961 100114 1212G7 84 83 35 69457 54644 75224 95804 116384 85 84 40 69615 52280 72544 92808 113073 85 85 45 70551 50591 70781 90972 111162 85 85 50 72313 49493 69848 90202 110557 85 85 55 74990 48928 6 % 94 90460 111226 84 85 60 78720 48863 70308 91754 113199 83 84 65 83710 49279 71708 94137 116567 82 83 70 90256 50171 73945 97719 121494 80 82 75 51540 77104 102668 128232 78 81 98792 80 109953 53385 81307 109229 137151 76 79 85 55678 86711 117744 148777 73 77 124695 90 58319 33499 128679 163859 70 74 144502
1820 SR4 Page 40 FIGURE 16B
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Pago 41 2.2.3.3 Operating Time Bench Test - The following is a summary of the operating times recorded during the operational test performed on each valve.
The tests were performed using a 100 psig air supply with a
- maximum flow rate of approximately 70 SCFM. There was no flow throug'h the valve during this test. I TABLE 3 Valve Bettis Opening Closing Mark no. Size Actuator Time
- Time of Valve (inch) Model No. Sec. Sec.
l A0-57-121 6 NT312-SR5 # #
A0-57-131 6 2.7 2.6 A0-57-104 18 NT-820-SR5 22.1 2.5 ,
A0-57-114 24 NT-820-SR4 # #
A0-57-123 24 26.9 1.8 A0-57-124 24 26.1 1.7
- 0pening times were restricted by Clow test set up (Air hose used had approximately 3/8" 1.D.)
- will be furnished at a later date For a description of operating time for valve Serial No. 80-8170-03-01 during a LOCA and Seismic Simulation Test refer to the Vought Corp. Report (reference 7.0 ). The Vought Test denonstrated whentherewasflowthroughthevalve,theaerodynamictorhueaided closure thus reducing closing time.
page 42 '
2.3 Actuator Design (Electric) 2.3.1 Basic Electric Actuator Description (Limitorque Design)
The electric actuator is a device by which electrical energy is converted to a controllable rotary (90 0) motion through the use of a motor, gearing, and electrical control elements.
This type actuator has a constant torque output capability over the entire output rotation of 90 0.
The actuator assembly consists of two major subassemblies, these are the manual type HBC and the electric type SMB. The HBC is directly coupled to the valve shaft and provides the output torque to position the valve disc in the required position. The SMB is coupled to the input shaft of the HBC and provides the required torque input to the HBC as well as containing the motor, electrical control elements, and a manual override feature.
A typical HBC is shown in Figure 17.
The HBC is a worm gear type unit that has a self locking gear set. A self locking gear set has a specified input shaft, and output drive. The output drive cannot be rotated to produce a rotation of the input shaft.
This self locking feature prevents the valve disc from
" driving" or rotating the HBC input shaft when fluid dynamics forces act upon the valve disc. The HBC also contains mechanical stops for the open and closed position limits of the actuator .
Since the Tricentric is a torque seated valve and the valve seat serves as the closed position stop, the mechanical closed position
Page 43 A TYPICAL HBC GEAROPERATOR I
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Page 44 l
l stop in the HBC is adjusted to be used as a backup stop only.
There are two types of mechanical stops used on HBC ectuators.
The type of stop used depends on the unit size, with the H0BC thru H3BC using the " Hex Nut" type stops and the H4BC thru H7BC using the "Stop Screw" type stops. The HBC units also have an indicator arrow on the pointer cap that is to show disc position.
A cutaway drawing showing the construction and component Parts for the H0BC thru the H3BC is shown in Figure 18 and the H4BC thru the H7BC is shown in figure 19.
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Page 45 l
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\. e e 'e 'e PARTS LIST PC. NO. DESCRIPTION PC. NO. DESCRIPTION PC. NO DESCRIPTION 1 HOUSING 8 POINTE R 15 WORM SHAFT O RING 2 HOUSING COVER 9 POINTER CAP 16 STOD SCREW COVE R 3 END CAP 10 END & THRU CAP GASKET 17 STOP SCREW & LOCv,5CREV.
4 TH AU CAP 11 HSG COVER GASKET 18 hex STOP NUT 5 DRIVE SLEEVE & WORM GEAR 12 DRIVE SLEEVE BUSHING 19 LIMIT STOP HOUSING 6 WORM SHAFT 13 WORM SHAFT BE ARING 20 CAP LIMIT STOP HSG 7 SPLINE ADAPTER 14 drive SLEEVE O RING 21 LivlT STOP HSG GASKET FIGURE 18 - H0BC THRU H3BC COMPONENT LISTING
Page 46 H4BC - H7BC O O 6 q g- ~ -
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PARTS LIST PC. NO DESCRIPTION PC. NO. DESCRIPTION 1 HOUSING PC NO. DESCRIPTION ;
B POINTER
_2 HOUSnNG COVE R 15 WORM SHACT O RtNG .
9 POINTE R CAP 16 STOP SCREW COVE R 3 ENO CAP 10 END a THRU CAP GASKET 17 STOP SCREW A LOCKSC AEW 4 THRU CAP 11 HSG COVER GASKET S DRIVE SLEEVE & WORM GEAR 12 DRIVE SLEEVE BUSHING 6 WORM SHAFT 13 WORM SHAFT BEARING 7 SPLINE ADAPTER 14 DRIVE SLEEVE O A I FIGURE 19 - H4BC THRU H7BC COMPONENT LISTING
Page 47 Since the HBC is a wom gear device,the output torque will be constant over the operating range of approximately
~
90 for a constant torque input. The output torque is directly proportional to the input torque, gear ratio, and efficiency of the unit. The formulas used for detemining torque output or required torque input are given below in formulas 1 and 2 j respectively.
NOMENCLATURE T out = Output torque (inch-pounds)
Tin = Input torque (inch-pounds)
R = Wom gear ratio (dimensionless)
E = Unit efficiency (dimensionless) l Formula 1 Tout =
Tin xRxE Formula 2 Tni,t T" I
=
RxE
- _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ - _ _ .__L
Page 48 The maximum output torque of an HBC is dependent on the structural strength of the HBC unit. Table 4 below lists the unit type and size, maximum torque output, gear ratio, and the efficiency. The efficiencies listed are the " Break-away" values since these represent a worst case value used in sizing the units.
TABLE 4 Maximum j Unit Type Output Torque Gear Efficiency j
& Size (in.lbs.) Ratio Break-away HBC-0 5,340 71:1 0.26 HBC-1 15,600 70:1 0.28 HBC-2 26,400 70:1 0.26 HBC-3 67,800 70:1 0.33 l
HBC-4 153,600 60:1 0.34 HBC-5 235,000 65:1 0.34 HBC-6 552,000 66:1 0.36 HBC-7 760,000 69:1 0.36 i
j A typical SMB is shown in Figure 20.
Page 49 M0T R OUTPUT pr. MANUAL DRIVE
,pen n .{
- DECLUTCH LEVER Eh r5 s f .
- x. *
= ,'
a.
(' \ j '
\
m . >h,
- m ts KANUAL HANDWHEEL MANUAL S TORQJE SWITCHES FIGURE 20 - SMB MAJOR COMPONENTS
Page 50 The electric motor on a SMB drives the worm shaft thru a helical gear set (Figure 21 ) with the pinion being fixed to the motor shaft. This helical gear set makes up the first reduction from the motor. The worm shaf t drives the worm which drives the .
worm gear. This worm gear set (Figure 22) which is usually self locking, makes the second reduction of the SMB. The worm gear drives the output of the SMB usually thru a " hammer blow" device to form the final drive output which is then coupled to the input shaft of the HBC. The torque switches and limit switches are driven off the worm shaf t and provide for both torque and position control of the actuator. The complete drive including the torque switches and limit switches, is shown in Figure 23 . The SMB also provides for manual operation in the event of an electrical power failure. The handwheel on the SMB can be engaged by pulling the declutch lever to the manual position. The SMS also has an electrical enclosure box were all electrical components, including torque switches, limit switches, space heaters, and terminal strips are located.
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MOTOR '
PINION ,,
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FIGURE 21 -Helical Gear Set ,
I WORM 33
-, e .,
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6 FIGURE 22 worm ce.r se
Page 52 l
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LIMIT l WORM SWITCH
~
TORQUE ~
=
SWITCH A tRPEl
(. -- , ,s a .
, HELICAL GEAR
- l 3' THAMMER BLOW FINAL I . ,
Wg j, DEVICE WORM GEAR T
FIGURE 23 - SMB -Internal assembly showing position of Limit Switches
- i l
l l
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. - i-_ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _
Page MM 2.3.2 Nuclear Electric Actuator Materials of Construction The Limitorque actuators furnished for this project are all SMB electric actuators with HBC manual units. These actuators were specified for nuclear service per Bechtel Spec. .
No. 8031-G-11 and qualified per IEEE 323-1974, IEEE 344-1975, and IEEE 382-1972. These actuators incorporate use of special materials for nuclear service as listed below. i l
SMB seals - Viton I HBC seals - Viton Lubricant - Exxon Nebula EP-0 or EP-1 Limitswitch gearing lubricant - Mobi. 28 Exterior paint - Carbozinc 11 Limitswitch rotor & base - Fribrite Torque switch - Fribrite Limitswitch gear box - Bronze Terminal strip - Marathon 300 Gaskets - Anchorite 2.3.3 Actuator And Valve Operation 2.3.3.1 Description The actuators supplied for this project are not self contained units, and therefore, require customer supplied control devices.
The customer supplied control system must be compatible with the actuators furnished to assure proper operation and prevent damage to the actuators. During operation, the valve must be able' to perform several functions that include closing tightly, opening fully, and stopping at any intermediate position. The control of these functions is accomplished thru the use of a motor control unit used in conjunction with the torque and limit switches.
l D
Page 54 Since the motor controller is a customer provided unit its selection and design cannot be covered in this report.
During operation of the valve from any position to the full
~
open position, the open limit switch will trip allowing the motor .
controller to stop the valve disc in the open position. Wired in series with the open limit switch is a torque switch (es) which protects the valve and actuator from damage due to overtorquing.
This torque switch is adjusted and set to meet the design require-ments of the valve and actuator. In the event of an open limit switch failure, the SMB would drive the HBC against it's mechanical stop. At that time, the torque would increase causing the torque switch to trip and stop the actuator. The mechanical stop is a protection means only and should not be used for normal operation.
When operating the valve from any position to the full closed position, two switches will trip unlike going to the full open position. The first switch that will trip is the closed limit switch. This switch will not stop the valve from closing, but serves another function. The closed limit switch is set to trip at approximately 3 degrees from the full closed position. When this switch trips. two of the contacts on the switch will parallel the open torque switch. While this switch is tripped, the open torque switch (es) is bypassed and allows full actuator output torque to be applied to the valve when opening the first 3 degrees.
At the closed position, the valve disc seal will make contact with
the seat. As this occurs, the actuator output torque will increase until the closed torque switch (es) setting is exceeded and causes the closed torque switch (es) to trip allowing the motor controller to stop the valve disc in the closed position. In the closing direction, the torque switch (es) is the only electrical device used to stop the actuator. In the event the closed torque switch (es) failed, the actuator motor would go to locked rotor until any thermal overload device would trip out. If any of the above failure conditions were to occur, it would indicate that severe damage had occurred to the valve or actuator.
Shown in Figure 24 is a general view of the torque switch and limit switch assemblies.
g . ..
Page 56
' TOROUE SMTCH -
(see front a bo right) opTA+
9 de:
m --
j qq
.. w j
1..WOAM OPEN TOE!
ADJUS?"ENT I, WOvivtNT 5
,7 N
BELL!viLLE <h Lau,5 FACK #
5 ! d Y CLOSE TCA3'JE ADJUSTVEN1 TOROUE S'MTCH (front vie
- c.a cetaa FIGURE 24 - TORQUE SWITCH ACTIVATION
Page 57 Several of the SMB units supplied for this project do have double torque switches, however, when the switches are wired in parallel per the customer specification, they do not provide re-dundancy. When the switches are wired parallel, both switches must trip to stop the actuator output. Wiring the torque switches in this manner may lead to potential problems. If one switch under-goes a mechanical failure and cannot trip, even though the other switch may trip, it will not stop the actuator and locked rotor will occur.
Another item, which must be addressed in the SMB units, is the non-self locking worm gear sets in the SMB-2 units. Although most SMB units are self locking, the SMB-2 units could not be made self locking and still meet the required operating times.
A problem may occu- in the SMB-2 when the torque switch trips in the closed position. Since the torque switch (es) is driven by the worm shaft, when the worm shaft moves back after applying the required torque, the closed torque switch (es) may remake contact.
When this occurs, the torque output of the HBC unit is not affected, since it has self locking gearing; therefore the valve will still be fully closed. It is possible, after the torque switch remakes contact, to electrically drive the HBC further closed. This could result in excessive torque output of the HBC, which could damage the actuator and/or the valve itself. It is therefore, necessary that the motor controller be designed with sufficient protection to prevent this from happening.
l l
1 I
l Page 58 At any time when the actuator is not being operated electric-ally, the manual override may be engaged. To engage the handwheel, it is necessary to pull the declutch lever to the manual position.
The handwheel may then be used to position the valve disc in the required position. The declutch lever will automatically disengage when any electrical actuation occurs.
2.3.3.2 Actuator Output Torques And Operating Times Design Requirements The design requirement for operating times was 5 seconds or less. This was achieved through the selection of the proper gear ratios and motor RPM used in the actuators.
The output torque of the actuators is dependent on several variables that include motor sizing, gear ratios, and torque switch setting. Of these three variables, two are fixed by the sizing of the actuator and provide no adjustment. The torque switch setting is an adjustable. variable that allows precise setting of the sensed output torque.
Due to the fast operating times required, one other effect must be considered and that is inertia. An article describing fast closing actuators is included in Appendix C that was prepared by the Limitorque Corp. Inertial effects continue after the torque switch (es) has tripped. Since the motor and gearing are moving at a relatively high speed with a given mass the actuator will continue to increase torque output even though the electric power to the actuator has stopped, a
Page 59 The inertial effects cannot be determined empirically, since the system rigidity plays an important factor. A torque magnification of 1 to 3 times the torque at which the torque switch (es) trips, may be produced. It is therefore, necessary to either design the entire system to withstand this higher torque output, or adjust the torque switch (es) to a lower setting so the net torque produced will not exceed the required amount.
Design and measured actuator output torques are shown in the table below. Subsequent sections describe how these torques relates to safe operation of the subject valves.
TABLE q, ELECTRIC ACTUATOR TORQUE CHARACTERISTICS Torque Torque Switch Torque Torque To Fail Valve Design Trip Switch Applied # Weakest
- Size Torque Torque Setting To Valve Component Sa fety (inches) in-lb Actuator Model (in-lb) Required (in-lb) (Key) Factor
- 4 2,112 SMB-00-10-HIBC 1.0 7,594 +
6 7,800 SMB-00-10-H2BC 1.5 27,585 +
18 63.120 SMB-1-60-H5BC 34,000 1.0 86,000 207,563 2.41 24 135,000 SMB-2-60-H5BC 1.0 362,496 +
(Based on key machanical properties measured by test. Valve operator compatibility j forms supplied separately indicate failure point for key based on mechanical l properties lower than actually used. l
- At normal voltage 460 AC (higher than design or torque switch trip due to inertia).
+uill be furnished at a later date 1
=___-______.
Page 60 2.3.3.3 Operating Time Bench Test - The following is a summary of the operating times recorded during the operational test performed on each valve.
The tests were performed using a minimum and normal voltage ,
s upply. There was no flow through the valve during this test.
TABLE 6 Valve Limitorque Opening Closing Mark No. Size Actuator Time Time of Valve (inch) Model No. Sec. Sec.
Voltage Voltage Min. Norm. Min. Norm.
MO-57-161 4 SMB-00-10-H1BC MO-57-163 4 SMB-00-10-H1BC M0-57-109 6 SMB-00-10-H2BC MO-57-162 6 SMB-00-10-H2BC M0-57-164 6 SMB-00-10-H2BC MO-57-112 18 SMB-1-60-H5BC M0-57-115 24 SMB-2-60-H5BC M0-57-135 24 SMB-2-60-H5BC MO-57-147 24 SMB-2-60-H5BC
- Will be furnished at a later date
'Page 61 3.0 VALVE OPERATING AND INSTALLATION REQUIREMENTS 3.1 Valve Operating Conditions The normal and accident operating conditions for the subject valves are taken from Bechtel Power Corp. Specification 8031-P-144 Rev. 1. Appendix 17. Paragraph 1A and 38. Leakage requirements are per spec. paragraph 10.0. This data is presented in summarized form in Tables 7 thru 9. ,
I TABLE 7 Seismic Loadings For All Valves Acceleration l Values (g) l Condition Loading Condition Horiz. Vert.
Normal operation gravity load only 0.0 1.0 (no seismic acceleration)
Upset All loads per 8031-P-144 4.5 4.5 Emergency Appendix 17. Table C-3 Faulted (worst case basis) g = Acceleration as a fraction of the acceleration due to gravity.
4
'Page 62 TABLE 8 Pressure Differentials Applied to Valves NORMAL OPER. DESIGN OPERATING TEMP. DIFFERENTIAL NORMAL RANGE PRESSURE FLOW FAILURE VALVE PRESSURE (PSIG) SCFM MODE SIZE VALVE MARK NO. (PSIG) (OF) 55 150 in position 4" M0-57-161,163 1.5 65-340 55 150 in position 6" MO-57-109,162,164 1.5 65-340 A0-57-121,131 1.5 65-340 55 150 closed 6"
55 4400 in position 18" MO-57-112 1.5 65-340 1.5 65-340 55 4400 closed 18" A0-57-104 MO-57-115,135,147 1.5 65-340 55 6600 in position 24" A0-57-114,123,124 1.5 65-340 55 6600 clos ed 24" TABLE 9 Allowed Seat Leakage Rates (Per Spec at 5, 25, and 55 PSIG Pneumatic)
VALVE SIZE VALVE MARK N0. ALLOWED LEAKAGE cc/ min 4" M0-57-161,163 . 133 6" MO-57-109,162,164 . 20 A0-57-121.131 20 6" .
M0-57-112 . 60 18" A0-57-104 . 60 18" 24" M0-57-115,135,147 . 80 24" A0-57-114.123,124 . 80 o
e O
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Page 63 3.2 Valve Installation Configurations In addition to the pressure and flow conditions specified in 3.0, the valve performance is effected by the as installed orientation. Upstream and downstream, tees, elbows, reducers, and other valves can effect the aerodynamic torque characteristics of butterfly valves. These effects are discussed in Section 5.0 The installed configurations for the subject valves as derived from Bechtel prints are summarized in Figures 25 thru 35.
with appropriate print references.
NOTE: All valve discs in the figures are shown in the partially open (approximately 20 off of seat) position.
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Page 64 FOR 82-2053(N)
-- 4 "-
MO-57-161
(
0 2-?>
LOCA FLOH 3
- o
= 7 5/8" = 4'5 3/8" (1.9 DIA) (13 DIA)
FIGURE 25 : v8LVE MO-57-161 AS INSTALLED PIPING CONFIGURATION PER DHG. HBB-127-2 REV.10.
DISC ORIENTATION SH0HN IS WORST CASE, NOT NECESSARILY AS INSTALLED.
~ 4 '"
w MD-57-163 I
n M
LOCA FLOW i
i v o
!= 6'4 7/8" (19 DIA)
FIGURE 26 : VALVE MO-57-163 AS INSTALLED PIPING CONFIGURATION PER DHG. HBB-124-2 REV.17.
. DISC ORIENTATION SH0HN IS HORST CASE, NOT NECESSARILY AS INSTALLED.
I i
fug- L DRAWINGS FOR QUALIFICATION REPORT p,g, 33 FOR 82-2053(N)
-24 " = 90* ELBOW m' A0-57-131 e [
(+j s-l 2-:> 1 l LOCA FLOH g l
o l m != 2'5 1/2" (4.9 DIA)
= -
5'3/8" (10 DIA)
FIGURE 27: VALVE A0-57-131 AS INSTALLED PIPING
[ CONFIGURATION PER DHG. HBB-124-1 REV.7.
DISC ORIENTATION SH0HN IS HORST CASE, NOT NECESSARILY AS INSTALLED.
- 6 "- 1 MO-57-109 O -
LOCA FLOW g
Cc>
=
10'9 5/8" = =
1 ' 2 1/ 4 "--
(21.6 DIA) (2.3 DIA)
FIGURE 28: VALVE MO-57-109 AS INSTALLED PIPING CONFIGURATION PER DHG. HBB-125-1 REV.7.
DISC ORIENTATION SH0HN IS HORST CASE, NOT NECESSARILY AS INSTALLED.
-2 4 " a AD-57-121 8
l @
M l 1 6"
l LOCA FLOW u l
Go!= 3'1/4" (6 DIA)
=
.=
7'10 5/8" (15.7 DIA)
=
FIGURE 29: VALVE A0-57-121 AS INSTALLED PIPING CONFIGURATION PER DHG. HBB-125-1 REV.7.
DISC ORIENTATIDN SH0HN IS HORST CASE, NOT NECESSARILY AS INSTALLED.
_= _ ______ _____ _____ _____________ ___
FOR 82-2053(N) Page 66
- 24'G T
90" ELBOW I
r MO-57-164
( -
O l
2-:>
LOCA FLON +j\, 6-u
(
od
=
2'1/4" = -
8'2 1/8" (4 DIA) (16 DIA)
FIGURE 30: VALVE MO-57-164 AS INSTALLED PIPING CONFIGURATION PER DHG. HBB-126-3 REV.7. .
DISC ORIENTATION SH0HN IS WORST CASE, NOT NECESSARILY AS INSTALLED.
-1 8 " =
T MO-57-162 N "
n l l-?> % 6" LOCA FLOW o
/
v0
2'1/4"
l (4 DIA) l FIGURE 31: VALVE MO-57-162 AS INSTALLED PIPING l CONFIGURATION PER DHG. HBB-128-3 REV.5.
l DISC ORIENTATION SHONN IS HORST CASE, l NOT NECESSARILY AS INSTALLED.
l A
FOR 82-2053(N) Page 67
-2 4 "--
Y A0-57-114 MO-57-115 n
O
>> /
LOCA FLOW e\
33'
- 3 ' 1 1 7 / 8 "--
(16.5 DIA) (2 DIA)
FIGURE 32 : VALVES A0-57-114 & MO-57-115 AS INSTALLED PIPING CONFIGURATION PER DHG. HBB-127-1 REV.12. .
DISC ORIENTATIONS SH0HN ARE HORST CASE, NOT NECESSARILY AS INSTALLED.
-1 8 "---
w AO-57-104 MO-57-112
(
LOCA FLOW O 3
- o
= 6' = -
5'S 7/8" -
11' =
(4 DIA) (3.6 DIA) (7.4 DIA)
FIGURE 33 : VALVES A0-57-104 & MO-57-112 AS INSTALLED PIPING CONFIGURATION PER DHG. HBB-128-1 REV.11.
DISC ORIENTATIONS SH0HN ARE HORST CASE, NOT NECESSARILY AS INSTALLED.
L
FOR 82-20b3(N) page 68
-2 4 L.
90" ELBOW A0-57-124 MO-57-147
's c
D-> O
+ 24" l l
LOCA FLOW l
) v d
= 3'9" -
- 3'11 7/8" . 6'0" (1.875 DIA) (2 DIA) (3 DIA)
FIGURE 34: VALVES A0-57-124 & MO-57-147 AS INSTALLED PIPING CONFIGURATION PER DWG. HBB-126-1 REV.7. '
DISC ORIENTATIONS SHOHN ARE WORST CASE, NOT NECESSARILY AS INSTALLED.
--24 b AO-57-123 MO-57-135
( ,
LOCA FLOH I O s
1
--3 ' 1 1 15/1G"- -
3'8 7/8" -
(2 DIA) (1.87 DIA)
FIGURE as: VALVES A0-57-123 & HO-57-135 AS INSTALLED PIPING CONFIGURATION PER DWG. HBB-124-2 REV.17.
DISC ORIENTATIONS SH0HN ARE HORST CASE, NOT NECESSARILY AS INSTALLED.
--l -- -__ . .__L____________________._..____________ _ _ _ _ , . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
lfage (69 4.0 VALVE STRUCTURAL INTEGRITY UNDER SEISMIC AND OPERATIONAL LOADINGS Operability of the subject valves has been verified by a combination of testing and analysis in accord with Bechtel Power Corp. Specification 8031-P-144, Rev.1. Separate reports have been prepared and provided demonstrating suitability of valve components and the assembly. A listing is provided in the references (7.0) at the end of the report. This section summaries the results of such tests and analyses in meeting the conditions as presented in Section 3.0, 4.1 Valve Frequency And Stress Analysis Valve frequency and stress analysis was performed by Patel Engineers, Huntsville, Alabama for each valve size. The analysis -
was made using the ANSYS finite element computer program developed by Swanson Analysis System, Inc., Houston, Pa. This public domain program has had a sufficient history of use to justify its applicability and validity. The analyses were made for the seismic conditions stated in Section 3.0 and for pressure and temperature as specified in Table 11. The obtained lowest l
resonant frequencies for the valve assembly are presented in Table 10. The lowest resonant frequency as determined by a conservative analysis approach is 85 Hz. For stress analysis, allowable stresses were in accord with ASME Section III require-ments and Table C-3 of 8031-P-144 Appendix 17. Table 13 thru 16
n summarize the maximum stresses in the valve elements and how these relate to allowed values.
4.2 Bettis Actuator Resonant Frequency Test A low-level Seismic Vibration T6-t was performed on a NT312-SR5 and NT820-SR4 actuator to determine resonant frequencies. The test .
was performed at NTS* Saugus, Ca. The test program consisted of uniaxial sine sweep testing in each of the three orthogonal axis.
The actuator was instrumented with accelerometers to measure input and response accelerations. The test identified the units structural resonances within the frequency range of 1 to 100 Hz.
This information is supplied by a report under separate cover (see references 7.0) 4.3 Asco Solenoid Valve Resonant Frequency Test A valve actuator solenoid valve Asco model 831664, was subjected to both a sine test and sine beat test in each of three orthogonal test orientations for a previous Clow contract.
In addition, the specimen was tested for leakage prior to and after each test segment (a segment being a test in one of the three orientations). Also, during the test, pressure was applied and measured, and functional operability was monitored.
- National Technical System a
.s _
Wage VB The test demonstrated no major resonance: between one and 130 Hz. One orientation showed a system resonan e between 130 and 140 Hz which was outside of the required operability range.
The sine beat test which consisted of 6260 beats per orientation ,
at 5 to 100 Hz and accelerations of 2.0 to 11.0 g (within test table acceleration limits) showed the solenoid valve to be operable before, during, and after the test. No detectable leakage occurred during any phase of the tests. Although these tests were not performed for the subject contract (no report is submitted for review or approval), the report is available for 1
review at Clow's Westmont Illinois facility.
4.4 Static Load Test During Simulated LOCA Flow .
As part of the operability test performed at Vought (see Reference 7.0 ) an 11.0 g load was applied in each of two orthagonal directions through the approximate center of gravity of the actuator. With the load applied and flow through the valve greater than expected in service, the valve operated within the required time period. This aspect to the test demonstrated that the actuator to valve connection was sufficiently rigid to remain in fully operable under this load. Further details are included in the subject report.
.. n.. . .
. n -n- - . . _
mm1 -
_'_nh.'_ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
Page 72 4.5 Bettis Actuator Seismic / Hydrodynamic Operability Test A Dynamic Test was performed on a NT312-SR5 and NT820-SR4 actuator to demonstrate operability under anticipated loadings which may be encountered in service. The test was performed at NTS, Saugus, Ca., in accord with NTS Test Procedure 528-0951.
The test demonstrated the units would operate as required before, during, and after the test. This information is supplied by a report under separate cover (see Reference 7.0).
i
Page 73 TABLE 10 Lowest Valve Resonant Frequencies (Per Analysis)
VALVE SIZE TYPE FREQUENCIES (Hz) 4 MO 87 6 MO 107 6 A0 105 ,
18 MO 139 l
18 A0 131 24 MO 85*
24 A0 86
- Analysis assumed use of SMB-3 electric actuator which ,
is a worse case than the SMS-2 shipped with the valves.
l TABLE 11 l
Condition Applied For Stress Analysis Body Disc Design -
i Design Seating Design Differential Pressure Pressure pmp. Torque Valve Size Valve Mark Nos. (PSIG) (PSI) F (in-lb) 4 MO-57-161.163 285 65 340 2,112 l 6 MO-57-109,162,164 285 65 340 7,800 l
A0-57-121,131 18 MO-57-112 285 65 340 63,300 A0-57-104 24 M0-57-115,135,147 285 65 340 135,000 A0-57-114.123,124 TABLE 12 Allowed Stress Applicable Membrane Membrane + Bending Valves Condition Stress Limit Stress Limit All valvas Upset 1.1 S 1.65 5 Emergency 1.5 5 1.8 S Faulted 2.0 S 2.40 S S = Stress allowed per ASME, Sect. III, Tables 1-7.1 thru l-7.3 and Appendix XVII 2460 (as applicable)
--- - - --9--------- -----o-- o
~ - -- - - ____-_-_
'f '
Table 13 4" Valve Sunenary of Allowable Stresses (Worst Case) . - - - _ ..-
.. . - . . ~ . . . . . _ . . - - .
El.F. MENT STRESS A1.LOWABl.E S1HESS STNESS 1
NATEMIAl. HATIO IDCATION ( ps i ) val.llE (PER ASHE SECTION (psi)
III, TABl.ES l-7.1 THROUGil 1-7.3) 61 0.24 17500 4109 Valve Body SA 516 GR.70 191 0.19 17500 3309 Disc SA 516 GR.70 202 0.81 34550 28020 Drive Shaft SA 564 Type 630 H-1045 .
212 0.11 17500 1847 Operator Adapter SA 516 P la t e GR.70 f N/A 0.08*
25000 92840 Adapter Plate SA 193 6198:
Bolts GR.87 N/A 0.04*
25000 5616 0 Ope r at or / Adapt er su 193 4871 l t iloi t s CN.87 N/A 0.19 (1.5)(17500) 3758 Cover Plate SA 516 GR. 70 =26250 N/A 0.003* y 25000 32860 m Cover Plate Bolts SA 193 18 i -
to
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'ec t ion III, Appemli x XVil , Subsul ar t icle 2460.
meansk w m m u b ask thmsk htwe. hasta, ihaan6. thdas hada, tahd. head has M m m W fo '
Table 14 6" Valve Sununary of Allowable Streases (W0rst case)
FIATERIAL ALIDWABl.E STRESS STRESS El.EMENT STRESS IDCATION
( ps i ) VAtt!E RAT 10 (PER ASHt. SECTION (psi)
III, TABl.ES 1-7.1 Ti!ROUGil 1-7.3)
Valve Body SA 516 17500 3836 81 0.22 GR.70 Disc SA 516 17500 8817 202 0.50 GR.70 Drive Shaft SA 564 34550 23038 216
- 0.67 rype 630 11-1045 Operator Adapter SA 516 17500 4377 226 0.25 Plate GR.70 Adapter Plate SA 193 25000 150900 N/A 0.15*
N Bolts GR.B7 77641 Operator / Adapter SR 193 25000 17284c N N/A 0.27*
Bolts ( flybr id GR.B7 111941 Worst Case Assembly) ,
Cover Plate SA 516 (1.5)(17500) 6613 N/A 0.39 GR. 70 =26250 ?
Cover Plate Bolts 'iA 193 25000 54900 N N/A 0.Ol* y J;R . B 7 21 1
- Per ASME,Section III, Appenelix XVII, Subsuharticle 2460
- - - - - h .
a .:- --.-. - -4 I t' Table 15 18" Valve Sunenary of Allowable Stresses (worst Case) .
ALIAWABLE STRESS STRESS El.EMENT STRESS IDCATION FIATERIAL (psi) VAttlE NATIO (PER ASME SECTION (psi)
III, TABt.ES I-7.1 TilROUGH I-7.3)
SA 516 17500 2156 81 0.12 Valve Body GR.70 t
SA 516 17500 5043 214 0.29 Disc GR.70 34550 22168 232 0.64 Drive Shaft SA 564 Type 630 11-1045 SA 516 17500 3053 261 0.18 Operator Adapter
< P la t e GR.70 SA 193 25000 9839 N/A 0.09*
Adapter Plate GR.B7 6645 Bolts SR 193 25000 10479 N/A 0.24*
Ope rator / Adapt e r 11806 Bolts (flybrid GR.B7 Worst Case Assembly)
(1.5)(17500) 17641. N/A 0.68 ,
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- Per ASME,Section III, Appendix XVII, Sisbsubartictii 2460
m .- _
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Table 16 e
24" Valve
! Sunenary of Allowable Stresses (Worst Case) e EI.EMENT STRESS ALL0tlABIE STRESS STRESS
! IDCATION HATERIAL RAT 10 (psi) VAI.llE (PER ASME SECTION (psi)
III, TABLES I-7.1 TilROUGil I-7. 3) 4089 106 0.23 valve Body SA 516 17500 i
CR.70 6930 244 0.40 SA $16 17500 i Disc CR.70 '
27263 267 O.79 SA 564 34550 Drive Shaft Type 630 11-1045 5891 296 0.34 SA 516 17500 Operator Adapter Plate CR.70 17215s N/A 0.19*
SA 193 25000 Adapter Plate 8884r N
- GR.B7 I
! Holts
! 155660 N/A 0.61*
SR 193 25000 i ope rator / Adapt er 19076i Bolt s (Limitorque ) CR.B7 1614to N N/A 0.57*
SA 193 25000 Oper ator / Adapt er 18276t Holts (Bet tis ) CR.B7 o 24052 N/A 0.92 SA $16 (1.5)(17500)
- Cover Plate =26250 CR. 70 U 70000 '
N/A 0.Ol*
SA 193 25000 Cover Plat e Bolts 302i GR.B7
- Per ASME, Sect ion III, Appendix VVII, Subsubartic le 2460
Page 78 5.0 VALVE AERODYNAMIC TORQUES Depending upon the valve design, actuator sizing, inplant installed configuration, and operating conditions, aerodynamic
~
torque may be of major concern to valve operability. The magnitude and direction of this torque, which is produced by flow of the medts over the disc, depends on several factors:
- 1. Disc shape
- 2. Pivot shaft location
- 3. Magnitude of differential pressure across the valve
- 4. As installed upstream piping elements (elbows, tees, etc.) including distance and orientation relative to these items.
- 5. As installed downstream piping elements (elbows, tees, length of pipe runs, etc.) including distance and orientation relative to these items.
- 6. Angle of the disc l Clow has done numerous tests of scale models of the Tricentric design and a test of a full size 12 inch production valve. The data obtained in these tests provide a substantial base for predicting aerodynamic torques in full size production valves under various operating conditions.
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Page 79 5.1 Model Tests In 1980, Clow established a program to determine mass flow and aerodynamic torques of the Tricentric design. Exact scale models (see Table 17 ) were designed and built of 150 lb class Tricentric valves of standard design. Scale models of a 12, 24, 48, and 96 inch valve were constructed and tested using University of Illinois facilities under the direction of A.L. Addy, Ph. D. (Engineering Consultant in Fluid Dynamics and Engineering and Associate Head, Department of Mechanical and Industrial Engineering U. of I. at Urbana, Champaign,111.).
The tests were made with air in accord with ISA standards for a straight pipe run flow test. The tests were run at various pressure ratios (upstream to downstream pressure) in" both the choked and non-choked pressure regimes. Very low pressure ratios were also applied to allow correlation to incompressible (liquid) flow in accord with ISA standards. Tests were made with flow in the normal direction for Tricentrics (shaft upstream) and for reverse flow (shaft downstream). Further, several pressure ratios near the choked flow point were applied to determine the point of choking. This test pointed out that the standard rule of thumb (downstream pressure / upstream pressure =
.528) for determining when choking occurs is not valid at all disc angles. The tests showed choking will occur at a ratio of .75 in the full open position and .54 in the near closed l
Page 80 l
position. The test also showed, that although choking prevents 1
the fluid velocity from increasing, aerodynamic torque will rise 1
in a linear fashion in accord with the pressure differential l across the valve in the choked flow regimes.
The models used for testing were made in accord with the Tricentric standard 150 lb class double flange design. This is a fabricated design in which the seat is at a 10 degree angle from a normal to the pipeline axis. Due to the seat i
position, this valve rotates only 800 from closed to full open.
The valves supplied for the subject job uses a similar geometry except the seat is normal to the pipeline axis making this a 900 (k turn) valve design. Therefore, at small open,ing angles ,
(00to 200) there are some differences in torque. For engles over this amount, the aerodynamics are the same. Also, at small angles the torque approaches the value of the pressure area torque (as explained in Section 2.1.3) thus, differences between the two designs are not significant. With reasonable similarity between the test models and the full size valves, the data may be used to predict torque characteristics for produced valves.
From the data base developed by the model tests a computer program CVAP (Clow Valve Analysis Program) was written for use in predicting valve operating characteristics. In this program, mass flow rates are predicted by standard equations for flow I F
~
Page 81 through an ideal converging nozzle adjusted with coefficients developed in the tests. Torques are predicted on the basis of the equation T=CT a P Dy 3 where T = predicted aerodynamic torque (in Ib)
CT = torque coefficient developed in model tests aP = pressure differential across the valve (lb/in )2 Dy = nominal valve diameter (in.)
The test performed on a full size 12" valve showed that the mass flow obtained was within approximately 10% of that predicted by the computer model while torques were much less t,han predicted. .
Torques were on the order of 65% of that predicted which could be correlated by changing the power of 3 to 2.84 in the above equation. The power of 3 used in the equation and in the Program CVAP is a derived value obtained by use of the equations for conservation of momentum for a general control volume.
Thus the program indicates torques which would be higher than those obtained in the actual situation.
Table 17 shows the dimension of critical (to torque conditions) elements of the double flange Tricentric 12, 24, 48, and 96 inch designs and their scaled down dimensions which were used for model construction. Table 18 shows a comparison I between the provided size valves and the interpolated sizes.
,_.--v - _ ______ _ _
Page 82 Linear interpolation was used to predict torque characteristics in Clow Program CVAP, thus a similar interpolation of sizes is applicable for size comparison purposes. It can be seen in the
~
i table that very good (less than 8% deviation) correlation was obtained for torque critical items for the 18" and 24" valves.
For the 6" valves correlation is good (less than 9% deviation) for all critical dimensions other than disc thickness. From test data, greater disc thickness would reduce the potential for torques tending to resist valve closure. For the 4" valves interpolated dimensions are larger than production dimensions for the I.D., A 2, and K 2. This would lead to higher calculated torques than would actually be experienced. The offset E used ,
for calculation is only 69% of the production size which would indicate a lower torque than in service. However, service torques would not be higher than 44% of those indicated in Tables 27 and 28.
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Page 83 TABLE 17 Test Valve Scaled Sizes (Critical Elements)
VALVE SIZE ELEMENT 12" 24" 48" 96" Full Model Full Model Full Model Full Model Size Size Size Size Size Size Size Size I.D. 11.04 3.07 22.62 3.07. 46.00 3.07 96.00 3.07 A2 11.33 2.91 21.89 2.97 45.59 3.04 96.20 3.07 K2 10.80 2.78 20.86 2.83 43.44 2.90 91.66 2.93 Shaft
- Dia. 2.25 .58 3.25 .44 6.0 .40 12.0 .38 Shaft CL to Seal q,, L 2.0 .51 2.69 .36 5.06 .34 7.51 .24 ,
Domed Disc Shape Thickness 1.5 .38 1.88 .25 3.75 .25 11.63 .37 Shaft Offset E + 1.25 .32 .81 .11 1.31 .09 1.18 .04 Shaft Offset LC + 1.67 .43 1.38 .19 2.31 .15 1.66 .05 Ear Width
- 2.25 .58 3.25 .44 6.0 .40 12.0 .38 Ear Height
- 3.38 .87 4.88 .66 9.0 .60 15.25 .49
+ E is offset from disc centerline LC is off from body centerline
- Ear is element welded to disc which shaft is -~ced to.
Note: Full size dimensions are for a Clow Tricentric 150 lb class double flange design.
A2 = Major axis of elliptical seal K2 = Minor axis of elliptical seal E = Offset between shaft axis and disc center (see Figure 2)
LC = Offset between shaft axis and pipe run centerline 1 All dimensions in inches l
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Page 84 TABLE 18 Comparison of Production Valves to Valve Model Sizes (Critical Elements)
VALVE SIZES ELEMENTS 4" 6" 18" 24" Size Ratio Size Ratio Size Ratio Size Ratio
- I.0, 4.026 1.20 6.07 1.09 16.88 1.02 22.62 1.0 3.289 1.15 5.33 1.07 16.08 1.02 21.71 1.01
- A2 3.40 1.12 5.20 1.05 15.70 1.01 21.25 .98
- K2 Shaft Dia. .75 1.17 1.25 1.02 2.5 1.10 3.0 1.08 Sha f t CL to 1.08 Seal Q ,L .75 1.07 1.0 1.16 2.19 1.07 2.50
- Disc Thickness .625 .99 .625 1.43 1.63 1.04 1.75 1.07
- Shaft Offset E .88 .69 .88 .96 .95 1.08 .69 1.07 Shaft Offset LC .91 NA .91 NA 1.00 NA .76 NA Ear Width 1.0 .88 1.5 .85 2.5 1.10 3.50 .93 Ear Height 1.06 1.25 1.56 1.23 3.25 1.27 3.75 1.30
- Elements considered important to torque characteristics
" #D ** * * * **
NOTE: RATIO = production valve size A2 = Major axis of elliptical seal K2 = fiinor axis of elliptical seal E = Offset between shaft axis and disc center (see Figure 2)
LC = Offset between shaf t axis and pipe run centerline All dimensions in inches l
Pago 85 5.1.2 Tests With An Upstream Elbow One element of piping system which has an effect on the .
_ aerodynamic torque of butterfly valves is a turn which may occur with a elbow or a tee. Since numerous types of elbows (short- and long radius, reducing, mitered, etc.) may exist in a particular piping system, it was necessary to determine a worst case condition for testing. It was determined use of a mitered elbow would be a worst case and that this configu-ration had applicability to flow through tees also.
The mitered elbow produces the greatest separated flow region at the inside of the turn and biases the flow to the
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outside corner to a maximum (see photo from water table study Figure 37 ).1 Further, the mitered elbow produces flow patterns f more severe than expected for tee flow (see Figures 37 and 40 ).
The testing performed has given added evidence in support of this assumption. (See report reference 7.0 C-3) Flow around the corner produces a lower local pressure around the inside of the turn and higher local pressure to the outside. This will oppose closure for geometry 1 (see Figure 41) and aid closure for geometry 2.
Based on these considerations, models of a 12", 24", and 48" valve (per Table 17) were tested for torque characteristics.
All valve models were tested for geometries 1, 2, and 3 at 2 diameters downstream from the mitered elbow. In addition, the 1 See reference 7.0 E-3 .
Page 86 12" model was tested at 4 and 8 diameters downstream. The test showed the greatest variation of torque from that obtained for straight-line flow occurred at 2 diameters downstream from the elbow. Differences due to valve orientation were small at 4 diameters downstream and were just detectable at 8 diameters downstream.
For the subject job some valves are installed closer than 2 diameters from an elbow. Since the mitered elbow used in the model tests is a worst case condition and radius type elbows are typically used for in plant installation, use of the test data for 2 diameters downstream for determining installed operability is considered reasonable. If torque operating ,
margins are adequate, this judgement is further justified.
5.1.3 Tests With Two Valves In Series When two valves are installed in series in a pipe run at a relatively close distance (less than 8 diameters) some level of interaction will occur. Several different orientations of the two valves relative to one another are possible as shown in Figure 36. Model tests performed to determine aerodynamic torque characteristics, indicate that orientation 2 with the upstream valve failed (stuck) at 60 open and the downstream valve at full open would represent a worst case condition (highest torques resisting downstream valve closure). These model tests are more A
fully described in a separate report indicated in the references (Section 7.0).
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Page 92 Geometry 1 A I E Offset
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Page 93 5.1.4 Downstream Piping Effects In various tests described in this section, it was necessary
_ to provide downstream piping to discharge the flow. In the conduct of these tests the effects of downstream piping were noted several times. In the straight line tests, a downstream valve was installed to vary back pressure. Any increase in back pressure lowered the torque values. In the elbow tests an elbow was installed 20 or more diameters downstream. It showed that for the 24" and 48" models in the full open position, the down-stream piping would choke before the valve model. This prevented any substantial increase in pressure differential across the valve model even with large increases in upstream pressure, thus the torque was limited. From the piping layouts provided down-stream, piping would provide some degree of back pressure making the assumption (atmospheric pressure downstream) used for calcu-lation of torques conservative.
1
vage 9xj 5.2 Model Data Verification A test of a full size 12" valve was run at Vought's High Speed Wind Tunnel in Dallas, Texas (see reference 7.0 B3) to
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demonstrate operability and substantiate model test data.
The tests demonstrated the valve would operate in the required 5 sccend period. It further showed that torque values weie less than predicted from model data. The valve used for the test incorporated a one piece thru shaft design while the model had a two piece shaft. To verify the torque effect due to this change, another test was made (data not put into a formal report form) in which a 2 piece shaft was installed in place of the thru shaft. The test was made with the disc held in a station-ary position by a manual worm gear type actuator. The result was that the peak torque was the same for both the one and two piece shaft design. The only difference was that the two piece shaft design showed a peak torque closer (by 5 to 10 degrees) to the full open position. A test was also run with the one piece shaft design with the disc held in a stationary position. This was done to provide direct correlation with the model tests which were done in this manner. It also allowed a comparison to the torques measured during the dynamic test with the shaft connected to the pneumatic actuator. A summary of the operability test is included in Appendix B.
i L
Page 95 5.3 Application of Model Aerodynamic Test To Full Size Valvc Operability 5.3.1 Valve Operating Times Expected In Service All valves were designed to close within 5 seconds for
~
flow conditions produced by maximum differential pressure (see 3.0, Table 8 ). These are the maximum conditions expected in the event of a LOCA. The valves were designed to fully open within 5 seconds for conditions of normal flow, though most are capable of opening fully within this time for maximum pressure differential.
All air actuated valves will fail closed through use of a return spring in the actuator. They will open within 5 seconds if the air supply to the actuator is adequate. All electric actuated valves will fail in position on loss of power. The electric ..
actuated valves may operate full open to closed in a pericd greater than 5 seconds due to the design provided by the actuator manu-facturer. Any such deviation has been approved by the customer prior to shipment.
In the Vought test, which used a pneumatic / spring return actuated valve, (Reference 7.0 ) closing times were shown to improve slightly with flow through the valve. Opening times were retarded on the order of 1/2 to 2 seconds depending on flow conditions.
These changes are of a conservative nature since it was necessary to restrict both the valve opening 'and closing air supplies to prevent pressure upstream of the valve from increasing to an unreasonable level during the test. The conduct of the test would suggest that l
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Page 96 opening times in actual service for similar valve / actuator assemblies might be retarded about .3 to .5 (since normal flows are much lower than tested flows) and closing times might be improved by the same amount under maximum differential pressure conditions relative to the Clow bench test data. Since the electric actuated valves have a self locking worm gear set, little change in operating time (.2 sec. or less) would be expected under full flow conditions as compared to no flow conditions.
5.3.2 Aerodynamic Torques For Valves As Installed As described in Section 5.1, torques from straight line model tests can be used to predict full size valve torques by D 3scaling.
Tables 19 thru 26 present torque and other data for the subject valves at various operating conditions. The item of concern for valve operability is TQ (for normal operating conditions, open cycle) and TQA (for maximum operating conditions, closing cycle). All torque values shown are positive, tending to close the valves. The meanings of the other listings can be found in 7.0 References C-1.
To obtain torque conditions for the as installed valves a judgement must be made as to what set of test data most nearly represents the actual conditions.
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, . . .- TABLE 19 - . Page 97 4.., . . .v - , NORMAL FLOW CALCULATIONS 4" VALVE ,
f CASE: BECHTEL. LIMERICK,82 W ~*f4
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- Page 98 i
TABLE 20
. NORMAL FLOW CALCULATIONS 6" VALVE CASE: BECHTEL. LIMERICK,82-2 UN1TS SYSTEM: ES...'
(+
ti. ' -
DATE: 4-19-83
- .' ' 4 - ,' .
SHAFTt US '4 1-PATH: 14.70(PSIA) -
T PSU = 15.70(PSIA) TSU = 609.67(R) 29.0 a;,
. MEDIUM:
CAS = A '
~
GAMMA =
OPTION = 1 1.40 MW =
. FLOW = UF .* .,-e,..
P' C
W80 = , 8/ 150.00(SCFM ) ,
., S, ;,
. l, DV = '6.000(IN) ..
a ;..... - .e s, n... .. . .. - -.
1 . .y%
- y. d., .Y . e n E P u t D4.-a ..i a , .- _ -
. . . .. . - p. .
~- - .
f l i . '- -
- n. : .
'.00(PSICi.
- ie:
g'.":.4.? ?z.> p ..c. SOLUTIbH's. ,.:,., DP80' =
...: .n.,
i , . n. .. . ,. .e. '-
~s~ ,.,
- - 3'l
.9996 , ' rn, .
3, , . . PSD/POU = . .
,t, TOR 1 ' -
DPS/PSU PSU/POU PSC/POU POD /POU ALPHA CF WR
?, ..
.1124
=.- .
.0003 .9999 .7446 .9980
!'g 90.0 .4970 1.0000 .9999 .7430 .9938 .1504 75.0 .4877 .9814 .0003 .1786 9.
.9444 .0003 .9999 .7393 .9904
- 70.0- '.4693 .0003 .9999 .7334 .9878 .1981 ,
.65.0 .4432 .8918 .7252 .9860 .2098 4107 . 8264 .0003 .9999
,. . 60.0
.9999 .7145 .9849 .2148
'55.0 '.3732 .7510 .0003
.9847 .2142 "i'", .6682 .0003 1.0000 .7012 50.0 .3321 1.0000 .6854 .9852 .2092 , ,.'
f[ 45.0 .2887 . 5808 .0003
.9866 .2008 .
.2444 .4917 .0003 .1.0000 .6674 '
} ", . 40.0 1.0000 .6477 .9887 .1905 35.0 .2006 .4036 .0003
.9916 .1794
.3193 .0004 1.0000 .6269 t' 30.0 .1587 .6059 .9953 .1688 25.0 .1200 .2414 .0004 1.0000
.1729 .0004 1.0000 .5860 .9998 .1603 ,
S ,20.0 .0859 .5683 1.0000 .1550 15.0 .0578 .1164 .0004 1.0000 ..l
.0747 .0004 1.0000 .5547 1.0000 .1546
10.0 .0371 1.0000 .1606
.5.0 .0251 .0505 .0004 1.0000 .5464 e
W I
'TO .. '. -.. #
. g ALPHA -
>- ' YCV
. '- (LBM/HR)" ('I. ,4 N-LBF ) (g.- , ., ' * /, f,,y
.' . r 4 ( DE G ) p, .up . '
- z. . ..~n i ;,;,. -( . .' . ), -
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~
5.10 h' . .
672.40
. ,' ~
616". ** *)
'E' ' ' ' -
"" "l$ , .
, ;' 8 0 .'0 .. " ' " ' ' . " 401.09
'. .o .659.89 .14 75.0 * - ~
572.54 '635.03 .17
~70.0 -
599.68 .20 65.0 532.40 484.10 555.70 .21 60.0 504.95 .23 55.0 432.60 .23 379.16 449.28 50.0 390.56 .24 45.0 324.59 270.93 330.65 .23 <-'
40.0 271.41 .22 35.0 220.24 .22 172.79 214.68 30.0 162.34 .20 25.0 129.92 .20 92.63 116.24 20.0 78.24 .19 15.0
- 62.25 .19 39.92 50.20 10.0 33.97 .20 5.0 27.03 . ,
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- Page 99 .
5 TABLE 21 - +
NORMAL FLOW CALCULATIONS 18". VALVE -
- CASE: BECHTEL,LIMERILK,82-2053(N) , ...
DATE: 4-7-83 ,
UNITS SYSTEM: ES <'
PATM: 14.70(PSIA).;. SHAFT: US s; i"
PSU = 15.70(PSIA) TSU = 609.67(R)
- MEDIUMS GAS = A .l.' GAMMA = 1.40 MW = 29.0
- FLOW = UF DPTION = 1 .
', . .'WBO s' 4400.00(SCFM ) ~ ' '
, : DV = 18.000(IN) <
" i .2 , - s g. Q. n . . . r;. .- > . . ,
.e___________ . .
8 .
- 4.7 ;; . .,0 U T P U T .D AT A . ;, .- . , 'i . .
. .; ., f
- .t . .
., i ~ SOLUTION: DP80 =, .03(PSIG) .
,e :
PSD/POU = .9971
- DPS/PSU PSU/POU PSC/POU P0D/P00 r
.9968 .0532 5 80.0 .5747 1.0000 .0019 .9991 .7558 ^
75.0 .5640 .9814 .0020 .9991 .7547 .9926 .1034 70.0 .5427 .9444 .0020 .9992 .7520 .9891 .1360 65.0 .5125 .8918 .0021 .9992 .74.74 .9064 .1546
- ~
60.0 .4749 .8264 .0022 .9993 .7405 .9844 .1627 1629 55.0 .4316 .7510 .0023 .9995 .7306 .9832 . . .
~, .9829 . 1580 50.0 .3840 .6682 . 0025 .9996 .7177
.0026 .9997 .7019 .9833 . 1499
. 45.0 .3338 .5808 *
.9998 . 1405 40.0 .2826 .4917 . .0027 .6831 .9845 ~'
35.0 .2320 .4036 .0027 .9998 .6620 .9865 .1310 : .
30.0 .1935 .3193 .00L .9999 .6394 .9893 .1224 '
25.0 .1387 .2414 .0028 .9999 .6163 .9929 .1152 20.0 .0993 .1729
.0029 1.0000 .5940 .'9973 .1095 15.0 .0669 .1164 .0029 1.0000 .5741 1.0000 .1051 10.0 .0429 .0747 .0029 1.0000 .5586 1.0000 .1014 E.5.0 .0290 .0505 ~.0029. 1.0000 .5491 1.0000 .0972 .
.s.
.',- .. ' ~, ..
- ALPHAl,7,,'., ,, . / '. . , ([ Y C V W
T Q * ,.-
- . ]. ;
j'.' ,
. ,( D E G )
, ...) (LBM/HR) (IN .LBF) . , ,$.
.- . . - . - .v. , . . . . . ,
~ 80.0 6789.91 19723.82 . 9.41
, 75.0 6606.13 19356.80 18.62 70.0 6248.45 18627.55 25.35 '
65.0 5771.90 17590.59 30.12 60.0 5220.21 16300.50 33.26 55.0 4627.45 14811.78 33.00
- 50.0 4022.69 13178.93 35.55 45.0 3425.04 11456.55 35.16 -
40.0 2849.53 9699.13 34.12 35.0 '2306.50 , 7961.22 . 32.71 30.0 1805.31 6297.34 31'.21 25.0 1355.14 4762.02 29.81 20.0 965.68 3409.78 28.61 15.0 648.21 2295.12 27.62 _
10.0 415.34 1472.56 26.70 '
- 25.62 5.0 ,
280.97 -
996.57
COMPUTATIONS COMPLETE-------- ,
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Page 100 M
.. TABLE 22 i.
NORMAL FLOW CALCULATIONS 24". VALVE
. CASE: DECHTEL, LIMERICK,82-2053(N) .
UNITS SYSTEN: ES DATE: 4-7-83 ,
~~
, PATH: '14.70(PSIA) SHAFT US PSU = 15.70(PSIA) TSU = 609.67(R) .
OPTION = 1 ,
W80 = 6600.00(SCFM )
. .s DV = 24.000(IN) l .
. .c..e . . . -v
- i +: . v . .- ;- -..........
.. .,,00TPUT DATA ,
l d.. : . .jl,~.--~~------
. .,' J., ; . r. e : .
- ' SOLUTION: DP80 = .02(PSIG)
PSD/POU = .9981
. ALPHA CF WR DPS/PSU PSU/POU PSC/POU POD /POU TQR1 B0.0 .5970 1.0000 .0012 .9993 .7578 .9973 .0237
. 75.0 .5859 .9814 .0012 .9993 .7569 .9931 .0800
- 70.0 .5638 .9444 .0013 .9994 .7546 .9896 .1148
. 65.0 .5324 .8918 .0014 .9995 .7506 .9869 .1331 60.0 .4934 .8264 .0014 .9995 .7434 .9850 .1392 55.0 .4483 .7510 .0015 .9996 .7346 .9839 .1371 50.0 .3989 .6602 .0016 .9997 .7220 .9836 .1300
- 45.0 .3467 .5808 .0017 .9998 .7062 .9840 .1205 ,
40.0 .2936 .4917 .0017 .9998 .6873 .9853 .1105 35.0 .2410 .4036 .0018 .9999 .6660 .9874 .1014 ,
30.0 .1906 .3193 .0018 .9999 .6429 .9902 .0941
-25.0 .1441 .2414 .0019 1.0000 .6192 .9939 .0885 20.0 .1032 .1729 .0019 1.0000 .5963 .9983 .0843 15.0 .0695 .1164 .0019 1.0000 .5757 1.0000 .0804 10.0 .0446 .0747 .0019 1.0000 .5597 1.0000 .0749
.5.0 .0302 .0505. .0019 1.0000 .5499 1.0000 .0657 ,
' ALPHA
~ ~ '
YCV -
W 2TQ
-(DEG) (...) (LBM/HR) (IN-LBF) 80.0 12000.11 29585.73 6.29 75.0 12435.19 29035.20 21.70 70.0 11738.00 27941.31 32.34 65.0 10818.02 26385.90 39.34 60.0 9758.69 24450.74 43.44 55.0 8629.82 22217.65 45.18 50.0 7483.72 19768.39 45.09 45.0 6359.47 17184.81 43.72 ;
40.0 5283.71 14548.70 41.63 35.0 4271.33 11941.82 39.40 30.0 3340.71 9446.01 -
37.37 25.0 2505.86 7143.03 35.74 20.0 1785.03 5114.67 34.40 15.0 1197.93 3442.68 32.97 10.0 767.38 2208.04 30.04 5.0 ,
519.04 1494.85 27.07
COMPUTATIONS COMPLETE-------- ,.
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TABLE 23 Page 101 EMERGENCY FLOW CALCULATIONS 4" VALVE CASE: BECHTEL, LIMERICK,82-2053(NG ,
DATE: 4-7-83 UNI,TS SYSTEM: ES ,
PATM: 14.70(PSIA) ,
SHAFT: US ,
PSU = 69.70(PSIA) TSU = 799.67(R)
FLOW = CF .
, OPTION = 2 *
, DV = 4.000(IH) ,
' 't .
- i .
- ' *- OUTPUT DATA ,. . 3 :
.c-------,. ,
. , ,.4 f.t .5
. 4. ,,.. . ,,*,, ,.
CHOKING PRESSURE RATIOS: PSC/POU =- .740 DPS/PSU = .217...n.
SOLUTION: W80 = 0.15(LBM/S) -
NOTE: TQ BASED ON DIFFERENTIAL PRESSURE AT ONSET OF CHOKED FLOW'
TQA BASED ON PSU UPSTREAM AND PATM DOWNSTREAM .
PSD/POU = .7397 ALPHA CF WR DPS/PSU PSU/POU PSC/POU POD /POU TQR1 80.0 4710 1.0000 .2171 .9447 .7397 .8491 .1046 75.0 .4623 .9814 .2188 .9469 .7378 .8381 .1404 .,
70.0 .4448 .9444 .2223 .9510 .7338 .8265 .1676 65.0 .4201 .8918 .2268 .9566 .7277 .8148 .1868 60.0 .3893 .8264 .2319 '.9630 .7193 .8032 .1987 '
55.0 .3537 .7510 .2372 .9697 .7084 .7919 .2041 50.0 .3;47 .6682 .2423 .9762 .6951 .7814 .2038 45.0 .2736 .5808 .2469 .9821 .6795 .7718 .1999 ,,
40.0 .2316 .4917 .2508 .9872 .6619 .7634 .1904 35.0 .1901 .4036 .2540 .9914 .6427 .7562 .1796 30.0 .1504 .3193 .2564 .9947 .6225 .7503 .1679 25.0 .1137 .2414 .2581 .9970 .6024 .7458 .1567 .
,, 240.0 .0814 .1729 .2592 .9984 .5832 .7426 .1475 15.0 .0548 .1164 .2598 .9993 .5664 .7406 . 1421 .. .
~
10.0 .0352 .0747 .2601 .9997 ' .5532
~
.7397 . 1423.,
.'5.0 '.0238 .0505 ..2602 .9999 .5456 .7398 .1499 ?fi.
.y
,c . ALPHA '. -
YCV .- W ,
~
TQ - TQA .
- 3,s*k
(DEG) (...) (LBM/HR) .(IN-LBF) (IN-LBF) -
80.0 247.86 29341.05 101.25 368.11 75.0 241.70 20795.05 137.32 495.17 70.0 229.79 27710.23 167.25 593.83 65.0 213.57 26167.67 191.36 665.86 60.0 194.40 24248.53 209.56 713.03 55.0 173.47 22033.90 221.64 737.34 "
50.0 151.70 19604.90 227.58 741.23 45.0 129.86 17042.70 227.64 727.70 40.0 100.51 "14428.37 222.58 700.39 35.0 88.13 11843.06 213.57 663.63 30.0 49.16 9367.89 202.19 622.36 25.0 52.00 7083.95 190.37 582.12 20.0 37.10 5072.37 180.28 548.90 ~
15.0 24.92 3414.21 174.24 529.22 10.0 15.97 2190.57 174.69 529.96 5.0 10.81 1482.49 184.15 558.41 '._
"9" TABLE 24 EMERGENCY TLOW CALCULATIONS 6" VALVE
. CASE: BECHTEL, LIMERICK,82-2053(N) .
DATE: 4-7-83 ,
UNITS SYSTEM: ES PATM 14.70(PSIA) SHAFT: US PSU = 69.70(PSIA) TSU = 799.67(R) .
l'EDIUM: CAS = A GAMMA = 1.40 MW = 29.0 FLOU = CF DPTION = 2 '.
DV = 6.000(IN) '
~
_c , . s ". * ' '
'0UTPUT DATA-
..4,*,.
- 4. _. . . .. -
- p. , CHOKING PRESSURE RATIOS: PSC/POU = .745 DPS/PSU = .206 .
SOLUTION: W80 = 19.47(LBM/S)
NOTE: TQ BASED ON DIFFERENTIAL PRESSURE AT ONSET OF CHOKED FLOU TQA BASED ON PSU UPSTREAM AND PATM DOWNSTREAM PSD/POU = .7446 ALPHA CF WR DPS/PSU PSU/POU PSC/POU POD /POU TQR1
. 80.0 .4970 1.0000 .2062 .9380 .7446 .8564 .0956 75.0 .4877 .9814 .2083 .9404 .7430 .8455 .1334
. 70.0 .4693 .9444 .2122 .9451 .7393 .8339 .1613 i 65.0 .4432 .8918 .2174 .9514 .7334 .8219 .1803 60.0 .4107 .8264 .2233 .9586 .7252 .8099 .1916 55.0 .3732 .7510 .2293 .9661 .7145 .7983 .1961 50.0 .3321 .6682 .2351 .9734 .7012 .7874 .1950 45.0 .2887 .5808 .2403 .9800 .6854 .7774 .1895 40.0 .2444 .4917 .2447 .9858 .6674 .7687 .1808 35.0 .2006 .4036 .2483 .9905 .6477 .7613 .1702 30.0 .1587 .3193 .2510 .9941 .6269 .7553 .1589 25.0 .1200 .2414 .2529 .9966 .6059 .7508 .1482 20.0 .0859 .1729 .2541 .9983 .5860 .7476 .1395 -
15.0 .0578 .1164 .2549 .9992 .5683 .7456 .'1342
! 10.0 .0371 .0747 .2552 .9997 .5547 .7448 .1338 , j. ~~
5.0. 0251 .0005 .2553 .9999 .5464 .7450 .1397
. c.
ALPHA YCV W TQ TQA .
(DEC) (...) (LBM/HR) (IN-LBF) (IN-LBF) 80.0 608.02 70082.34 296.73 1135.55 75.0 592.18 68778.28 419.32 1588.72 70.0 561.76 66187.12 519.23 1930.79 65.0 520.63 62502.62 598.72 2173.08 60.0 472.48 57918.69 658.14 2325.86 55.0 420.36 52628.94 697.25 2399.28 50.0 366.65 46827.16 716.21 2404.18 45.0 313.13 40707.25 716.31 2352.59 40.0 261.16 34462.80 700.17 2257.96 35.0 211.81 28287.68 671.75 2135.15 30.0 166.03 22375.62 636.20 2000.28 25.0 124.76 16920.34 599.48 1970.53 20.0 ,88.96 12115.57 568.05 1763.84 15.0 59.74 0154.99 548.61 1698.72 10.0 38.29 5232.27 547.87 1694.14 5.0 25.90 3540.99 572.52 1769.48
COMPUTATIONS COMPLETE-------- .
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Page 103 TABLE 25 EMERGENCYF, LOW CALCULATIONS 18 VALVE CASE: BECHTEL. LIMERICK,82-2053(N) -
DATE: 4-7-83 ,
UNITS SYSTEN: , ES "-
PATM 14.70(PSIA) SHAFT: US PSU = 69.70(PSIA) TSU = 799.67(R) ,
CAS = A GAMMA = 1.40 MW = 29.0 MEDIUM:
FLOW = CF OPTIDF = 2 ~
DV = 18.000(IN) .
- ~- * '
c -
OUTPUT DATA ,
'( ..
, 'g
.. ~-
.? -
DPS/PSU = .174 CHOKING PRESSURE RATIOS: PSC/POU =
.756 SOLUTION: WBO = 207.00(LBM/S) '
NOTE: TQ BASED ON DIFFERENTIAL PRESSURE AT ONSET OF CHOKED FLOW TQA BASED ON PSU UPSTREAM AND PATM DOWNSTREAM PSD/POU = .7558 .
~
ALPHA CF WR DPS/PSU PSU/POU PSC/POU POD /POU TQR1 -
- 80.0 .5747 1.0000' .1738 .9148 .7558 .8785 .0390 75.0 .5640 .9814 .1769 .9183 .7547 .8676 .0890 .
70.0 .5427 .9444 .1828 .9249 .7520 .8553 .1211 65.0 .5125 .8918 .1905 .9337 .7474 .8422 .1391
- 60.0 .4749 .8264 .1991 .9437 .7405 .8288 .1464 55.0 .4316 .7510 .2078 .9541 .7306 .8156 .1460 50.0 .3840 .6682 .2160 .9640 .7177 .8032 .1404 '
45.0 .3338 .5008 .2233 .9731 .7018 .7920 .1317 40.0 .2826 .4917 .2295 .9809 .6831 .7821 .1218 35.0 .2320 .4036 .2344 .9872 .6620 .7739 .1119 30.0 .1835 .3193 .2381 .9920 .6394 .7673 .1030 25.0 .1387 .2414 .2407 .9955 .6163 .7624 .0956
.5940 , .7591 20.0 .0993 .1729 .2424 .9977 .0898
'. 15.0 .0669 .1164 .2434 .9989 .5741 .7572 ~.0853
- i: 10.0 .0429 .0747 .2439 ' 9996 .5586 .7566 . 0815 ,, *[,
5.0 .0290 .0505 .2440 .999G .5491 .7570 .'0773 ,
it
- =:
~,
g n
ALPHA YCV W TQ TQA 4 (IN-LBF)
(DEG) (...) (LBM/HR) (IN-LBF) l 80.0 7067.39 745186.25 2754.78 12508.30 75.0 6848.29 731319.75 6423.24 28649.73 70.0 6436.71 703767.75 9097.80 39270.39 65.0 5897.96 664591.00 10996.10 45543.58 60.0 5289.47 615849.75 12225.10 48449.50 55.0 4653.79 559604.00 12860.17 48835.41 50.0 4019.45 497913.62 12987.55 47448.19 45.0 3404.63 432840.25 12717.72 44944.75 40.0 2820.83 366443.25 12179.94 41888.17 l l
35.0 2276.20 300783.00 11505.42 38735.11 30.0 1777.64 237919.94 10809.42 35822.37 25.0 1332.31 179914.06 10175.66 33355.00 '
20.0 9,48.55 128824.97 9645.87 3139E.53 15.0 436.39 86712.09 9214.48 29875.41 10.0 407.67 55634.77 8826.97 28564.29 5.0 275.73 37651.42 8381.03 27101.45
COMPUTATIONS COMPLETE-------- 2 ,.. .
. v.4 . * .-; - - ..g :~ : ,. u : . . .q ,
TABLE 26 EMERGENCY FLOW CALCULATIONS 24 VALVE i CASE: BECHTEL, LIMERICK,82-2053(N)
UNITS SYSTEM: ES DATE: 4-7-83 ~~
PATM 14.70(PSIA) SHAFT: US PSU = 69.70(PSIA) TSU = 799.67(R) .
CAS = A GAMMA = 1.40 MW = 29.0 MEDIUM: ,. .
~" OPTION = 2 FLOW = CF
'. DV = 24.000(IN) -
=L -. . .
O' -
0UTPUT DATA e
'.4?! ' l'.Ts... ' ~
~
". g, g; CHOKING' PRESSURE RATIOS: PSC/POU = .758 DPS/PSU = .165 SOLUTION: W80 = 384.99(LBM/S)
>r~
NOTE: TQ BASED ON DIFFERENTIAL PRESSURE AT DNSET OF CHOKED FLOW -
TQA BASED ON PSU UPSTREAM AND PATM DOWNSTREAM PSD/POU = .7578 ALPHA CF WR DPS/PSU PSU/POU PSC/POU POD /POV TQR1 80.0 .5970 1.0000 .1647 .9072 .7578 .8646 .0101 75.0 .5859 .9814 .1683 .9111 .7569 .8737 .0662 70.0 .5638 .9444 .1748 .9103 .7546 .8612 .1005 65.0 .5324 .8?l8 .1834 .9280 .7506 .8476 .1180 60.0 .4934 .8264 .1930 .9389 .7434 .8336 .1234 55.0 .4483 .7510 .2025 .9502 .7346 .8199 .1205 50.0 .3989 .6682 .2115 .9611 .7220 .8069 .1126 ..
45.0 .3467 .5808 .2195 .9709 .7062 .7952 .1024 40.0 .2936 .4917 .2262 .9793 .6873 .7850 .0919 35.0 .2410 .4036 .2316 .9862 .6660 .7764 .0824 30.0 .1906 .3193 .2357 .9914 .6429 .7696 .0747 25.0 .1441 .2414 .2385 .9951 .6192 .7646 .0609 20.0 .1032 .1729 .2403 .9975 .5963 .7612 .0646 15.0 . 0695 .1164 .2414 .9999 .5757 .7592 .0605 10.0 .0446 .0747 .2419 ,.9995 .5597 .7586 ".0551 5.0 .0302 .0505 .2421 .9998 .5499 .7591 .0458 . ,
ALPHA YCV W TQ TQA
.(DEG) (...) (LBM/HR) (IN-LBF) (IN-LBF) 80.0 13516.89 1385976.00 1607.16 7698.29 75.0 13071.36 1360185.50 10779.35 50557.17 70.0 12241.69 1308941.50 17128.80 77308.28 65.0 11169.06 1236076.00 21328.98 91760.06 60.0 9973.56 1145422.00 23735.84 97074.44 55.0 8740.61 1040810.50 24626.73 95952.09 50.0 7523.94 926072.00 24318.80 90721.25 45.0 6355.61 805042.00 23186.74 83352.94 40.0 5254.50 681549.50 21626.09 75433.00 35.0 4233.11 559427.75 19991.84 68115.01 30.0 3302.06 442508.62 18538.34 62079.27 25.0 2472.85 334623.00 17374.76 57489.56 20.0 17.59.68 239602.19 16440.28 53984.20 15.0 1180.26 161276.25 15497.00 50667.03 10.0 755.96 103475.44 14133.72 46113.09 5.0 511.27 70028.09 11771.11 38374.39
CDMPUTATIONS COMPLETE-------- . s, ,
. .t. . . . :" . . .
Page 105 For Figures 26, 28, 29 and 32 (Section 3.2), the configurations for M0-57-163(4"), MO-57-109(6"), A0-57-121(6"), and A0-57-114(24")
indicate upstream piping is at a sufficient distance so that all the valves will respond as if under fully developed straight run pipe flow. Thus, for all these valves the torque modification factor comparing straight line flow to actual flow is 1.0 as indicated in Tables 29, 30. and 25. For Figure 25 (M0-57-161) an upstream radius elbow at 1.9 diameters is indicated. This is reasonably represented 1
in the calculations for the case of an upstream mitered elbow at 2 diameters. A worst case orientation (flow tends to minimize torque 1
aiding closure) was used in performing calculations even though the l valve may be installed in a different orientation. For Figure 27, 30, and 31 an upstream tee is the major element of concern (reducer j ahead of M0-57-162 tends to straighten flow into valve). These valves (M0-57-162, 164, A0-57-131) are best represented by a 1
mitered elbow upstream at 4 diameters. In reality, use of a mitered elbow for calculation here is more severe than would occur in actual service. Also a worst case orientation was assumed. For Figure 33, valve A0-57-104 and upstream mitered elbow at 4 diameters with worst case valve orientation is considered. For Figures 34 and 35, valves A0-57-123 and 124, an upstream mitered elbow at 2 diameters is con-1 sidered. For Figures 32, 33, 35, and 34 valves M0-57-112,115, 135, and 147 a worst case fe.ilure is assumed. Valve orientations relative to one another are assumed worst case a'nd it is assumed 0
the upstreara valve has failed and stopped at the 60 position
Page 106 (a worst case for the downstream valve). This will result in an aerodynamic torque on the downstream valve which tends to resist closure. Such a torque must be overcome by the actuator for the valve to close. For all the subject cases, the resultant torques are summarized in Tables 32 and 35.
The tables show model test valve angle and actual valve angle for the supplied units. There is a 100 difference between these due to the seat angle design differences explained in previous sections. It is reasonable to expect all angles over 20 to be a
~
proper representation of the magnitude and direction of torques.
At 20 or below, the magnitude may differ but the direction is correctly indicated. Since peak torques occur in the 60 to' 80 range, these low end torques are of no consequence.
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Table 27 Valve No. M0-57-l61 (4")
Model Data For Torque Modification: Mitered elbow 2 diameter All torques in In-lbs. upstream Geometry 2 Model Test Actual Torque for Torque Torque for Valve Valve Straight Flow Modification Installed Condition Angle Anale Normal Maximum Factor Normal Maximum 80 90 1 370 .45 1 167 70 80 1 595 .84 1 500 60 70 1 715 '. 9 4 1 672 50 60 1 740 1.0 1 740 40 50 1 700 1.0 1 700 30 40 1 625 1.0 1 625 20 30 1 550 1.0 1 550 10 20 1 530 1.0 1- ' 530 Table 28
'. al ve No. M0-57-163 (4")
Model Data For Torque Modification: Valves under straight line All torques in In-lbs. flow conditions -
Model Test Actual Torque for Torque Torque for Valve Valve Straight Flow Modification Installed Condition Anale Anale Nornal Maximum Factor Normal Maximun 80 90 1 373 1.0 1 320 70 80 1 595 1.0 1 595 60 70 1 715 1.0 1 715 50 60 1 740 1.0 1 740 40 50 1 700 1.0 1 700 30 40 1 625 1.0 1 625 20 30 1 550 1.0 1 550 10 20 1 530 1.0 1 530 k_m
Table 29 Page 108 Valve No. MO-57-109 & A0-57-121' (6")
Model Data For Torque Modification:
All torques in In-lbs, Valves under straight line flow conditions Model Test Actual Torque for Torque Torque for Valve Valve Straight Flow Modification Anale Anale Normal Installed Condition Maximum Factor Normal Maximum 80 90 1 1135 1.0 1 1135 70 80 1 1930 1.0 1 1930 60 70 2325 1 1.0 1 2325 50 60 2400 1
1.0 1 2400 40 50 2260 1 1.0 1 2260 30 40 2000 1 1.0 1 2000 20 30 1 1765 1.0 1 1765 10 20 1 1695 1.0 1 1695 Table 30 Valve No. M0-57-162, M0-57-164, A0-57-131 (6")
Model Data For Torque Modification: Mitered elbow 4 diameter All torques in In-lbs. upstream Geometry 1 Model Test Actual Torque for Torque Torque for Valve Valve Straight Flow Modification Installed Condition Angle Anole Normal Maximum Factor Normal Maximum 80 90 1 1135 .75 1 850 70 80 1 1930 .95 1 1835 60 70 1 2325 1.0 1 2325 50 60 1 2400 1.0 1 2400 40 50 1 2260 1.0 1 2260 30 40 1 2000 1.0 1 2000 20 30 1 1765 1.0 1 1765 10 20 1 1695 1.0 1 1695
Table 31 Page 109 Valve No. A0-57-104 (18")
Model Data For Torque Modification: Mitered elbow 4 diameter All torques in In-lbs. upstream Geometry 2 Model Test Actual Torque for Torque Torque for Valve Valve Straight Flow Modification Installed Condition Anole Anole Normal Maximum Factor Normal Maximum 80 90 9 12,510 .36 -3 -4,503 70 80 25 39,270 .84 +21 32,990 60 70 33 48,450 1.0 33 48,450 50 60 35 47,450 1.0 35 47,450 40 50 34 41,890 1.0 34 41,890 30 40 31 35,820 1.0 31 35,820 20 30 28 31,400 1.0 28 31,400 10 20 27 28,565 1.0 27 28,565 Tabl e 32 Valve No. MO-57-112 (18")
Model Data For Torque Modification: 2 valves in series All torques in In-lbs. upstream valve assumed failed at 600 open position Model Test Actual Torque for Torque Torque for ,
Valve Valve Straight Flow Modification Installed Condition Angle Anale Normal Maximum Factor Normal Maximum
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80 90 9 12,510 .17 -2 -2,127 70 80 25 39,270 +.67 +17 +26,310 60 70 33 48,450 1.0 33 48,450 50 60 35 47,450 1.0 35 47,450 40 50 34 41,890 1.0 34 41,890 30 40' 31 35,820 1.0 31 35,820 20 30 28 31,400 1.0 28 31,400 10 20 27 28,565 1.0 27 28,565
Valve No. A0-57-114 (24")
Model Data For Torque Modification: Straight pipe run All torques in In-lbs. (Upstream elbow at more than 8 dia.)
Model Test Actual Torque for Torque Torque for Valve Valve Straight Flow Modification Installed Condition Angle Anale Nomal Maximum Factor Normal Maximum 80 90 7 7,698 1.0 7 7,698 70 80 32 77,310 1.0 32 77,310 60 70 43 97,075 1.0 43 97,075 50 60 45 90.720 1.0 45 90,720 40 50 42 75,430 1.0 42 75,430 30 40 37 62,080 1.0 37 62,080 20 30 34 53,980 1.0 34 53,980 10 20 30 46,100 1.0 30 46,100 Table 34 Valve No. A0-57-123 and 124 (24")
Model Data For Torque Modification: Mitered elbow 2 diameter All torques in In-lbs. upstream Geometry 1 Model Test Actual Torque for Torque Torque for Valve Valve Straight Flow Modification Installed Condition Angle anale Nomal Maximum Factor Nomal Maximum 80 90 7 7,698 -6.0 -42 -46,188 70 80 32 77,310 1.52 +48 115,970 60 70 43 97,075 1.18 51 114.550 50 60 45 90,720 1.20 54 108,860 40 50 42 75,430 1.16 49 87,500 30 40 37 62,080 1.0 37 62,080 20 30 34 53,980 1.0 34 53,980 10 70 30 46,100 1.0 30 - 46,100
Table 35 Page 111 Valve No. MO-57-115, 135, and 147 (24")
Model Data For Torque Modification: 2 valves in series upstream All torques in In-lbs. valve assumed failed at 600 open position Model
_ Test Actual Torque for Torque Torque for Valve Valve Straight Flow Modification Installed Condition Angle Anale Normal Maximum Fa ctor Normal Maximum 80 90 7 7,698 -1.5 -11 -11,550 70 80 32 77,310 + .48 15 +37,110 60 70 43 97,075 1.0 43 97,075 50 60 45 90.720 1.0 45 90,720 40 50 42 75,430 1.0 42 75,430 30 40 37 62,030 1.0 37 62,080 20 30 34 53,980 1.0 34 53,980 10 20 30 46,100 1.0 30 46,100 1
e
Page 11@
5.3.3 Conclusions Concerning Valve Operability To determine whether a given valve actuator assembly will operate under the required flow conditions, two sets of criteria must be applied; one for pneumatic actuated valves and one for
~
electric actuated valves. The following criteria apply for pneumatic and electric actuated valves:
- 1. Actuator torque output must overcome with sufficient margin the worst case torque resisting valve closure.
- 2. Peak aerodynamic induced closing torques must not exceed actuator or valve design torques The following additional criteria apply for electric actuated valves only:
- 1. The worst case torque resisting valve closure must be less by a sufficient margin than the closed torque switch trip point.
For LOCA flow conditions it can be seen in Tables for A0-57-121, 131, 114 that all aerodynamic torques tend to aid valve closure. For valves A0-57-104, 123, 124 aerodynamic torques for the first 5 degrees from full open resist closure. Pertinent torques for air operated valves are listed in Table 36.
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Page 113 TABLE 36 Pneumatic Actuated Valve Torques Torques (in-lb)
Valve Valve Design Size Valve No. Torque 1 2 3 4 5 .
6" A0-57-121 7,800 none none req'd 2,400 27,585 26,668 6" A0-57-131 7,800 none none req'd 2,400 27,585 26,668 !
l 24" A0-57-114 135,000 none none req'd 97,075 362,496 211,000 l
18" A0-57-104 63,300 4,503 76,096 48,450 207,563 174,000 i
24" A0-57-123 135,000 46,188 124,695 115,970 362,496 211,000 24" A0-57-124 135,000 46,188 124,695 115,970 362,496 211,000 1
1 Worst case closure resisting aerodynamic torque 2 Actuator torque use to overcome aerodynamic torque 3 Maximum aerodynamic torque 4 Torque to yield actuator key
)
5 Actuator safe structural torque l
Page 114 For valves MO-57-109,161,162,163,164 for LOCA flow conditions, it can be seen in Tables 29, 27, 28, 30 that all aero-dynamic torques are in the closed direction. Due to the use of an - i electric actuator, this torque will not help the valve close.
For valves M0-57-112, 115, 135, 147, aerodynamic torques for the first 50 from full open resist closure. For these units the closed torque switch trip point must be greater than the maximum aerodynamic torque resisting closure.
TABLE 37 Electric Actuated Valve Torques Torque (in-lb)
Valve Valve Design Size Valve No. Torque 1 2 3 4 5 4" M0-57-161 2,112 none none req'd 740 7,594 15,600 4" M0-57-163 2,112 none none reg'd 740 7,594 15,600 6" M0-57-109 7,800 none none reg'd 2,400 27,585 26,400 6" M0-57-162 7,800 none none req'd 2,400 27,585 26,400 6" MO-57-164 7,800 none none req'd 2,400 27,585 26,400 18" MO-57-112 63,300 2,127 34,000 48,450 207,563 235,000 24" MG-57-115 135,000 11,550 97,075 362,496 235,000
! 24" M0-57-135 135,000 11.550 97,075 362,496 235,000 24" M0-57-147 135,000 11,550 97,075 362,496 235,000 1 Worst case closure resisting aerodynamic torque 2 Torque switch trip torque (Actuator torque used to overcome aercdynamic torque) 3 Maximum aerodynamic torque 4 Torque to yield actuctor key 5 Actuator safe structural torque
Page 115 From the preceding data it can be seen that the minimum actuator torque margin over that required to overcome worst case aerodynamic torque is better than 2.7 The safety factor
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is obtained even after full containment pressure has been developed!
- From the presented data and supplemental test reports, it has been shown that the valves will operate as designed under the prescribed conditions. This has been shown using the conser-vative assumption of no credit taken for pressure ramp in containment and no credit taken for back pressure due to down-stream piping. Further, no credit has been taken for activation of the first valve under back pressure conditions produced on the second valve due to closure and pressure drop across the first valve.
6.0 VALVE SEALING CHARACTERISTICS 6.1 Normal Sealing The following chart shows the sealing ability of the valves as they were shop tested for record. The tests were performed with pressure on the indicated side of the disc and the opposite side open to atmosphere. The normal recommended flow direction for these valves is with pressure on the shaft side, so when !
pressure is applied to the clamp ring side, it is considered to be the reverse flow direction. The test performed was an air under water test in which the smallest detectable leakage was a bubble (.15cc).
- GD___
Ma m .--
Page 116 l
TABLE 38 -
VALVE SEALING CHARACTERISTICS LEAKAGE (Bubbles / min)*
VALVE Pressurized Side ;
VALVE SIZE TEST PRESSURE Shaft Clamp Ring !
MARK NO. (IN.) PSIG Side Side l
- 1 M0-57-161 4 5, 25, & 55 0 0 M0-57-163 4 5, 25, & 55 0 0 .
MO-57-109 6 5, 25, & 55 M0-57-162 -
6 5, 25, & 55 0 0 M0-57-164 6 5, 25, & 55 0 0 A0-57-121 6 5, 25, & 55 A0-57-131 5 5, 25, & 55 0 0 M0-57-112 18 5, 25, & 55 A0-57-104 18 5, 25, & 55 0 0 MO-57-115 24 5, 25, & 55 M0-57-135 24 5, 25, & 55 MO-57-147 24 5, 25, & 55 A0-57-114 24 5, 25, & 55 A0-57-123 24 5, 25, & 55 0 0 A0-57-124 24 5, 25, & 55 0 0
- Worst case pressure c , - -, , ,-
Page 117 l 1
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6.2 1.ong Term Sealing The conical seal / seat design of the Tricentric valve in combination with the laminated metal / asbestos seal offers good long term sealing characteristics. When the seal and seat are machined a certain surface finish is obtained. With this finish ;
certain leak rates are obtained during a bench test (see 6.1).
On a microscopic scale these surfaces contain peaks and valleys.
When the disc is seated, these surfaces mate and high local (above yield) stresses are induced at the peaks. The peaks will yield and deform and form a match between the seat and seal.
As the valve is cycled throughout its life, this match tends to improve end a visual seating pattern appears. This results in improved sealing as the valve ages.
This has been verified by experience and is documented in the Shell Internatio Cycling Test (reference 7.0 D3). This A
test was performed by Gebruder Adams of Bochum, West Germany.
Clow's Engineered Products Division produces the Tricentric design under license of Gebruder Adams. The test showed -
sealing improved continuously up to 41,000 cycles, the limit of the test.
Page 33@
6.3 Debris Effects On Sealing A test was performed to determine the effect on sealing capability of a Tricentric valve if a foreign object became trapped between the seat and seal. As with any valve, if the object is large enough and hard enough and happens to be caught between the sealing surfaces, the valve will fail to close completely and the valve will leak. Leakage will be dependent 1
on the size and s' ape of the object and open gap size which remains when the valve does not fully close. Since no stand-ards as to debris size exist, the test made determined leakage due to object damage after the object was removed. For in plant operation this would represent leakage after recycling of the , ,
j valve if the object was blown out of the way during recycling.
The object selected was a cooling tray liner used in the petrochemical industry. It's dimensions were approximately 1/8" x 1" x 6" and was a filled polyvinyl chloride plastic of 80 shore D hardness. The valve was closed upon this material, opened to remove the material, then closed again to measure -
leakage. Depending on the applied seating torque, a leakage 1 of .015 SCFM to .333 SCFM was measured. This test showed the valve could tolerate some large debris and still maintain a relatively low leakage even with a damaged seal (See reference 7.0 D-2).
Page 119 6.4 Sealing Under Temperature Variations The Tricentric design has been used successfully for sealing applications from cryogenic to 9000F. The Shell
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International Cycling Test describes sealing characteristics for a media operating temperature of 8420F when the body reached a temperature of 7160F.
The Tricentric conical seal / seat design lends itself.
well to accommodating temperature changes in the body and resultant size variation of the sealing components. Due to the torque seating design and some seal flexibility, the valve will self adjust to the small dimensional variations which could be anticipated for the subject valves. Of course, if large thermal gradients (very unlikely from information provided to Clow) existed around the body circumference higher levels of leakage could be expected. Again no standards exist to the knowledge of Clow personnel which could become a basis for preuiction or a test of such leakage.
O
Page 120
7.0 REFERENCES
A. Seismic Analysis Reports prepared by: Patel Engineers Huntsville, Alabama The following include stress and frequency analysis for the subject valves:
- 1. Report PEI-TR-83-16 Rev. A for Clow 4" Wafer Stop Valve (April 83)
Mark Nos. 4"-HBB-BF-M0-57-161 and 163 Clow Job No. 82-2053-01(N)-01 & 02
- 2. Report PEI-TR-83-15 Rev. A for Clow 6" Wafer Stop Valve (April 83)
Mark Nos. 6"-HBB-BF-M0-57-109,162, and 164 6"-HBB-BF-AO-57-121 and 131 Clow Job No. 82-2053-02(N)-01, 02, & 03 82-2053-03(N)-01 & 02
- 3. Report PEI-TR-83-14 Rev. A for Clow 18" Wafer Stop Valve (April 83)
Mark Nos .18"-HBB-BF-MO-57-104 18"-HBB-B F- AO 112 Clow Job No. 82-2053-04(N)-01 82-2053-05(N)-01
- 4. Report PEI-TR-83-13 Rev. A for Clow 24" Wafer Stop Valve (April 83)
Mark Nos. 24"-HBB-BF-M0-57-115, 135, and 147 24"-HBB-BF-AO-57-114,123, and 124 Clow Job No. 82-2053-06(N)-01, 02, & 03 82-2053-07(N)-01, 02, & 03 B. Seismic Qualification Test Plans and Reports prepared by: National Technical Systems Testing Division Saugus, California
- 1. Nuclear Test Procedure No. 528-0951 Rev. A (Jan. 83)
" Seismic Qualification Testing of Butterfly Valves and '
Actuators"
Page 121 REFERENCES (con't)
- 2. Nuclear Test Report 528-0951 Rev. A (May 83)
" Seismic Qualification Testing of Butterfly Valves and Actuators"
~
prepared by: Wyle Laboratories Huntsville, Alabama
- 3. Test Procedure 541/0465/WB (May 83)
" Static Load Test Procedure For An 18-Inch Valve Assembly"
- 4. Test Procedure 46823-1 (June 83)
" Static Load Test Program on an 18" Butterfly Valve Assembly With Limitorque Operator"
- 5. Report No. 45832-1 " Low Level Seismic Vibration Test Program on a 12" Butterfly Valve Assembly" (Nov. 23,1981).
Low level biaxial sine sweep resonant search.
- 6. Report No. 45828-1 " Seismic Sime tion Test Program on a Valve Actuator Solenoid Valve" pav. 22,1981). Low level sine sweep resonant search and sine beat test (to 11.0 g max.) for Asco solenoid valve. ,
prepared by: Vought Corp.
High Speed Wind Tunnel Facility Dallas, Texas
- 7. Report No. 2-59700/1R-52972 " Simultaneous Static Seismic Load of Flow Interruption Capability Tests of a 12 Inch Valve for the Clow Corporation" (Dec. 15. 1981).
Application of 11.0 g biaxial static load to valve actuator during operation with choked air flow thru the valve.
C. Air Flow Tests prepared by: A.L. Addy, Ph.D.
Urbana Illinois (Engineering Consultant in Fluid Dynamics)
- 1. Final report on the Clow Valve Analysis Program CVAP (Oct.1981). Report covers methods of analysis, develop-ment of data base from model tests, and set-up of computer program to predict characteristics of full size valves.
l
Page 122 REFERENCES (con't)
- 2. Report on " Aerodynamic Torque and Mass Flowrate For Compressible Flow Through Three Geometrically Simifar Scale-Model Clow Valves Located Downstream of a 90 ,
Mitered Elbow" .
- 3. Report on " Aerodynamic Torque and Mass Flowrate For Compressible Flow Through Geometrically Similar Scale- -
Model Clow Valves in Series" (October 1982)
- 4. Report on " Water Table Investigation of a Two-Dimensional Scale-Model of a 24-Inch Clow Tricentric Butterfly Valve" (November 1982)
D. Other Reports and Information
- 1. Operating Instructions for Clow Tricentric Wafer Stop Valve covers installation, maintenance, and operating instructions for 82-2053(N) valves.
- 2. Clow Test Report Project No.82-003 " Effects of Foreign Bodies on Tricentric Sealing" by Robert Sansone.
- 3. Shell International Cycling Test (2/6/72) by M. Nijenhuis (Note; Clow produces Tricentric valves under license of Gebruder Adams of Bochum, West Germany.)
E. Other References
- 1. Bechtel Power Corp. Design Specification 8031-P-144, Rev. 2
- 2. "A Water Table Investigation of Two-Dimensional Models of the Clow Corporation Tricentric Valve" by Dr. Robert F.
Hurt, Engineering Consultant, Professor of Mechanical Engineering, Bradley University, Peoria, Illinois, Sept.14, 1979.
- 3. "A Parametric Study of a Butterfly Valve Utilizing the Hydraulic Analogy" by Bruce A. Coers; Thesis for fulfill-ment of Master of Science in Mechanical Engineering require-ments, Graduate School of Bradley University, Peoria Ill.,
1983.
e APPENDIX A NUCLEAR REGULATORY PURGE VALVE OPERABILITY GUIDE LIflES l
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. - - , . . _ _ - = = - - - * * - - .
Page A-1 BRANCH TECHNICAL POSITION CSB 6-4
- CONTAINMENT PURGING DURIttG NORMAL PLAi4T OPERATIONS A. BACKGROUND This branch technical position pertains to system lines which -
can provide an open path from the containment to the environs during normal plant operation; e.g., the purge and vent lines of the containment purge system. It supplements the position taken in SRP section 6.2.4.
While the containment purge system provides plant operational flexibility, its design must consider the importance of mini-mizing the release of containment atmosphere to the environs following a postulated loss-of-coolant accident. Therefore, plant -
designs must not rely on its use on a routine oasis.
The need for purging has not always been anticipated in the design of plants, and therefore, design criteria for the contain-ment purge system have not been fully developed. The purging experience at operating plants varies considerably from plant to plant. Some plants do not purge during reactor operation, some purge intermittently for short periods and some purge continuously.
The containment purge system has been used in a variety of
( ways, for example, to alleviate certain operational problems, I
such as excess air leakage into the containment from pneumatic controllers, for reducing the airborne activity within the contain-ment to facilitate personnel access during reactor power operation,
- Note: This paper is retyped for legibility fron paper supplied by NRC.
Page A-2 and for controlling the containment pressure, temperature and relative humidity. However, the purge and vent lines provide an open path from the containment to the environs. Should a LOCA occur during containment purging when the reactor is at ptder, '
the calculated accident doses should be within 10 CFR 100 guide-line values.
The sizing of the purge and vent lines in most plants has been based on the need to control the containment atmosphere during refueling operations. This need has resulted in very large lines penetrating the containment (about 42 inches in diameter). Since these lines are normally the only ones provided that will permit some degree of control over the containment atmosphere to f'acilitate personnel access, some plants have l used them for containment purging during normal plant operation.
Under such conditions, calculated accident doses could be signif-icant. Therefore, the use of these large containment purge and vent lines should be restricted to cold shutdown conditions and refueling operations.
The design and use of the purge and vent lines should be based on the premise of achieving acceptable calculated offsite radio-logical consequences and assuring that emergency care cooling (ECCS) effectiveness is not degraded by a reduction in the contain-ment pressure.
Purge system designs that are acceptable for' use on non-routine basis during normal plant operation can be achieved by
_.-__________.x_a --
N .h.mm____ O
Page A-3 providing additional purge and vent lines. The size of these lines should be limited such that in the event of a loss-of-coolant accident, assuning the purge and vent valves are open and subsequently close, the radiological consequences calculated in accordance with Regulatory Guides 1.3 and 1.4 would not exceed the 10 CFR l'00 guideline values. Also, the maximum time for valve closure should not exceed five seconds to assure that the purge and vent valves would be closed before the onset of fuel failures following a LOCA.
Thc size of the purge and vent lines should be about eight inches in diameter for PWR plants. This line size may be overly conservative from a radiological viewpoint for the flark III BWR plants and the HTGR plants because of containment and/or core design features. Therefore, larger line sizes may be justified.
However, for any proposed line size, the applicant must demon-strate that the radiological consequences following a loss-of-coolant accident would be within 10 CFR 100 guideline values.
In suceary, the acceptability of a specific line size is a function of the site meteorology, containment design, and radio-logical source term for the reactor type; e.g., BUR, PUR or HTGR.
B. BRANCH TECHNICAL POSITION The system used to purge the containment for the reactor operational modes of power operation, startup. hot standby and hot shutdown; i.e., the on-line purge system, should be indepen-dent of the purge systen used for the reactor operation nodes of cold shut down and refueling.
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page g,g \
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- 1. ' The on-line purge systen should be designed in accordance with the following criteria:
- a. The performance and reliability of the purge system isolation valves should be consistent with the oper-ability assurant.e program outlined i.n MEB Branch Technical Position MEB-2, Pump and Valve Operability Assurance Program. (Also,see,SRPSection3.9.3) The design basis for the valves and actuators should include the buildup-of containment pressure' for the 'LOCA. break spectrum, and l
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the purge' line and ven+ line floyrs as a function of time up to and during valve closure. l
- b. The number of purge and vent lines that may be used should ,
be limited to one purge line and one vent line.
- c. The size of the purge and vent lines should not exceed.
about eight inches in diameter unless detailed justifi-
! cation for larger line sizes is provided. ,.[
l The containment; isolation _ provisions.for the purge',ystem
~
2 d.
! lines should meet-the standards appropriate to engineered l
l safety features; e.e., quality, redundancy, reliability and other appropriate criteria.
- e. The inst'rumentation and control systems provided to isolate the purge system lines should be independent and actuate'd by diverse parameters; e.g.. containment pressure, safety injection actuation, and containment radiation level.
If energy is required to close, the valves, at least two diverse sources of energysshal$'he provided, either of ..,
which can affect the isolation Yunstion. . .
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4x , J .. ,
< r
- a. , ,
, wy, L - _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
Page A-5
- f. Purge system isolation valve closure times, including instrumentation delays, should not exceed five seconds.
9 Provisions should be made to ensure that isolation valve closure will not be prevented by debris which could .
potentially become entrained in the escaping air and steam.
- 2. The purge system should not be relied on for temperature and humidity control within the containment.
- 3. Provisions should be made to minimize the need for purging of the containment by installing containment atmosphere cleanup systems within the containment.
- 4. Provisions should be made for testing the availability of .
the isolation function and leakage rate of the isolation valves, individually, during reactor operation.
- 5. The following analyses should be performed to justify the
, containment purge system.
- a. An analysis of the radiological consequences of a loss-of-coolant accident. An analysis should be done for a
spectrum of break sizes, and the instrumentation and l setpoints that will actuate the vent and purge valves closed should be specified. The source term used in the radiological calculations should be based on a calcul-ation under the terms of' Appendix K to determine the extent of a failure and the concomitant ' release of fission products, and the fission product activity in the primary coolant. A pre-existing iodine spike should
+ ;
,k
Page A-6 l
l be considered in determining primary coolant activity.
The volume of containment in which fission products are mixed should be justified, and the fission products from the above sources should be assumed to be released through the open purge valves during the maximun interval required for valve closure. The radiological conseq-uences should be within 10 CFR 100 guideline values.
- b. An analysis which demonstrates the acceptability of the provisions made to protect structures and safety-related equipment; e.g., fans, filters ar.d ducting located beyond the purge system isolation valves against loss of function to control the environment created by the escaping air ,
and steam,
- c. An analysis of the reduction in the containment pressure resulting from the partial loss of containment atmosphere during the accident for ECCS backpressure determination.
- d. The allowable leak rates of the purge ar.d vent isolation valves should be specified for the spectrum of design basis pressures and flows against which the valves must close.
Fage A-7 GUIDELINES FOR DEMONSTRATION OF OPERABILITY OF PURGE AND VENT VALVES OPERABILITY l
In order to establish operability it must be shown that the L
valve actuator's torque capability has sufficient margin to over-l come or resist the torques and/or forces (i.e., fluid dynanic, 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.
- 2. Flow direction through valve; AP across valve.
- 3. Single valve closure (inside containment or outside containment valve) or simultaneous closure. Establish worst case.
- 4. Containment back prtissure effect on closing torque margins of air operated valve which vent pilot air inside contain-ment.
S. Adequacy of accumulator (when used) sizing and initial charge for valve closure requirements.
- 6. For valve operators using torque liniting devices - are the settings of the devices compatible with the torques required to operate the valve during the design basis condition.
N _ _ . _ _ _ _ _ _ . _ _ _ _ _ _ . _ _ _ _ _ _
Page A-8
- 7. The effect of the piping systen (turns, branches) up-stream and downstream of all valve installations. .
- 8. The effect of butterfly valve disc and shaft orientation to the fluid nixture egressing from containment.
DEMONSTRATION Demo'nstration of the various aspects of operability of purge and vent valves may be by analysis, bench testing, insitu testing or a combination of these neans.
Purge and vent valve structural elements (valve / actuator assembly) must be evaluated to have sufficient stress margins to withstand loads imposed while valve closes durin2 a design basis accident. Torsional shear, shear, bending, tension and compression ,
loads / stresses should be considered. Seismic loadings should be addressed.
Once valve closure and structural integrity are assured by analysis, testing or a suitable combination, a deternination 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:
- ~~ -
Page A-9 Bench Testing A. Bench testing can be used to demonstrate suitability of the in-service valve by reason of its tracibility 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 assenbly or by extrapolation of data from a similarly designed valve.
- 2. Llhether measures were taken to assure that piping up-stream and downstream and valve orientation are siraulated.
- 3. Whether the following load and environmental factors were considered .
- a. Simulation of LOCA l
- b. Seismic loading
- c. Temperature soak l
- d. Radiation exposure
- e. Chemical exposure
- f. Debris B. Bench testing of installed valves to demonstrate the suitability of the specific valve to perform its required function during the postulated design basis accident is acceptable. i l
- 1. The factors listed in -items A.2 and A.3 should be considered when taking this approach.
In-Situ Testing In-situ testing of purge and vent valves may be performed to confirm the suitability of the valve under actual conditions.
Page A-10 l! hen performing such test, 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.
i' End CSB 6-4 1
Page A-Il CLARIFICATION OF SEPT. 27 LETTER TO LICENSEES REGARDING DEMONSTRATION OF OPERABILITY OF PURGE AND VENT VALVES
- 1. The LP across the valve is in part predicated on the contain-ment pressure and gas density conditions. What were the containment conditions used to determine the AP's across the valve at the incremental angle positions during the closure cycle?
- 2. Were the dynamic torque coefficients used for the deter-mination of torques developed, based on data resulting from actual flow tests conducted on the particular disc shape /
design / size? What was the basis used to predict torques developed in valve sizes different (especially larger valves) .
than the sizes known to have undergone flow tests?
- 3. Were ins'.a11ation effects accounted for in the determination of dynamic torques developed? Dynamic torques are known to be affected for example, by flow direction through valves with off-set discs, by downstream piping backpressure, by shaft orientation relative to elbows, etc. What was the basis (test data or other) used to predict dynamic torques for the particular valve installation?
- 4. When comparing the containme't pressure response profile against the valve position at a given instant of time, was the valve closure rate vs. time (i.e. contstant or other) taken into account? For air operated valves equipped with spring return operators, has the lag time from the time the
- Note: This paper is retyped for legibility from paper supplied by NRC.
Page A-22 valve receives a signal to the time the valve starts to stroke been ac:cented for? .
NOTE: Where a butterfly valve assembly is equipped with spring to close air operators (cylinder, diaphragm, etc.), there typically is a lag time from the time the isolation signal is received (solenoid valve usually deenergized) to the tine the operator starts to move the valve. In the case of an air cylinder, the pilot air on the opening side of the cylinder is approximately 90 psig when the valve is open, and the spring force available may not start to move the piston until the air on this opening side is vented (solenoid valve de-energizes) below about 65 psig, thus the lag time. ,
- 5. Provide the necessary information for the table shown below for valve positions from the initial open position to the seated position (10 increments if practical).
Valve Position 0 (in degrees - 90 Predicted aP Maximum t.P
= full open) (across valve) (capability)
- 6. What Code, standards or other criteria, was the valve designed to? What are the stress allowables (tension, shear, torsion, etc.) used for critical elements such as disc, pins, shaft yoke, etc. in the valve assembly? What load combinations were used?
- 9. For those valve assemblies (with air operators) inside contain-nent, has the containment pressure rise (backpressure) been considered as to its effect on torque margins available (to close and seat the valve) from the actuator? During the closure period, air must be vented from the actuators opening i
Page A-13 side through the solenoid valve into this backpressure.
Discuss the installed actuator bleed configuration and provide .
basis for not considering this backpressure effect a problem on torque margin. Valve assembly using 4 way solenoid valve ,
should especially be reviewed.
- 10. Where air operated valve assemblies use accumulators as the fail-saft feature, describe the accumulator air systen config-uration and its operation. Previde necessary information to show the adequacy ofthe accumulator to stroke the valve i.e.
sizing and operation starting from lower limits of initial air pressure charge. Discuss active electrical components in the accumulator system, and the basis used to determine their qualification for the environmental conditions exper-ienced. Is the accumulator system seismically designed?
- 11. For valve assemblies requiring a seal pressurization system (inflatable main seal) describe the air pressurization system configuration and operation including means used to determine that valve closure and seal pressurization have taken place. Discuss active electrical components in this system, and the basis used to deternine their qualification Is this system for the environmental condition experienced.
seismically designed.
For this type valve, has it been determined that the " valve travel stops" (closed position) are capable 'of withstanding the loads imposed at closure during the CBA-LOCA conditions.
9
Page A-14 p
- 12. Describe the modification made to the valve assembly to limit i
the opening angle. With this modification, is there sufficient f torque margin available from the operator to overcome any dynamic torques developed that tend to oppose valve closure, start.ing from the valve's initial open position? Is there sufficient torque margin available from the operator to fully seat the valve? Consider seating torques required with seats tnat have been at low ambient temperatures.
- 13. Does the maximum torque developed by the valve during closure exceed the maximum torque rating of the operators? Could this affect operability?
14 Has the maximum torque value determined in =12 been found to ,
be compatible with torque limiting settings where applicable?
- 15. Where electric motor operators are used, has the mininum avail-able 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 h5A conditions with these lower limit voltages available. Does this reduced voltage operation result in any significant change in stroke timing? Describe the emergency mode power source used.
- 16. Where electric 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
Page A-15 the valve being left in the handwheel mode following some maintenance, test etc. type operation.
- 17. Describe the tests and/or analysis performed to establish the qualification of the valve to perform its intended function
- under the environmental conditions exposed to during and after the' DEA following its long term exposure to the normal plant environment.
- 18. What basis is used to establish the qualification of the valve, operators, solenoids, valves? How was the valve assembly (valve / operators) seismically qualified (test, analysis, etc.)?
- 19. Where testing was accomplished, describe the type tests per-formed conditions used etc. Tests (where applicable) such as flow tests, aging simulation (thermal, radiation, wear, I vibration endurance, seismic) LOCA-DBA environment.(radiation, steam, chemeials) should be pointed out.
- 20. Where analysis was used, provide the rationals used to reach the decision that analysis could be used in lieu of testing.
Discuss conditions, assumptions, other test data, handbook data, and classical problems as they may apply.
- 21. Have the preventive maintenance instructions (part replace-ment, lubrication, periodic cycling, etc.) established by the manufacturer been reviewed, and are they being followed?
Consideration should especially be given to elastomeric com-1 ponents in valve body, operators, solenoids', etc. where this hardware is installed inside containers.
i
APPENDIX B
SUMMARY
OF 12" CLOW TRICENTRIC CHOKED FLOM/ STATIC SEISMIC OPERABILITY TEST (Refer to Vought Corp. Report No. 2-59700/1R-52972) 1
4 .
APPENDIX B DESCRIPTION OF OPERATIONAL TESTS OF A 12 INCH CLOW TRICEtiTRIC VALVE FOR i NUCLEAR PURGE SYSTEM SERVICE BY J. E. KRUEGER
' NUCLEAR VALVE DESIGN Et4GIr4EER NOVEMBER 30, 1981 i
4 23 I a
- ~,_.....
Page B-1 INTRODUCTION -
A test was performed at Vought Corp., Dallas, Texas, on November 16, 1981, to demonstrate operability of a 12 inch Tricentric valve for flow and load conditions possible in case of a LOCA (Loss of Coolant Accident) in a nuclear plant. The test was run with a valve to be used in Jersey Central Power and Light's Oyster Creek Plant. The test was performed by Vought personnel under the direction of a Clow Engineer.
Witnesses to the tests included representatives of GPU Nuclear of New Jersey and Bechtel of San Francisco.
OBJECTIVE -
The test was performed to demonstrate that the valve would operate under pressure, flow, and loadings simulating operating and seismic conditions possible during a LOCA. It was also desired that the open to close cycle be demonstrated to occur in less than 5 seconds. A secondary objective was to show aerodynamic torques produced by air flow over the disc were equal or less than those predicted and used in designing the valve and selecting the actuator. (Predicted torques used in design derived from previous air flow test performed with 3 inch scale models.)
TEST SET-UP -
The valve was installed in a straight pipe run with a stagnation chamber upstream approximately 6 feet. Downstream 3 feet was a diverging nozzle to prevent downstream pressure
Page B-3 VALVE AND ACTUATOR DESIGN PARAtlETERS -
The valve tested was designed for a differential operating i pressure of 65 psi and combined operating and seismic loads of (I
11.0 g's. The seal was of laminated 316 SST and asbestos. l The body design was 150 lb. class per ANSI B16.34. The shaft '
, f used for transmitting torque to close and seal the valve was of f a 17-4 PH age hardenable stainless steel, heat treated to condition H-1100. The actuator used was a Bettis NT-316B-SR2 pneumatic spring return actuator. The actuator was of a fail closed design with the spring supplying the closing and seating torque (Note: Tricentric valves are designed for torque l
seating). The actuator was qualified for nuclear service.
CONDUCT OF TEST -
The test consisted of applying the static loads to the actuator and establishing a 65 psig upstream pressure with the Clow valve closed. A signal was then initiated to open the valve.
The valve then cycled full open against flow and remained open until a signal to close the valve was provided. The valve then cycled to the closed position and seated. During this period data was taken automatically at.10 measurements per second at all sensors. This test was repeated 4 additional times at 65 psig and once at 35 psig. Note: These upstream pressures produced choked (flow at sonic velocity) flow through the valve during the valve open period.
Page B R RESULTS OF TESTS -
The tests demonstrated the following: .
- 1. The Clow disc and shaft geometry provides for a positive aerodynamic closing torque for all angles from full open to full closed.
- 2. The aerodynamic torque values used for design of
'the Clow valve are conservative relative to measured torques. (Design torques were based on previous 3" scale model tests.)
- 3. The construction of the valve is rigid in its design such that no binding resulted under an 11.0 g load applied in two directions simultaneously.
- 4. The valve will cycle from full open to full closed in less than 5 seconds with any amount of flow from none to the maximum tested (10B lb/sec of air).
Any value of flow above zero tended to close the valve faster (the valve closed in 3.6 sec. for a no flow condition).
- 5. Operator sizing was sufficient to cycle the valve from full closed to full open in less than 5 seconds for any tested flow rate.
CONCLUSION -
Clow has demonstrated that their nuclear purge valve design can meet and exceed typical specifications for this type of service. It was further shown that the valve will function as
^
Page B-5 required regardless of the LOCA pressure ramp curve (assumes lower pressures upstream at start of valve closure) often .
used by other valve manufacturers to show operability. In conjunction with other tests (now in progress) to show opera- ,
bility under many installed piping configurations, Clow valves can allow full open purge function during shutdown for refueling as opposed to the partially open position now allowed by the NRC. Further, it has been shown that the Tricentric can meet tight leak rate requirements with a metal to metal sealing .
which is more reliable and less costly in maintenance than sealing with elastomers.
0 4
~
.e--..e.4
Page C-1
' APPENDIX C ,
o C
srn FAST CLOSING VALVE 0?EMT038 i=
An Industry Problem... Speed VersusTorque/ Thrust Prepared by 1.lMITOROUE CORPOR ATION a torque sensing device called the torque switch.
PREFACE This switch (shown in Fig.1) operates by measuring the amount of compression of a fixed pre-calibrated Since the advent of nuclear power, there has been an As a load ever increasing demand on valve operator manufac- belleville spring (the torque spring).
turers to build electric actuators with faster oper- developes (either a seating load or an obstruction in the valve) and the output worm gear is presented ating times. This demand has, for the most part, been satisfied with very slight moc'ifications to exist- fr m rotating, a thrust is generated on the worm ing, popular models, such as the Limitorque SMB orcausing it to compress the torque spring.This action takes place in either direction of rotation of the out-SMC' The use of existing and proven equipment incor- put worm gear.
potating these minor modifications has given allof the, , g ,,
valve coinpanies, engineering consultants and end g, have reduced the forces to a linear spring force (F)
B users the confidence necessary toacting continue toarmplaceever ; increasing de on a torque (TA) to produce a ggtorque (T) g and development to find the limits to which present ssb d h w sW imi% d designs might be utilized to obtam the faster closingbrated in pounds foot and a nameplate is affixed to speeds of tomorrow, each switch compartment denoting the calculated torque output at that setting. The maximum setting APPLICATION allowed by Limitorque for the particular application Most valve applications are either torque seated or at is also marked on the torque switch calibration tag.
lent protected in the closing direction by the use of Figure 1 :
TOROUE SwlTCH (see trom new Dottom tqN) 0O%
\ .
WA .
It ,*~~~' OoEN TOR 3UE t
p h ..' worm A3xsTUI*.T
- yTA py
'a 3 f N0vtWEt,T F , TORCE OnWORV h 1A r00'.E ARM S etC . -
T TOROUE DUTPUT tF T A, ~
m 7 O _
~
.e.
y EELitCtti < h ,
er'
$PRihG PACA CLOSE TOROUE g W
ASJUS1NEh1 10MUE salCM tiro"1 v'tw 3 48 fM3'O
Page C-2 I = i Hgur:2 34 _i.i _ 4. 1 Fig. 2 shows an examplCot a torque w och uhbr.- !
I I
! ,, ' I tion tag for a valve operator designed tu dehwr
J,,
g { l [ lbs ft. of torque with a mmmum safe op ratur use. p<
abihiy of 105 lbs. ft. Limitorque al>o m> tait, a int. 'r ,
!l que hmiter plate on the torque switch to phyucati.
f7 l '
I prevent someone from setting the torque switch to a
[ setting higher than the maximum computed capabil.
/ ity of the umt.
r ,
EFFECTS OF STARTER DROPOUT
- b t
The torque switch, although a mechanical device,is
" * " " ' "
- an electrical control element. It is wired in series
- ""C*"' "
with either a pushbutton or relay and the reversm;
__sd7%
vi- t e.
[
i starter coils (Fig. 3). To stop the motor in the vveientonour swnew ,ss ff { j u cgg$e" direction, the torque switch contacts must CM5n& TION _ ,
open and the motor starter coil de-energizes. This
- W_-*'.Y"'* 'C'.5 N_NPa brerks the starter contacts and disconnects the elec-trical power to the motor. Tests have indicated that the starter dropout time (the time from the torque switch trip point to zero voltage on the motor) aver-Ugure 3 a es between 25 to 50 milliseconds depending on opti. the size of the reversing starter and its physical con-1
- -i FA struction.
In the time between the torque switch trip and qh H a2
" the motor starter drop-out (25 to 50 milliseconds),
cu s* the motor develops torque proportional to the resis-
_l_ -1 tance of the valve. This means that the more rigid nest the valve system, the faster the motor approaches its {
to
~'~*
1t "r A
^ "
pullout torque rating. At normal stem speeds of 12"/ min. on gate valves and 4"/ min. on globe valves
{- gg , or less and an output of 50 rpm or less of the valve N, s operator, the 25 to 50 milliseconds drop-out time of O c,,w cu the motor starter has a negligible effect on the total 4 , c,,,,n,, ,,,,e sc
- 4 %,
output torque of the valve operator. As the speed of ot o the operator increases, the starter drop-out time be-g [ w % ,cem.n. comes a very appreciable factor in the f' mal output torque at any given torque switch setting as the unit travels further in this time interval.
Figure 4 Fig. 4 illustrates a gate valve (Curve 1) at 2 imenmemvt 6 ir, . 12"/ min, and 36 rpm output of the valve operator
- L*'a cuivr Q, e x ,a against Curve 2 for the same unit at 36"/ min. and 10S
@ rpm output of the valve operator. Similar curves for f"'
g" 'f "-
globe valve operators (Curves 3 &4) at stem speeds of 6" and 12"/ min.respectively are shown.The tor-
$ que switch in all four cases trips at the predeter-E,_._
j 75 p! ;
l mined output torque (~15 lbs.-ft.) shown as (* *). The motor starter in each case requires 35 milliseconds Y$ j .i '- ,,l l
to break the motor voltage. The output torque twe m m @ # "*
stem mvt plotted at the motor startea trip point is indicated teascunvi @ e r'" with an (*).
[ EFFECTS OF EFFICIENCY
-l i l- ie. =..-,- e n u =
g The efficiency variation of valve actuators de
- ', ["ylj'y [,/,l upon two primary factors, gear ratio and worm f,,/ '. l l ..a 5'
'; t Jj
> L ,.4. caernsat t Lenn.torew corner. ten d
Tsant I" -m
Page C-3 speed. The lower the gear ratio, the higher the effi. T he stand.ird A6:nc 'thrw utslu..- i- oh ciency and also the higher the worm speed the companies today ahu v.ines in ein,.i,na a %...,3,.'
O higher the efficiency. Fig. 5 demonstrates the effici- creases.
by surface finish, Acme thrcad and cleanhnew ella.nency lut rn.. nonnli .gedig jij','[,
d ency vanation ot valve actuator gearms illustratirig
~
the difference between selflocking(high ratios)and not uncommon to find an increase in stem cliiuenn of 10-15% at yoke nut speeds greater th.an 100 rpnh n3n locking (Iow ratios). It can be seen from Fig. 5 that the efficiency of actuators at higher worm speeds is much greater than at lower worm speeds.
EFFECTS OF INERTI A The rigidity of the system determines the slope Fast operating times necessarily bring with them in.
cf the load curve (toroue) as plotted in Fig. 4. It can ersial loads which are always additive to the in-be seen that the more rigid the system. the steeper creased load due to the motor starter drop-out lag.
the slope, as illustrated by the difference between The magnitude of this inertialload,like operator ef.
gate (Curves 1 & 2) and globe (Curves 3 & 4) valve ficiency, depends on the operator speed and the curves where the average globe valve is substantially rigidity of the system. It is possible to have a valve more ngid than the comparable gate valve. This operator traveling at 120 to 144"/ min. in a rigid rigidity determines the speed of the worm existing system to experience a 1-1/2 to 2-1/2 load magnifi ct the time the torque switch activates (the steeper cation factor due to inertia alone. Unfortunately,it the curve, the higher the speed of the worm). For is impossible to determine this load macnification example, if a valve operator having a self-locking factor empirically before the operator is mstalled on worm gear set was used to compress a set of springs the valve.
until the torque switch tripped (at 75 lbs. ft.), the load curve would look like Curve A in Fig. 6.lf the Suggestions have been retei.ed thtough the same unit was operated against a rigid structure in-stead of the springs, the load curve would look like years as to the best method of handling the increases in efficiency and inertia on high speed actuators. A Curve B in Fig. 6. The difference in the motor cur- tempting solution is to reduce the size of the actua-rent draw is attnbutable to the fact that the motor tor and utilize some of the highei efficiencies and on Curve A had begun to be loaded gradually, ap- inertial loads during seating. This solution is, from O preaching its torque rating, causing it to decelerate experience, a fruitless venture as it has been painful-and thereby decrease the worm speed, thus lowering ly proven that the additive factors of efficiency and the unit efficiency. On the other hand, the operator inertia being relied upon so heavily are not always in Curve B had its motor instantly loaded and before present. The best, and applications proven method the motor speed dropped any appreciable amount, to handle high speed actuators is to control these the unit tripped the torque switch, hence, the higher magnification factors to the point of predictability.
efficiency due to higher worm speeds.
l
. Figure 6 Figure 5 EFFECT OF SYSTEM RIGIDITY TYPICAL UNIT EFFICIENCIES ON OPERATOR DUTPUT TORQUE
,d j j {
t i 6
, 50ei-Lettait i
GIARAG f W .
l BAT 10tCWin l nu n toi - '
g.3 5,. $
\ \ - - - , i ,,. .: ~ ,, ..
l / Erf ! A,_, L, , . ,.-
- b. i i i .eistw cs u.
B .,
% ,m ..
. % .s
% a seo y .
3J00 t 900 3 000 b D 5 WORW SPttOIAPWI TIME (Mal.amoansis) * * '.*(~u"ime.conost T1Mr (Interna 4 W,then actenseeri v
_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ . _ _ _ . _ . _ _ _ . . _ _ . . _ _ _ ._. _ .________m
Page C-4 compared to the valve flexibility).
Since 1956. Limitorque has studied ways in (2) Increased operator and stem efficiencies are wluch torque spnng calibration and inertial loading greatly reduced as the loads are gradually apphed ..
for high speed operators could be handled in a man- because of the spring compression and begin to '
ner in which they would be similar in application to approach the minimal efficiencies used in I slow speed operators. The S A type unit, which was a selections.
modified SMA incorporating inertia absort. 7g belle- (3) Inertia is greatly minimized in most applications ville spnngs in the dnve sleeve assembly, was :he cul- and completely eliminated in the remamder.
mmation of these studies.The SA unit along with its Research is continuing in an effort to improve SB and SC counterparts allows the stem nut to float b one or both directions to absorb the impact of the predictability of the many variables in high seating (or backscating) by reducing rigidity in the speed actuator applications. Many valve companies * '
are contributing to this program to incrca'.e the system. If a valve system was defined as Curve B, compatibility of the valve actuator to the valve. A Fig. 6, with the use of the standard SMB or SMC notable advance which has developed from the re-actuator,it would become Curve A with the use of search ori high speed is that the use of the energy an SB or SC type unit. absorbing SB or SC type unit even in standard speed The SB or SC actuators provide a way to in- applications increases the valve life substantially.
crease the controlability and predictability of high This life increase is due to the soft, even application I speed applications. This is accomplished by minimiz- of seating thrust when closing the valve.
ing, if not eliminating n!! of the adverse elements which have been discussed previously:
(1) Motor starter drop-out time has minimal effect Information concerning the Limitorque SB or
) with an SB or SC unit as the load generated by SC units may be obtained from any of the local I
I the overtravel of the actuator during this time Limitorque sales offices or by writing to Limitorque period is primarily absorbed by the springs Corporation Lynchburg. VA 24502.
rather than the valve (because of its flexibihty as i LIMITORQUE TYPE SB @ g r y t-
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