ML18031A359
| ML18031A359 | |
| Person / Time | |
|---|---|
| Site: | Susquehanna  |
| Issue date: | 03/11/1982 |
| From: | Kaza R, Wrona T BECHTEL GROUP, INC. |
| To: | |
| Shared Package | |
| ML17139A731 | List: |
| References | |
| RTR-NUREG-0737, RTR-NUREG-737, TASK-2.E.4.2, TASK-TM D-0026-2, D-26-2, NUDOCS 8206160252 | |
| Download: ML18031A359 (139) | |
Text
ISOLATION/PURGE VALVE ANALYSIS FOR 18"-1200 BUTTERFLY VALVE Project Site Sus uehanna Steam Electric Station Berwick Penns lvania Customer Penns lvania Power & Light Engineer Bechtel Power Corporation Specification No. 8856 Original Purchase Order 8856-P-31-AC Original Pratt 'Job No. D-002'6-2 Valve Tag Nos. HBB-BF-AO-5724, HBB-BF-AO-5725 HBB-BF-AO-5703'BB-BF-AO-5704 General Arrangement Drawings C-2599 Rev. 6 Cross Section Drawing C-2987 Rev. 2 Prepared by:
Date:
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'Reviewed by:
Date:
Certified'by:
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820hi60252 820&10 PDR ADOCK OS000387 PDR
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CONTENTS Pacae I. Introduction II. Considerations.
III. Method of Analysis
~
A. Torque Calculation B. Valve Stress Analysis C. Operator Evaluation
'IV. Conclusion 10 V. Attachments (1) Input Documents (A) Pressure vs. Time Graph (B) Customer/Engineer Response to Request for Information'2)
Valve Assembly Stress Report (3) General Arrangement and Cross-Section Drawings
I. Introduction This investigation has been made in response to a request by the customer/engineer for evaluation of containment isolation/purge valves during a faulted condition arising from a loss of coolant accident (LOCA).
The analysis of the structural and operational adequacy of the valve assembly under such conditions is based principally upon containment pressure vs. time data, system response (delay) time,.
piping geometry upstream of, the valve, back pressure due to ventilation components downstieam of the valve, valve orientation and direction of valve closure.
The above data as furnished by the customer/engineer forms the basis for the analysis. Worst case conditions have been applied in the absence of'efinitive input.
ZI. Considerations The NRC guidelines for demonstration of operability of purge and vent valves, dated 9/27/79, have been incorporated in this 1
evaluation as follows:
A.l. Valve .closure time during a LOCA will be less than or equal to the no-flow time demonstrated during shop tests, since fluid dynamic effec'ts tend to close a butterfly valve. Valve closure rate vs. time is based on a sinusoidal function.
I
- 2. Flow direction through valve, contributing to highest torque; namely, flow toward the hub side of disc if asymmetric, is used in this analysis. Pressure'on upstream side of valve as furnished by customer/engineer is utilized in calculations.
Downstream pressure vs. LOCA time is assumed to be worst case.
- 3. Worst case is determined as a single valve closure of the inside containment valve, with the outside containment valve fixed at the fully open position.
- 4. Containment back pressure will have no effect on cylinder operation since the same back pressure will also be present at the inlet side of the cylinder and differential pressure will be the same during operation.
- 5. Purge valves supplied by Henry Pratt Company do not normally include accumulators. Accumulators, when used, are for opening the valve rather than closing.
- 6. Torque limiting devices apply only to electric motor operators which were not furnished with purge valves evaluated in this report.
Drawings or written description of valve orientation with
'respect to piping immediately upstream, as well as direction of valve closure, are furnished by customer/engineer. In this report, worst case conditions have been applied to the analysis; namely, 90 elbow (upstream) oriented 90 out-of-plane with respect to valve shaft, and leading edge of disc closing toward outer wall of elbow. Effects of downstream piping on system back pressure have been covered in paragraph A.2. (above).
This analysis consists of a static analysis of the valve components indicating if the stress levels under combined seismic and LOCA conditions are less than 90% of yield strength of the materials used.
A valve operator evaluation is'resented based on the operators ability to resist the reaction of LOCA-induced fluid dynamic torques.
Sealing integrity can be evaluated as follows:,
Decontamination chemicals have very little ef fect on EPT and stainless steel seats. Molded EPT seats are generically known to have a cumulative readiation resistance of 1 x 10 rads at a maximum incidence temperature of 350 F. It is recommended that seats be visually inspected every 18 months and be replaced periodically as required.
Valves at outside ambient temperatures below 0 F, if not properly adjusted, may have leakage due to thermal contraction of the elastomer, however, during a LOCA, the valve internal temperature would be expected to be higher than ambient which tends to increase sealing capability after valve closure. The presence of debris or damage to the seats would obviously impair sealing.
III. Method of Analysis Determination of the structural and operational adequacy of the valve assembly is based on the calculation of LOCA-induced torque, valve stress analysis and operator evaluation.
A. Torque calculation The torque of any open butterfly valve is the summation of fluid .dynamic torque and bearing friction torque at any given disc angle.
Bearing friction torque is calculated from the following equation:
TB=P xAXUxd r
where P =pressure differential, psi A = projected disc area normal to flow, in 2 U = bearing coefficient of friction d = shaft diameter, in.
Fluid dynamic torque is calculated from the following equations:
For subsonic flow CR
~ 1 ~ l. 07 (approx. )
a P2 T
D
=D 3 XC lxp x K x F RE For sonic flow P2 D'.4 1
Where T
CR T,=D3 xCT2xP2x D
=
K xF~
fluid dynamic torque, in-lbs.
(F~ 1)
F~ = Reynold number factor RCR
= critical pressure ratio, (f (~) )
Pl = upstream static pressure't flow condition, psia I
P2, = downstream static pressure at flow condition, psia D = disc diameter, in.
CT 1 subsonic torque coef ficient C
T22
= sonic torque coefficient K = isentropic gas exponent ( 1.2 for air/steam mix)
= disc'ngle, such that 90 0 = fully open; 0 0 = fully.
~
c'losed Note. that CTl and CT2 are a function of disc angle, an exponential function of pressure ratio, and are adjusted to a 5" test model using a function of Reynolds number.
Torque coefficients and exponential factors are derived from analysis of experimental .test data and corr'elated with analytically predicted behavior of airfoils in compressible media.
Empirical and analytical findings confirm that subsonic and sonic flow conditions across the valve disc have an unequal and opposite effect on dynamic torque. Specifically, increases in up-stream'pressure in the subsonic 'range result in higher torque values, while increasing Pl in the sonic range results in lower torques.
Therefore, the point of greatest concern is the condition of initial sonic flow, which occurs at a critical pressure ratio.
The effect of valve closure during the transition from subsonic to sonic flow is to greatly amplify the resulting torques. In fact, the maximum dynamic torque occurs when initial sonic flow occurs coincident with a disc angle of 72 (symmetric) or 68 (asymmetric) from the fully closed position.
The following computer output summarizes calculation data and torque results for valve opening angles of 90 0 to 0 o.
D-34933(D-0026-2) TORQUE TABLE f 3/ 9/82 JOB:SUSQUEHAHA/BECHTEL SAT.STEAM/AIR MIXTURE MITH 1.4 LBS STEAM PER 1-LBS AIR SPEC.GR.= .738255 MOL.IJT.= 21.3872 KAPA(ISEHT.EXP.)= 1.19775 R= 72.1972 GAS'OHSTAHT-CALC.
SOHIC SPEED'(MOVIHG MIXTR.)= 1354.57 FEET/SEC AT 265 DEG.
ABSOL.M'AX.TORQUE(FIRST SOHIC)AT 72-68 DG.VLV.AHG.= 23211 IH"LBS e 68 DEG.
MAX.TORQUE IHCLUDES SIZE EFFECT(REYHOLDS HO.ETC)APPX. X 1.25425 FOR 18 IH CH BASIC LIHE I.D.
ALL PRESSURES USED:STATIC(TAP)PRESS.-ABSOLUTE;P2 IHCL.RECOVERY PRESSp (TORQUE)CALC'S VALIDITY!Pf/P2>f.07; VALVE TYPE: f8"-1200 CLASS 150 DISC SIZE 15.7 IHCHES OFFSET ASYMMETRIC DISC SHAFT DIA.: 2.25 IHCHES BEARIHG TYPE: BROHZE SEATIHG FACTOR: 15 INLET PRESS.VAR.MAX.: 48.2 PSIA OUTLET PRESSURE(P6): 26.8 PSIA (72 DEG. ACTUAL PRESS.OHLY(VAR.))
MAX ~ AHG ~ FLOff RATE: 64400 ' CFM> 156772 ~ SCFMr 8618 ~ 18 LB/MIH CRIT.SOHIC FLOff-90DG! 7663.63 LB/MIH AT 30.416 IHLET PSIA VALVE IHLET DEHSITY! . 1 33821 LB/FT"3-MIH..129262 LB/FT 3-MAX.
FULL OPEH DELTA P: 21.9507 PSI SYSTEM COHDITIOHS:
PIPE IH-PIPE-OUT -'AHD- AIR/STEAM MIXTURE SERVICE e 265 DEG.F MIHIMUM 0.75 DIAM. PIPE DO(fHSTREAM FROM CEHT.LIHE SHAFT.
'f ABS. PRESSURE(ADJ.)FOLLOffS TIME/PRESS.TRAHSIEHT CURVE.
&&5 IH.MODEL EQUIV.VALUES --""ACTUAL SIZE VALUES" AHGLE Pf P2 DELP PRESS. FLOW FLOU TD TB+TH TIME (LOCA)
APPRX.PSIA PSIA PSI RATIO (SCFM)
(LB/MIH) "-IHCHLBS"" TD"TB-TH SEC.
90 48.01 24.00 24.01 .500 CR 156771 8618 9866 642 9223 5.00 85 48.02 24.04 23.98 .501 CR 186963 10277 12884 839 12045 6.48 80 48.03 23.66 24.37 .493 181649 9985 11220 730 10489 7.91 75 48.04 22.76 25.28 .474 CR 172296 9471 17901 f166 16735 9.25 72 51.6& 21.57 30.11 .417 CR 157649 8666 23205 1511 21694 10.00 70 48.05 21.12 26.93 .440 CR 'f51432 8324 20940 1364 19576 10.46 65 48.06 19.53 28.53 .406 CR 1290&1 7095 19239 1253 17985 11.5f 60 48.07 l&.21 29.86 .379 CR 1'10966 6100 15463 1007 14455 12.36 55 48.08 17.03 31.05 .354 CR 92052 5060 13725 936 1278& 12.99 50 48.09 lb.30 31.79 .339 CR 74535 4097 11007 1057 9949 13.37 45 4'0 15.79 32-31 .328 72597 3990 9572 1162 8410 13.50 40 48.1 1 15.45 .32.66 .321 49987 2747 6832 1273 5559 13.63 35 48.12 15.11 33'.01 .314 37994 20&& 4300 1351 2949 1 4.01 30 48.13 13.71 34.42 .285 27852 1531 2616 . 1499 f116 14.64 25 48-14 14.80 33.34 .308 19155 1053 1520 1544 "24 15.4.9 20 48.15 'l4.74 33.41 .306 11715 644 997 1724 -726 16.54 15 4'6 14.71 33.45 .305 6782 372 , 358 1940 -1582 17.75
-1880 19.09 10 48.17 12.99 35.18 .270 3169 174 207 2087 5 48.18 11.31 36.87 .235 1039 57. 136 2216 "2080 20.52 0 48.20 14.70 33.50 .305 0 0 6684 1824 4859 22.00-SEATIHG + BEARIHG + HUB SEAL TORQUE (M/M)= 6684 IH-LBS e 0 DEG.
MAX.DYH. - BEARIHG - HUB SEA'L TORQUE (M/M) = 23205 IH-LBS 8 70 DEG.
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B. Valve Stress Analysis The Pratt butterfly valve furnished was specifically designed for the requirements of the original order which did not include specific LOCA conditions.
The valve stress analysis consists of two major sections: 1) the body analysis, and 2) all other components.
The body is analyzed per rules and equations given in paragraph NB 3545 of Section III of the ASME Boiler and Pressure Vessel Code.
The other components are analyzed per a basic strength of materials type of approach. For each component of interest, tensile and shear stress levels are calculated. They are then combined using the formula:
Smax = 3: (Tl+T2) + 1- (Tl+T2) + 4 (Sl+S2) 2 2 where Smax = maximum combined stress, psi Tl = direct tensile stress, psi T2 = tensile stress due to bending, psi Sl = direct shear stress, psi S2 = shear stress due.to torsion, psi The calculated maximum valve torque resulting from LOCA conditions is used in the seismic stress analysis, attachment 02, along with "G" loads per design specification. The calculated stress values are compared to code allowables if possible, or LOCA allowables of 90%.
of the yield strength of the material used.
C. Operator Evaluation Model: Bettis T312-SR3 Rating: 52,300 in-lbs at full open and closed positions only.
34,500 in-lbs at 68 29,J00 in-lbs at 45 (minimum rating) .
Maximum Valve Torque: '3,211 in-lbs at 68 0'he maximum torque generated during a LOCA induces reactive forces in the load carrying components of the actuator.
Since the LOCA induced torque derived in this analysis is less than the maximum absorption rating of the operator, it is concluded that the Bettis models furnished are structurally suitable to withstand combined LOCA and seismic loads.
10
~
IV. Conclusion It is concluded that the valve structure and the valve actuator are both capable of withstanding'ombined seismic and LOCA-induced loads based on the calculated torques developed in this analysis.
ATTACHMENT 1A PRATT PROPOSAL LETTER
) PHATT, HHNHY PHATT COlvlPANY
(:I;C:Xtix( ( t. "it>(( I it> i'()I I'l(ti(l ~.iit(.'ItIS 401 Sotfl H HIGIILANDAVEYUE AURORA. ILLla40)S 6(MOF
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April 16, 1981 Bechtel Power Corporation P.O. Box 3965 San Francisco, CA 94119 Attention: Mr. E.B. Poser Project Engineer
SUBJECT:
Susquehanna Steam Electric Station Containment Isolation/Purge Valve Analysis Gentlemen:
'ith reference the valves and to your recent inquiry regarding suitability of actuators to withstand aerodynamic LOCA conditions, please note the following:
Torque calculations will be performed for aerodynamic torque generated as a result of LOCA. These calcula-tions will be pe formed using the following data to be furnished by you.
A. Containment Pressure Time Curves.
B Containment Temperature Time Curves C The combined resistance coe ficient for all "
ventilation system components downstream of the valve (one for'each valve size) or A graph of back pressure vs. LOCA time at a distance 10-12 diameters downstream of the valve. Consider also the capacity of the piping, filter and duct work to resist increases in back pressure.
D Maximum and minimum delay times from LOCA to initiation of valve rotation.
Eo Drawings or written description of'valve orientation with respect to elbow immediately upstream of valve (within 6 diameter's), as well as direction of valve closure (clock-wise or counterclockwise) as viewed from operator end.
Bechtel Power Cc.. adoration Page April 16, 2
1981 5'HATTY In the absence of the above information, the following assump-tions will apply to the purge valve analysis.
- 1. Back pressure of 19.7 psia throughout valve closing cycle. Higher back pressure increases maximum dynamic torque and valve stresses.
- 2. Delay time from LOCA to initiation of valve rotation shall be chosen to permit initial sonic flow condi-tion and critical valve disc angle to coincide, resulting in maximum possible dynamic torque.
3 ~ 90o elbow immediately upstream, oriented 90 qf-plane with respect to valve shaft, with leading edge of disc closing away from 'outside radius of elbow. Such orientation and closure will increase torque values by 20% or more.
- 2. Based on the above results, a static load stress analysis will be provided for valve components affected by the dynamic torque loadings in combination with pressure and seismic loads.
The actuator supplier will be asked to verify the suitability of the actuator or the reaction or back drive force resulting from aerodynamic torque conditions.
The cost of performing the evaluation of valve components will be $ 12,800 each size "or 6", 18" and the 3~
24" valves.
4 ~ The completion of this analysis is projectedto be twenty-six (26) weeks after receipt of purchase order and data requested.
above based on availability of engineering schedule.
- 5. Our response to NRC's criteria for demonstrating operability
~
'of purge valves is included in the analysis.
. 'his proposal is for investigative analysis only and is not intended to guarantee the adequancy of the equipment as fur-nished when subjected to LO A loads currently being defined.
The proposal is valid for thirty (30) days. The terms of payment will be'Net 30 Days.
We hope you will find the proposal responsive to your needs. If we can be of any additional assistance in this matter, please advise.
Very truly yours, HENRY PRATT COMPANY
~~/
Glenn L. Beanc GLB/tl ~
Manager-Application Engineering
ATTACHMENT 1B CUSTOMER/ENGINEER RESPONSE TO REQUEST FOR INFORMATION
~ppUPg t(flJ Pt"",K
)9 f Bechtel Power Corporation Engineers Constructors Henry Pratt Company Fifty Beale Street
.401 South Highland Avenue San Francisco, California Aurora, Illinois 60507 Mail Address: p.o. Box 3965. san Francisco. GA 94 t 19 Attention: Mr. G. L. Beane
Subject:
Susquehanna Steam Electric Station Units gQ 1 and 2 Job 8856 P.O. 8856-P-31-AC, Containment Isolation/Pu e Valve Anal sis Ke Gentlemen.
In order to perform the analysis Henry Pratt requested certain information.
The following is our reply:
A. Containment pressure time curve is attached.
B. Containment temperature time curve is attached.
C. A back pressure of 19.7 psia should be used in this analysis. &is back pressure is per the assumptions in your letter of April 16, 1981.
Do Minimum delay time is O.l seconds. Maximum delay time is 5 seconds.
E. Isometric drawings for both units are attached. We believe that Hemy Pratt is in a better position to determine the direction of valve closure as viewed from the operator end. This information is not apoarent on the drawings you submitted to Bechtel.
In addition, if Henry Pratt's 16 week analysis report shpws the valves to be unqualified, Henry Pratt will state at what angle the valves must be blocked open in order to meet the NRC's interim position. Henry Pratt will also make recomtendations on how to block the valves and to provide a detailed drawing of the stop.
We trust that the foregoing information is satisfactory and will enable you to complete the qualification of the subject valves.. If you have any questions, please contact Al Daily at (415) 768-9235 or A. Tiongson at (415) 768-7770.
Very truly yours, E. B. Poser Project Engineer Written Response Req'd: No Design Document Changes: "No CHN/APT/cgs WP30/3-1 cc Mr. T. M. Crimmins, Jr. (PL) w/a 5025 Qp 0 ~g Vw/0 ~ (Wgl
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ATTACHMENT 2 Hu~lf
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SEISMIC 'ANALYSIS FOR 18 INCH NUCLEAR PURGE VALVE
TAHJ,E OF CONTI'.ITS
~Pa e
.List: oF. Fiq>>rcs 2 Nomenclature Summar Tables 19 Stress Level Summary 21 I:requcncy Analysis Summary 26 Valve Dimensional Data 27 Stress Ana1isis 29 Introduction 30 End Connect:ion Analys is 3'4 Body Analysis 35 Disc Analysis 41 S1iaft Ana1ysis '44 Disc Pin Analysis 46 Shaft Bearing Analysis 48 Cover Cap Analysis TllrQs t Searing Ana lysis 52 Opera t or I lount hi g An a lys i s 53 ):req>>cncy A! a
~ li s i s 64
LIST OF FIGURES. LXo Title ~Pa e Valve Body Spatial Orientati o>> 31 Valve Cross-Section 3 Pressu're Area Analysis Cross; 36 Section in Crotch Region
- 4 Pressure Area Analysis Cross-. 39 Section in Bod'y Di'sc 42 Bottom. Trunnion Assembly 50 Top T'runnion bfounting 54 Trunnion Bolt Pattern 57 Bonnet Bolt Pattern 59 Adapter Plate Bolt Pattern 60
NOh/Eb:CLA'I URE The nomenclature for this analysis is 'based upon I the nomenclature established in Paragraph NB-3534 of
'ection III of the ASME Boiler and Pressure Vessel Code. lthere the nomenc lature comes directly from the. code, the re fcrence paragraph or figure for that
~
symbol is given With th e definition. For symbols C not defined in the code the definition is that assigned by Henry Pratt Company for use in this analys is'. p):
~ ~ ~ ~
ANAI.YSIS NOhlENCLATURE Af Effective fluid prcssure area based on fully corroded interior contour for calculating crotch primary mem-brane stress (NB-3545. 1 (a) ), in2' Am hfetal area based on fully corroded interior contour effective in resisting fluid force on Af (NB-354S. (a) ), in2 A3 Tensile area of cover cap bolt,'n2 A4 Shear area of cover cap bolt, in . A"5 . Tensile area of trunnion bolt, in2 A6 Shear area of trunnion bolt, in A7 Tensile area of. operator bolt, in2 A8 Shear area of operator bolt, in Bl Unsupported shaft length, in. Bearing bore diam'eter., in. B3 Bonnet bolt tensile .area, in2 B4 Bonnet. bolt shear area, in2 BS Bonnet body cross-sectional area, in2 B6 Top bonnet wold size, in. B7 Bottom bonnet weld. size, in. I B8 Distance to outer fiber of bonnet from shaft on y axis, in. Bg 'istance to outer fiber of bonnet from shaft on x axis, in." A factor depending upon the method of attachment of head, shell dimensions, and other, items as listed in NC-3225.2, dimcnsionless (Fig. NC-322S.1 thru Fig. NC-.3225.3) 13 ANALYSIS NOMENCLATURE Cb Stress index for body bending seconda'ry stress re-sulting- from moment in connected pipe (NB-3545.2(b)) Cp Stress index for body primary plus secondary stress, i>aside surface, resulting from internal pressure (NB-.3545.2(a)) C2 index for thermal secondary membrane stress
'tress resulting from structural discontinuity '
C3 Stress index for maximum secondary membrane plus bending stress resulting from structural discontinuity C6 Product of Young's modulus and coefficient of linear thermal expansion, at 500oF, psi/oF(NB-3550) C7 .Distance to outer fibe'r of disc for bending alon'g the shaft, in. cs Distance to outer fiber of disc for bending about the shaft, in. C9 Distance'to outer fiber of flat plate of disc for
'bending of unsupported flat plate, in.
I'nside diameter of-,body neck at crotch region (NB-3545.1(a)), in. Inside diameter used as basis 'for determining body minimum wall thi'ckness, . (NB-3541), in. Dl Valve nominal diametor, zn. D2 Shaft diameter, in. D3 Disc,'in diameter, in. D4 Thrust collar outside d i ameter, in. D5 Spring pin diameter,. in. D6 Cover cap bolt diameter, in. D7 Trunnion bolt diameter, in. 1.3 ANALYSIS NOMENCLATURE Ds Operator bolt diameter, in. Dg Bonnet bolt diameter, in.'. E hfodulus of elasticity, psi Fb Bending modulus -of stand'ard connected pipe, as given by Figs. NB-3545.2-4 and NB-3545.2-5, in.> Fd 1/2 x cross-sectional area of standard connected pipe, as given by Figs. NB-3545.2-2 and NB-3545.2-3, in,2 FN Natural frequency 'of respective assembly,.hertz Fx W3gx--Seismic force along x axis due to seismic acceleration acting on operator extended mass, pounds Fy W3gy--Seismic force along y axis due to seismic acceleration acting on operator extended mass, pounds Fz W3gz--Seismic force along z axis due to seismic acceleration acting on operator extended mass, pounds Gravitational accelerate.'on constant, inch-per-second Valve body section bendi:lg modulus at crotch region (NB-3545.2(b)), in> Gd Valve body section. area at crotch region (NB-3545.2 (b)), in2 Gt Valve body section tor'sional modulus. at crotch region (NB-3545.2(b)), in>. gx Seismic acceleration constant along x axis'eismic acceleration constant along y axis gz Seismic acceleration constant along z axis
.L3 ANALYSIS NOi~t)!NCLATURE 2 Top trunnioh bolt square, in. H3 Bottom trunnion bolt square, in.. H4 Bonnet bolt square, in. Operator bolt square, in. H6 Bonnet bolt circle, in. H7 Operato'r bolt circle, in. H8 .Bonnet height,. in. H9 Actual body wall thickness, in.. I 'Bonnet body moment of inertia about x axis, in4 l'2 Bonnet body moment of inertia about y axis, in" I3 'isc shaft, area moment of inertia for bending about tLe in4 Disc area moment, of inertia for bending along the shaft, in4 hfomcnt of iI crtia of valve body, in4 hfoment of inertia of shaft, in< Disc area moment of inertia for bending of unsupported Q.at plate, in4 Distance to neutral bending axis for top trunnion bolt patter>> along x axis, in... J2 Distance to neutral bending axis for top'runnion bolt p attcrn alon g y axis in. J3 .Distance to neutral bending axis for bonnet bolt pattern along x axis, in.
'istance I
J4. ~ to neutral bending axis for bonnet'olt pattern a3ong y axis, in. Distance to neutral b'ending axis for operator bolt pattern along x axis, in. Distance to >>cutr;ll bending axis for operator bolt P:lttcrll;llollg y axis 1Il,
ANALYSIS NOMENCLATURE
. Spring constant Kl Distance of bonnet leg from shaft centerline, in.
Kp Thickness of disc above shaft, in. Kg Length along z axis to c. g. of bonnet plus adapter plate assembly, in. K4 Top trunnion width, in. Kg Top trunnion depth, in. K6 IIei g ht of to P trunnion in. Ll Valve body face-to-face dimension, in. L2. Thickness of operator housirig under trunnion bolt, in. Lg', Length of engagement of cover c'ap bolts. in bottom. trunnion, in.. $ '~ Length of engagement of trunnion bolts in top trun'nion, in. r LS Bearing .length, in. L6 Length of structural .disc hub welds, in. L7 Length of engagement. of bon'net bolts in'adapter plate, in. L8 Length of engagement of b'onnet bolts in bonnet, in. r
" L9 Length of engagement of stub shaft in disc, in.
Reciprocal of Poisson's ratio Mass of component Mx Ng(gyZo+gzYo), operator extended mass seismic bending moment about the x axis, acting at the base of the operator, in-lbs.
13 ANALYSIS NOi~ll!NCLATURL'y N3 (gxZp+g Xp), operator- extended mass seismic bending. moment about the y axis, acting at the base of the operator, in-lbs. Mz, tt3 (gxYp+gyXp), oPerator extended mass seismic bending moment about the - axis,. in-lbs. C
. Ffx hIx+F>T5, operator extended mass spismic bonding moment abou't the x axis,'cting at the'ottom of the adapter plate, in-lbs.
, Hy My+FxT5, operator extended ma'ss sei'smic bending moment about the y axles, acting at the bottom of the adapter plate, in.-lbs.
Mx+F (T5+Ha)+gylf4K3, . operator'xtend'ed mass seismic bending moment about the x- axis, acting at the base of the bonnet, in-lbs. hly+Fx(T5+HB)+gxl'4K3, operator extended mass seismic bending moment about the y axis, acting at the base of tho bonnet, in-lbs. 4
'. hip Bending moment at joint of flat plate to disc hub,.
in-lbs.
~ ~
Na number I'ermissible of complete start-up/shut-do>>m
.cycles at hr/100oF/hr/hr fluid temperature change
,rate (NB-3545.3)
NA Not applicable to the .analysis of the 'system
'I Nl Number of top disc pins
,Number of operator bolts Number'of trunnion bolts Pd Design prcssure, psi pr Primary pressure rating, pound ps Standard calcu'lation .pressure from Fig. NB-3545..1-1, ps1 pe Largest value among Pcb, Ped, Pot, psi I
ANA;LYS S NOMENCLATURE'eb Secondary'tress in crotch region of valve body caused by bending of connected standard pipe, calculated according to NB-3545.2(b), psi ' Ped Secondary stress in crotch region of valve,'body caused by direct or axial load imposed by connected standard piping, calculated according to N8-3545.2(b), psi Pet Secondar> stress in crotch region of valve body .caused by twisting of con'nocted standard pipe, calculated
.'according to NB-3545.2(b), psi Pm 'General primary membrane stress intensity at crotch region, calculated according to NB-3545.1(a), psi Pm'p Primary membrane*stress intensity in body wall, psi Sum of primary plus secondary stresses a't crotch resulting"from*internal pressure, (NB-3545.2(a)), ...
psi
- QT Thermal stress in crotch region resulting from'00oF/
hr fluid temperature .change rate,'psi Maximum thermal stress. component caused by through wall temperature, gradient associated with 100oF/hr fluid tempera'ture change rate'(NB-3545.2(c)), psi . Maximum thermal secondary membrane stress resulting f'rom 100oF/hr fluid'emperature change rate, psi Maximum thermal secondary membrane plus bending stress resulting from structural d~.scontinuity and 100oF/hr fluid temperature change m< te, psi Mean radius of body wall at crotch region (Fig. NB-3545.2(c)-X), 'in. Inside radius of body at crotc)i reiiion for calculating Qp (NB-3545.2(a)), in. r2 Fillet radius of oxtcrna) surface n~ crotch (NB-3545.2(4)), in.
13 ANALYSIS.NOllENCLATURE Disc radius, in. R5 Shaft radius, in. Rm Mean radius of body wal3, in. Rg Radius to 0-ring in cover cap, in.
~
Assumed maximum stress in'onnected pipe for cal-culating Pe (NB-3545. 2 (b) ), 30,000. psi Sm Design stress intensity, (NB-3533), psi Sn Sum of primary plus secondary stress intensit'ies at crotch region resulting from 100oF/hr temperature change rate (NB-3545.2), psi Spl . Fatigue stress'intensity'at inside surface in crotch region. resulting from 100oF/hr fluid temperature change rate (NB-3545;3), psi
. Sp2 Fatigue stress intensity at outside surface in crotch region resulting from 100 F/hr fluid temp-erature change rate (NB-3545.3), psi S(l) through S( 83) are listed after the alphabetical section.
~
Minimum body wall thickness adjacent to crotch for calculating thermal stresses (Fig. NB-3545.2(c)-1), in. Minimum body wall thickness as determined by NB-3541, ln Te Maximum effective metal thickness in crotch region for calculating thermal, stresses, (Fig. NB-3545.2 (cg -1), in. 4T2 Maximum magnitude of the difference in average wall temperatures for walls of thicknesses te, Te, resulting from 100oF/hr fluid temperature change rate, oF Thick>>ess of cover cap behind bolt .head, in. T2 Thickness .of shaft behind spring pin, in.
ANAl.YSI S NOSER'Cl.ATURE ~ T3 Thrust collar .thickness, in.'over T4 cap thickness, in. T5 Adapter plate thickness, in. Thickness of bottom bonnet plate, in. Thickness of top bonnet plate, in. T.8 maximum required operating torque for valve, in-lbs Ul Area of bottom bonnet weld, in
. Area of..top bonnet. weld, in U2'3
.Shaft bearing coefficient of friction Bearing friction torque due to pressure ~
loading U4'5. (shaft jo'urnal bearings), in-lbs. Bearing friction .torque due t'o pressure loading plus seismic loading (shaft j.ournal bearings), in-lbs. U6 Thrust bearing'riction torque, in-lbs. 4 Distances to bolts in bolt pattern on adapter plate, i.n i Distances to bolts in bolt pattern on adapter plate, in. V3 Distances to bolts in bolt pattern on adapter plate, in. V4 Distances to bolts in bolt pattern on adapter plate, 'V5 Distance to bolts in bolt pattern on bonnet, .in. V6 Distance to bolts in bolt pattern on bonnet', in.' V7 Distance to bolts in bolt pattern on bonnet, in. V8 Distance to bolts in bolt pattern on bonnet, in. Total bolt load, pounds Valve weight, pounds
'Banjo weight, pounds'perator weight, pounds N4 Bonnet and.adapter plate assembly weight, pounds 13
. ANALYSIS NOMENCLATURE Weld size of disc structural welds, inches Weight of disc, pounds ws Length of weld around perimeter of bonnet, in. X( Eccentricity of center of gravity of operator extended
- mass along x axis, inches E
Yo, Eccentricity of center of gravity of operator extended mass along y axis, i'nches
,Z o Eccentricity of center of gravity of operator extend'ed ~
mass along z axis, inches Zl Bending section modulus of bottom bonnet welds, in Z2 . 'ending section modulus of top bonnet welds, in> Torsional section m'odulus of bottom bonnet welds,'n> '. Z3'7 Torsional section modulus of top bonnet welds, in3 ffaxi'mum static deflection of component, inches Distance to edge of disc hub, inches
13 ANALYSIS NOMENCLATURE S (1) Combined bending stress in disc, psi S(2) . Bending stress in disc"due to bending along the shaft, psi S(3) Bonding stress in disc duc to bending about the shaft, psi S(4) Bending tensile stress in unsupported flat plate, psi S(~) Shear tear out of shaft through disc, psi S(6) Shear stress across structural hub welds of disc, psi S(7) Combined stress in shaft, psi, S(8) Combined bending stress .in shaft; psi S(9) 'ombined shear stress ir. shaft, psi
'S(10) Bending stress in shaft due to seismic and pressu're loads along.x axis, psi S'(1 1) Bending stress in shaft due to seismic load along y axis, psi S(12) . Torsional shear stress in top shaft due to operating torque, psi J
S(13) Direct shear stress in shaft due to seismic and prcssure loads,.psi S (1'4) Torsional shear stress at reduced disc pin cross-section, psi
. S(15) Shear stress acro'ss top disc'pin due to operating tolque psi S (16) Bearing stress on top disc pin, psi S(57) Combined shear stress across. bottom d3.sc p1n> ps1 S(18) She:l'r stress across bottom disc pin due to tor-sional load, psi S(19) Shear stress across bottom disc pin due to seismic load,.
Collpressivc stress s}laft bearing due'to seismic ps'(20) on and pl'css'ul c I Gads > ps 1
13 ANALYSIS NOh<I!NCLATURH S(21) Shear tear out of cover cap bolts through tapped holes in bottom trunnion, psi S(22) Shear tear out of cover cap bolt head through bottom
~ 'over cap, psi S(23) 'Combined stress in cover cap 'bolts, psi S(24) 'Direct tensile stress in cover cap bolts, psi S(25). Shoar s.ress in cover cap bolts due to torsional loads, psi S(26) Combined stress in cover cap, psi S(27) Radial stress in cover cap, .psi S(2S), Tangential stress in cover cap, ps'i S(29) Shear stress in cover cap, psi S(30) Bearing stress on thrust collar, psi S(31) Shear load on thrust collar s p rin g p in p ounds
~
S(3?) . Bearin'g stress of spring pin on thrust collar, psi I S(33) Shear tear out. of spring pin thorough thrust collar, psi S(34) Shear tear out of spring pin through bottom shaft, psi
13 ANALYSIS NOi'IEWCLATURE -W..-:s(ss) Shear tear out of trunnion bolt through tapped hole ln trunnion psl s(s63 Ecarinj stress of trunnion bolt on tapped hole in trunni'on, stress of, trunnion bblt on through hole in ps'(37) 13caring bonnet plate, psi s(~s) Sh ar tear out of trunnion bolt head through bonnet plato, psi s(~9) Combined stress in trunnion bolt, psi S ('AO) Direct tensile stress in .trunnion bolt, psi S(4l) Tensile stress in trunnion bolt due to bending . moment, psi
.S(42) 'Direct shear stress in trunnion bolt, psi.
s(4~) Shear stress in trunnion, bolt due to torsional load, ps1
'S (,44) .Shear'ear out of bonnet bolt through tapped hole in bonnet, psi S(.4S) Bearing stress of bonnet bolt on tapped hole in bonnet., psi C
s('46) ~ Bearing stress of bonnet bolt on through hole. in adapter plate, psi S( 47) 'hear tear out of bonnet bo'lt head through adapter plate, psi S( 48) 'ombined stress in bonnet bolts, psi s(49) Direct tensile stress in bonnet bolts,.psi S(SO) ...Tensile stress in bonnet bolts duc to bending'moment, psl S(Sl) ~ - Dirac". shear stress in bonnet bolts, psi I'
ANAJ.YSIS NOWENCJ.ATURJ-' Shear stress in bonnet bolts due to torsional. loads, psi S(53) . Shear t'ear out of operator bolt'head th>ough,adap'ter
..plate, psi
'(54) ~
Bearing stress'of operator bolt on through hole in adapter plate S(55) Combined stress in operator bolt, psi S (56) , Direct tcnsi'le stress in operator bolt, .psi S(57) Tensile stress in opera'tor bolt 'due to bending moment; PS 3. S(53) Direct shear stress in op'erator.bolt, psi S(59) Shear stress in operator bolt due to bending moment','S1 S(60) Combined stress in. bonnet body, psi . ..:.. S(61) Direct tensile str'ess in bonnet body, psi S (6Z) ~ ..'I'ensile 'stress in bo.>net body duc to bcndi'ng moment,,
~PS3. *
,S ( 63) Direct shear stress in 'bonnet body, psi S( 64) Shear stress in bonne't body due to torsional load, PS 3.
S(65) Combined shear stress in bottom bonnet weld; psi S( 66) Total tensile.stress in bottom bonnet weld, psi. S( 67) Direct tensile stress in bot'tom bonnet weld, psi S(68) Tensile stress in bottom. bonnet weld due to bc>>ding moment', psi S(69) Total 'shear stress in bottom bonnet weld, psi S(7o) ~ Direct shear str'ess in bottom bonnet weld, psi
1 3 ANAJ.YS IS NO.'1EflCI.ATUI<F. S(71 ) Shear stress in'ot'tom bonnet weld duc to torsional load, psi s(72 ) Combi.ndd shear stress in top bonnet weld, psi .
'(7j) Total..tensile stress in top bonnet weld, psi S(74 ) Direct tensile stress in top bonnet weld, psi s(7s ) Ten'sile stress. in top bonnet weld due to bending
~
moment, psi . 0 s(76) . Total shear stress in top bonnet weld, psi s(77 ) Direct shear stress in top bonnet weld, psi s(78) Shear st'ress. in top bonnet'weld due to torsional load, psi S(79 ) Combined stress'n trunnion'ody,. ps'i s(8o ) Direct tensile stress in, trunnion body,. psi s(81) Tensile stress in trunnion body due to bending moment, ps1 I
"'S'(82 ) -Direct shear. stress in .trunnion body, psi s(8s ) Shear stress in trunnion. body due to torsional'oad q ps1
A 13 SUMMARy TABLE I !TRODUCTION
'In the folio>>ing pages, the pertinent data for the butter-fly valve stress.analysi's is tabulated in" three categories:
l.. Stress Levels for Valve Components
- 2. Natural Frequencies. of Components
- 3. Valve Dimensional Data In. Table 1', Stress Levels for Valve Components, the following data is tabulated:.
Component Name Code Reference (when applicable) Stress Level Name. 'and Symbol Analysis 'Refer'ence Page Material Specification, Actual Stress Level Allowable Stres's Level The material specifications, are taken from Section II of the code when applicablc. Allowable stress levels are Sm for tc>>.,ilc stresses and .6 Sm for shear stresses. Thc allowable J levc)s are t)>c same whether thc calculated stress is a combined stress or results from a single load condition. Sm is the clesign stress intensity value as defined in Appendix I, Tables I-7.1 of Section III of the cocle. In Tab)c 2,'atural Frequencies of Valve Components, the folio>>'ing .data is tabulated:
Summar Table Introduction Component Name Natural Frequency Symbol Analysis Reference Pago
.Component bfatcrial Natura.l Frequency In Tal>lc 3, Valve Dimensional Data, the values for the pertinent valve dimensions and parameters are. given.
~
~
~ ~ ~
~ ~ ~ ~ ~ ~ ~ ~ I~ ~
~ ~
~ ~
~ ~
~ ~ ~
~ ~
~ ~
~ ~ ~
~
~ ~ I ~ ~ ~ ~
~
~ I - ~
~ ~ ~
~ ~ ~
~ ~
~ ~
~ ~ ~ ~ ~ ~ ~ I I ~
~ ~
~ ~ ~ I
~ 0 ~ ~
I ~ I ~
~
~
~ ~
~
~
I I
~ ~ \
~ ~
~
2 '
~ ~
~
~
0
~
~
~ ~ ~ ~ ~
~ ~ ~
~
I I I
~ ~
~ ~ ~ ~
~ ~ ~
eel
~ ~ ~
owl
~ ~
~ ~
~ ~ I ~ ~
~ ~ ~
0
~ ~
~ ~
~ ~ ~
~ ~ ~
~ ~ ~
~ ~
~ ~ 4
~ ~ ~
0 ~ ~ ~
~ ~
~ ~ ~ ~ ~ 0 ~ ~ ~ ~
~ ~
TABLE 1 'TRESS LEVELS FOR VALVE COMPONENTS I II ALLONABLE, CODE REFT REF STRESS STRESS LEVE I COMPONENT PARAGRAPH SYMBOL 6 NAME PAGE MATERIAL LEVEL, PSI PSI Operator Shear tearout of Mounting Bonnet bolt head .6 Sm ~ (Cont'd) hrough adapter. plate 56', A-36 7560 S (47.) . ASTM ombined stress in Bonnet bolt S (48) 56 ASTM A-193 GR 87 .9X yield 94 5'00 Shear tear out of perator bolt head SM through adapter plate S (S3) 'S8 ASTM A-36. 12600 Bearing stress of SM perator bolt on S (54) 58 ASTM A-36. 12600 dapter plate I OMBINED STRESS IN .9Xyield perator bolt S (SS) . ASTM A-193 GR B7 94500 ombined stress in onnet body S (60) 61 ASTM A-36 ) hi~3 .9Xyield 32400 ombined shear stress S(6S) .6 SM in bottom bonnet .weld l 61 7560 ombined shear stress .6 SM.' in .top bonnet .welds . S .(72) 62 )zoo 7560 ombined..stress in 3.2 1+ SM trunnion 'body S (79) 63 ASME SA-516 GR.55 13700
P ~
~ ~
0 Table 2 NATURAL WRBQUENCIBS OP. VALVE COMPONENTS Component Name - . Natural Frequency Symbol
'ef. Page Material Natural Frequency (Hertz) .
58 ASWB,Sg-516 FN1 Gr.55 ~S w9 Banjo AShlL Sh-564 TVV<<30 Cond.. ll-1 1 50 Cover Cap FNS S9 AShfI! SA-516 Gr'. 70 Bonnet FN4 60 AS'l'hl A-36
Y 0
30 TABLE 3 DIMENSIONAL DATA
'ob Number: D- Valve Size.' .
)2.cc Operator Mounting: Operator:
Ag C3 C6 A3 A4 CS GT Ag. <x A6 d 7g Sy Ay dm Sz As H2 ~
~Q By Dg H3 P
.Bg D3 H4
~
Hg H6 BS D6 Hg Dy H8 Da Hg-B8 Dg Bg
~
i Fb Cb Fd I4 Fx Co Fy I6 J,a< Cp i' \ l
'I
<<27>>
30 Mz hT2 Jz ~, ~ M Jwb7 T2 J4 "x ~ a My T4 M8 TS Ko 3 ooa Tg Kl . Nl T7 .6a5 K2 NZ TS N3 Ul. U2 U3 ps Vl le c V2 L2 5 r5l V3 Lg V4 L4 l.o Vg Lj R4 L6 Rg V7 L7 VS LS R6 3 ~ %7 Lg S, Rd rod Mx )ao . tm
30 N ' X Y Z p Zg Z2 Z3 Z7 28a
Qi Pages 21 "28, Stress Level Summary, Frequency Analysis Summary, and Valve Dimensional'Data sheets have been assembled at the beginning of the report submittal. They are located directly behind the design review 'record for the corresponding'roduction order. 21-28
Stand'ard Stress Report
. for NRS Butterfly Valve'1th Bonnet h1ounted Cylinder Operator ANAI,YSIS INTRODUCTION Described in the following pages is the analysis used in verifying the structural adequacy of the main elements of the NRS butterfly valve. ~
The analysis is structured to comply with Paragraph N)3-3SSO of Section III of the ASME Boiler and Pressure Vessel Code (hereafter referred'to as the code). In the analysis, the design rules for Class 1 valves are used, since the requirements for this class of valve is much more 'xplicit than for either Class 2 or 3 design rules. The de-
~ ~
sign rules for Class 2 and 3 are exceeded by the rules for Class 1 valves. Valve components are analyzed under the assumption that
'the valve is either at ~
maximum fl.kaid dynamic torque or seating
~ ~
against the maximum design pressure. ~ Analysis temperature is t 300 F. Seismic. accelerations are simultaneously appl'ied
~
'in each of three mutually perpendicular directions, Seismic loads are made an integral part of the analysis by the inclusion of the accelerati.on constants gx, gy gz. The symbols gx, gy, gz represent accelerations in the x, y and z directions respectively. 'These directions are defined with respect to the valve body centered co-ordinate system as illus-tratcd in Figure 1. Specifically, the x axis is along the pipe axis, the z axis is .along the shaLt axis, and tlute y axis. is mutually perp'endicular to the x and z axes, forming a right hand triad witl> them. 1'inure 1 VAl.Yl 1/ODY Si'ATIAL 0]~11!NTATIO
'I
0 13 Anil+1'5 is I)) tl'oductiotl V~lve orientation with respect to g)':)v)ty is taken into account by adding the appropriate quantity to the seismic loads. The justification for doing this is that a gravita-tional load is completely .equiva'lent to ;) 1g seismic load.
'I The analysis of each main element of sub'-assembly of the butterfly. valve is described separately in an appropri-ately titled section. In 'addition to containing sketches where appropriate, each section contains an explanation of t'e basis for each calculation. Where applicable, it also contains an interpretation of code requirements as they
\
apply .to the analysis. Figure 2 is a cross.-section view of the butterfly valve, and its associated components. Detailed sket'ches are pro-vidcd throughout the report to clearly define the. geometry. 2
TOP SHAFT SHAFT PACKING
-SHAFT BEARING TOP TRUNNION VALVE SEAT PISC PINS DISC
'BOTTOM SIIAFT BOTTOM TRUNNION COVER CAP
<<33
I:AD CONNECTION ANALYSIS The NRS butterfly valve is a uniflange design. Rather than having flanges that age external to and distinct from the body, the body shell is fabricated so that the end connections aro mach,inod directly into the body shell. The outside and inside diameter of the body shell conform to the requirements of the American National Standards Inst'itute
'I (ANSI) standard B16. 5. The end connections, either flanged or weld end,. also conform to this standar'd.
BODY ANALYSXS
.The body analysis consists of calculations .as detailed in Paragraph NB-3540 of Section ?EX of the Code. Paragraph NB-3540 is not highly oriented to butterfly valves as related to various design and shape rules.r Therefore, certain of the r
design equations cannot be. directly applied for butterfly valves. Where interpretation unique to the calculation is necessary, it is explained in the subsecti,on containing that calculation description. 1 Figure 3 illustrates the essential features of the body geometry through the trunnion area of the valve. The symbols used to'define specific dimensions-:are-consistent with ~ those used in the analysis and with the nomenclature used in the Code.
- 1. Minimum Body. Wall Thickness Paragraph NB-3542 gives minimum body wall thickness re-quirements for standard pressure rated valves.
The actual minimum wall thickness in the NRS valve, occurs between the flange bolt holes and body bore. Ag I?RESSURE -AREA ANALYSIS BODY CROSS-SECTION Figure 3,
<<36-
~lip<1 Analyst.s 2.. Bod Sha o Rules The NRS valve meets the requirements of Paragraph NB-3544 of the code for body shape rules. The ex-ternal fillet at trunnion to body intersection must be greater than thirty percent og the minimum body wall thickness. 3 ~ Primary i~lembrane Stress Due to Internal Pressure Paragraph NB-3545.1 defines the maximum allowable stress in the. neck to flow passage junction. In a butterfly valv~', this corresponds with the trunnion to 'body shell junction.. Figure 3 shows the geometry throu'gh this section. The code defines the stresses in this area using the pressur'e area method. As seen ip Figure 3, certain code-defin'ed dimensions are not applicable to this style of butterfly valve. For example, there is no radius at the crotch. when seen in a view along the flow pa'ttern, as the neck extends to the face of the body. To comply with .the intent of the code, the areas Af and Am are interpreted as shown in th'c cross-section (Figure 3). Using these areas, the priinary
.membrane stress is then calculated.
Pm = (AF/Am+'5) Ps 37>>
As an alternate method of determining the primary' I meriibrane stress, an equivalent analysis for primary membrane stress is applied to an area away from. the. trunnions. In these areas, .the metal area 'and fluid area ar'e as shown in Figurc 4. Since the depth of the metal area. is equal to the depth of the fluid area, the ratio Af/Am is equivalent to the mean radius of the- body over ~ the thickness of the body shell,'. Rm/119.. The primary membrane str'ess through this .section is then: m'=('m/'9+ ) ps
~S A. Body Primary plus secondary stress due t'o internal pressure.
III of, the
~ ~
Paragraph NB-3545.2(a) of Section code defines the formulas used in calculating this stress'. ri + .5 ps te B. Scco>>dary Stress due to.pipe reaction. Paragraph. NB-3545. 2(b) gives the formulas for finding stress due to pipe reaction. Ped. = I gS (Direct or Ariel I.oad. Effect) Gd Peb = CbF1)S (Bonding l,oad':ffcct) 2FI.,S (Torsional l.oad Effect)' bg
Ag PRESSURE AREA ANALYSIS CROSS-SECTIOif Figure 4 39
13
~B(
J ~ Jg C. Thermal Second<<ry Stress.. Paragraph NB-3545.2(c) of Section III of the code gives formulas for determining the thermal secondary strcs.,cs in thc pipo. I QT QT1 + QT2 iIth'ore
= CGC2~T2 QT2 I
J); Primary .Plus .Secondary Stresses. II This ca3culation is per Paragraph NB-3545;2 and is the sum of. the throe previous secondary stresses. n" Qp+ e. + Qt2 m III of 1 Paragraph NB-3545.3 of .Section the code defi'nos rcquircments for normal duty valve fatigue.
, Thc allowablo stress lovel is found from Figure I.-9.0.
Since the n'umber of cycles is. unl;nown, a maximum value Il of 2,000 is <<ssuiI}ed.
~
The allowable stress can t1>en be from Figure I-9.1
'ound for carbon steel. This then gives an allowable stress of 65,000 psi.
Spl 2/3 Q + P b/2 + QT3 +'.3Q Sp2 Qp + ,Pob + 2QT3' i'(herc: QT3 = C6C" hT2
0 30-
- e. DISC ANA1,YSIS Section ?AB-3546.2 defines the design requirements of the valve disc. Both primary bending and primary membrane stress are mentioned in this section. For a flat plate such as the butterfly valve disc, membrane st ress is not defined until the deflection of the disc reaches one-half the dirac thief'.ness.
Sin'ce total d'eflection of the disc is much less than one-half the thic1'ness, membrane stresses are not. applicable to the analysis. Figure 5 shows the disc for the'RS butterfly valv'es; The disc is designed to proVide a structurally sound pressure retaining component while providing the. least interference to the. flow. Primar Bendin Stress l Due to .the manner. in which the'disc is supported, the disc experiences bonding both along the shaft axis and about 'th' shaft axis. The combined bending stress is maximized't the disc center where. the maximum moment occ'urs.. The moment is a result of a uniform pressure load. Combined bending stress in disc: S(~) = (S(2) '+ S(3)2)~ LUhere: S(2) .90413 PsR4 C7 Bending stress due to moment along shaft axis, psi I4 S (3.) . 6666 Ps R4 C8 Bending stress due to moment about shaFt axis, psi
30 DISC PINS SHAFT BORE UB BLOCK 1IUB 'l(ELDS SEAT
~FLAT PLATE RETAINER SCREl('S NRS VALVI! DIS.C Figure 5 13
~*""""
t 'I Bending stress of unsupported flat plate: S(<) = MZC< I7 Shear Tear Out of Shaft I The disc is'designed so the minimum thic];ness of material surrounding the shaft extension in the disc .is above the shaft on the ar'ch side. The loading is 'clue to both seismic and pressure loads. 4 S(5) = <PsK4 2 +4'2gx Shear tear. out of shaft - s
=
(~ ~ 5D ) th ough dxsc, ps'hear Stress in Hub lYelds S(6) '.r l<~4 2P
- s. + T88 2 283(g 2~g 2)~ 2 li'a z i j.6176
SIIAFT ANAI.YSIS Thc shaft is analyzed in acco dance with Para NB-3546.3 of Section III oE the Code., The shaft. loading is a combination of seismic, pressure, and operating loads. Maximum torsional loading is either a combi.nation of seating and bearing torque or bearing and dynamic torque. Columnar stress in not con-side'red in thc shaft loading due to its'egligible effect on the stress levels. Figure 2 sl>ows the banjo assembly with the through shaft. Shaft stresses due to pressure, seismic and op~rating 'loads: S(7) =
~S(S +,(S 8) +4 S(9) 2) <
2= ~ ~ ~ i li'here: S (S) = (S (10) 2+S (11) 2) --' Combined bending stress, psi, S(10) = (~R4 2 Ps+4'2gx) ~ 25 BIR5 = Bending tensile stress m.25R lT 4 due to pressure:and seismic
~
loads along x axis, psi S (1,1) ~it
.25 R54 Bending tensile. stress due to seismic loads, along y axis, psi
~.
S(9) (S(12) 2+S(13) 2) C Combined shear'stress, psi TSRr, Torsional shear stress, psi
.SIRS4 S(13) = 1.333 5R42P.+ 5Ã2(gx2+g 2)< Direct shear mR5 stress, psi Also worthy of attention is the torsional shear stress at thc rcduccd cross-scc tion whore t'hc disc pin passe's tlirough the
- shaft, nR~4 S(r4) = S(i2>
~Rg - DZDZZ - DZDZ.
2 12 12
0 0
~ ~
DISC PIN AW:"iSIS As scen in 1'igure 2, there are two stub shafts to the disc pin. The top pins arc subjected to torsional load as they tran mit C. e operating. torque. Thc boy tom pin i.s ub-ject to shear loads duc to seismic'nd torsional loads.. Shear stre 'n top disc pin: T- 5U S (15) 2141R5. 785D3 Bearirig stres on top disc pin: T8 "5U5 S'('16) (R5 + . 5K2) 2K2D3N1 P Combined shear stress bottom d'isc pin: S (17) = S (18) + S (Z9) shear stress I'orsional in bottom disc pin: (. 5U5 + U6) S (18) D2.785D3 Shear tress in bottom pin due to seismic acceleration
+ prcssure on'nd of shaf t: .
J H>g + R52PO S(19) =- 2 (.705) D3 46-
DISC PIN'NALYSIS Wlhere: U4 .7S5(2R4) POU3R5 U5 U4 + N29 U3R5
~
U H2g + mR5 2 PQ ~ 25 (D4 +D2 2 z I PO. Actual shut-off pressure r
13 SIIAFT BEARING ANALYSIS The sleeve bearings in the trunnion (Figure 2) are subjected to both seismic and pressure loads. S(20) = ~pdR4 +N2(g~ +q ) < = Compressive stress on shait bearing, psi
2,3
'OVER'AP ANALYSIS Figure 6 shows the bottom trunnion assembly, including
'the cover cap and cover cap bolts.
- 1. Cover Cn i Bo]t Stresses The cover cap experiences loading from the weight of the banjo and from pressure loads. In determining stress leve3.s,, the bolts are assumed to share tor-sional and tensile loading equally.
Shead tear out of bolts through tapped holes in trunnion: ' (21) = N2gz++psR62 4L3 2. 83 D6 S)lear tear out of bolt. heads through cover cap, psi S(22) = N2gz+"psR62
'Tl 5.2 D6 Combined stress in bolts, psi S(2i} = S(25) +(25j + 4S( 24)
- 2. '2 Nhere:
S(24)' .25N2gz(D2+.66.(D4-D2)). Shear stress in bolts
~
.707 H3 4 A4 due to torsional -load I
S(25) .= N2gz+"PsR62 Tensile stress in bolts 4A3. due to seismic and pre~sure loads, psi
- 2. Cover Ca Stresses The. combined stre'ss in the= cover cap is calculated using the following formulas:
S(26) = S(27)+S(28) + ('(S(27)+S(28)) +4 S(29) ') ~ 2
8 8 .3
/5
/y BOTTOM 'TRUNNION
/ >>
SHAFT BEARING. t, 1, 4 BOTTOM SHAFT HRUST'OLLAR SPRING PIN
- COVER CAP I
I COVER CAP BOLTS I I SHINS
~ i f
BOTTOM TRUNNION AND THRUST BEARING ASSEMBLY
}inure 6 .
50
~ ~
'13
'Cover Ca A>>al sis TVhcrc:
S(27) = 3 (. 7S5 (D4+. 25) Ps+NZgz) = Radial stress 4%T42 (28) - 3( 785(D4'+ 25.)2Ps+TY2gz) ='angential stress 4"T4 m S (29) 7SS (D4+ 25) Ps+Tlf2gz Shear s'tress m(D4+.25)T4 51
13'IIIHIST BEARING ANALYSIS As seen in Figure. 6, thc thrust bearing assembly is located in the bottom trunnion., It provides restraint, for the banjo in the z direction, assuring that the disc 'edge remains b. correctly positioned to maintain optimum sealing. Formulas used to analyze tho assembly are given below;
- l. Bearing stress on thrust collar 'due to seismic 1 and L
pressure loads: S(30) = '2gz+mpsR52
. 785(D4" (D2+.25)"')
- 2. She'ar load on thrust collar spring pin due to seismic, pressure and torsional loads:
S(31) (4'2gz+~ sRS ) + .2514t2gz(D2+,0833+.66(D4-D2)) RS
- 3. Beqring stress of spring pin ~ on thrust collar:
S(32>> = ((1'2gz+" PsR52) 2. + ( 251<2gz)
~ )"
DS (D4-D2 )
"4. Shear tear out of spring pin through .thrust collar:
S(.33) = 102gz+" PsR T3(D4-DZ) S. Sh'car tear out of spring pin through. bottom of shaft: S(34) 1~'2gz+"PsR52 2D2(T2+ ~ -" D5) 52
'Y wc@
gV OPERA'f OR hiOUNT ING AHAJ,YS I S
~ I The'perator mounting consists of the top trunnion, the I
bonnet, the operator housing, and the bolt connections. of the assembly are sho>>'n in Figure 7. Tlie'lements II'
- l. Bolt stresses and localized stress due to bolt loads. ~ ~
~
'fhe folio>>'i,ng assumptions are used in the development of the equations:
A. 'forsional, direct shear, and direct tensile loads are sha'red equally by all bolts. in the pattern.
~
B. hfoments across..the bolt pattern are opposed in such a way that the load in each bolt is proporti'onal to its distance from the neutral bending axis. a' Shear tear out of trunnion bolt through tapped hole 1 4
.in top trunnion S< > =, l,a,+>~(r.
4 - '" ( 9m I.4D7
'.")'
rrtZ "f>,"")'.
- b. Henri>>g stress on tapped holes in trunnion.
S{36) (M,+vz) t>*~'> )'g(a,~'i
.707 il 1'71.4
- c. Hearing stress on through hole in b oil no t ~
{37) = '(Mz S 1 8)
~4( ~ 707)l )
>>7T6 53
TQI 7 RLINiVIVI'r'OUhlTIIVG I='/6 LIRE AOAPTFR PLATE I IILEI. lJELD T5 AlL APii3lliVD.
/'OAS/E7 TOP 7 RLbVN/0/v'ALVE BODY 7RDNNIQN BQL T5
~ ~
54
13 Operator tfou>>ti>>q Analysis r
- d. Shear tear out of trunnion bolt heads through bonnet.
r)))=r')))J)))M~tJ l)' 'I
~+2 (J 1+1 f2)
J2 + 2 (J2+H2) 2 J1
- 5. 2 .I)7'f'(
Combined stress in trunnion bolts (See Fig.') S(39) ='(40 +S(41 + (S 40)+S(413)-+4 (S(423+S(43) 3
~ I 2
l'I'herc i S(40) I:.z+4'4jz = uircct Tensile Stress ps 1 r 4 A5 S.(4l) = 5fx (JZ+112)
. hf~ (Jl+H2) =. Tensile' tre s s 2J2 +2(J2+H2) 2Jl +2(Jl+H2) due to evten<led'ass bending A5 ~
mome';t, psi S(423 = (>- +F )'14(g +g ) "- Direct shear stress, psi. s(43) = (w,+Ta) Shear stress due to (.,707 kl2) l A6 torsional load, psi Shear tear out of bonnet. bolt through tapped hole in adapter plate. ~ ~ 1 ~ S(44) " Fg + hip(J)h+Hg) ~ I) (Jr+hi I)
)I 2J)i +'2(JP+iii) 2JJ +2(JS+flq)
~ 9>< D9 f8 Bearing stress on'apped ho3cs in adapter plate.
)r.
)
I"'gf.8 55
Operator hIounting Analysis
~
~
~
1 h.'earing stress on through 1>olcs in 'adapter plato. s)t46)' M~+Ts +(Fz +F ~)< D9T5
- i. Shear tear out of bonnet bolt head through adapter
'I plate.
) )) = )*'*<JJ')
J ) )J
): )
- J)
J
< '.1'.
<J HJ) 5.2 D9T5 j; Combined stress in bonnet bolts (See Fig. 9) .
S (48) = S(49)+S (50) + ((S(49)+'S(50) ) 2+4 (S(5l)+S(52) ) 2) 4
.Nhcrc:
E S(49) =.F. = Direct Tensile Stress, psi 4 Bz S(50) = hl<(J4+I34) (Jg+II4) p M = Tensile stress
~JJ' '>>) 2jz +2 (J3+II4) 2 due to bending, psi
's(sx) - 2
~(sx +r.
2 ~ Direct shear 4 34 stress s(52) =. hI>+T8 Shear stress due (.707114) 4 B4 to torsion, psi 56
4 I 1
~
l
~ e I
.X
~ ~
~ ~
- J2 ~
'K4 TOP TRUNNIO'8 BOLTIii'G Figurc S 57
<< ~
.13 0 >orator lfounaina'An~el si s
- k. Shear 'tear out of operator bol't, head'hrough adapter plate.
S (53) = (hf) +M ) V4 + Fz 2(Vl +V2 +V32+V4 ) 4
,5. 2 D8T5 Bearing stress of operator bolt on hole in adapter p'late. P S ( 54) .(hfz+T8)
.5 H7 8 T5D8
- m. Combined stress in operator bolts..
S(55). = S(56) +S(57) g. ((S(56)'+S(57 ) 2+4 (S(58)+S.(59)) 2) ~ Nhere: S( 56) Fz ~ Di'rect tensile stress, psi 4 A7 S (57) (hf -+hf )V4 = Tensile stress due to 2 (Vl +V2 +V3 +V4 ) A7 bending moment, psi I"
~,
S( 58) o ~ Direct shear stress, psi 4 S(59) - (W,+T8) Shear stress due
.5 H7 8 A8 to torsion, psi
'2. Bonnet Stresses.
Thc bonnet strcsscs are calculated with the assumption that loading is thr'ough the bolt connections as previously defined. 58
13
~ ~
l hl
~ ~
~ ~
II4 BONNET BOLT PATTERN
.Figurc 9 59 4 ~
V2 V .ADAPTER PLATE BOLT PATTERN Figure 10
13
~0>crcror blonnrinn A~n.".l -is
- a. The maximum combined stress in "hc bo>>nct. was calculated using thc following formulas:
S(60) = S(613+8(623 ~ ((S(613+S(62332+4 (S(633+S(6433 "3) <, 2 2 Combined stress i>> bonnet legs ,~ S(61) = Fz+h(4gz = Direct Tensile, Stress, psi Bg S(62) .~fxBS + h~1B9 = .Tensile stress
'duc to bc>>ding Il momcn't, psi Nhcre:
S(63) (Fx2'+F>2)g + N4(g 2+g 2)q . '= Direct snear
'BS stress, psi S(64) = ("4z+Ts) (B82+B92)'~ , =.,Torsional 'shea
. stress I2+Il
- b. The maximum combined shear stress in thc bonnet mounting plate to body welds was calculated using. thc following formulas:
Bottom Bonnet Neld
.. (S(66) +4(S(693)2)
.S(6Z) =
-'ombined shear stress in bottom weld Nhcre S(66) = S(67)+S(68) ~ Total tensile stress, psi S(g 7) -
l:q+lll4gq = Direct tensile .stress, Ug psi
13 0 erato> llonnf,in, An~al ala S(68) = hf+hf =.Bending tensile stress Zl S(69) = S(70)+S(71) = Total shear stress S(70) = (F 2+F 2)4+~i~4(g,.2+g 2)~ ='Direct shear stress, psi Ul., 1 (7.3.) = t! z+T 8 = Torsional shear stress, psi Z3 Top Bonnet 1<eld S(72) =
. (S(>>) + 4(S(74)) ) = Combined 'shear stress
- in top borinet weld Nhere S (73) = S P
(75)+S (76) = Total tensile stress, psi Direct tensile stress, psi 5). S(75) = U2 F. 1 S(76) . = h~fx+hi =. Bending tensile stress, psi Z2 S(74) = S ~ (77)+S (78) = Total shear stress, psi S (77) ~P'I')."'-- = Dire'ct shear stress, psi S(78) = iha+T8 Torsio~>al shear stress; psi Z4
- c. Trunnion Body. Stress The trunnion body stresses are calculated. using the iollo>>ing assumptions:
- 1. Operator loading is through the bolt conhections.
13 Q)orator'htountinv, Anal 'sis ~ ~
~ I
- 2. Thoro is an equal and opposite reaction to the bolt loads at the 'body.
The combined stress in the, trunnion body was calculated using the folloiling formulas: S(79) = S(80 +S(81) ((S(80)+S(81) ) +4 (S(82 )+S(83 )') ) ~ 2 ~ 2. Nhere S(80) = F +N4g = Direct Tensi'le Stress, psi K4KS-,785822 P S(8j) = '(hfdf.+F Kl)) . SK4 (hf +FXK6).SKS . = Bending tensile.
'.0833KSK4 -B2 .0833K' 3-~B 4 4 5 2
$ (82) = (5 2+P 2)4+/l/4 (g 2/g 2)g Direct'hear OII stress, psi K4K5-.285B22 S (83) = (hlz+T8) 5 (K42+K52) Torsional shear
. 0833 (K4KS +KSK4 ) "'B2 stress, psi 32
13 FRE UENCY.ANALYSES Introduction To calculate the natural .frequency 'of the various com-ponents of the Triton NXI. valve, a mcrdcl system with a single degree of freedom is constructed. The individual P components and groups of .components are modeled and analyzed restoring spring forces which act to oppose the re-1's spective >>eight forces they are subjected tb. The static deflection of the component is calculated and is'e'lated to natural frequency as.: n or Fn=18 2m or Fn -" .9.8 Qy The analysis details the equations and Issamptions used in dote'rmining the natural frcquenc'ies listed in the summary table. S1'etchcs are provided where appropriate. B. Valve Bod 'ssembl The body shell, as seen in Figurc 1, is assumed to cxpcriencc loading duc to thc entire val've weight. Natural Frequency of thc body shell: Nl 64
.Frc uencv Anal.s is l'/here bye = Ã1L13 = Maximum deflection of body shell 48E Ig due to valve weight, inches C. '~Ab Figure 2 shows the banjo assembly 'in the body. 'The natural frequency of the banjo assembly is calculated using the following: 1 FNg = 9 8
~y2 li'here
~y2 lf7B1 .
= Maximum .deflection of'shaft, inches 12E IG D. Cover Ca Assembl in Figure 6, tl>e cover cap supports the banjo..
~
As seen ~
~
The'natural frequency. of the pover cap is calculated as follows: FN P h lt'here
.hy3 = 3(m2-1) N2 (.5D4+ .125)2 = Maximum deflection of cove'r cap 16trE T43m2 E ~ 8 0 n 11 c t. A s s c mb ]
'Fi'gure 7 shows the top trunnion assembly. The following assumptions are made in calculating the bonnet natural irequcncy:
65
- 1. The worst valve 'assembly mounting'osition is where the bending moment is predominant in'roducing deflection.
- 2. The bonnet is assumed fixed at the top trunnion.
~ 3. The adapter plate is assumed to bc integral with and have a cross'-section the same as the comp'onent, it mounts to, Natural frequency. of bonnet: Fg4 = 9.8 Nhere hy4 = lU3H8 +!U4K3 + lU3ZoH8 3EI1 . 2EI1
0 ATTACHMENT 3 GENERAL ARRANGEMENT AND CROSS-SECTION DRAWINGS
i
LIATS I . ~ o H Zl L-'--: I
~ 's nor-PI A~<<lr II I Inn C.t hr= on IAlf.l IA P Y VALV'c. U re I. I ." J ej r'-
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.SIZE G
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" I+-X Ifl C7 ITI p XJ I
2'1 kY'- I
.-+EMOTE.: ALL DIMENSIONS ARE SHOWN IN INCHES. bK- PZ -I ..A
~ =X C}+-, I/lfo 'THRLI IO" VALVES, )
PEKOE
..... ADJUSTABLE SCREW STCIP % D+ I/& 'OR I L" VALVE& AND LARGER, ~
~~
0
2 NPT Qy+QAIL QAAPt+ARA~iurhrS HA LE rrrP7Arv)err~~~
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.'TYP EAC,H gND THIS E.NO I Ce Z
OPERATOR u TO QAVE, BREAtV;ER PLUG,. SIZE p +.- 'P~ 0'l z- C
'g 17 IAIAX.
0+
~O
'X P
T312-5R2 59ri p p A l
'7 IE P 2 T312-5R3 /~
.t i, .rrt~~', c.~~-.t hark T312-5R'I 14 ze'/z 0 k-I
---- =.: -.'v~~ g '~i ~l Pi m i Z PI L ACCESSOR IES o K
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f H- - a DI
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I- CIRCLE.DEAL SrrroLE soLEArorD VALVE >b
~
IAI S g 0'0 N IfEI g3 MANUAL!n
~
gfTI. fr'ID& 'IEIOD'EL 5.V 3 I 5 9 IOI-I 120 V.A.C. NOH~ IPIIAPJED+ t
-'.- "'- --.: '=,'..-'-..-',.'-,
- 2) 2. TIIAHCO EA=I lO-2OOCIO LIMNI'T SIRE lTCHES tv- r c.. .
FABRICATED REPLACEABLE = -: I2OY.A..C. IO AMP. D.P. C0. T PACKING BONN'ET- jr -..;;-' +- Q) I- tlOFFMAhl JONCT IOhl BOX Ttf-'A-I'Or'IIAI=.. -PO p
'V'2" M.P.T. FOR LANTERN GLAND -,
BLEED-GFF (CAFrPEQ W/ Pl PE. '.-- PLUG). I.E.7A6 l8.-II88-8F-A0-$ 724. -.'- TAG= IB =HBB BF-AO-570.4
'8- M88-8F-A o-5'72$ f IS=II 88-Bi'-Ao S703 r- qp
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;.==POSITIOW ~ -'POS IT ION, B I FLANGE I-7'n'. x V4ELD - END Q0 0 I lO B.tS hi 7 7-7to CEIAA" a I Sf B Qo .' .
O h I1-2 ; <-3-7:o Rt . Lrt LI~ ir R! i t;Att ter LVP
~r' V';I I f L,H T trtttlt t CCrAIPALLT
~F=NO. A.ND SIZE OF I-IOLES x EMERALD A'RRANGEMENT E DEEP. STRADDLE',ENTERLIhk NUCLEAR .N.R.,S. VALVE. Vl/EON.
NOIYIINALVALVE SIZE DD ~R.F. DIA. ISO LB. 0 SAS FLANGE BOLT LAYOUT. .
." " '..POSITION ',2 '" ' . t",: 'O'SlTIO $ BETTIS SPA,INC -RETURN OPER, I
V5' P.T. BLEED-OFF CONNECTION 6< Lt I NOI'lK jrLtt R .I 6= BOLT.CIRCLE '(CAPPED W/PIPE PLUCr) ~ t.IIAI'rid ltt~~gg~o, .<f C ALL! Br "C. I ~O I C D-002t'o-Z 7
PARTS AND MATERIALS OF CONSTRUCTION ; PARTS AND MATERIALS~CONSTRUCTIOh '
,I I h I. BODY: MAT'L.l SA-SI6 GR. SS I. THRUST, COLLAR PIN: MAT'LD'D AISI 420 STN. STL J
- 2. SEATING SURFACE: MAT'L.D SA-479 TYPE 304 I2. GREASE: DOW CORNING I I I
)
DISC: 13. 0-RING: MAT'L.D E. P. T. 3A. HUB: MAT'L.g
/
SA-SI6 GR. SS '4..BOTTOM I COVER: MAT.'L.D I SA-5llc2 GR,.70 l J
- 38. FLAT PLATE: HAT'Lu GR.70.
r'A-SI6 IS. COVER BOLTS: MAT'L.> SA-I93 GR. 8-7
/
- 4. SEAT: MAT'L.; E.P. T. lb. LOCKWASHER: NAT'L CARBON STEEL I
CLAMP SEGMENT RING: MATJL.> SA-28S GR. C, l7. BOTTOM BEARING: MAT'L.> ASTN B-438 GR.I TYPE 2 BRONZE
- 6. CLAMP SEGMENT SCREWS: MAT'L.l SA-193 GR. B-7 IB. TOP BEARING: MAT'L.e ASTM 8&38 GR.I TYPE 2 BRONZE SHAFT: MAT'LD .SA-S64 TYPE 630 COND. HI ISO 19. SHAFT SEAL:, MAT'L.e E. P.T.
7,. I SB-l44 ALLOY 38 Ql S. PINS: MAT'L.'A-320GR. BSM 20. PACKING RETAINER RING: MATJL.'D
- 9. THRUST COLLAR SHIMSt MATtL.i HARD BRASS 2I. PACKING: MAT'L.D E. P. T. V-RINGS J
IO. THRUST COLLAR: MAT'L.l SAE 660 BRONZ+ 22. LANTERN GLAND RING: MAT'Lp ASIA "A-269 I
, I 7 STUB SHAFT COATED WITH SILICONE LUBRICANT FOR ASSEMBLY PURPOSES ONLY.
NOTE: I-PIN TOP & I%IN BOTTOM FOR l4" VALVES. I
'. '-PINS TOP & I-PIN BOTTOM FOR l6" THRU 24" VALVES.
l3 ) I IP 3A v I I 22 I4 I P.. I/2"'NPT FCR LANT D ~' GLANO BLESO-OFF (CAPPEO M/PIPE PLUG) IS
'N IIXLVK BOOV ~
lb 12 P 'l. I I 20 - 2I 9 6 IB t ', t",", J 2 4 CONNECTION I STUB SHAFT COATED WITH cAPPEO IJ/PIPE BODY SEATING SURFACE SEAL WEL'D WILI: ,,'i,'ILTCOh!E LUBRICANT FOR ASS'Y. BE "PT" EXAMINED.',, ',
'URPOSES ONLY, i, DISC HUB WELDS WILL BE TJPTJi OR) I "NT", EXAMINED..
I I
~ I I
I AND NDE STANDA SHALL BE IN ACCORDANCE IAjITH ASME SECTION LA5 QUIR n
'MATERIAI r~ I
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CUSTONER P. O,: .PBB56 n
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