ML20092B788
| ML20092B788 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 04/24/1984 |
| From: | Williams N CYGNA ENERGY SERVICES |
| To: | |
| References | |
| I-001, I-1, NUDOCS 8406200366 | |
| Download: ML20092B788 (169) | |
Text
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~ 00CMETED 00(.KETED USNRC USNPC UNf8tDSTAESgy:NAERICA
'84 MY -2 A9:34 NUCLEAR REGULATORY COMMISSION OFRCE CF SECFiir f
BEFORE THE AEdkIO,5/d'dTS AND LICENSING BOARD e m cq 00Cr.Eijtjb}.yEPV-in the Motter of )
TEXAS UTILITIES ELECTRIC ) Docket Nos. S0-44S COMPANY, elo_1. ) ,.
S0-446
)
(Cornonche Peak Steam ) (Application for Electric Station, ) Operating Licenses)
Units I and 2) ) ,
)
TESTIMONY OF NANCY H. WILLIAMS IN RESPONSE TO CASE QUESTIONS OF FEB. 22,1984 TO CYGNA ENERGY SERVfCES
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h.2G7 DAVID R. PIGOTT of ORRICK, HEP,RINGTON & SUTCLIFFE A Professional Corporation 600 Montgomery Street Son Francisco, CA 9411I Telephone: (415) 392-1122
~1 NUCLEAR REGULATORY COMV:sst0N M 'd'[
Docket No. of_ d'm O -M474 Official Esh, No. f in the matter sL ru f.,t/44/
Staff IDENil4ED J [![/
Apowat PEcEwED v/tv Interwnc' ~ REJECTED April 12,1984 cone, ern contrator p
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. . . , wt 9406200366 840424 -
- l. Question: Please state your nome, current business position.
Answer: I am Noney H. williams, Project manager, Cygno Energy Services,101 California Street, Suite 1000, Son Francisco, California
- 2. Question: What is the purpose of the testimony being presented at this time?
Answer: During hearings of February 20 through February 24,1984 in this proceeding, Board Exhibit No. I " Independent Assessment Report," Volumes I and 2 were introduced into evidence. During those some hearings I testified in support of the report and was cross-examined by parties to this proceeding. At the conclusion of that set of hearings, it was agreed that intervenor CASE would provide Cygna with its cross-examination questions in writing. Attached hereto os " Attachment I" is o copy of the written questions submitted to Cygna by CASE.
Subsequent to receipt of " Attachment I," Cygno formulated its responses and informally circulated those responses to the Board and the parties in a document entitled " Testimony of Noney H. Williams in Response to CASE Questions of C
February 22,1984 to Cyano Energy Services" and dated March 18,1984. As a result of conferences between Cygno and CASE on March 21,1984, March 27,1984 and April 3,1984, correspondence from CASE, and guidance provided in the Board's
" Memorandum (Clarification of Open items)" dated March 15,1984 Cygno has reformulated its responses to the questions contained in Attachment 1.
Attached hereto and inccrporated herein ore copies of Cygno's responses to the CASE questions mentioned in Attachment No.1.
TESTIMONY OF NANCY H. WILLIAMS osa e
9 UNITED STATES OF AMERICA 4
- NUCLEAR REGULATORY COMMISSION ;
l
- l
' BEFORE THE ATOMIC SAFETY AND LICENSING BOARD In the Motter of' )
l ) Docket Nos. 50-445 TEXAS UTILITIES ELECTRIC 50-446 COMPANY, e_t t aj. )
)
) (Application for
- (Cc T. cide Peak Steam Operating Licenses)
Electric Station, )
Units I and 2) )
)
i i
CERTIFICATE OF SERVICE i
I hereby certify that copies of the foregoing Testimony of "Noney H. Williams in Response to CASE Questions'of Febroory 22,1984 to Cygno Energy Services" in the obove-captioned matter were served upon the following persons by overnight delivery (*),
l
==. or deposit in the United States mail, first class, postoge prepaid, this 12th doy of April, D7 I984.
- Peter B. Bloch, Esq. Chairman, Atomic and Sofety Licensing Chairman, Atomic Sciety and Licensing AppealPr.el
,, U.S. Nuclear Regulatory Commission Doord U.S. Nuclear Regulatory Commission Washington, D.C. 20555 Washington, D.C. 20555 Mr. Williom L. Clements
- Dr. Walter H. Jordon Docketing & Service Bronch U.S. Nxteor Reguiotory Commission 801 W Osic: Drive Washington, D.C. 20555 Ook Ridge, Tennessee 34830
- Nicholas S. Reynolds
- Dr. Kenneth A. McCollom Dean, Division of Engineering William A. Horin Architecture and Technology Bishop, Liberman, Cook, Purcell &
Oklahoma State University Reynolds Stillwater, Oklahoma - 74074 1200 Seventeenth Street, N.W.
Washington, D.C. 20036 Mr. John Collins 'Stuart A. Treby, Esq.
Regional Administrator, Region IV U.S. Nuclear Regulatory Commission Richard G. Bochmann, Esq.
611 Ryan Plaza Drive, Suite 1000 Office of the Executive Legal Director i
Arlington, Texas 76011 ' U.S. Nuclear Regulatory Commission Washington, D.C. 20555
- ;+==
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- Chairman, Atomic Safety and Licensing I
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Board Panel U.S. Nuclear Regulatory Commission Washington, D.C. 20555
- Renea Hicks, Esq.
Assist'ont Attorney General Environmental Prctection Division P.O. Box 12548, Capitol Station Austin, Texas 787II Mr. Lanny A. Sinkin iI4 W. 7th, Suite 220 '
Austin, Texas 78701
- Mrs. Juanita Ellis President, CASE 1426 S. Polk Street Dallas, Texas 75224 -
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AT*"ACHMENT 1 From CASE Witness Jack Doyle to CYGNA Sumary of Cross-Examiriation Questions 2/22/84 V h . BRIEF
SUMMARY
OF GENERIC PROBLEMS Omitted from Calculations and Omitted from Checklists
- 1. Cinched up U-bolts:
o Not in compliance with Cygna criteria o Not in compliance with NRC criteria o Stresses of unknown quantity due to pre-stress, themal and design loads , _ , , _ _ _
o Effects on pipe not shown on calculations ' , ;_ _ __.___ _ . -- - r- 8 Y9.y ::. . - - .
o Not in compliance with Board Notification. 3 / v / /*.'
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- 2. Local effects on tube walls: '- UU d .3 , % fa /A%i o Punching shear FILE: y p,w t Dg, Cross REr. FILE o Effect on welds l o Resultant effect due to wall flexibility on moment at tube weld.
51 5 3. Dead weight of structure not included in calculations.
- 4. Weight of support masses as they affect pipe stress.
- 5. Inaccurate conclusions as relate to KL/R for pinned columns:
o If a column fixed at its base and free at the top has an effective K of 2.0 cutting at some point up from the base and adding a pin does not address the problem.
6.16-inch pipe with about 20 kip load along 31/2-inch length induces high bearing stresses which require pads. This is not addressed, o ASME Code against flattening. .
- 7. Clip angle 4x4x1/2 which supports U-bolt not addressed (critical to maintainingstability):
o Section modulus .04 in cube o Moment am at least 2 inches
" o .1100# load exceeds Code allowables, o Pre-tensioning to obtain a clamping force required could exceed
Jack Doyle to Cygna 2/22/84 -
tgty Page 2 this (not including thennal constraint and design loads) o Clamping force with no margin of safety for single degree system (not point contact or line contact) is force / coefficient of friction or about 4 times what is required for clamping force.
- 8. There is no documentation in calculations to support the conclusion that. flair weld is stronger than fillet weld--no calculations, there-fore why did Cygna accept this statement?
o Flair weld strength depends on radius of flair (depth).
- 9. The reduction rf weld capacity in the calcuation is based on 1350 Actual tangental angle is 150.30 Therefore, an error exists. Did Cygna take note of this?
{gf o More stress in weld than stated, o Wide / thin ratio induces cracking as well as the 1:4:1 ratio width to
. depth.
- 10. Changing from fit.ir weld tc fillet weld induces flange bending. Has this -
been addressed by Cygna?
- 11. Effects of cut-of-plane seismic excitation of support handware not included in calculation. Did Cygna address this point?
o Additional loads on support o Additional loads on pipe
- 12. Restraint of rotation by the pipe because of coupling effect of hardware on both sides of a pipe:
o load increase in 1 of 2 snubbers / struts o Alteration of dynamics of pipe system during seismic event kh 13. In Note 2 following page PS-01-4 of 4, Cygna decided to eliminate their stiffness criteria based on their knowledge that a report existed to ad-dress the problem (but without personal knowledge of what was contained
- i Jack Doyle to Cygna 2/22/84
=- Page 3
'GF in the document in detail). Why didn't Cygna consult with their experts--
for example Eric von Strijgeren (who was the editor on a paper by T.Y.
Cb, C.H. Chen and O. Bilgen)--in reference to deviations from generic stiffnesses in pipe supports and the effects on piping systems.
o Third paragraph introduction et. seq. (CM56 h,9ff9]
- 14. In Note 1, same source, did Cygna consider the additive . effects of self weight' excitation if the stiffness is considered from node point to hard point as opposed to the stiffness of the frame independent of hardware, local effects, base plate and anchor bolts?
o Spring rate of base plate / anchor bolts (particularly bearing-type joints) can be considerable (observation of base plate II finite analysis).
Em b- 15. Was themal lockup considered for anchors which restrain pipe radial growth?
o Induces frame moments
- 16. The base plate analysis is based on distribution of shear relative to Did Cygna address load path / stiffness for all bolts in the pattern.
this problem? ,
o With oversized holes and the inability to eliminate construction tolerances (. location of the bolts combined with localtion of the bolt holes), it is not possible for all of the bolts in the system to be active. (See CASE Exhibit 906).
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l o The stiffness of the bolts is such that deflection cannot be counted
' ca. as a means to achieve full pattern participation l
- o Even if deflection could result in full activity, the first bolts
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" deflecting would receive the larger portion of the load in an ideal symetrical and systems.
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Jack Doyle to Cygna
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m 2/22/84 ay Page 4 o For non-synnetrical system and systems of variable stiffness, the inactivity of a number of the bolts will alter the accuracy of the computerized analysis.
- 17. Has Cygna- verified the statement: "No 2-inch topping"?
o This affects the calculations for Hiltis relative to embedoent, since a non-monolithic shear plane has been established.
- 18. The base plate analysis performed without including stiffeners alters the stiffness matrix of the base plate and consequently the distribution of moments and tension to the bolts. Beyond this point, stiffeners remain unqualified. Has Cygna addressed this?
g.3 The preceeding questions are the primary areas in which I will be cross-examining Cygna witnesses. (Additional questions may be triggered by Cygna witnesses' answers.)
In addition, CASE has not yet received all of the documents which it requested from Applicants' on the Cygna report. Therefore, additional questions may be triggered from these documents (if and when they are supplied).
L__
From CASE witness Jack Doyle Sunnary of Cross-Examination Questions _ ~ 2/22/84
$i5.
MATRIX OF EXHIBITS AND DOCUMENTS CASE Exhibit Concerns 891 1, 3, 4, 5, 6, 7, 11, 13, 14, 15, 16 892 9, minor question relative to pad width diameter + (Rt)b 893 8, 10, 14 894 1, 4, 5, 11, 14 895 14, 16 896 12, 14 897 1, 2, 3, 4, 5, 11, 14, 16 898 14, 15, 16, 18 899 14, 15, 16, 18
- gh-900 14, 15, 16 901 Has minimum weld violation (walk-down) 902 Has support completely rebuilt on CMC and then calculated This retrix has been compiled to the best of our ability due to time constraints. (It is from notes, etc.)
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BRIEF SUPMARY OF CROSS-EXAMINATION OUESTIONS job No . TVGt/L BY CASE WITNESS MARK WALSH TO CYGNA, log No. . #* 2 O Fli.E ; /f3/64 A4 /Aq Appendix E of Cyona Report cross REr. RILE /L 3 A I d 4,/png j, ,
Section DC-2.4.4. What was the yield point used for A500 Gra'de B tube steel? .',. ~
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- J., Observation Record PS-02-01
- The Applicants did not consider shear cone interaction of adjacent bolts.
3 PI-01-01. There has been no detailed computer-analysis performed to consider the concentrated loads.(valves, etc.) and their effect on dead weight and seis-ic.
[" Also, the seismic analysis wi'll not be linerally proportional.
i / P!-02. Is there an er r in the table shown?
1 1 f CTS-00-03: See CASE Exhibit 889, sheet 129. Fbx = should be 21.2, not 23.2 or 22. The length is 6' not 5.5'.' .
See CASE Exhibit 890: 1) Why was only 1/2 SEE considered?
- 2) Why was 4% damping used; not consistent with FSAR? 3) Assumed cable tray l was rigid when lumping the mass; this resulted in not combining the dynamic
. effects of,the cable tray itself to the support; did not include effect on welds.
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- 4) The validity that the nble , trays have the capacity to transfer a load JJ
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around a corner when one run of cable tray has no axial restraint, as se.n '
5 E on drawing 2323 El-0601-01. (NOTE: We only have a '36"x48 drawing; please E
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g m let us know when you want to look at it.). 5) What documentation did Cy)e -
G see that justified the hangers' receiving a latical, load around corners h,t 3 <l M
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resist the axial , load from the tray segment that'contains no axial restragg.s; O
WALSH TO CYGNA (2) san. ,
er how did Cygna evaluate it? It appears the axial loa . has not been taken into account. 6) CASE Exhibit 902. Did not consider base plate flexibility.
(, , CTS-00-05: In the description, it discusses a channel bent about its weak axis. The resolution does not consider this problem nor does the document CASE requested on discovery; see CASE Exhibit 907. On CMC 88306, are the originator and coprover the same person?
7, CTS-00-06. What is the "significant design margin" as shown in the resolution?
The analysis that included the beam element did not consider j3, CTS-00-07:
g==i prying action and the flexibility of the base plate to determine the center
=
of compression.
O. WD-03-01: What documentation was there that " accept as is" was valid? Were there calculations to support this? ,
i 10 WD-07-02: What documentation did Cygna see that showed the temperature indicator would be installed at a later date?
I:, Pipe stress checklist, note 3 item a: 1) What is the basis for considering that the effects were negligible? 2) What pipe stress run did Cygna look at, since the inclined load was used in the design of support RH-1-010-003-522R?
ch ar e
WALSH TO CYGNA (3)
Cable Tray Check List: CTS-ll , Item .6,' probl en 4. This was not discussed
- 11. .
in CTS-00-07.
The preceeding ' questions are the primary areas in which I will be cross-examining Cygna witnesses. (Additional questions may be triggered by Cygna witnesses' answers.)
In addition, CASE has not yet received all of the documents which it requested from Applicants' on the Cygna report. Therefore, additional que.;tions may be triggered from these documents (if and when they are supplied).
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C ComanchePeak ASLB Hearings Response to CASE Questions Question No.: Doyle #1 Exhibit No.: 891, 894, 897 1.0 CASE Question Cinched up U-bolts:
o Not in compliance with Cygno criterio o Not in compliance with NRC criteria o Stresses of unknown quantity due to pre-stress, thermal and design loads o Effects on pipe not shown on calculations o Not in compliance with Board Notification 2.0 Cygno Interpretation N/A 3.0 Response
- Section 4.1.2 of the Cygno review criterio document, DC-2, states the following:
6 "A cap shall be provided to occommodate radial expansion and construction tolerances.
The maximum total gap allowed in the restrained direction is I/8". In unrestrained directions, the support design shall allow clearances for the most severe thermal plus seismic movements of the pipe. Proper installation tolerances shall be provided where thermal movement cannot be accommodated within the specified gap minus 1/16"."
This criteria is intended to apply only to pipe supports which do not require physical contact with the pipe to insure that the require restraining forces are developed.
Supports which require physical contact with the pipe in order to develop the proper restraining forces, such as pipe clamps and cinched U-bolts, cannot have gaps and therefore are not required to satisfy the conditions of DC-2, Section 4.1.2.
The NRC Information Notice No. 83-80 identifies potential significant problems that may exist with the usage of specialized " stiff" clamps. Under certain conditions, these clamps may induce high local stresses in the pipe. Cygno did not encounter any " stiff" ,
l clamps during the Cygno IAP review.
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fT Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Doyle #1 Page 2 As defined in the Independent Assessment Program Plon, review of the RHR System included design criterio, analyses, design and drawings. It did not include installation specifications, . where torqueing requirements such as cinching, would normally be defined. Cinching was not required or defined in any of the documents reviewed by Cygno. Accordingly, cinching loads were not known and were not considered in the design assessment.
Loads on the pipe due to cinching were not assessed for the reasons discussed above.
Pipe loads due to the zero gap were judged to be negligible. The conclusions in the IAP Draft Report are based on that engineering judgment.
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Comanche Peak ASLB Hearings
( Response to CASE Questions Question No.: Doyle #2 Exhibit No.: 897 1.0 CASE Question Local effects on tube walls:
, o Punching shear o Effect on welds o Resultant effect due to wall flexibility on moment at tube weld 2.0 Cygno Interpretation When tube sections are employed in the design of pipe supports, how were the following local ef fects considered:
- a. Punching shear?
- b. Effect on welds?
( c. Resultant effect due to wall flexibility on moment at tube weld?
3.0 Response Pipe support RH-1-062-002-S22R (CASE Exhibit 897) is designed using a tube section, TS 4" x 6" x l/2", welded to a baseplate at one end and to a strut clevis at the other end.
Punching shear and welding stresses are discussed below:
- a. Punching shear stresses are within allowable for o!! supports reviewed by Cygno.
This is evidenced by the punching shear check provided in Attachment D2-1.
Adequacy with respect to punching shear con also be determined by inspection through a simple comparison of fillet weld size and tube wall thickness. The basic relationship for this comparison is established by considering a unit length of weld and tube wall as a freebody and equating the allowable force in the weld to the allowable shear force throdgh the thickness of the tube wall.
The allowable force in the weld is P,=Fw
- 1 * .707
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Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Doyle //2 Page 2 i
The allowable shear force in the plate is Pc = Fe
- 1
- t e where t e = tube wall thickness, inches F = allowable weld shear stress, ksi (use 18 ksi) e t, = fillet weld leg size, inches F, = allowoble tube shear stress, ksi (use 0.4
- 31 ks0 and substituting the proper values for allowable stress, the By equating P, to Pc following relationship is established:
t c = (18 * .707
- t,)/(0.4
- 31) te = 1.0 t, Therefore, ossuming the fillet weld is properly sized, i the iftube the tube wa (t equal to or greater than the fillet weldk size, (1/2") is punchin twice the attached fillet weld (1/4").
b.
Eoch welded connection in support RH-1-064-011-522R is discussed be and
- c. Tube-to-Baseplate _
This connection is a standard beam-to-column detail, as evidenc Manual, Port 4. Furthermore, the flore-bevel weld detail has be evoluoted and sized by the designer.
Tube-to-Clevis Attaching the strut clevis to the tube flange introduces no adverse the connecting fillet weld.
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j Comanche Peak ASLB Hearings
( Response to CASE Questions Guestion No.: Doyle #2 Page 3 4
The flexibility of the tube wall produces no significant additional foods on the weld. This welded connection compares favorably with certain stonderd weldments shown in Blodgett's Design of Welded Structures (see Attachments D2-1; D2-2 Figure 9; and D2-3, Figure 12). The connections shown in these attachrnents are more " flexible" than the tube-to-clevis detail in support RH-1-064-522R, and are not evoluoted for added weld stresses due to diaphragm action or plate flexibility.
- c. AWS Section 10.5 specifically addresses stepped tube connections cad the evoluotion of tube wcil capacity for the cose where the connecting tube transmits both oxiot and bending loads to the tube wall. The design equation (Section 10.5.1) used in the evoluotion is a function of both the ratio of the tube widths (Beto) and the tube wall thickness. It seems implicit that by satisfying the design equation the local stresses within the tube wall are within acceptoble limits at the des *gn load.
in addition it should be emphasized that the Beto parameter alone is not sufficient to evoluote the serviceability or strength of stepped tube connections. The Beto For
- parameter must ,be considered in conjunction with the tube wall thickness.
example, a connection having a Beto = 0.4 will possess approximately the some ultimate moment capacity and punching shear capacity (as well as the some moment-rotation and oxial lood-deflection chorocteristics) as a connection having a Beto = 0.8, if the connection with Beto = 0.4 has a wall thickness one-third greater than the wall thickness of the connection with Beto = 0.8 (see Korol & Mirzo paper, ASCE, Journal of the Structural Division, September 1982, Figures 7,8, t I and 12 and Tables 2 and 3). Thus, o tube (or clevis) welded to a 3/8"-thick tube for which Beto = 0.8 will behave opproximately the some with respect to deflection, rotation,
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punching shear and ultimate moment as when the tube (or clevis) is welded to o 1/2"-thick tube wall for which Beto = 0.4.
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1 ATTACHMENT D2-1
-( (Page 1 of 3)
Punching shear check for Support No. RH-1-062-002-522R.
I
Reference:
American Welding Society (AWS), D1.1, Section 10.5.
T
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h n k h Attachment 3 :: 3 6
V k E l *-t i,________ __
1 s - 1/4
'l1/2 >
FIGURE D2-1 APPLIED AXIAL LOAD = 5092 LBS.
Since the attachment is not a tube and only welded on the 3" side, the calculation of f, in the following equation for Acting Vp ( AWS Section 10.5.1) will be conservatively high, because the loads shared by the 1-1/2" sides of the tube are being neglected.
f sin e f Acting V p =t (a K ,
a b 2:
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7 ATTACHMENT D2-1 (continued) ~
- (Page 2 of 3) where f, = 5092/(3+3) bt = 849/t b f
b
=0 0 = 90 degrees i = t b/tc t
B = b/D K
a
= 1.0 Acting V p = 1698 Basic V p = Fy /(0.6y) where y = D/2t
, c
= 31350/(0.6)(6) = 6/2(1/2) = 6 U = (f a + I b)/0.6 Fy (see Note 1, Table 10.5.1) f a = 849/tb = 849/(1/4) = 3395 psi .
(Note: f ais conservatively calculated using tb of 1/4", i .e. , the weld size).
U = (3395 + 0)/(.6)(31350) = 0.18 Since U is less than 0.44, Qf = 1.0; and, since beta (0.5) is less than 0.6, Qb = 1.0.
Allowable Vp =QObgf (Basic Yp )
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= (1)(1)(8708 psi) = 8708 psi O
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ATTACIDENT D2-1 (continued)
.( (Page 3 of 3)
This is considerably greater than the Acting Ap= 1698 psi .
Design margin = (8708/1698) - 1 = 4.12 = 412%
4 OK.
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ATTACHMENT D2-3 Web Froming Angles / 5.49 (Page 1 of 1 )
of 38.4 kips for a weld size of w = " and angle p
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length of 1, = 10" slightly exceeds the reaction. He corresponding (Field) Weld B, using u = %", also is b
- sausfactory. Since the beam's required web thickness is 0.al" while the actual web thickness is 0.25", the indi. ly a cated 3" x 3" x ~ is all right. l ;
,' l 3
--+
If the beam is made of A36 steel, this connection's l
capacity will be reduced in the ratio of 0.25/0.29 of l l l
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actual to required web thickness. The resulting capacity v a of 33.1 kips is less than the reaction. The next larger connection with apparently sufBeient capacity shows that (Shop) Weld A's capacity is 47 kips, using same FIG.10-Double. web froming ongle.
angle section but an angle length of L, = 12". Apply.
ing the multiplier of 0.25/0.29 reduces the capacity of the connection to 40.5 kips, which exceeds the end rt 8' e _
reaetion.
'e l 81
- 5. SINGl.E. PLATE OR TEE CONNECTION lt ON BEAM WEB l!
In the previous design of the Geld weld, connecting a pair of web framing angles to the supporting column , l or girder,it was assumed that the reaction (R) applied y" eccentric to each angle, resulted in a tendency for the
( angles to twist or rotate, in doing su, they would press i . H-Single plate se Tee.
\, together at the top and swing away from each other at the bottom, this being resisted by the welds. These forces are in addition to the vertical forces caused by column would be designed then for just the vertical the reaction (R); see Figure 10. reaction (R); see Figure 11.
However, in both the single plate web connection In the shop weld of the singic plate to the web and the Tee.section type, this portion of the connection of the beam, Figure 12. this double vertical weld would welded to the column is solid. Thus, there is no be designed for just the vertical reaction (R). There tendency for this spreading action which must be re. is not enough eccentricity to comisider any bending sisted by the welds. These vertical Seld welds to the action.
Field weld to supporting column or web of supporting girder ,p
' Shop weld _
f r.. ........ !
% 5eom we "[
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, Flat plaie used for flesible ?
connection on web of boom f; .
FIG.12-Flot plate wed for floalbfe connection on web of boom.
ATTACHMENT D2-2 5.2-4 / Wcided.Cxrectlen Deelga (Page 1 Ot 1)
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Don't hook wefd around corner; Hook weld oround will not have full throat corner of sect ongle Seat Angle Width Sect Angle Width Greater than Column Flonge Las than Column Flonge FIGURE 8
- 5. HoltlZONTAL STABILITY greater than usually required under normal load con.
ditions.
C~ A Hexible top angle is usually used to give suScient Notice in the following Sgure, that the greatest horizontal stability to the beam. It is not assumed to movement or rotation occurs in the Ellet weld connect.
carry any of the beam reaction. The most common is a ing the upper leg of the angle to the column. It is 4" x 4" x %" angle, which will not restrain the beam important that this weld be made full size.
end from rotating under load. After the beam is erected, This test also indicated that a return of the Ellet j this top angle is Seld welded only along its two toes. weld around the ends of the angle at the column equal For beam Banges 4" and less in width, the top angle to about % of the leg length resulted in the greatest is usually cut 4" longi for beam Banges over 4"in width, strength and movement before failure.
the angle is usually cut 6"long.
In straight tension tests of top connecting angles at Lehigh University, the 4" x 4" x %" angle pulled out
' as much as 1.98" before failure, which is about 20 times Hook around L )g
/
% Column flonge G rototion occurs
% 4 rootest .t l near vpper weld y 5, '. Top Horizontal (connecting ongle movement t','....._.. FIGURE 10 m..._m- s
- - - I
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- ***1 Pull on e Problem I l
e boom {. connection Design a Hexible seat angle to support a 12" WF 27#
[\ beam, having an end reaction of R = 30 kips. Use AS6 FIGURE 9 steel. E70 welds.
( Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Doyle #3 Exhibit No.: 891, 897 1.0 CASE Ouestion Dead weight of structure not included in calculations.
i 2.0 Cygno Interpretation N/A 3.0 Response General purpose structural design codes specify that dead load shall be considered in the design of structures. The significance of the various components of deed load in the design of a structure varies with the type of structure. In the case of a piping system, dead load is considered in the design of pipe supports. The dead lood lincluded in the design of a pipe support consists of the piping dead weight and the weight of all material attached to or integral with the piping, such as insulation, volves, etc. Since the dead weight of the pipe support itself is generally very small compared to the piping dead
(- load, thermal load and seismic load for which the support is designed, it can usually be neglected. Cygna believes that neglecting this specific component of dead load (i.e.,
support dead weight), except in the case of very unusual supports, is consistent with industry practice.
l With respect to the specific supports cited, the total dead weight of the support in CASE Exhibit 891 and 897 is 715 lbs and 82 lbs, respectively. This amounts to 4% and 2% of the design load for these supports. These percentages will be even smaller winen compared to the support capacities.
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( Comanche Peak ASLB Hearings !
Response to CASE Questions '
Question No.: Doyle #4 Exhibit No.: 891,894,897 1.0 CASE Question Weight of support masses os they affect pipe stress.
2.0 Cygno Interpretation What is the effeet of support weights (masses) on the pipe stress analysis?
3.0 Response Standard industry practice is not to include support mosses in the analysis of pipe stresses. This practice, which was employed on the RHR system Train B at Conmnche
- Peak, produces satisfactory results for the following reasons: 1) support weights are 4
relatively small; 2) support stif fnesses are relatively high; 3) support domping is typically higher than piping system damping; 4) standord analysis techniques are structured to envelope minor vorlations such as those ossociated with support masses.
The importance of each item is discussed in detail below, in order to help place these discussions in perspective, the following basic equation of motion may be useful.
M*x" + Cx + Kx = -Ug (1) where
) M = mass
- C = domping
'K = stiffness
'x = occeleration x = velocity x = displacement -
-Ug = input motion 3 J
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( Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Doyle #4
- Pope 2 .
Equotion (1) describes the response of a systwm (left hand side of the equation) to o l
particular input motion. If the input motion is set to zero and system domping is small, the response tendencies of the system con be colculated os, 4
3 (2) f =h h where l f = fundamental system respcase (frequency) ,
11 = 3.1416 1
3 Equation (2) links the system response (f) to basic system characteristics expressed in [
l terms of stiffness and mass. From this equation, it con be seen that on increase in
' stiffness will tend to increase the frequency, while on increase in mass will decrease the i frequency.
^(" Standard response spectrum techniques are founded on Equations (t) and (2), such that l i the system response con be directly related to accelerations plotted on a response l i
spectrum. Domping effects are normally included in this process by developing sets of ;
l response spectro for vorlous stondord domping values.
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Except for particularly unbalonced and massive supiwrt configurations,, which were not observed in the RHR reviewed by Cygna of Comanche Peak, support enosses are small i
relative to the piping system mosses that drive the overall response. ,
l In order to test this offact on Cwr.ad.e Peak, Cygno performed on analysis of a segment l of piping within our scope of review, using the ANSYS code. As litustrated in Attoch-i ment D4-1, the moln piping from the RHR pump to'the heat exchanger was studied. ,
Bronch lines, including the safety injection lines,'were omitted to make the model more i
manageable for this test, Basicolly, the test model contoins about 75 feet of moln piping i
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( Comanche Peok ASLB Hearings Response to CASE Ouestions ,
Question No.: Doyle //4 Page 3 with 16 supports. This model was anolyzed with and without support mosses using the some standard onolysis techniques employed on Comanche Peak. The only difference between the two onolyses was support mosses, which are listed below Support Masses Support Number .Weight (Ibs)
RH-l-010-003 72 RH-t -010-004 42 RH-1-010-002 II RH-l-010-001 87 RH-l 064-010 41 RH-t -064-004 77 RH-1-064-011 25 RH-l-064-003 15
('_ RH-l .064-005 26 RH-l-064-009 24 RH-l-064-002 27 RH-l-064-006 50 RH-l .064-007 56 RH-1-064-008 122 RH-l-064-001 31 RH-l.010 005 30 The results of this test are contained in Attachments D4-2 (colculation package), D4-3 (computer output without support mosses), and D4-4 (computer output with support mosses). A summary of the system frequencies and pipe stresses of the most massive support (RHR-l.064-008) is provided below -
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Comanche Peak ASLB Hearings Response to CASE Questions
(- Ouestion No.: Doyle #4 '
Page 4 ,
reW esi Without With Support Mass Support Mass Difference Mode (%)
No. (hertz) (hertz) 7.7 7.6 1.1 decreme i I 10.3 10.1 1.9 decrease 2
12.3 11 A 7.5 decrease 3 2.9 decrease 4 20.0 19A 22.1 21.9 1.0 decrease 5 '
1.3 decrease d 23.2 22.9 28.2 27.4 2.3 decrease 7 5.0 deerease 31A 8 33.0 Pipe Stress (at Support RH-t-064 008)
Without(I) WithU)
Difference
( Sup(port hertt) Mass Sup(port hertz) Mass (%)
(, Stress as th7 2Ak l A decrease
-2.32 -2.29 i A decrease
- a l A decrease 51 2.50 .. 2A5 2A8 2A5 I A decreme
'E where '
a, a maximum principal stress Fa = minimum principal stress 51 = stress intensity = maximum of g-a,, g g , a,.a, ~
= s .
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aE equivalent stress ,y, = y-(q-a,ina,- a,i'+ eg-a,.; ,
' (1) From computer output dated 4/10/84 @ 10:29 for element II, node 4 (2) From computer output dated 4/10/84 e 10:21 for element 13, node 14.
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( Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Doyle #4 PageS Interpreting these results, it con be seen that the added mass results in only a minor decrease in system frequencies. Pipe stresses actually decrease slightly, but the changes are negligible.
Pipe Stiffness Pipe supports are normally designed to be rigid in their support direction relative to the ottoched piping. This design method tends to uncouple support response from overall system response.
In the off (non-support) direction, the support stiffness normally has no effect. Unless gaps are provided to uncouple the support mass in the off-direction, the mass will participate with the piping. The effect of this interaction has already been shown to be negligible in the " Relative Support Weights" discussion.
Sepport Dampino Domping directly associated with pipe supports is not considered on Comanche Peak.
However, if support mosses and stiffnesses are included in the analysis, there support domping should also be included.
As shown in Equation (I), the accelerotion and displacement terms will tend to decrease inr a given input motion as domping (velocity term) increases. USNRC Regulatory Guide 1.61 recommends domping values up to ~% for structures and 3% for piping under SSE loodings. Therefore, if support response is a significant contributor to overoH system response, then the overall system domping will fall somewhere between the individual domping values for piping and supports.
Standard Analysis Techniques There are many conservatisms built into the' standard onolysis techniques that are intended to simplify the onelysis and focus on the most significant mechanisms. A few of these ore briefly discussed beloW:
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Comanche Peak ASLB Hearings
( Response to CASE Questions Question No.: Doyle #4 Page 6
- c. Low System Damping Researchers have shown that piping systems exhibit domping values greater than those allowed by USNRC Regulatory Guide 1.61. For example, the Pressure Vessel Research Council (PVRC) hos proposed the domping values shown in Attachment D4-5.
- b. Modal Response Method This method combines Individual system responses (modes) without regard for
- direction (or signs). For example, even though responses rnay be either left or
! right, this technique assumes that all responses act to the right. A more refined onelysis would circumvent this combination technique, but the costs are not practical for production analyses.
- c. Spectro Broodening Motions input to the piping analyses in the form of response spectro contain two significant conservatisms: (1) the rough (sow-toothed) spectro are broadened, usucIly f.15%, and (2) the rough shape is enveloped by a smooth curve.
- d. Ground Spectro The shape of the ground spectrum is generically defined per USNRC Regulatory Guide ' I .60. A site-specific spectrum would normally impose significantly less demands on the structures, systems and ccmponents. The. peak ground occelerations are also based on conservative interpretations of the geotechnical conditions,
- e. Elastic Analyses Pipe systems have considerable inherent strength that is not topped by the standard analysis / design technioues. Being constructed of steel, these systems exhibit This is of ten defined as ductility. In strength well bey (ond .he yield point.Attochment D4-6 Appendix A to St O
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( Comanche Peak ASLB Hearings Response to CASE Questions '
Question No.: Doyle #4 Page 7 recognizes this fact. Although Attachment D4-6 is intended for impact or impulse loods, it shows that steel members in tension con resist strains up to 10 times yield and still perform their intended function.
In conclusion, Cygno does not recomrnend that the conservatisms noted above be deleted from the Comanche Peak analyses. But, on the other hand, the presence of these conservatisms should be recolled whenever minor offects are considered, such as the offeet of support masses on pipe stress analyses. Regarding the Comanche Peak practice of not including support mosses in the piping onalysis, Cygno considers this practice to be consistent with industry proctice and with the degree of refinement of the onolysis techniques. Furthermore, the test problem results show that support masses have a negligible effect on pipe stresses in a system similar to the one reviewed by Cygno.
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ATTACHMENT D4-6 (Paga 1 of 1)
APPENDIX A
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STANDARD REVIEW PLAN SECTION 3.5.3 PERMISSIBLE DUCTILITY RATIO
_FOR OVERALL DAMAGE PREDICTIdN I.
INTRODUCTION elements (e.g. , missile barriers, columns, .
and steel slabs, structural etcI
' response (i.e. , ductility ratios greaterof nonlinear is generally acceptable provided that the intended than unity) o ructural elements or protected by the elements are maintained. structural onents supported elem positions for review and acceptance of ductility ratios forThe following summ and steel structural elements subjected to impactive andconcrete reinforced ve loads.
impulsi II.
SPECIFIC POSITIONS 1.
Reinforced Concrete Members ductility ratios is stated in Regulatory Guide 1 142T o permissible of Revision 1 of Regulatory Guide 1342, the staff Prior to publication ductility will be provided to applicants on a caseposition regarding
( 2.
Structural Steel Members by-case basis.
- a.
For tension due to flexure pd i 10 0 b.
For columns with slenderness ratio (1/r) equal to of less th an 20 pd I 1'3 Where 1 = effective length of the member r = the least radius of gyration For columns with slenderness ratio greater than 20 pd i 1.0 M c.
For members sutrfected to tension f i
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u pd < 0.5 y7
~
Where e = Ultimate strain
'y = Yield strain ,
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3.5.3-7 mew. 0 - July 1981
Comanche Peak ASLB Hearings '
( Response to CASE Questions Question No.: Doyle #5 .
Exhibit No.: 891, 894, 897 1.0 CASE Question I
Inoccurate conclusions os related to KL/R for pinned columns:
o if a column fixed at its base and free at the top hos on effective K of 2.0, cutting at some point up from the base and adding a pin does root address the problem.
2.0 Cygno Interpretation ,
Does a stability problem exist for CASE Exhibits 891,894 and 8977 3.0 Response The stability chorocteristics of a structure under the action of compressive loads con l generally be separated into three categories. These include rigid body modes of instability, Euler column buckling, and beam-column effects. For time purposes of discussion, the three support configurations in question (CASE Exhibits 891,894 and 897)
(. con all be idealized to the basic configuration shown in Figure I, wherein the x ,
component of reaction at A is provided by frictional clamping forces. For this basic l configuration, the rigid body modes of instability generally occount for the overall stability chorocteristics of the entire structure, while Euler column buckEing and beam-column effects are confined to the individual members.
. PIPE (
A ny B
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Comanche Peak ASLB Hearings lI Response to CASE Questions Question No.: Doyle //S Page 2 The rigid body mode of instability con be initiated in three ways: (1) when the clamping force at A is insufficient to develop the lateral (x) component of frictional force necessary to prevent sliding; (2) when the clamping force at A is insufficient to develop the resisting torque necessary to prevent the clamp from rotating; and (3) for the '
specific cose of alpha equals theta, when the contilevered member BC does not provide sufficient laterol stiffness at point B to prevent rigid body rotation of member AB.
Euler column buckling of member AB con occur for all values of alpha and theto given in the three exhibits. The correct value of K to be used in evoluoting the stability of member AB is 1.0, since the member is pinned at both ends and con therefore only develop exial loads. Similarly, Euler column buckling con eccur in member BC but only when alpha equals theta. The correct value of K to be used in evoluoting the stability of this member is 2.0 since the member is fixed at one end and free at the other.
Beam-column effects account for the fact that the bending stresses produced by laterol
, loads on a column are amplified by the presence of the oxial isood. What this means is that the maximum stress in a laterally loaded column is not siemply the sum of the oxial
(' stress and bending stress, but is in fact the sum of the axioil stress and on omplified
( bending stress. This amplified bending stress is the product of the bending stress produced by the lateral loed and on amplification factor which i:s given by the expression I I- /Fcr where P is the oxial load in the column and Per is the Euler buckling load for the column. Only member BC is influenced by the beam-colurnn effect. Obviously beorn-column effects have no influence on member BC when members AB ond BC are either co-linear or perpendicular.
Each of the three CASE exhibits con now be briefly discussed with respect to each of these three categories of instability.
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Comanche Peok ASLB Hearings e 4 ( Response to CASE Ouestions Question No.: Doyle #5 .
Page 3 TABLE D5-1 CASE Exhibits i
891 894 897 800 M) N/A Required Lateral Stiffness of point B (Ibs/in) 700,000 2,D00 N/A Actual Laterol Stif fness of point B (Ibs/in)
TABLE DS-2 Required Recpuired '
Type of Clamping Belt -
Force Tosrque l CASE Clamping Force Exhibit Resistance (Ibs) (ft-lbs)
Sliding 150 2 891 107 I Rotation 96 1 Sliding 894 16
. Rotation Sliding N/A N/A 897 122 I Rotation Euler Buckling of member Bb has been accounted for in the calculation and is not a problem. Member AB'fs a pre. qualified component and as such is stable with respect to Euler Buckling.
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- Comanche Peak ASLB Hearings
( Response to CASE Questions Question No.: Doyle #5 Page 4 The only support for which beam-column effects are opplicable is CASE Exhibit 891.
Since the critical buckling lood for member BC is so large the amplification factor is 1.00. Therefore, beam-column effects have no influence.
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l Response to CASE Questions' ,,
Question No.: Doyle #6 '
i l Exhibit No.: 891 ,
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1.0 CASE Question r 1
16-inch pipe with about 20 kips,lood clN 3-l/2 inchri of length induces high bearing l stresses which require pods. This is swt addressed. ,
i o ASME Code against flottening. . j j j
_ t 2.0 Cygno interpretation How did Cygno evoluote the stresses induced into the piping by the following, os related to the ASME Code coution ogoinst'insseing excessive flottening into the pipe wolli l . .
- a. U4 cit?
l b. 5" x 5" x l/2" tube' steel frome?
3.0 Resporse l (.. l Cygno originally evoluoted the general code requirements for ottochmerits to piping and i Temos Utilities' opplication of the. code. l In Section ill, the ASME B&PV C$de provides the fol8owing caution Subsection NS-3645 (Class I Component:),
" Lugs, brockets, stiffeners, and other ottochments may be welded, bolted, or studded to t
the dutside or inside of piping. The offects of ottochments in producing shermal stresses, stress concentrations, and restroints on pressure retolning members shall be taken into occount in checking for enmpliance with stress criterio."
Subsections t4C N45 (Class 2) and ND-3645 (Class 3)
/
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' External on/ hternal attochments to piping shall be designed se es not to cause i l flottening of'the pipe, emesssive locollaed be4 ding stresses, or harmful thermol podients in the pipe wel(. It is*Impertant that such ottochments be designed to minimize stress concentrations in applications 'where the number of stfeds cycles, due either to pressure l or thermol offact,is relotlysiyje70e for the expected life of the equipment."
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Comanche Peok ASLB Hearir,gs
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Response to CASE Ouestions Question No.: Doyle #6 Page 2 The Code statement for Class I components specifies thct local effects due to ottochments shall be token into occount for compliance with the stress criteria. The Code statement for Class 2 and 3 components, such as those associated with RHR systems, specifies that ottochments shall be designed to minimize localized stresses of
- the pipe. It doe *, not define the term "flottening." A reasonable interpretation would be that the designer of Class 2 and 3 piping should consider the significance of any odditional stress Induced in the pipe due to ottochments. Such a consideration does not imply a requirement for calculations in all Instances depending upon the method of attachment.
The Comanche Peak project did use CYLNOZ, a local stress analysis program, when welded ottochments were mode to the RHR system. It is not common practice to onolyze the effects of bearing or clamping except where judgement Indicates the need for such on evoluotlon based on the specifics of a porticular design.
In its original review of the adequacy of the loods introduced into the pipe wall by support Sl-1-325 002-532R (CASE Exhibit 891), Cygno considered the following:
( c. U-bolt. Cygno judged that the loods introduced into the piping due to design loods would not prevent the piping from perform.1ng its intended function.
U-bolts are frequently used in the industry for simllor opplicottons. Further discussions on U-bolt opplicottons are provided in response to Doyle Question
- 1.
- b. 5" x 5" x I/2" Tube. Although on unlikely ochievement, the drawing detall speelfles o 0" gop of all four bearing points. Cygno reviewers concluded significant stresses would not develop in the pipe. It should be noted that rodlot thermal growth for such a 16" pipe would be I/50", obout the width of two business cords. An onolysis of these oflects on the pipe was performed to substantiate our judgement on the worst cose effects and is contained in Attachment D6-l. The results show that the stresses are acceptoble. It is important to note that this is not a typical detall.
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4 Comanche Peak ASLB Hearings -
( Response to CASE Questions Ovestion No.: Doyle #7 .
Exhibit No.: 891 ,
1 i
l 1.0 CASE Question ,
i Clip ongle 4" x 4" x I/2" wSch supports U-bolt not addressed (critical to maintaining stability):
o Section modulus .04 in cube o Moment arm at least 2 inches o 1100# load exceeds Code allowobles.
o Pre-tensioning to obtain a clamping force required could exceed this (not including thermal constraint and design loads) o Clamping force with no margin of safety for single degree system (not point contact or line contact) is force / coefficient of friction or about 4 times what is required for clamping force. -
s 2.0 Cygno interpretotion Did Cygno check the clip ongles (item 15 on support no. Si-i-325-002-S32R) for a
~
potential overstress condition due to: U-bolt torquing, thermai' toads, and mechanical
(- loads?
3.0 Response During the original Cygno review of this pipe support,'o judgement swas mcde that the friction forces necessary to resist sliding of the support along the lengith of the pipe were minimal. At first, these small resisting forces were assumed to be developed by the U-bolt while the mech'anical loads, those resulting from static, the:rmal, and dynamic analyses would be resisted by the box frame. Cygno believes that thiis was a reasonable assumption to make, given that the support drawing calls for F' ciecronce between the pipe and the box frame. However, because the U-bolt was connected to the support through clip ongles that were not considered substantial, a theoretical loss of U-bolt copobility was assumed. The reviewer assessed that given this possible loss of U-bolt function and capabilities, sufficient friction ~ forces to resist siisng would still be developed between the box frame and the pipe. These frictional forces were calculated as part of the response to Doyle Question #5 and found to be suffscient to resist the sliding effects required'io maintain stability, l 1
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- Comanche Peak ASLB Hearings
( Response to CASE Questions Question No.: Doyle #7 Page 2 ,
A review of the installation instructions (not within the scope of the Cygno audit) indicates that the torque placed on the U-bolt not in the regular course of installation would theoretically overstress the clip ongles. Although the installation procedures were not considered in the Cygno review, the correct conclusion was reached since the reviewer ossumed a loss of U-bolt capability.
Cygno does consider this support to be o poor detail if significant cinching loads have been applied to the U-bolt. Installation practice is o new consideration which will be accounted for os part of the on-going Phase 3 review.
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Comanche Peak ASLB Hearings R:sponse to CASE Questions Question No.: Doyle f/8 Exhibit No.: 893 t the conclusion that flore this welds are t
1.0 CASE Ouestion l
There is no documentation i in calculations to supportherefore, wh t strcnger than fillet welds -no calculat ons; statement? th).
o Flore weld strength depends on radius of flore (dep 2.0 Cygno Interpretation fi.iet welds when no calculations wer :
Why did Cygno consider flore welds stronger than !,
mode? !
3.0 Response Pipe Support Design Review be Criterio,"
As specified in Cygno Design Criterio DC-2, "ith AW3 D.I.I.
welds were reviewed for compliance wAs shown below, in the cose o "unsatisfoetory."SI-l.079 001 5325, flore welds are stronger than a ,
Minimum effective throat thickness (t,) is greater
- 1) !
o For flore weld:
t, = 5/16 R =' 5/16 (5/8a) = 0.20" di where R = minimum weld groove ra us
= inside radius + thickness
= 1/8" + 1/2" = 5/8" o For fillet weld:
t, = 0.707 (1/4") = 0.18" t- 0% stronger than o I/4a fillet weld.
' since 0.20">0.18", a flore weld is 1 l; -
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Comanche Peak ASLB Hearings
(
\ Response to CASE Questions Question No.: Doyle #8 Page 2
- 2) More weld length For the welded beam attachment considered, the weld length is 2" along the square side versus 3" along the beveled side. Consequently the installed flore weld along the bevel will give this support 50% more capacity for the some te.
Therefore, changing from o I/4" fillet weld to o minimum flere bevel groove weld increases the copocity of the weld by 65%.
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l Comanche Peak ASLB Hearings '
( Response to CASE Questions Question No.: Doyle #9 Exhibit No.: 892 I
1.0 CASE Ouestion l 1
The reduction of weld capacity in the calculation is based on 135 degrees. Actual !
tangential angle is 150.3 degrees. Therefore, on error exists. Did Cygno take note of l this?
e More stress in weld than stated.
. Wide / thin ratio induces cracking as well as the 1.4:1 ratio of width to depth.
2.0 Cygno interpretation What was the basis for concluding that the stanchion-to-pipe weld shown in CASE Exhibit 892 is adequate?
3.0 Response C' ITT Grinnell design procedure, SA 3912, (Attachment D9-1) states that credit shall only be taken for the portion of the weld up to 135 degrees. Cygno concurred with this procedure and confirmed that it was properly employed on the subject support.
Attachment D9-2 shows that the weld length included in the strength calculatica was only that portion where the angle between the stanchion and the pipe was less than or equal to 1350
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ATTACHMENT D9-1 (Page 1 of 4) SA 3912 Rev. A Page 1 of 33
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WELD PROPERTIES FOR PIPE / STANCHION AND EL50'.1/ STANCHION CONNECTION FOR COMANCHE PEAK PROJECT PROCEDURE SA 3912 C
703INctayATDNCNLY Prepared By b
Checked By_ [h44 ) /" ,,,, , 4 ) ,ig, 27y/gy ApprovedBy;yffc. "G >
Revisien A 02/08/83 c mP- G Re v'- l CC t
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PAGE 110F 33
!.T.T. GRINNELL PIPE HANGER OlVISION S A-3 913 REY. A
( ATTACHMENT D9-1 (Cont.) .(
.(Page 2 of 4)
WELD ANGLES FOR STRAIGHT ~:! PES WITH STANCHION ATTACHMENTS
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FIG. 1
PAGE .,o c; 33 1.T. T. GRINNELL PIPE HANGER CIVI!!ON s A-3 912 R EV. A ATTACHMENT D9-1 (Pa9e 3 of 4)
(Cont.)
4.1 TABLE 1 WELD PROPERTIES OF STR AIGHT PIPES WITH STANCHION ATTACHMENTS (Ref. Fig. 1 & 4)
LIMITING WELD ANGLE = 135 i
WELD PROPERTIES NOM. NOM. OVERALL PIPE STANCH. WELDED size SIZE LENGTH L, gy S, Jw Ls 1
1 4.24 4.24 1.36 1.36 1.79 4.24 i 6.24 2.84 2.84 5.39 6.24 2 1/2 1 1/2 6.24 1
8.04 5.36 4.17 1.74 , 7.01 5.36 2 ;
2 1/2 11.08 6.16 5.64 1.57 10.37 6.16 6.16 2.84 2.84 5.39 6.16 1 1/2 6.16 , _
7.82 4.43 4.43 10.52 7.82 3 2 7.82 6.12 2.54 12 . A r. 6.48 ;
2 1/2 9.72 6.48 7.50 8.36 2.32 18.71 7.50 3 13.49 7.69 4.43 4.43 10.52 7.69 2 7.69 9.42 9.42 6.49 6.49 18.66 9.42 l 4 2 1/2 l 11.72 8.46 9.29 4.60 24.32 8.46 3
17.35 9.64 13 . P ". ; 3.85 39.76 l 9.64 4
3., . 11.34 11.34 9.62 9.62 33.67 11.34 l 14.82 15.90 15.90 71.57 14.82 !f G 4 14.82
}
6 25.54 14.19 29.96 8.34 126.87 14.19 ]
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, 1.T. T. GRINNELL PIPE HANGER DIVISION SA-391: RIEV. A PAGE 710F 33 ATTACHMENT D9-1 (Page 4 of 4)
~(Cont.)
4.1 TABLE 1
. WELD PROPERTIES OF STR AIGHT P1 PES WITH STANCHION ATTACHMENE (Ref. Fig. 1 & 4)
LIMITING WELD ANGLE = 135 N OM. NOM OVERALL WELD PRCPERTIES PIPE STANCH. WELDED SIZE SlIE LENGTH Lw Sr Sr Jw Ls 4 14.57 14.57 15.90 1 5 .90 71.57 14.57 8 6 22.14 17.22 33.86 159.76 177.62 17.22 i 8 33.25 18.47 50.77 14 .14 279.96 19.47 4 4-14.46 14.46 15.90 15 .90 71.57 14.46 10 6 21.65 21.65 34.47 344.47 228.37 21.G5 l f
( 8 29.05 20.97 56.44 2 7.95 363.95 20.97 i 10 41.44 23.02 78.88 21 M 542.05 23.02 4 14.41 14.41 15.90 15 ,
71.57 14.41 ,
r; 6 21.45 21.45 34.47 34. 47 228.37 21.45 ;
12 8 28.40 28.40 58.43 58.43 503.93 28.40 10 36.56 24.37 85.53, 3!5.49 650.46 l 24.37 1
12 49.15 27.31 110.95 3 CD . 91 904.37 27.31 !
6 21.37 21.37 34.47 3-0 228.37 21.37 8 .,. 28.18 28.18 58.43 5e .43 503.93 22.16 .i i i t 14 10 35.93 27 84 E9.16 52.02 l755.87i 27.E4 [
FOR ! UOR MATLON ON.y
ATTACEMENT D9-2 (Page 1 of 1)
(
PIPE ANGLE SUBTENDED BY THE PAD j WELD AT THE PAD /STAtiCHI0li f INTERSECTIOtt
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WELD METAL INCLUDED IN
( c m /- m l STRENGTH CALCULATION l WELD METAL EXCLUDED FROM STRENGTH CALCULATION
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SECTION k-A STAtlCHION - TO - PIPE WELD
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Comanche Peak ASLB Hearings
~\ Response to CASE Ovestions I l
Question No.: Doyle #10 .
i Exhibit No.: 893 )
1.0 CASE Ovestion Hos this been Chonging from o flore weld to o fillet weld induces flange bending.
addressed by Cygno?
2.0 Cygna interpretation The calculation sheet attached to Exhibit 893 states that the we bracket and the beam flange was changed from o fillet to o florea/ bevel weld. T the weld lines, from to-flore change results in a 90 degree re-orientation of[Did Cygno evoluote the perpendicular-to-porallel to the web of the wide flange.
additional loads on the flange?
3.0 Response Cygna judged that this re-orientation would not cause on oversstress in following calculation verifies that judgement:
hD siNooi.532s
"'""'fg""""
.s I : weed Lenom . 3*
Consilever Etf active width
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y 3* . 2(l.375") . 5.75" g ,. _
/criesame el section es web eterseccian. face fienge Conservatively p - N" W ""
5 . Y (.345)2 , ,ggg g,3 .3 3. ..
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Mio able stress = .75F, e 75(32,350 poifs 24.262 psi g g., ,L.23* . f.375*
18,382 (, 26,262 psi ,
uS2s* v Cygno ogrees that the maximum stress condition is due to flanage bending.
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Comanche Peak ASLB Hearings )
f Response to CASE Questions ;
i Question No.: Doyle #1I l Exhibit No.: 891, 894, 897 l l
1.0 CASE Question Effects of out-of-plane seismic excitation of support hardware not included in calculation. Did Cygno address this point?
e Additional loads on support?
e Additional loods on pipe?
2.0 Cygno Interpretation Did Cygno evoluote the effects of support self-weight excitation in the off-direction, os related to:
- a. support design?
- b. pipe design?
3.0 Response
- o. During Phase 2 of the Independent Assessment Program Cygno did identify this os o potential problem. In the IAP Report, Cygno noted that self-weight excitation was not included in the support design. Note I to Checklist PS-01 states:
" Support Self-Weight Excitation in general, pipe support vendors have not included support loads due to self-weight excitation in their fooding. Texas Utilities hos done a generic study in response to Walsh/Doyle ollegations which shows the effects are negligible.
The NRC Site Inspection Team (SIT) has reviewed and accepted this evoluotion in item 3.h of Inspection Reports S0-44S/82-26 and S0-446/82-14."
Since the IAP was performed for the NRC Stoff, further evoluotion of on issue ofready identified onCreviewed by the Staff would have been redundant.
Accordingly, Cygno noted the potential deficiency on the oppropriate checklist and deferred to the Stoff evoluotion. .
b.- The effect of support masses on the piping onalysis is discussed in the response to Doyle Question #4. 1 I
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Comanche Peak ASLB Hearings '
(- Response to CASE Questions Question No.: Doyle #12 '
Exhibit No.: None l.0 CASE Question Restraint of rotation by the pipe because of coupling effect of hardware on both sides of a pipe:
e Load increase in I of 2 snubbers / struts e Alteration of dynamics of pipe system during seismic event 2.0 Cygno Interpretation Support RH-l-010-003-S22R consists of two struts attached to trunnions, which are welded to the pipe at diametrically opposing points. How was this considered in the piping and support evoluotions?
3.0 Response Cygno reviewed the pipe stress analysis to determine whether or not accepted modeling techniques were employed. Cygno determined that the RHR pipe stress model used by
( Gibbs & Hill was acceptable when compared to general practice. CASE has proposed the need to model certain pipe support configurations into the stress onalysis which is different than the existing approach. Gibbs & Hill reron the onelysis of piping segment AB-l-70 (see Walsh Question #11) using the CASE technique The results, when compared to the original analysis, were different, however, there 3 no reason to believe the Gibbs & Hill model is inappropriate.
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Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Doyle #13 Exhibit No.: 891 1.0 CASE Question in Note 2 following pages PS-01-4 of 4 , Cygno decided to eliminate their stiffness i
criteria based on their knowledge that a report existed to oddress the problem (but without personal knowledge of what was contained in the document in detail). Why didn't
! Cygno consult with their experts - for example Eric von Stijgeren (who was the editor on a paper by T.Y. Chow, C.H. Chen and O. Bilgen) - in reference to deviations from generic stiffnesses in pipe supports and the ef fects on piping systems.
? e Third paragraph introduction et. seq. (CASE Ex. 884).
2.0 Cygno interpretation Did Cygno evoluote the effects of support stiffnesses on the piping analyses?
3.0 Response During Phase 2 of the independent Assessment Program Cygno did identify this issue os a potential problem. As stated in the IAP Draft Report, Cygno questioned the pipe support stiffnesses utilized on Comanche Peak. Note 2 to Checklist PS-01 states:
" Pipe Support Stiffnesses The' NRC SlT raised the issue of support stiffness in item 3.j of the above referenced reports. Gibbs & Hill has performed a generic study for review by on NRC consultant. The study shows that using 1/16" deflection criterio on support design provides acceptable stiffnesses for the piping onelysis (chonges in support stiffness do not greatly offect piping results). The NRC review results were not available at the time of the Cygno review."
Since the IAP was performed for the NRC Stoff, further evolvation of on issue already identified and reviewed byithe Staff would have been redundant. Accordingly, Cygno recorded the potenti,ql deficiency on the oppropriate checklist and deferred to the Staff evoluotion.
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Comanche Peak ASLB Hearings e
( Response to CASE Questions Question No.: Doyle #14 Exhibit No.: 891, 893, 894, 895, 896, 897, 898, 899, 900 1.0 CASE Question in Note I, the some source, did Cygno consider the additive effects of self-weight excitation if the stiffness is considered from node point to hard point as opposed to the stiffness of the frame independent of hardware, local effects, baseplate and anchor bolts?
o Spring rate of boseplate/onchor bolts (particularly bearing-type joints) con be considerable (observation of baseplate 11 finite analysis).
2.0 Cyr Interpretation Did Cygno consider the following:
- a. The effect of support s'tiffnes.s on the evoluotion of self-weight excitation?
- b. The flexibility of each element in the support load path?
(.. 3.0 Response
- o. In order to evolute the influence of self-weight excitation on support design, one must opply the oppropriate dynamic loads and then calculate the induced stresses and deformations. The opplied load, in this case, is the support self-weight.
Support stiffness is effectively considered twice in this process. First, it is included in calculating the opplied dynamic load. This con be illustrated by the following elementory formulos:
- 1. Load = function (freq)
- 2. freq = (I/6.28) SORT (Kg/F) where freq = support fundamental frequency K ~
- t. = support stiffness ,
F = self-weight l "g = gravity l
^
- Comanche Peok ASLB Hearings
( Response to CASE Questions Question No.: Doyle #I4 Page 2 Secondly, the determination of support stresses and deflections involves a structural evoluotion which considers the support stiffness.
For o further description of Cygna's review process relative to support self-weight excitation, see the Cygno response to Doyle Question #11.
- b. As stated in the response to Doyle question #13, Cygno recorded that support stiffness calculations on Comanche Peak were potentially deficient. When it was learned that the NRC Staff had evoluoted this issue, Cygno deferred to the Staff evoluotion rather than performing a redundant review.
- Regarding the effects of component flexibilities on the overall support stiffness, current standard practice is not to include the baseplate connection. These effects are being studied by various industry groups. One such group is the Structural Engineers Association of California (SEAOC). An update on their activities is provided in Attachment Wl4-1. Until resolution is reached on the relative merits of considering the baseplate connection in the stiffness calculation, Cygno does not consider it reasonable to evoluote Comanche Peak against a requirement to include these effect.
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Communications i AL% i Report i
llifillllllllilillllllllllllli ATTACHMENT Wl4-1 l
(Page 1 of 2) l(~ cornpanr
- Teiec n conference nepon Texas Utilities Proje*t Joe No Comanche Peak Steam Electric Station 84042 Independent Assessment Program - Phase 3 4/10/84
Subject:
Time:
3:00 pm Column Base Plate Flexibility ****
- Sp Participants' of Helmut Krawinkler (415) 497-4124 SEA 0C and Stanford Universitw T. Wittio (415) 397-5600 Cyana F.cogired j ltem Comments Action By
~
Reference:
" Recommended lateral Force Requirements and Commentary," SEAOC,1980.
Mr. Krawinkler chairs a Structural Engineer Association of California (SEAOC) subcomittee on " Steel."
I asked for an update on activities related to the following excerpt from Commentary Section 4 of the referenced publication:
" Column base connection performance is of particular concern where a fixed base is assumed in design. The effects of inelastic extension of anchor' bolts on column moments, frame drift and stability need investigation;"
Mr. Krawinkler noted that this question is complex and that SEAOC has not established a position. Furthermore, there will be no position stated in the upcoming revision to the referenced document. ,,
~
Regarding the,, application of this question to pipe supports, he emphasized that Section 4 is-titled " Steel Ductile Moment Resisting Frames." The comentary note was added because hinge formation needed to develop ductile behavior in steel framed buildings could conceivably occur within the column base plate
~
connection. Since information on the ductile behavior of such connections is insufficient, tha issue was identified as Signeo. 1 1J ' Page o8 YkTk,hAAA4 lom 1 2 H. Krawinkler (Stanh Univ.), D. Wade, N. Williams, G. Grace, T. Wittig,
' m e'. Project File
Communications ALn i Report l1111111111lll111111lll1111111 ATTACHMENT Wl4-1 (Page 2 of 2)
( comments [c"e4 iiem requiring study. Applying this question to pipe supports is clearly inappropriate, because they are not designed as ductile moment resisting space frames.
I told Mr. Krawinkler that our conversation would be reported during the hearings on Comanche Peak.
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Comanche Peak ASLB Hearings ,
Response to CASE Questions
( Question No.: Doyle #1S i Exhibit No.: 891, 898, 899, 900 1.0 CASE Question hos thermal lockup considered for anchors which restrain pipe radial growth?
o Induces frame moments.
2.0 Cygno interpretation o How was the effect of thermal radial pipe growth considered in the review of CASE Exhibits 891,898,899, and 900?
3.0 Response Exhibit 891 (Suppcrt SI-l-325 002-522R)
Exhibit 891 shows a box-frame enclosing a 16" diameter pipe. The design details specify o zero gap between the pipe and frame at the four points of contoct. Cygno reviewers evoluoted this configuration and judged that the thermal stresses would be acceptoble.
To oddress some concerns raised during the ASLB hearing regarding this issue, Cygno performed a finite element analysis of the frame / pipe with zero gops. Figure Dl5-1 shows the model. The pipe was heated to 3S00F ( T = 2800F) and the flexibility of both the pipe and frame were considered.
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Comanche Peov. ASLB Hearings
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Response to CASE Questions Question No.: Doyle #15 Page 2 The results are summarized below:
Thermot Only Stress Allowable % Allowable Element (psi) (psi)
Pipe 37,700 64,800(I) 58%
38,300 56,400(2) 68%
Frome Thermal + Mechanical Stress Allowable % Allowable Element (psi) (psi)
Pipe 39,300 64,800 61 %
39,800 56,400 71%
Frame
(' Notes:
(1) 3Sm per ASME B&PV Code, Section 111, Figure NB-3222-1. Sm = 19,300 psi per Appendix 1 for SA376, Type 304, material at 350 F.
(2) 35 per ASME B&PV Code, Section !!!, Porographs NF3213.10 and NF3231.lo.
S = 0.6 Sy, where Sy = 36,000 psi per Appendix I for AS00, Grade B, tube steel of 700F.
Note that the element stress allowables are cosed on membrone plus bending stresses defined in the ASME code. This is appropriate because the rnodel employed discrete, shell elements.
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( Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Doyle #15 ,
Page 3 Exhibits 898 and 900 (Supports SI-t-037-005-S22A and SI I 030 003-S32A, respectively) diametricolly opposed trunions DIS-2 illustrates this configuration.
wherewhich the pipe is form welded to the h member of the frame. Figure PIPE 7 I
{ ] i
.-- FRAME S-+ '
? fBOLTED ANCHORACE i .
FIGURE 015-2 impoet on the support design constroint of free end displocements.
expansionsince the to have stresses negligible uced in the support result from the pipe thermal expansion. increase in allowoble stress when th evels are combined with the effects of To demonstrate this conclusion, Cygno performed a h 898 (See Attachment DIS-1) which incorporate the effand calculation f with mechanical loads. The results ects of of pipe thisthermol expansion calculation h '
ond baseplate are below the allowables.
I s ow that all stresses in the frame !
Exhibit 899 -
Figure DIS-3 illustrates this configuration.I i
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i Comanche Peak ASLB Hearings
( Response to CASE Questions Question No.: Doyle #15 Page 4 By inspection, thermal radial growth of the pipe is primarily unrestrained. A secondary restraint will develop at the bimetallic weld due to thermal gradients and the material differences. Cygna's reviewers judged the effect of this secondary restraint to be negligible.
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Comanche Peak ASLB Hearings .
(. Response to CASE Ouestions Question No.: Doyle #16 Exhibit No.: 891,897,898,899,900,906 l
1.0 CASE Question - )
i i
The baseplate analysis is based on distribution of shear relative to load path / stiffness for l
oil bolts in the pattern. Did Cygno address this problem?
l e With oversized holes and the inability to eliminate construction tolerances (location of the bolts combined with location of the bolt holes), it is not possible for all of the bolts in the system to be active (see CASE Exhibit 900.
e The stiffness of the bolts is such that deflection connot be counted on as a means to achieve full pottern participation.
e Even if deflection could result in full activity, the first bolts deflecting would receive the larger portion of the load in on ideal symmetrical and . ystems.
. For non-symmetrical system and systems of variable stiffness, the inoctivity of a number of the bolts will alter the occuracy of the computerized onelysis.
2.0 Cygno interpretation N/A
> 3.0 Response a
The determination of the distribution of snear forces to the anchor bolts of a baseplate is based upon the some methodology which hos for decades been sucessfully used for the design of bolted connections of both bearing and friction type. In this " conventional method" of bolted connection design it is assumed that all bolts in the pottern are active to one degree or another depending upon the location of the pottern center of twist relative to coch bolt. Should the center of twist lie within the bolt pottern, some bolts may be completely inoctive compared to others in the pattern. Where the pattern center of twist is for exterior to thehlt pottern it is more likely that all bolts will be equally octive in resisting sheor-forces. Using this method the forces on the most highly stressed bolt within the pottern then determines the bolt size to be used for the entire pottern.
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4 Comanche Peak ASLB Hearings
( Response to CASE Questions ,
Question No.: Doyle #16 l Page 2 Cygno finds no problem with this standard design methodology which is referenced in oil standard textbooks which deal with the design of bolted connections.
In responding to the question, it will be assumed (conservatively) that no friction whatsoever con be developed between either the baseplate and the concrete or the anchor bolt nut / washer and the baseptote. For this extreme cose it will be explained how full baseplate functionality to resist the ultimate design shear forces is rnaintained.
Construction tolerances associated with either locating the bolt hole in the baseplate or the bolt hole in the concrete have obsolutely no influence on the distance that a baseplate must move before it bears directly on on anchor bolt. The only thing that offects the maximum distance that a baseplate must move until it bears directly against the bolt is the difference between the diameters of the bolt hole and the bolt. At Comanche Peak this maximum distance is 1/16" for bolts less than la and I/8" for bolts I" and greater, although most baseplates with I" holes which have the lesser oversize of I/16" specified. Oversized holes is a fact of life in connection design. Codes specify the allowable oversize for various types of connections.
With oversized holes (and again conservatively neglecting friction) it is not possible for all bolts to be initially active. Even offer all bolts become active some bolts will be resisting much higher forces than others. This is a well recognized fact in any bearing connection. What is essential for a bearing connection is that it be able to reach its design ultimate capacity. It is not important that all bolts be stressed to the some level.
In the design of a connection oversized holes would never be specified in a connection constructed from brittle material or from materials which exhibit non-ductile behavior.
Connections must be made of materials which exhibit relatively ductile behavior so that shear force redistribution con occur among the bolts in the pattern.
For a bearing connection a relationship exists among the size of the hole oversize,.the ultimate shear displacement of the bolts, the stiffness chorocteristics of the bolts, the percentage of bolts not initially in bearing and the desired baseplate safety factor. This relationship is derived below.
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l Comanche Peak ASLB Hearings ,
( , Response to CASE Questions Question No.: Doyle #16 '
Page 3 i l
1 Consider a baseplate with N total bolts of the some diameter and embedment. X of the N bolts are in immediate bearing with the baseplate. Therefore, N-X bolts are not in immediate bearing and are all (conservatively) assumed to have o maximum gap of oo (the hole oversize). Thus (N-X) bolts will log the response of the X bolts by a displacement of Let Po= Total Design Shear Load on Boseplate SF = Baseplate Sofety Factor Desired Pu= Ultimate Boseplate Load = (SF)P o Fu= Ultimate Bolt Shear Force F
FD = Allowable Bolt Shear Force = [ per Design Criterio The octual bolt shear force-displacement curves con be closely approximated by a bilinear force - displacement curve such as the one shown below.
C F
u LKo To s
I Tangent Stiffness I
I
---oo = Hole Oversize Secant Stiffness b
b
- Ou U1timate Disp 1acement Now Pu= XFu+ (N-X)(Fu - Kya,) (1)
Po= NFD = NFu5 (2)
' P =u (SF)PD = (SF)NF /u5 (3) 27 5
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Comanche Peak ASLB Hearings .
(. Response to CASE Questions Question No.: Doyle #16 ,
Page 5 This is a sufficiently high safety factor for a baseplate. It con be seen that the 1/8" oversize hole only reduced the overall baseplate factor of safety below the bolt factor of safety by 4%.
The " conventional method" is the basis for both hand analysis and computerized onelysis of baseplates to determine the relative distribution of shear forces within a bolt group.
The " conventional method" is o design tool, it is not a rigorous nonlinear analytical technique. Where used for connection design with sufficiently ductile materials it guarantees that the required ultimate shear capacity of the baseplate will be reached.
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s Enclosuro D16-1 Sheet 1 of 2
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ABBOT A.HANKh5
- - asteaus'nsoases j 1115 lNDI ANA WTR E ET. P. O. sOK rrass File No. 412189 61 SAN FR ANCISCO. CA 94107 * , .
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(4 t s) 2st.aeco
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January 30,i1974 ..
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. HILTI - FANT7NING SYSTDtS. 'INCg
- *e # " 7 *"**. ' ' i'. *.*I
- i 360 Fairftold Avenue. -
Stassford. Connecticut 06904 Sun.7ECT: . lot!L'.,ot.T'TtisTINC Ha M - thAD hS. DISPtaCD eWT CRAPHS At your request. iwe have conducted a cergrehensive' piossram of testing of .
. the seven their performace differentcharacteristics diameters of 941k*-Bolts in 2.000.- 4,000 .(1/4" throussh and A0002 pai 1/4".) to determine concrete.
- The
- results obtained ifrom this program are as. noted on the settached graphs. '
' Anchors,.dri1G and drili bits were furnishad by H1LT1 strom 'regutar production runs and are considered to be indLcative of that osatoriar.1 notwelly used for installations of this type.
- Hanks Cencrete was supplied by a local batch plant and p!sced' sesser Abbot A.tdots reinforced slambs w supervasion by a general contractor. used 11tnesstone aggrwgatm in teatinc. The concrete mix(3/4" the test slabs for mm.nimm) and Type 11. comment. '!he concrete accordance with ASUt C.33 was pinced in typica2 c9nstnaction manne r and fini shed **dith 4 000 aand t>ull-float.
6 000 poi.
Test s 2 sbs6 were designe.. for '2R de.y strengths of ::,000 standard a x I-2 Ench cylida from Cosupre each slab s s ive s trengths we re ve rt fied f romin scenrdance with ASTM C.31 and tes.ted in accordance with pronared l [ ASTN C-39.
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- suns 11e and shoor testina w us
- crforri rd us1Er '/or a hollow-e.
tensile core testing tyydreulte the testing j ack equiplied with a calibrated pr. essure 1. tree legN g auge. ibuted the equiprwnt was supported 24" af.byt aa ete ' circ 2 = .e I'. Janchors reaction 3/4" tripo.1 s:- w+.sch ' . :er or distr less ,and .
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Comanche Peak ASLB Hearings s C- Response to CASE Questions ,
Question No.: Doyle #17 N ,
l Exhibit No.: None - s
% g i
1.0 CASE Question Hos Cygno verified the statement: "No 2-inch topping"?
Y e This offects the calculation for Hiltis relative to embedment, since a non-monolithi,c shear plane has been established.
2.0 Cygno Interpretation ?
Three support drawings within Cygna'sI scope of review contain a note regarding the 2" topping., These are:
e RH-I-010-002-S225, Rev. S' -
e RH-1-02/4 -011-S22A, Rev. I e, SI-I-038-Of 3-S22A, Rev. 2 On the'first stwo drawings, tbn note states "No 2-inch topping". On the other drawing, C 2 inches of topping is specified, What credit was take for this t$pping sin the calculation of minimum expansion anchor embedment? ,
,s 3.0 Response To verify the adequac'y'of expansion anchor embedment leng- Comanche Peak began with the full length of the anchor and then subtracted items : os the plate thickness, thread length, grouting and topping. Therefore, in calculating minimum embedment length, no credit was taken for the strength of the topping.
2- ,
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11lll1111!I111111111111111111-I i
-- 1
Comanche Peak ASLB Hearings Response to CASE Questions .
{
Question No.: Doyle #18 Exhibit No.: 898,899 1.0 CASE Question baseplate onalysis was performed without including stiffeners alters' the stiffness matrix of the baseplate and consequently the distribution of moments and tension to the bolts. Beyond this point, stiffeners remain unqualified. Hos Cygno addressed this?
2.0 Cygno Interpretation Did Cygno consider the bolt loading for the baseplate in the stif fened condition? Also, did Cygno qualify the stiffeners?
3.0 Response it is a conservative approach to ignore the ef fects of stif feners on plate or bolt design.
When stiffeners are added a redistribution of forces to the bolts does occur in the presence of stif feners. More importantly, this redistribution of forces is favorable since it will produce higher bolt forces and therefore o more conservative design for the stif fened baseplate.
It is important to recognize that the most critical elements in the design of the baseplate ore the anchor bolts. The high degree of indeterminoney of the plate portion of the baseplate combined with the significont membrone resistance (in addition to bending) which must develop prior to failure of the plate material, makes the overall failure of the baseplate by failure of the plate material very unlikely. The more likely failure mode for o boseplate results from bolt failure since the bolt system is generally less 4
indeterminate and does not possess the alternate load carrying mechonism that mernbrone action provides for the plate. Recognizing this, baseptote analysts tend to moke assumptions which maximize the tensile forces in the most highly stressed bolt (s).
One such assumption is to neglect the presence of stiffeners.
Stiffeners make o flexible baseplate behave more like a rigid plate. By making the plate more rigid, the internal moment arm, create'd in the plate by the compressive force in the concrete and the tensile (orce in the bolts, becomes a mcximum. Therefore, to resist !
a given applied external moment, the maximum bolt tension will be smaller in a rigid I
(stiffened) plate thoniii a flexible (unstiffened) plate.
f/
es 1111111ll1llllllllllllllll1111
l l ( Comanche Peak ASLB Hearings
\ Response to CASE Questions Question No.: Doyle #18 Page 2 On the other hand, stiffeners have no effect on bolt shear forces. This is because the in-plane stiffness of a baseplate is already very large and the addition of stiffeners do little to increase this already high stiffness. Well proportioned stiffeners (relatively thick and deep with length to depth ratio < 3) are generally not a problem in baseplate design. A simple and conservative stiffener analysis shows stresses well below allowables.
Detailed baseplate calculation = for Sl-1-037-005-532A ond RH-1-024-011-S22A (Attachments Dl8-1 and Dl8-2) ior the stiffened and unstiffened cases support the above observations in a general way. The tables on the next page show that the maximum bolt tensile forces and plate stresses are greater for the cases without stiffeners than they are with stiffeners.
From these tables it con also be observed that for bolts with a Jorger provision ratio, the bolt loading for the unstiffened condition is greater. Bolt provision ratio is defined as follows:
B P ratio = +
{
where:
T = cetual tension TA = oil w ble tension V = actual shear VA = allow ble shear t
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- (Ibs)
Ratio Wi t.' - Without With Bolt Without Stiffeners Stiffeners Stiffeners Stiffeners i,700 0.27 0.25 I i,900 240 0.13 0.15 2 0 460 0.11 0.14 3 60 2,000 0.27 0.25 4 2,260 Table 2 - Support RH-1-024-011-522A (Cose 1)
BoltForce Provision (Ibs)
Ratio With Without With Bolt Without Stiffeners Stiffeners Stiffeners Stiffeners 1,580 0.40 0.43 i 1,170 800 0.35 0.31
-- ,2 1,260 0 0.45 0.46 3 0 770 0.41 0.45 4 240 2,100 0.35 0.23 5 3,660 E 2,710 0.40 0.42 6 2,510 l
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Provision
- Bolt Force Ratio (Ibs)
With Without With Bolt Without Stiffeners Stiffeners Stiffeners Stiffeners 0.31 0.31 0 0 1
0.29 0.29 560 610 2
0.56 0.59 1,610 1,670 3
0.63 0.62 4 3,I40 2,930 0.31 0.23 3,050 2,070 5
0.23 0.28 6 250 870 Toble 4 Maximurn Plate Stress (psi)
Without With Stiffeners Stiffeners Support. Number 9300 6600 SI-l-037-005-532A 8500 3600 RH-1-024-011-522A (Case 1) 9800 3800 RH-t-024-011-522A (Case 2) t D
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- 13.t7e .f.501C i s .E 7e 12.C01 6.3753 12.C01 13.f7e 5Cir 6.3753 * ? .C t 1 E.3753 f.
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L C A D I t. F C K F. A T I C A (FCECFS -LES) (P'0Ffbl 11h-LkS) Y-PCP EhT ZEIST SEC y-$ FEAR Y- S P E A F. 7-FGELE N X-FCPifhT 2 - T C R O L' E
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kFC 7.0 (PRIPE) HEC, C7 FAE 1964 14:u0: 17 bSik: p.C.JCFh5Cb FAGE CLIErl: 1U51 111LE: "10d1" EASEFLATE *NALYSIS CTGh8 EAEEGY SERVICFS CA1E : 1LE, Ct "AE 1St FFCCRAF FFLATE bER$1Ch 2.C 1C1 CALIFCEh1A S1hEET 11:24:15 SAh FRAbCISCC, CALIF. 54111
. !!-1-n37-r05-53?A .._
- PLATE S1RESS SUFFAST .
PIh!FLP Z-CISF=-C.S?3E-C4 (bcCf= 34) VAXIFUP l-EISFr C.16tE-C1 (tCDE 35 FAX ECUIVALEAT $1RE!! FCF E75EFLA1E : C.9aCF 04 Fil (AT ELEFEAT = 35)
A L L C L A E L E E C U I V A L E t.1 STRLS$ FCE EA$rFLATEz 0.2 7f 0F C5 C.929(.E C4 EASEPLATE FECVISICP. PATIC * ----------- =\ C.3443E CC C.27CCE C5
- E C L.1 SUP P./ E Y
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T V ECL1 FR(VISICh RATI('* ----- 4 -----
1A VA khEEE T AAC 1A AGE ACTLAL AhD ALL0kAELF TLhSIOP. F C F C E S /51R E ES E S AP O V AtD VA ARE AC101.L t t. C ALLCbAElf SHEAR f05CE!/51EESFE5 ECL1 F0bCE! At.0 ETEi!!!!
t.U F I E P T lt V VA P F CV 3 51C h kATIO 1
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RFC 1.C (PRIFE) hEt, C7 FA> 1974 14:uC: 17 LEEE: 6.C.JCth5Ch FAGE CLIEkT: TUSI TI1L E : " tub!" EASEFLATE ANALYSIS
( 7. TUSI FLATE hC. 2 FCF A F. 7 0 PEccFaP E PLA TE CYGhA EEErGY SFPv'ILES CAT 5 : TUE , Cd F AR 19tt VERSICA 2.C ifi CALIfCEh!A 51RLE1 11: 15:3(
SAh FRAhCISCO, CALIF. 94111
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............ ? n IhFU1 bALUES IhFL1 EY: .d... ".~..... [All: 3.~. 3 . .".
ChECrEC NY:
p ., tall: 5-M.-Ed ,
PLATI : 3-CIF = 25.44LC Ih. Y-DI =iv.5bOC1h. 1FICKhE55 = 1.(,COC Ihc
\ E = C.2CCE LF F51 khu = C.ICCC ALLOW AELE S T FE SS= C.27CCE C5 F.E. FESF: hEL-X = t hEl-Y = F FEGLLAk PL5'? bC FLEFthi TYFE =PL/1E FLCCF : !TIF F /8P F A/Lhll CISF = 0.c25F Cc POI!!Ch85 RA110 =C.3f ( I r C h r. = C D
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3.219
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=
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GEAEFt1El ELFrFN1 DIPEt!!Ch! (1-DIRECT!Chs ELEPfhl Y- ~
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NLPPFI X-CCORD!t;Alf T-CCC6DthA1C A -G F l l' 4.GEIC h0 PEER IT F(
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(FFIFE ) kEC, C7 FAs 1914 14:uC:17 LSEE: F.C.JCth!CA F AGE i NFC ?.C "1Uhl" E75EFLAlt ahAL1515 CLIFFT: 1053 111LE:
t F 71 1
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= 1 NLFRfR Cf SEC170h!
SEC110h FFFEPEhCE FC1h1 --G F I D F C I h15 - -
SECTICA ----- GftrR/1ED 4ALUF5 -----
X-GFIC 1 -G R I D FIGIC T YTO ALPBER 3-COCECIt.81E 1-CCCCDIhATE 5 1 95 1 11.13 1C.25 5 L!hr SEGFEh15 CF Thl! SEC11Ch ARL LCCATEC............- A1 Eh* J -------------
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y-LOCRL1hA1E Y-CCCRDI AA TE n-CCCRCIhATE t-CCCRCIhATE 13.675 E.5C1C N E.3757 6.5C1C 12.CC1
' P.501C 1.5.P75 13.E75 F. 3757 12.CC1
' 13.f75 12.CC1 6.3757 E.501G
- F.3757 12.CC1 11.12? C.CCCC0
% 11.1?F 5'.5C1C 25.440 1C.251
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- 12. CT 1 11.126 N 11.125 C.uGC.00 1C.251 N e.3757 1C.211 LCAO IhFCRPATICA (FCECES -LFS) (PCFEk1 - I t. - L E S ) ZC151 SEC X-5& EAR ,
1-Ski /E Z.f0FCE *-PChiti. Y-PCP Eh1 2 - 1 C S C L' E C.CCDE is (.2(Al C4 NC. tit t C3 *C.15tf C4 %C.ct2h C4 %C.612E C5 .C.1C3E C5 8:
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FRCGRAF EFLATE 101 CALIFCRh1A fikEET VFESTCP e o . .
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SAh FF AhC IS CC , CALII. 94111 51-1-037-005-53?t........................
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FAX E C L'1 V A t r h 1 STRESS FCF EASEFLATE ' i ALLCbAEL.E EC('IV A L E r.T ST 04 R E S S FCE BASFFt.ATEs C.27
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532 G Rg ' 17.2.39 Q=
Fg,s 414 3 Kg,= 12'J 7 / 2.
8Ast y tc. losd ots AT kTTAckwr.>rt, Loe.$3,ad :-
rou Ow) bcm, (#%h Fx g e n = +123 c)
M(g = /189 ~h2 N y g g e. - \ M U L hg=r 414 3 '
FiL3 4
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GEhEkATEC ELFFEAT DIFEh!!Ch! ( 7 - D I F E C T I C t. )
ELEFEh1 E-huFfER C IP E t.! ! C A 1 3.rco i ? 2.9?*
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3.752 it ,
2.?46 11 2.249 12 2 . C ' e.
1? 2.e,46 , .
14 2 . e. ! 0 15 2.rCC GEhfrATEr ELEFENT DIPft.!!Ch! (1-DIPLC11Ch)
ELEFEh1 Y- 2:
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1- 3.ritV . . .
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LIAP
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- 5l1 a 1 LP
. C .C C t t: th. IILEhe 1.CCCC 16. Fft 4= C.5CCE (t FS) PLCats C.CCC C+
)LtCF BCL15 =
hCRFAL STET 55/LCAts 0.1301 C5 ILLOWAPLt A hcot SifAS F AtteI A
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--GEle FCIh11.--
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, 3.000 3.r00 2
> ? 24 1 1
12.00 3 .C C C' 2 71 1 ,
36.25 3.000 if il 155 1 17.CC 15 Ic .2 5 152 1 17.00 t is i t. 00 ? 16 14e 1 l 3.CCO 17.00 l
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A FEffGihfE PCIN1 --G F ! r F CI AT S --
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1-CCCFDIhATE X-GFIC Y-66]D 3-CCCE0!hATI 4 4 1 99 2 r. . C (, 1C.CC i
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t .2 5CC 1C.CCC
?C v.25CC 23.612 D2 23.112 13.7!0
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13.78.0 10.312 ~~
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CA1E : bEC, 07 FAE 19F4 FFCGR AF EFLATE CYGh8 EhERGY SERVICES 11: 31: 35 VERSICA 2.C 1C1 CALIFCRFIA SThEET SAh FEAhCISCC, CALIF. 54111 F H-1-0 2 4 -011 -E 2 2 A
. FLATF STRESS $UFPAEY
..... ......... 4.........
FINIMLF ?-CISF=-C.??6E-C3 (hCDE: 11) MAXIrur Z-CISF= C .15 5 E-C 1 (h0DE: 166)
FAX ECUIVALENT STR;!S FCE EASEFIATE = 0.F46F G4'FFI (AT ELEFEb1 = 96)
ALLCbAElE ECLIVALt;.1 STRESS FCF BASEFLATE= C.27CCE CS C.F461E 04 N
DASEFLA1E FPCVISICh F. A T I C = ----------- = 0.3114E CC C.27CCE C5
- EOL1 SLMhARY C
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ECLT FRCVISICA F. A T I C = ----- '%r ~
TA khEFE T Aht TA APE ACTLAL Akt A L L 0 k A E1L E TtP510h FCFCES/51RESSES AhD V A r. D VA.APF ACTUAL Aht ALLCLAELE SHEAR F0 E C E S IS T RE 51E S ECL1 FCFCES Aho STrrSSES VA FFCVISICf; FATIC 1A V
hLFEER T C.4C32t OC 1 C.1172E 04 C.13CCE C5 C.2592E 04 0.t3CCE C4 C.13CCE C5 0.2CFFE 04 0.f3CCE C4 C.34F7E CC 2 C.12(4r 04 0.23CCE C4 C.454(E CC 0.GCCCr Or C.13CCE C5 C.3773E C4 Y 3
C.32t1E 04 0.53CCE C4 0.411?E CC 4 C.23628 C.' (.13CCE C5 C.3523E CC C.36(36 C4 f.13CCE C5 C.5F5cE C3 0.63CEF C4 5
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91 92 93 94 95 96 97 98 99 100 101 102 103 109 10@
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-76 77 78 79 80 81 82 83 89 85 86 87 88 89 90' 61 62 63 69 65 66 67 68 69 70 71 72 73 74 75, e 46 97 48 99 50 51 52 53 54 55 56 57 58 59 60
\' 31 32 33 34 35 36 37 38 39 90 91 42 43 49 45 16 17 18 19 20 21 22 23 29 25 26 27 28 29 30 1 2 3 4 5 6 7 8 9 -10 11 12 13 14 15
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131 132 ,133 ,134,,13[ 136, 137 138, 139 190 191 , 192 ,193,,18 l 113 ,119 115 ,
,116 ,117, 118 ,11f 120 121 ,122 ,123 129 125 ,126 ,12(1- !
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101 102 100 109 105 106 107 108 109 110 1 1 e .
81 82 ,
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C i G h' A ' E F E R G Y SERVILES CAli -
kEt, C7 FAR 1914 FFCGRsP EPLA'TE' 1C: 4?: 21 VERSI0h 2.C if1 CALitCRh!# STRLE1 SAh FRANCIECC, C ALI F .,94111 , ,
. RF-1-f.24-C11-522A WlT@ STIFF 6 O E125
. . . . . . . . . .............................o IhFU1 VALCFS 1hFL1 Et: . m . E. ...... IAlt: ........
.. C H E C r E.C. b.Y..: i . fg, ,l C A.T E. : .T..12 -A .4 Y y-DIP = ?c.25LC Ih. 'Y-DIF = ?u.000C 1h.'TFICKNE!5 = 1.5COC Itc FLATT E*= C.29CF (P FSI thU = C.2000 ALLOWAELE STGESir C.27CCE C5 15 hEL-Y
- if REGLLAR FtSb? bC E LEF EN T 1YFE =FLATE F.E. FESF: hEL-X =
FLCCF : 511 F F / A RE A /Lh! T C]SF z C./.25C Co PCliSCh*E RATIO = C.2C (]rChr =01 GEhERATEL ELEFENT DIPLt.EICt.5 ( ?-D I F f CT ICP:)
ILEPEh1 X-hurtIF CIFEhCIch 1 ?. . C f f ? . .
2.9?b
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7 1.t55 . .
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CLIEhl: 105I k
BCLTS : SE1 = 1 CIAF 0.EGCC IN. EFF Y.P. = 0 ?35E Ce PSI TCISF= 0.0CCO Th. EILEh: 1.CCCC IA. EFF b= 0.5fCE C6 FCI FLCAC: C.COC LE ALLCkAELE hCRPAL STRESS /LCAC: 0.13CE ( 5 / LL0k AF L E STEAR SikE55 /LCAC= 0.E30E C NUFSEr CF BCLTS = t ECLT ----- CEhfRA1EC VALUFF -- .-- .-C F I C FCINTS-- hCDE ,__MATEFIA 3-CCCRDIhATE Y-CCCLDIhATE X-GFIt Y-Gh1C kUPPER TYFE NLFPFF 3.CCD 3 .f'C 0 ? ? 1F 1 1
2 18.CC 3.CCC '
2 24 1 36.25 3.000 15 ! ?1 1 3
4 36.?S 17.CC 15 1r 159 1 15.Cr 17.00 - F 1' 152 1 5
3.000
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17.CC ? it 1A6 1 6
NUFBFR CF SECTICh5 = 1 SECTIOh F E F E E E t.' C E F C II. T SECTICF ----- GEFERATEC VALUES ----- --G F I C FCIATS--
Y-COCFCIhA1E X -G F I C Y-GRID EICIt T tP{
hlFBER X-COCEDINATE 1C.0C r 6 1 95 1 20.06 LIhF SECFEhT5 CF 1115 SEC11Ch.ARE LCCATED AT -------------
...________.. gne 3 .....________ _______ ...._ Ehc J X-COCRCIhATE Y-CCCRCIFATE X-LOCRDIhA1E Y-CCCbDIAATE 16.311 6.2500 2u.0(0 6.25CC 20.06C 6.250C 23.t12 4.25CC
(' 23.f12 (.25CC 1C.CLC 22.81?
?S.F12 1C.CCC 13.75G 23.512 23.F12 13.75C ?u.0t0 13.750 20.06C 13.750 1o.311 13.750 .
16.311 13.750 10.311 10.000 16.311 10.00C 1o.311 6.25CC 10.0C C 11.F11 1C.CCC
,N16.311 1C.CCC N 11.111 1C.CCC r.6CrCD N23.t12 1C.CCC 2c.310 1C.CCC N2P.310 1C.CCC 3c.?50 it.CCC
^ \ IF. 31 C 10.CCC 2e.310 PC.CCG
\ 2E.31C 1C.CCC ?o.310 C.CCCCC
\11.t11 1C.CCC 11.611 2C.CCC
- 11.F11 1C.CLC 11.611 C.CCCCC
. LCAD IhFLPFATICti ( F C 5 (. E S -LFE) ( P C F E A .1. - I t.- L F S ) ..
2-TCECUE ZCIST SEC X-SFEAR Y-SMEAR Z-FCFCE. X -F OP. E h i Y-PCFEh1 1N C.533F 04 NC.414E C4 -^
NC'.'? 6 ? E 04 \ C .7 ? 2(5t.\C.13CE Ct so.ZC7E Ce NC.CCOE l
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$Ah IRAhC15CC, CAL]f. 94111
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ALLC.AELF E 0 llI V A l f 6 1 STEESS ICF DASEftATEa t.27COE C5
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C.341tt C4 EASIILATE FELVISICA PA11C = ----------- = C.1325E CC ,
C.27CCL C5 ,
. FCLT SUPhARY *
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? C.6C4FL C3 C .13 Cl. E C5 0.?( <4E 04 0.73CCE C4 C.3CSEE OC 3 C.CCCfr Cr f.13CCE CE C.3797L C4 0.f3CCE C4 C.45745 CC g 4 n.7734F 03 r;.13rCE C5 C.3??7E 04 0.F3CCf C4 C.4543F CC 5 C.?1Ce.t et r .13 C r. E C5 C . 5 5 7 f.E 03 0.>1CCE C4 C.2241E CC 6 C.2711I C4 C.13CCE ,C5 0.1741E 04 0.t 3CCE- C4 0.41f3E CC
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106 107 108 109 110 til 112 113 98 99 100 101 102 103 104 105.
91 92 93 99 95 96 97 83 89 85 86 87 88 89 90 76 77 78 79 80 81 82 H
68 69 70 71 y 72 73 79 75 61 62 63 59 65 66 67 52 53 59 55 56 57 58 59 60 4 46 97 98 49 50 51 36 37 38 39 90 41 , 42 43 94 45 35
( 31 lb 32 17 33 18 39 19 20 21 22 23 24 25 26 27 28 29 30 '
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113 119 115 116 117 118 11'1120 121 122 123 129 125 126 127 1*
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45 96 d41
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SFG ?.C (FEIFE) kEC, C7 FAE 19t4 14:00:17 LSER: F.C.JCthSCb FAGF 1.
IITLE: "1051" EASEFLA1E ANAL 1EIS CLIEhl: TUSI TUSI PLAlf hC.'3 F05 AE-1-71A
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CYChe EkrRGY SEPVILES CATE : kEC, C7 FAR 1914 FFCGFAP E'PLA1E 12:C?: 11 vrkSIch 2.C 1(1 CALIF 0ENIA S T F.L E T SAh F F. A h C I S C C, CALIF. 94111 , _ _ , ,
. Ft-1-024-C11-522A L.IO ST I P F E t.Wil.c.,
IhFU1 tituEE IhFt1 EY: .b . .....q.. LA1E: D.~.D
.,. Calf: 4.-12.-kd C H E C r.E Cff IsY ;}lgt.
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PLATE : )-CIF = 3t.25CC Ih. Y-D F = 2u.CCOC 16.
E = C.2COF f .* FSI t h l' = C.3000 ALLCWALLE hc ETFE5!= C.2?CLE C5 E LE"F h T TYFE =FLATE F.E. FESF: hEL-v = 15 hEL-Y = 1C PEGLLAR VsSH?
FtC06 : STIF F /t R F A/LhIl CISF = C.(25E Co POIS50h*S RATIO = C.3C (IKONK =C)
GEhEEATEL ELEVEAT DIFEASIch5 (y-DIF.ECTICh>
FLEFEAT X-AUFEER DIPEh5ICA 1 3.CC0 .
I 2.936
? 2.C3t 4 2.935
- 2.25"
- 2.2iC 7 1.t>!
.' 2.Ct1 4 ?.75?
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11 2.249 12 2.64f 17 . 2.d46 14 2.t50 15 2.rro GEhEFL1EL ELLFEN1 DIPfh51Ch! (T-DIFEC12 Chi LLEFEh1 Y- ~
u ht*P!rF DIFEhSICh 1 . 3.f f(1- .___
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FIC i .C (F R I P E ) kEL, C7 FA6 1964 14:00: 17 USER: k . (: . J C t h ! C h PAGE 14 CLIEti: 1t15 2 lITLE: "Tuh!" EtSEFLAlf %ALYSIS
( tcli!
IllSt =
- SF1 = 1 C.CCCr Ih. EFLEh=
CIAF = 0.LCCC IN.
1.00C0 Ih. EFF b=
EFF Y.P. = C.235E CC FSI C.5CCE CL FSI PLCAC: 0.000 tLi AlttbAFLF ACC'FAL 51FE55/tCAC: (. 13CE C5 ILL0b/DLE 5FEAF STRES! /LCAC= C.E30E C5 t-j hlFIth 0F ECL15 :
FCL1 ----- GEhEPAlEP VALUE5 ----- --GFIC FCIh15-- ACDE MATEFIAE hlPPEF X-CCCFDIAATE 1-CCChDINATE y-GFIC t-CkID kUFRER TYFE
~
1 3.CCC 3.f CC 2 1E 1 2 1E.CC 3.CCO F i 24 1 3 36.25 3.r(c 15 2 31 1 4 3/. 25 17.00 15 1 '- 159 1 5 11t .00 17.( C P 1~ 152 1
& 3.CCC 17.CC i il 146 1 htFEFR CF SECTIchs = 1 SECTIOh FEFEFEhCE FCIh1 SECTIOk ----- C E t. E u A T E t; bALUE! ----- --G F I t. F C I h l .5 - -
hlF0fD X-C00ECIhtlE Y-CCCECIhATE ' X -G F I D Y-Gkit FIGIC TYPL 1 20.06 1C.00 c. e 3 99 LIhE SEGFEb1: CF TFIS SECTICb ikt'LOCATEC AT -
............. Egg y ............. ............. the J X-COCECIIATE Y-COCFCIhATE ) -L O C F D I f. A l f Y-CCCp cIh A TE 16.312 6.25CC 2u.060 6.25CO 2C.Cer o.25CC 2s.512 (. 25CO 23.E12 6.25CC 23.512 10.CCC
{. 23.512 1C.CCC 25.Pi? 13.750 23.E12 13.750 20.0(0 13.759 20.(cf 13.750 10.312 13.75D 1c.312 13.75C 1o.312 1C.CCC 1/. . ? 12 1C.frC 10.312 (.25CC LCAD Ibfr4PATICh (FCE(E! -LFE) (PCFFb1 -IA-LES)
SEC x-56FAR Y-564 err Z-F0FCE X-FChEh1 Y-PC P Eh1 2-TCFCLE ZC151 N
1 T.53?I r4 sr.414r C4 NC.2c2E C4 Nr.722L r:5 DC .1?CL Cr C.2C7E Cc 0.CCCf T t
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C LIE F T : TUS) 111LE: "TUd!" Et5FFLAlE ANALYSIS 5
( FECGFAP EFLATE CYGhA EhEE6Y SEEVICES DA1E : kEC, C7 F AE 14t 6 VERSICh 2.C 101 CALIFCRh1A 51kEET 12:03: 11 S A P. F E A h C I C. C C , CALIF. 94111
- F F 0 2 4 -01.1 -! ? ? / .
- FLA1E STEFSS SUPPARY
- FINIFlF Z-C]SF=-0.3(bF-(3 (hCCF: 5) PAXIPUP Z-t1SF: C.225E-C1 (hCDEr 171 PAX E C L'1 V A L E t 1 51PESS FC' EASEFLA1E : C.90GE 0 4 FSI (si ELEPEh1 = 100)
ALLC6.a6LF EGL'IVALEh1 516t55 FCR EASEPLATE: f.27CCE ! CS C.9797E T4 B A S E r- L A T E FFCVIS10h RATIC = ----------- : C.3(2VE CC \
C.27C(E C5
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- ECLT SU'FhARY
- c; ..........................
1 i ECLT F F C V I S I r t. PATIL = -- .-- + -----
it VA kbEFE T Ahc T A AEE ACTLAL AhD ALLCkAILE ltNSIOP FCFCFS/ STRESSES Ahc V t r. D VA AFF ACTUAt Aht ILLC'.ABLF SrEAR F0 E C ES IS1 EES SE S EOL1 FCFCES Ai,C ETFISS!S F U P r F F. T lt t VF PRCVISIch RATIC 1 0.CCCrf CC C.13CrE C5 C.25SfE C4 0.73CCE C4 f.3131E OC 2 0.555/L 03 C.13CCE C5 C.2(tPE 0* O.63CCE C4 C.7945[ CC 3 C.1614r C4 C.13CCE C5 C.3773E 04 C.F3CCE C4 C.57FFt CC 4 C.313Ft 04 r.13(Cr C5 0.32e1C C4 0.53CCE C4 C.634?E CC \
5 C.304PI 04 C.13CCE C5 C.5t!6i (3 0.'3CCf C4 C.205rt CC 6 0.24E5t C3 (.13CCE f i5 C.173BE 04 0.t3CCF C4 C.22f5E CC
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Comanche Peak ASLB Hearings
( Response to CASE Questions Question No.: Walsh #1 Exhibit No.: None 1.0 CASE Question Appendix E of Cygna Report. Section DC-2.2.4. What was the yield point used for A500.,
Grade B tube steel?
2.0 Cygno Interpretation During the course of the pipe support design effort at Comanche Peak, on ASME Code Case was issued (N-71-10) which reduced the allowable yield stress from that stated im the Code Case being employed in the design (N-71-9). What code case was used in the design?
3.0 Response Comanche Peak typically used a yield strength equal to 42 ksi os required by ASME Code J Case N-71-9. The value for yield strength based on ASME Code Case N-71-10 is 36 ksE.
Cygna's original audit accepted calculations based on ASME Code Case N-71-9. Cygrwa later checked the calculations within the review scope to verify that the tube steel
~
design stresses did not exceed 36 ksi set forth in Code Case N-71-10. In each case, tive existing design met the 36 ksi allowable. (See Attachment Wi-l for list of supportts checked.)
The ASME has since provided a response to Texas Utilities' inquiry into the need to adoprt the lower yield strength values. A copy of this letter is provided in Attachment Wi-2.
Part of the response states that "...the provisions of later revisions to Code Cases are neither mandatory or retroactive." Further, based on the ASME review and notificationes, and as stated in the letter, the change from 42 ksi to 36 ksi is not considered a sofer:y Concern.
S I
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ATTACHMENT Wi-l (Page 1 of 1)
List of Supports Reviewed for Tube Steel Allowable SI-I-075-002-S22K RH-1 -064-008-S22K RH-1-010-004-522S RH-l-010-002-S22K RH-t -064-010-522R SI-1-075-001-S22R RH-t -064-007-S22R SI-t -075-003-S22R RH-1 -064-011-S22R SI-t -325-001-S32R SI-l-042-002-S22K SI-l-073-700-532R RH-1-008-007-S22R RH-l-064-001-S22R RH-t -010-001 -S22R RH-l-064-009-S22R
- SI-l-325-002-S32R SI-1-037-005-S32A
{- 51-1-070-007-S22A RH-t -024-Ol l-S22 A 1:
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iho Arnarican Socioty of Mochanical Engincors
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RECEIVED ' (Page 1 of 2) se semess
- Merritt Norman ATTACHMENT Wl-2
(- November 18, 1983 NOV 2 31983 Hall Johnson McBay poppiewer Texas U!!!!!iet Services, Inc.
CPSES Cont. Office Calder Cecamer Deem Kissinger
. Texas Utilities Services Inc. kJ W .* Fmneran PO Box 1002 Glen Rose, TX 76043 Stebaugh ! Murray 9, .,
, Attn: M. R. McBay Hk:xs
Subject:
Section III, Division 1 Gentry R. Bak.tr Code Case N-71-9 & N-71-10 ASTM A-500 Tubular Shapes .e .
Reference:
Your letter of October 25, 1983 ASF2 File i NI 83-101 Gentlemen:
Our understanding of the quections in your inquiry, and our replies are as follows:
Ouestion 1: An owner has contracted for construction of
( component supports under the provisions of Case N-71-9.
Must component supports constructed from ASTM A-500 tubular shapes under the provisions of Case N-71-9 be redesigned or re-analyzed using the lower yield(e.g strength
. , N-71-10)valuesfor published the same in a later revision of the Case material?
Reply'l: No, the provisions of later revisions to Code Cases are neither mandatory or retroactive.
Question 2: Why were the yield strength values for A-500 tubular shapes published in Case 1644-3 through N-71-9 re-duced in N-71-10?
Reply 2: The Committee recognized that the yield strength of A-500 in the cold wrought condition may be slightly re-The revised ~
duced in the heat df fected zone of weldments.
values, given.4n N-71-10, for A-500 were those used dition.
. values for A-500 tubular shapes in the welded is conThe r data for the welded condition l as required by theThe Code, higher presented to the Committee for consideration.
i ASME procedures provide for reconsideration of this interpretation when or if additional in available which the inquirer believes rmght affect the interpretation. fr
( Interpretation may appear to the cognizant ASME commmee or subcommittee. As stated l code documents ASME does not" approve,""certty,*"reie," or enoorse" anyitem, constructso l device or actmty.
6
\
NI 83-101 TGxas Utilition Ssrvices Inc. '
PO Box 1002 Glen Rose, TX 76043 ATTACHMENT W1-2 (Cont.)(Page 2 of 2)
( Attn: M. R. McBay ,
Page 2 of 2 .
of the many safety factors and desien constraints applied the yield strength in the design of piping supports.
Question 3: If a component support is ordered under a Design Specification which required compliance with an Edition and Addenda of the Code which was issued prior to final approval of Case N-71-10, and the contract date for the support is af ter the date of Council approval of Case N-71-10, does the Code allow the construction of the support under the pro-visions of Case N-71-9?
Reply 3: Yes,. in accordance with NA/NCA-114 0.
We note that when, in the opinion of the Commit-tee, a review of current code provisions indicate a potential safety con-cern there are established means of notifying c:rganizations and individuals who may be affected. These men:ns include notification through Mechanical Engineering mag: azine and letters to holders of Certificates of Authoriza-tion and jurisdictional and regulatory authorities. Theese measures were determined not to be necessary in the case of the yield strength values for A-500 tubular shapes in Cas.e 1644-3 through N-71-9.
Yours truly, . .
s
/
/Xevin BPVC Ennis Assistant
,4 Secretary (212) 705-7643^
XE/dp .
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Comanche Peak ASLB Hearings
(-
Response to CASE Questions r, Questjon No.: Walsh f/2 '
, Exhibit No.: None .
, r
, j l
1.0 CASE /
Question .
Observation Record PS-02-01. The opplicant did not consider' shear cone interaction of adjocent bolts.
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2.0 Cypo interpretation
- /
Cygno Observation PS-02-01 was written to evoluote on apparent discrepancy between drawing information and calculations, as related to anchor bolt embedment lengths. Was she~or cone interaction also addressed?
3.0 Response 'f f
Yes. Observation PS-02-01 identifies a concern witb the calculation of bolt embedment lengths. Investigation revealed that the embedmentWas provided to the constructor as a function of total bolt length which is specifiej'on the drawing. In addition, the greater
( of the two embedments derived from either thi consfruction specification or the drawing governs.
Although not related to this concern, Cygno did ch'eck both th analyses and construction to ensure that bolt spacing requirements viere met. Minimum bolt spacing criterio are necessary to assure f"Il development of bolt copocity as specified by the manufacturer.
Maximum bolt capacity is realized when the concrete shear cone is fully developed i
without, interferences. Interaction or overlopping between odjacent bolt shear cones will f
reduce bolt capacity as a functioq.of bolt dicmeters. The applicant properly considered these effects as stated in the'Hilti Menufacturers catalog (see Attachment W2-1).
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i ATTACllMENT W2-1 (Page 1 of 1) l N ! bYI KWIK-BOLT TECHNICAL INFORMATION l
- 1. Anchor Spacing The minimum anchor spacing and edge distance for 100% effective anchor performance ac-cording to EAMI (Expansion Anchor Manufacturers institute) are as follows:
Minimum Anchor Spacing = 10 hole diameters Minimum Edge Distance = 5 hole diameters According to EAMI, anchor efficiency is reduced on a straight-line basis down to 50% at 5 diameters center to-center anchor spacing.
- 2. Minimum Embedmont .
The minimum embedment for satisfactory anchor performance is 4% bolt diameters (6% bolt bolt diameters for the Super Kwik Bolt). Deeper embedmonts will yieki higher tension and shear capacity as indicated in the TR 111: " Kwik-Bolt Testing Program _ Embedment depths indicated in all test reports are before setting (tightening).
- 3. Maximum Working Loads The maximum working loads should not exceed % of the average ultiersate values for a spe-cific anchor size. Actual factor of safety to be used depends on the appiscation and should be selected by the designer on this basis.
- 4. Combined Loading Combined loading should calculated on a straight litie interaction diagram of pure shear (S) and pure tension (T).
S applied + T applied S allowable Tallowable
- 5. Stan'dard Kwik Bolt Materials
- a. Stud (bolt material is AISI 11L41 for bolt diameters %"-%" and At:St 1144 for diameters
%"-1%", meeting the chemical requirements for ASTM specification # 108.
- b. The two independent expansion wedges are made from AISI 1050 spring steel,
- d. Washers are fabricated from'S' AE standard material in accordance with ASA standard
- B27.2-1949. ..-
- e. Kwik Bolts are plated in accordance with the raquirements of Federma Specification 00 Z-325C, Type 11. Class 3, (clear chromate treatmunt),
- f. The Kwik Bolt meets the dimensional reqJirements of Federal Specification FF S 325, Group II, Type 4. Class 1.
3
Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Walsh #3 Exhibit No.: None 1.0 CASE Question PI-01-10. There has been no detailed computer analysis performed to consider the concentrated loads (volves, etc.) and their effect on dead weight and seismic. Also, the seismic analysis will not be linearly proportional.
2.0 C' ygno Interpretation Observation PI-01-01 states:
s "The wall thickness used for the computer analysis piping segments 16"-SI-074-151R-2 and 16"-SI-073-151R-2 'wos 0.5 inches. The correct value is 0.375 inches."
To evoluote the impact of this error in wall thickness, Cygno increased the pipe stresses by the linear proportion.0.5/0.375. Please address the following:
- o. The effect on thermal, pressure and deadweight . stresses as the pipe wall
(
( thickness decreases.
- b. The effect on seismic stresses, which are not linearly proportional to the change in wall thickness.
3.0 Response Jl,
- c. Figure W3-1 con be used to illustrate the effect on thermal stresses due to o local decrease in pipe wa'il thickness.
Pipe A s
Pipe B Pipe C
. . . wr7 1717 3
Figure W3-1 s y .g
~ Asume'that the thickness of Pipe A' reduces from 0.5"(to ) to 0.375"(tg). As shown
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Comanche Peak ASLB Hearings
- Response to CASE Questions Question No.: Walsh #3 Page 2 below, the axial thermal stresses developed within Pipe A cre unchanged as the wall thickness decreases:
6, = thermal stress = E a T (I) where E = modulus of elasticity a = coefficient of thermal expansion oT = temperature change The oxial thermal force in Pipe A actually decreases cas the wall thickness Acreases, since F, = thermal force = 6,A (2) where A = pipe crea = n Dt (3)
{-, D = pipe diameter t = wall thickness Any reduction in the oxio! force within Pipe A will ofsc) reduce the moments induced at the connection to Pipes B and C. So, the thermal ement in P!pe A will decrease os the wall thickness decreases. Since the bendin, rength of Pipe A is also decreasing along with the wall thickness, the net effe on thermal bending stresses depends upon the piping configuration and is not prec..ctable. However, the upper bound change in thermal bending stresses is ot /tg, the wolue used by Cygno.
Pressure stresses in the piping are also a linear function of wall thickness:
&, = circumferential (hoop) stress = n pD/LAt' (4) c=
2 longitudinal stress = n pD/2 t (5) where p = internal pressure
- Dead weight stresses due to the pipe itself are unaffectesi os its wall thickness decreases. This is shown below for o simple beam:
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Comanche Peak ASLB Hearings f'
\ Response to CASE Ouestions Question No.: Walsh #3 Page 3 uniform load, W w ///////A A A
i length, L Figure W3-2 The maximum deadweight bending stress (oD) in Figure W3-2 is:
2 g _ WL O (7)
D- 81 where W = AP= nDt P (8)
( F I
=
=
volume weight of steel moment of inertio = ndt/64 (9)
Inserting equations (3) and (9), shows that the wall thickness drops out of equation (7):
2 a (nDtP)LD ,
8 PL2 (10)
O= 3 D
8 n D t/64 Equation (7) shows that deadweight stresses due to other dead loods, such as insulation, would increase os the wall thickness decreases. This is because only the moment of inertio would changes.
In summary, pipe stresse,s induced by thermal, pressure and dead weight loadings are related as follows to d decrease in pipe wall thickness.
6 C
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Comanche Peak ASLB Hearings .
Response to CASE Questions Question No.: Walsh #3 .
Page 4 o Thermal Thermal stresses developed within the thinner pipe section are unchanged. Stresses induced by the thermal grczwth of ottoched piping will increase stresses by on amount linearly proportional to the pipe wall thickness:
- (temp) a f (!!)
o Pressure Pressure induced stresses will increase as the wcall thickness decreases..
The increase is linearly proportional:
a(pressure)af (12) o Deadweight Pipe stresses induced by pipe deadweight are unachonged by a change in wall thickness.
- (pipe deadweight) - unchanged (13) a(other deodweight) a f (14)
Deadweight stresses due to foods other than the; pipe itself will increase
, os the pipe wall thickness decreases.
Based on the above, the simplified procedure employed v Cygno to evoluote thermal, pressure and deodweight effects related to Ot cervation PI-01-01 is.
reasonable and in fact conservative.
- b. Figure W3-2 will also be used of illu' strate the effect of e decrease in pipe wall thickness-on seismic induced stresses.
For o simply supported pipe loaded by a uniform weight,, the fundamental pipe
- frequency is:
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/ Comanche Peak ASLB Hearings
\ Response to CASE Questions Question No.: Walsh //3 Page 5 f=h 2L E
(15) where i
f = fundamental frequency L = span length E = modulus of elasticity I = moment of inertio (Equation (9))
g = gravity W = weight per unit length Only two terms, W ond I, depend upon the wall thickness, therefore the frequency change due to o thickness change con be expressed as follows:
N I of = (f o -f)= I g2(W +Wo Ws +Wgj (16) s where of = frequency change fo = frequency associated with thickness t o fg = frequency associated with thickness tg Ws = t t I weight - unloaded weight
- Wo = unloaded weight for t o Wi = unloaded weight for ti l
o
= moment of inertio for t o li = moment of inertio for tg Substituting the equations for W and I, Equation (16) becomes:
g 3 (nD t )/64 (nD3 t )/64 Af * -
(II) 7 ( W, [(n,o t,y) 2 W, +( nDt gy))
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Comanche Peok ASLB Hearings
. Response to CASE Questions Question No.: Walsh I/3 Page 6 The following conclusions con be reached from Equation (17h o When "other" loads (W3 ) are zero, Equation (17) reduces to:
=0 (I8) of = constant x -
Therefore, the acceleration and stresses weill be unchanged (see Equation (13)).
o When W is greater than zero, its influence is srnali. Per Brown & Root s
drawing, BRHL-SI-t-RB-061, Rev. O, pipe sesgment S1-1-073 has the following properties:
D = 16 in.
L = 14.5 f t. = 147 in.(I)
Using these properties, Equation (17) reduces ten
{
/ 0.5_ j 0.375 ) (g 9)
Af = 1091 (YW3 +7 YW3 + 5 j; where E = 29,000,000 psi
- g = 386 in/sec2 1' = (490/1728) Ibs/cu. in.
to = 0.5 in.
tg = 0.375 in.
E Notes (I) The distance from the containment flued head to support S1-1-073-700-
. S32R is 14'-6 3/8".
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( Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Walsh #3 Page 7 Toble W3-1 lists the results of a sensitivity analysis performed using Equation (19). It shows that the maximum frequency change inthe simple model of line S1-1-073 is one hertz for all values of Ws . For the soke of comparison, the weight of water in a 16-inch diameter pipe is
. 7.3 lbs/in.
Table W3-1 f, hertz W3 , Ibs/in 0.46 2 0.99 4 1.25 6 1.39 8 1.46 10 1.49 12 1.51 14
(. l.51 1.50 16 I8 1.48 20 0.94 100 0.32 1,000 0.10 10,000 The small changes in frequency shown above have negligible effect on pipe stresses.
Therefore, the simplified procedure employed by Cygno to evoluote dynamic stress effects related to Observation PI-01-01 is conservative. The octual effect on pipe stresses will be less than the ratioot /ti .
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( Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Walsh //4 Exhibit No.: None ;
i 1.0 CASE Ouestion PI-02. Is there an error in the table shown?
2.0 Cygna Interpretation Referring to Observation PI-02-03, Attachment A, is there on error in the calculated table?
3.0 Response There is a typographical error in the calculated table. The allowable for restraint RH-1-064-007-S22R should be "44000", rather than "4400".
As shown on the attached Table, Attachment W4-1, this correction puts the allowable for the aforementioned restraint into line with the other restraints tabulated.
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[4 m'TJ("", """" ATTACHMENT W4-1
< rage 1 or 1 3 Record Review
( unununuintununnu Attachment A
""*"N'- 0 Checkflet No. PI-02 s h..: j of y Observation No.PI-02-03 Yes No Valid Observation X Closed X Comments 1.0 Root Cause Possible misunderstanding of the Gibbs and Hill procedure 2.0 Resolution Using the range for the 3 rigid restraints, Cygna calculated the following:
Load CYLNCZ General Support Range Stress Stress Total Allow 2700 10362 6763 17125 45 SI-1-032-003-532R 400 44000 RH-1-064-007-522R 1300 5172 5128 10300 8615 11225 9328 20555 44000 RH-1-016-001-532R The remaining 4 restraints are springs or snubbers and have no thermal load. Thus, there is no increase in stress above allowables.
Cygna also noted that the correct method was used for the welded attachments in anchors of Problem 1-70 and in all supports in Problem 1-69. Based on this, Cygna considers the eeror isolated. In addition, the RHR system will probably show the largest percentage difference (between maximtsn :oad and range), since it has many modes of operation. Thus, Cygna expects the error would have the most impact on this system. As the new calculations show, the impact on design is negligible and
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Comanche Peak ASLB Hearings
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Response to CASE Questions Question No.: Walsh (/5 Exhibit No.: 889 1.0 CASE Question CTS-00-03. Fbx = should be 21.2, not 23.2 or 22. The length is 6' not 5.5'.
e Why was only I /2 SSE considered?
e Why was 4% domping used; not consistent with FSAR?
e Assumed cable troy was rigid when lumping the moss; this resulted in not combining the dynamic effects of the cable troy itself to the support; did not include ef fect on welds.
e The validity that the cable trays have the copocity to transfer o load around a corner when one run of cable troy has no oxial restroint, as shown on drawing 2323 El-0601-01.
e What documentation did Cygno see that justified the hangers' receiving a lateral load around corners that resist the axial load from the troy segment that contains no oxial restraints.
.. 2.0 Cygno Interpretatien .
~
In Observation CTS-00-03, Cygno discusses several apparent deficiencies in,the modeling ossumptions associated with the frame analyses for cable iroys. As related to CASE Exhibits 889,890 and 902, please address the following:
Exhibit 889 Why was on allowable bending stress (Fbx) = 22 ksi used?
Exhib'it 890
- c. Why was only I/2 SSE considered?
- b. Why was 4% domping used?
- c. How were the dynamic effects of the trays included in the analysis?
n
- d. . On Drawing 2323-El-0601-01, there oppears to be no means for transferring load around the cable tray bend. Please discuss this.
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Comanche Peak ASLB Hearings
{ Response to CASE Questions
. -Question No.: Walsh #5 Page 2
- e. What documentation formed the basis for accepting the condition mentioned in item (d)?
Exhibit 902 How was baseplate flexibility considered?
3.0 Response
. Exhibit 889 The Gibbs & Hill calculation to determine the allowable bending stress for the channel section was performed in accordance with the guidelines set forth in the AISC Monval, Equation 1.5-7. This equation provides a method for calculating Fb nd also states that F shall not exceed 0.6 F perkquotion 1.5-7 to be 2k.i ksi, compared thatlue vd.toThe designer first calculated 0.6 F , and then selecte F[Ihe lesser value.
Section
. l.5.1.3.4 of the AISC Manual specifies that Fbx f r 36 ksi steel equals 22 ksi.
A direct calculation of 0.6 F yfor 36 ksi material would of course produce a value for Fbx equal to 21.6 ksi, rather than 22 ksi. As illustrated in AISC Section 1.5.1.3.4, this 1.8%
- dif ference is not considered significant. 22 ksi was used in the design.
If 6'-0" is used in equation 1.5-7 rather than 5'-6", as properly chosen by the designer, Fbx would equal 21.2 ksi. S'-6" is correct based on the definitions provided in the AISC code
- where
f, " = distance between cross-sections broced against twist or lateral displacements of the compression flange."
As shown on Attachment W5-1, this dimension is the clear span. Resistance to twist or !
laterol displacement is supplied by the welded connection to the vertical members.
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l Comanche Peak ASLB Hearings -
(s Response to CASE Questions ;
Ouestion No.: Walsh #5 l J
Page 3 l Exhibit 890
- c. Gibbs & Hill calculation SCS-10le, set 5, derives the opplicable load combinations and shows that, for seismic loadings, the 1/2 SSE (OBE) condition controls-Attachment W5-2 summarizes how that conclusion was reached. Since the supports were designed to OBE loods, the members were checked against the normal allowables with no increase for seismic loads. Inherent in this normalization is the fact that normal strength allowables may be increased for SSE loadings. Since anchor bolt allowables remain constant (i.e., no increase) for SSE loadings, unlike structural members, Cygno questioned the acceptability of this design approach.
The attached calculation (W5-3) was performed by Cygno to evoluote this situation. Gibbs & Hill had also evoluoted this concern in 1979 and arrived at a similar conclusion.
- b. USNRC Regulatory Guide 1,61 specifies that bolted structures, such as this, should be evoluoted using 4% of critical domping. Although some connections in the cable troy support system are welded, Cygno concurs with the designer's selection of 4%
domping, rather than the 2% domping value specified in R.G.1,61 for welded
(._- structures. The designer's selection is appropriate for the following reasons:
o The lower domping value for all welded structures recognizes that such a structure will dissipate less energy than structures with mechanicot connections. In the cose of the cable trays, there are many significant mechanisms for dissipating energy, e.g., the cables are loosely
' connected to the. trays, the trays are connected mechanically to the structural frames, and the frames are bolted to the concrete.
o Various papers on cable troy behavior illustrate that cable troy systems exhibit domping values greater than 4%. Attachment W5-4 is one such paper (see page 181),
- c. Gibbs & Hill designed the coble troy syste'm for peak spectral accelerations. Since 100% of the tray weight _was included and peak occelerations were employed, any
, influences due to troy flexibility have been conservatively incorporated.
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Comanche Peak ASLB Hearings
( Response to CASE Questions Question No.: Walsh //5 Page 4
- d. As shown in Attachment W5-5, the troy system in question is adequately supported. An oxial restroint is provided near each bend. The schematic in Attachment W5-5 is taken directly from Drawing 2323-El 0601-1 (CASE Exhibit 957).
- e. Gibbs & Hill calculation SCS-Il3c, set 3, addresses the longitudinal restraints discussed in item (d).
Exhibit 902 The calculation in question concerns the analysis of the two-bolt baseptote of a Detail "E" which was installed on a riser and is utilized as a three-way restroint.
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( Comanche Peak ASLB Hearings Response to CASE Questions Ovestion No.: Walsh //5 Page 5 Gibbs & Hill's analysis considered a rigid baseplate which was analyzed to resist rotations about the Y and Z oxis. Gibbs & Hill's analysis showed that compression against the concrete provides sufficient resistance in conjunction with the tension in the anchor
. bolt s.
A subsequent analyses by Cygno, using the baseplate !! program of CDC, verified the Gibbs & Hill results. ,
Gibbs & Hill Reanalysis Calculation:
SCS-146C, Set 8, Sheets 6S-69 (also Tech. File 11.2.l.S0, Sheets 15/81-19/81)
Cygno Baseplate Analysis:
Calculation Binder 83090/1-F, Section A Computer Binder 83090/l.1-F, Sections A, B ond C Interaction Ratios:
Gibbs & Hill Revised.Colculation = .584 Cygno Baseplate Analysis = .464 1
I 15
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(' (Page 1 of 1)
The loading combinations are:
(1) Operating Condition: S = D+L+F EQO (2) Safe Shutdown Condition: 1.6S = D+L+F EQS The earthquake loads are:
EQUIVALENT STATIC LOADINGS (G's)
Earthquake Intensity Seismic Direction SSE 1/2 SSE*
Horz. 4.0 G 2.67 G Vert. 2.5 G 1.67 G
!
- Numerically equal to 2/3 SSE values And the sign convention for vertical loads is:
Positive: in the gravitational direction (down)
' Negative: opposite gravitational direction (up)
Now, by substituting in equation (1) above the Operating Condition may be calculated as follows:
(Horz) (a) S = 2.67 (D + L)
(Vert) (b) S = (D + L) + 1.67 (D + L) = 2.67 (D + L) (down)
-0.67 (0 + L) (up)
And by substituting in equation (2) above, the Safe Shutdown Condition may be calculated as follows:
(Horz) (c) S = (D + L) = 2.5 (D + L)
(Vert) (d) S = 1*6 I
- 6 (D + L)
=
2.19 (D + L) (down)
=
-0.94 (D + L) (up)
Then,by comparison, the governing load cases are:
(Horz) Equation (a) S =
2.67 (D + L)
(Vert) Equation (b) S =
2.67 (D + L) 1 (down)
Equation (d) S =
-0.94 (D + L) = 1.0 (0 + L) (up) 1111111!!ll111111lll1lllllllll
ATTACHMENT WS-3
( (Page 1 of 6)
Gibbs & Hill considered two loading cases in the design of their cable tray systems:
Normal + Severe (called "0BE"): S = 0 + L + OBE Normal + Extreme (called "SSE): 1.6 S = D + L + SSE where, D = Dead Load L = Live Load OBE, SSE = Loads due to that Earthquake By normalizing the equations with regard to S, the governing load case was determined to be "0BE" for which the members were checked against the normal allowables with no increase for seismic loads. Pages 3 through 6 show clearly that the ratio of "SSE" to "0BE" is always less than 1.6, so all members and welds are acceptable.
For anchor bolts, Gibbs & Hill checked "0BE" loads against Hilti bolt allowable loads based on a minimum factor of safety of 4. As the loads increased to SSE levels, the bolt allowables, using IEB 79-02 as a guide, remain constant at a safety factor of 4. Therefore, the Hilti bolts may not meet a safety factor of 4 under "SSE" loading.
In response to Cygna's question, Gibbs & Hill stated that the factor of safety will not fall below 3 and quoted NRC Document MS 129-4 on the acceptability of
{. a safety factor of 3.
Cygna's Approach To accept this, Cygna must show that the increase in loads does not reduce the safety factor for "SSE" below 3.
Load Increases The attached tables show the effective "0BE" and "SSE" G-levels for all buildings. The G-levels are determined from ARS peak values and combined in the fashion on Gibbs & Hill's position calculation.
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( ATTACHMENT W5-3 (continued)
(Page 2 of 6)
Effective G Values Elevation OBE (G)* SSE (G)* SSE/0BE Reactor Internal Structure ,
905.75 5.447 6.799 1.25 885.50 4.704 5.882 1.25 860.00 3.790 4.772 1.26 832.50 2.864 3.681 1.29 808.00 2.372 3.108 1.31 783.58 2.251 2.932 1.30 Safeguards Building 896.5 4.560 5.948 1.30 873.5 4.365 5.790 1.33 852.5 3.698 4.956 1.34 831.5 3.072 4.158 1.35 810.5 2.603 3.698 1.42 790.5 2.212 3.056 1.38 785.5 2.158 2.967 1.37 1
773.5 2.056 2.790 1.36 Electrical Building 873.33 3.855 4.944 1.28 854.33 3.578 4.606 1.29 380.00 2.988 3.893 1.30 807.00 2.620 3.638 1.39 j 778.00 2.452 3.385 1.38 Auxiliary Building 899.50 5.132 6.446 1.26 886.50 4.664 5.948 1.28
< 873.50 4.255 5.501 1.29 852.50 3.864 5.003 1.29 831.50 3.339 -
4.451 1.33
- 810.50 2.788 3.731 1.34 790.50 S 2.535 3.560 1.40
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l ATTACHMENT WS-3 (continued)
(Page 3 of 6)
{ Effective G Values (continued) f Elevation OBE (G)* SSE (G)* SSEJOBE
- Fuel Building 918.00 4.397 6.058 1.38' 899.50 4.033 5.695 1.41 860.00 2.866 3.980 1.39 841.00 2.630 3.646 1.39 825.00 2.455 3.367 1.37 810.50 2.271 3.099 1.36 Interaction Diagram n () *( *I Equation (2)
( )+( )=1 [m ation (1)
,c l Definitions: x X = Tensile Load Y = Shear Load i Tg = T/FS, where T = Tensile Ultimate and FS = Factor of Safety V
A
= V/FS, where Y = Shear Ultimate x, = "0BE" tensile load Y = "0BE" shea,* load o
l A =
LoadIncreaseFactor(""SSE") 0BE" i
The exponential curve is based on the plots of shear tension loadings found in i the Teledyne response to IEB;;79-02. The document MS 129-4 is only a guide and does state that "0BE" safety f actor should be 5. We can. however, argue that a safety factor of 4 for the bolts is adequate based on IES 79-02.
y - -
ATTACHMENT WS-3 (continued)
( (Page 4 of 6)
I. Using Equation (1) - Linear Relationship:
To use this relationship to reach a factor of safety = 3, we must determine what the allowable load increase is above "0BE" loads.
- a. Assume pure tension, with the "0BE" load just meeting the criteria aX T
= 1.0 X, = 7 T
3 T
A.7 = 1.0 T
T
. . a = 4/3 = 1.33 The same result will be true for pure shear.
- b. For intermediate values of tension and shear ratios assume that the increase in the tensile and shear loads (in going from "0BE" to "SSE") are equal.
( Assume x=.7,5 o
1 Y,=.25h (4).75h+(a).25f "I T X 7 3 a .75 (3/4) + a .25 (3/4) = 1 a = 4/3
. . A load increase of 1.33 is allowed for the linear interaction equation oder the range of values for X, and Y . As can be seen,Jrom"EffectiveGValueTables,"some,buEnotall, areas of the plant would meet this criteria.
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l' ATTACHMENT WS-3 (continued)
(~ (Page 5 of 6)
(
II. , Equation (2) - Exponential Relationship Using the relationship from the Teledyne paper, calculate the allowable load increase, A.
- a. At the endpoints, the allowable load increase is 1.33 because the linear and exponential curves are coincident here.
i b. At an intermediate value an X,=.75h Y,=.25f a .75 h 5/3
+
a.25{5/3 "I T V 7 7 (a (3/4)-( .75) )S/3 + (a (.25) ( 3/4))S/3 , y 44 AS/3 , 1 AS/3 = 2.25
(~-- a = (2.25).6 = 1.63 > 1.42, so there are values of X, and oY which will give a safety factor of 4 in "08E" and 3 in "SSE"
. *. We must determine for what values of the tensile and sheer ratio that a = 1.42. For tensile and shear ratios between these values, the safety factor of 3 will be met for "SSE" loads.
- c. Assume a linear "0BE" relationship such that 1=RT+Ry where RT
- percent tension allowable using a safety factor equal to 4 Ry = percent shear allowable Ry =1PRT -
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ATTACHMENT WS-3 (continued)
(Page 6 of 6)
\
Substituting into the exponential relationship above:
[1.42R T ) 5/3 [1.42 (1-RT) --) S/3
+ = 1.01 N Y Y )
/ / = 1.0
(.75)(1.42)R T + (.75)(1.42)(1-RT)
R T 5/3 + (1-RT )S/3 = 1/(1.07)6/3 = .900 RT 6/ + (1-RT)6/ .900 = 0 Solving numerically en an HP-150:
R .93 RT = .07 T
Ry = .07 Ry = .93 Therefore, for "0BE" loads within the above range of ratios the safety factor of 4 is mat using a linear relationship and, for the maximum increase of 1.42 to "SSE" loads, the safety factor of 3 is C met using the Teledyne interaction method.
Based on Cygna's review of 43% of the cable trays, all shear / tension ratios fall within the above range, so there is no safety impact.
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ATTACHMENT W5-4 (Page 1 of 11) f SEISMIC TESTING 0F ELECTRIC CABLE SUPPORT SYSTEMS
- 6- by Paul Koss Bechtel. Power Corporation i Los Angeles Power Division I
- INTRODUCTION i i
- i Over the past two decades, many earthquakes have occurred within the United States. Of these, several were of sufficient magnitude to cause structura.1 damage to industrial facilities. Following such strong earthquakes, inspection of power generation and distribution facilities has offered valuable inform-scion as to the overall performance of engineered structures. The 1971 San Fernando earthquake has been of particular interest in this regard. It tems one of the most severe earthquakes Southern California has experienced in recent history. A survey of structural damage to the Sylmar Converter Station.
located within a few miles of the epicenter, provided data relative to the '
l behavior of electrical distribution equipment and electrical raceway sya eas j
- when excited by strong ground motion. Of special interest was the fact that simple unbraced raceway hanger systems were able to survive the earthquake i without major structural failures. Another finding was that even at locations l where a minor amount of structural distressing occurred, the cables within
the tray systems did not lose their functional integrity. The fact that the ,
converter station's unbraced support system survived the San Ternando earth-j quake generated interest regarding the practicality of using similar systems in nuclear power plants.
J In the years following the San Fernando earthquake, an increasing emphasis has been put into the design of earthquake resistant structures. This has been particularly true of structural systems in nuclear power plants. As early as 1971, design standards were developed in the industry that out1.1med l methodologies for the seismic design of raceway supports. In addition,
- USNRC regulatory guides and standard review plans were also being developed i
during the same period of time. Designs based upon these criteria have tended to require substantial amounts of bracing. By contrast, the Sylmar Statism support systems were essentially uebraced. Consequently, it appeared that l
l either the design methods or the design criteria, or possibly both, were unnecessarily conservative.
< 1 In order to bridge thetgap between design procedures and observed behavies-of these systems, a plan was initiated to test electrical raceway systems.
The goal of the" testing was to establish the best possible approach to create
, an economical, yet adequate, support system for electrical cabling withis j
- nuclear plants. By the first part of 1977, a clearly defined program that outlined the types and sizes of raceway systems that would be tested was l established. This Cable Tray and Conduit Raceway Test Program was initiated and managed by the Los Angeles Power Division of Bechtel Power Corporation. I
- The testing was conducted, and related consulting services were provided by ANCO Engineering, Inc., Santa Monica California. In the last months of 1977 l testing began. Full scale installatione of both cable tray and conduit raceway systems were tested. By the end of 1978, over 200 individual dynamic tests j had been' performed, generating over 50 volumes of raw data.
-174-
... .- D TTACHMENT WS-4 (Cont.) (page 2 of 11) F# -
DEVELOPMENT OF THE TEST FACILITY _ i G
(, , A typical raceway system consists of cable These supports take the form trays of a and conduits Ni wh 5
ical tiers. The various from overhead by threaded rod or strut. ugh (see figure 1) .
suspended trapeze and may support several trays in vertcoco p
1 N .
TYPICAL TROUGH TYPE TRAT
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TYPICAL LADDER
- TYPE TRAY
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Trough and Ladder Tray Figure 1.
-175-
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! ATTACHMENT WS-4 (Cont.) (Page 3 of 11)
Typically these suspended systems may extend vertically in excess of 10 feet, may be very long and may weigh up to 250 pounds per foot of length (multi-tier '
1 1
. (- systems). In view of these unusual characteristics, it was decided to design and construct a special test table capable of input to long suspended systems.
l ANCO engineers undertook the design and construction of a shake table capable of random and steady state input to raceway systems.
!~
The shake table was designed as an open steel frame, consisting of two
! parallel trusses interconnected by cross trusses and diagonal diaphragm i bracing at the top (see figure 2). The bracing was sized to prevent resonance i below 20 Hz. The frame is supported by five linkages which form an inverted 4 pendulum. The angle of the linkage determines the relative amounts of vertical j and horizontal table motion. The table can develop input either parallel or j perpendicular to its length. The vertical component will act simultaneously
- and be a scaler of the horizontal, depending upon the angle of the linkages.
1 In addition, the table can be rigidly fixed so that forces or displacements may be applied directly to the test specimen.
j The table was designed to accommodate a test setup of 40'-0" in length with j five vertical tiers. This required a clear height of 14'-0". The total.
- estimated weight of the heaviest test setup was 10,000 lbs. This weight was used to design a servo actuator system capable of achieving a maximum input j to the fully loaded five tier system of 1.13 The servo actuators are driven by high pressure hydraulic fluid stored in an i accumulator and released through control valves whose setting can be varied l in proportion to any arbitrary time varying electronic signal. Output from l (.
j the test was recorded in the form of time histories on strip charts or tape, spectral plots from a real time analyzer and as response spectra from a l digital computer.
j TEST PROGRAM SCOPE l
j The first step in defining the scope of the testing program was the identi-
! fication of possible significant variables in raceway system design. The l following'are the potentially significant variables that were identified in Pl anning the test program scope:
l l e Tray and conduit types j e Tray and conduit loading i
4 e Hanger types i
e Hanger length s Connection det ils
'e Number of trays ,
l .e Number of conduits l
e Conduit sizes e Conduit clamps l
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- In order to evaluate the effects of these various parameters, more than I \ 200 test setups were tested. These fell into three categories:
h (1) cable tray systems,'(*.)' ccnduit systems, and (3) combined tra) and conduit systems. Vfchin'ecch of chose three categories of testing :ast setups were developed to evaluate damping, frequency, and other significant
'c.hsracteristics for varying support types, connection details, and bracing.
In aJdition to dynamic testing with the s, hake table, a series of cyclie
- /Iacigue tests were perforved on connection details. The purpose of these
, ' tests was to determine the resilience of these connections and establish a fatigue criterion for use in design.
TEST SEQUENCE AND SEISHIC INFITT A typical test sequence consisted of up to ten individual tests. Initially a test setup would be, subjected to snapback tests (with the table fixed rigidly). These tests were used to determine resonant frequencies and anode shapes. Next, a series of increasing amplitude sinusoidal tests were performed to establish a reference relationship between desping and ampli-fication ratio at various output points. Finally, a series of simulated earthquake inputs were applied. These tests were used to determine bow seismic input amplitude affects frequency and damping.
The earthquake time . history used to formulate the majority of shake table input motions was a syntherb time history. This record was selected due to its conformance with USNR Regulatory Guide 1.60. In addition, a group of
- .four historical earthquake records was used during a limited group of tests.
- s. - 1Never, the actual input motion to the shake table was not the input ansstion
, ' corresponding to any one of the recoids mentioned. Rather, a modificardon to each record was made to account for effects of building amplificatiosa for i the purpose of creating a " worst case" shake table input notic.n.
In addition to the synthetic time history, historical recordings of actual earthquakes were used to drive the systems. The following four earthqssakes were used: i
- 1. San Fernando 2/9/71, Hollywood. Storage F.E. Lot, Comp N90E.
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- 2. San Francisco 6/22/$7J Go10en Gate Park, Comp N10E. '
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- 3. Kern County 7/21/52, Taft Lincoln Schok Tunnel, comp F21E. .
El Centro 6/18/40 Imperial Valley Irril gat';on fa District, Comp 500E.
4.
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. s' The process for selecting earthquakes was based upon the inspet: tion of l f .pproximately teihistorical recordings. Typically, each earthquake had three recorded components: two horizontal and one vertical. z The goal of the
[r selection process was to pick a nominal number of recordings that displayed p
different characteristics.
The'ayntheti,cr earthquake was selected because of its codforminee with Regulatory Guide 1.60. The response spectra shape s'as created to agrear with USNRC guidelines.
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i d seismic evenes The San Fernando earthquake is one of the f itsbent significance.
documente ever recorded and was selected prinarily because o [
selected because of t
Most of the activity The San Francisco earthquake, d one ofThethe f the shaking. shortest, wa response was over within the first two or three secon s odistinct peaks at 4.0 h and cter-spectra, depicting acceleration, shows two veryOf the earthquake 7 0 hertz.
1stics. d upon its frequency8.0 The Kern County (Taf t) earthquake was selected ike baseThere existl characteristics.
In addition, the Kern County earthquake has a predominant sp hertz.
around 3.0 hertz. historical significance The El Centro earthquake was selected based upon its in the field of earthquake engineering. ,
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TEST RESULTS_ tisfactorily at '
In general, rod supported raceway systems did not perform saOverall c input levels in excess of 0.5g.
in excess of 0.75g. d all testing without The strut supported systems that were tested i s surviveThe This damage was type of loss of function. l stic deformation
[- consisted mostly of fracturing of strut type litude loading. angle ficc l (i.e., two h that occurs at connections during large amp ver did
)
fittings that were used to attach the hanger ding to the more than one fitting of the four fracture during an amplitude test.
Most of the systems were tested at input levels corre to 1.0 to 3.0g's maximum acceleration. 0.75g f ree-field l to be equivalent to ground motion levelsNor of was0.25 toNever there any in the acceleration.
struc'tural collapse of a strut-supported raceway occur. d Specific h
loss of function in the electrical circuits results of the tests are described in the following paragrap that were s. monitor Damping ities associated During the cable tray test program,These two weret distinct nonlinear (1) inelasticity ble vibration.
with tray system dynamics were observed.
joints and (2) amplitude dependent frictionalide range due losses of ampli-to ca Despite these nonline'arities, observed responses over a wfrequencies d tudes indicated distinct vibrational modes whose ficant number of be cycles of with substantial changes in amplitude loading. d and a signiConseq l
accounted damping.
for by selecting an appropriate amplitu eThe dam
,{. greater the trays.
than bolted steel structures due to the motion oT
-179- *
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ATTACHMENT WS-4 (Cont.) (Page 7 of 11) x d were these losses.
That is, the greater the input level the more by measurement pronounce Equating these losses to an equivalent viscous dampingculted in pred techniques _ developed during the test program reA typical example of test viscous damping of up to 50% An some cases.Af ter tabulating the results of the seve results is shown in figure 3.hundred dearthquake as shown in figure 5.
type vibration tes) a conservative lower bound curve representing equ an value ac a function of input floor spectrum ZPA was plotte h ,
This curve was plotted at two standard deviations below t e me l 1 each amplitude.
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- e. I 0.16 0.18 0.20 0.22 0.24 0.26 02 030
. 1 0.14 0.02 0J4 0.06 0.08 0.10 0.12 INPUT ACCELERATION (s) t Figure 3. .Typicnl Test Results from Conduit and Tray T(.sts
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1.0 1.1 1.2 T1.3 0.5 0.6 0.7 0.8 0.9 0.1 0.2 0.3 0.4 INPUT (g) BRACEO HANGERS Tigure 4. Damping vs. Input Level for Braced Hanger Systems 24 50% TO FULLY LOADED T AY 20 3 16 /
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Figure 5. Reconsnended Damping for the Design of Raceway Systems
-181-3
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ATTACHMENT W5-4 (Contu) (Pnge 9 of 11)
It should be noted that system damping varies from the above values wesen l cable trays are lightly loaded. Specifically an unloaded tray will have an l
- associated lower bound damping value of about 7%, which is more or less consistent with recommended values for bolted structures (see USNRC Reg.
l Cuide 1.61). For a 24-inch tray, the damping will be in accordance with figure 5 (50% to fully loaded) as long as the tray has more than 20 lb/f t of cable.
The level of damping observed in supports that carry only conduit remains was generally about one half that observed in equivalent tray systems (see figure 3). Damping for such systems should be assumed at 7% of critical for input amplitudes in excess of 0.2g.
t The overall behavior of combined systems does generally trace the behavior of its components; however, not all the specific characteristics of esunduit carry through to the combination. The damping ratio of the segregatedconduit system is on the order of 7% of critical. When this same conduit is added to
, the combined system, the 'overall system damping is equivalent to the damping of the cable tray system (i.e., 20% of critical).
Frequency The testing of trapeze supports that are made from strut, and use predominantly strut type bolted fittings, demcastrated that these systems have fundamental frequencies falling between 2 and 5 hertz. The addition of heavy bracing, or the substitution of structural shapes in lieu of strut, or the attaching of
[_ support.s directly to walls or columns, or the use of many welded connection
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details, will increase the frequencies somewhat. However, it is highly unlikely that the fundamental frequency of a raceway supported in any combination of the above methods would ever be above 10 hertz.
The dependency noted in the rate-of increase of damping with respect to input has also been observed in cable tray system frequency characteristics. Gener-ally speaking, resonant frequencies were found to be dependent upon the level
! of tray response. Typically, the frequency might be expected to decrease by 30 percent as input levels increase from 0.05g to 0.50g.
Connections stiffness is. a major factor in determining the stiffness of a hanger system. The connections are either located where the hanger is attached to overhead supporting members or at the various joints within the hasager i itself. Strut type connections do not act as a pure pin, nor do they maintain l
! infinite rigidity. For partially braced or totally unbraced hanger systems, i j the moment-carrying capabilities of strut connections creates a mechosaism l l through which initial loads may be distributed to flexible supports. The modeling of strut;j:onnections with rotational springs is a prerequisite to correct prediction of frequency -characteristics, stress distribution., and deflection.
j The quantity and size distribution of electrical cables that fill cable trays I vary from tray to tray within a power plant. These variables were studied-
'to assess their effects upon tray frequency. The testing demonstrated that type of size of cables do not influence overall system stiffness. The mass of the cable is the only factor that need be considered in computing cable tray system dynamic responses.
-182-
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ATTACHMENT W5-4 (Cont.) (Page10 of 11)
I Fatigue Strength of Connections
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- In addition to the dynamic testing of support systems on the shake tab @ye th tests.
connection details were subjected to cyclic f atigue and streng linear Lg The purpose of these tests was to determine the extent to which noncl behavior of standard hot rolledThe primary interest was to establish low cycle fatigueQ@
of support systems. In general, for less than 250 stress reversals, i these the elastic g information. 3 connectors were capable of displacements of three to four t mesa typical cyclicSte limit, which was defined by a static strength test.The correlation between the elast N result is shown in figure 6. i f the fatigue curve was 0 the static strength test and the horizontal lim t oThese results indicated thatg a i
generally quite good. ratio for earthquake loadings was three to four. j
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l l l l l 1 1000 500 4 100 CYCLESTO CONNECTION FAILURE Figure 6. Typical Fatigue Curve for Hot Rolled A36 Clip Angle Ccznnectors CONCLUSION 3 f data
-The cable tray and raceway test program This in developed turn resultedainsubstantial some amount from over 2000 individual dynamic tests. Among these was the p spe,cific reconsnendations regarding design practice. ilience j equivalent viscous damping in excess of 20% and the significant resOf parti of hot rolled clip angles under low cycle fatigue. be expected is.the general conclusion that lightly braced raceway systems camsloss of function i to survive severe earthquakes (in excess of 0.5g) with no
[.
the circuits they support.
-183-l
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ATTACHMENT WS-4 (Cont. IP_ age 11 of 11-) - - - - - - - - li ACKNOWLEDGEMENTS _
)
The Cable Tray and Conduit Raceway Test Program has been supported financially by contributions from the Arizona Nuclear Power Project participants. Bechtel
- Power Corporation, Boston Edison Co., Georgia Power Co., Mississippi Power and Light Co., Pennsylvania Power and Light Co., Philadelphia Electric Co.,
Public Service Electrica and Gas Co., Puget Sound Power and Light Company, Standardized Nuclear Unit Power Plant System (SNUPPS) participants, Southern California Edison Co., and the Tennessee Valley Authority.
1 REFERENCES (1) Cable Tray and Conduit Raceway Test Program Test Report, ANCO Engineers and Bechtel Power Corporation, December 1978.
(2) USNRC Regulatory Guide No.1.61, " Damping Values for Seismic Analysis for Nuclear Power Plants."
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DETAIL "H* W/ BRAC.E 3 3 - WAY _
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ISOMETRIC VlE VV OF CABLE OrTAlt "F., 2. - W A.Y TRAY 4 SUPPORTS FROM i
DR AWI NG 2 323 - E l - O601 - 1 k' ,A DETAIL-" E" J-WAy REF: CA.SE. EXHIBIT 890 w
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Comanche Peak ASLB Hearings
(. Response to CASE Ouestions Question No.: Walsh #6 Exhibit No.: 907 1.0 CASE Ouestion CTS-00-05. In the description, it discusses a channel bent about its weok oxis. The resolution does not consider this problem nor does the document CASE requested on discovery; see CASE Exhibit 907. On CMC 88306, are the originator and opprover the some person?
2.0 Cygno Interpretation Please discuss the following:
- a. How did the resolution to Observation CTS-00-05 address the channel bent about its weak oxis?
- b. Are the signatures on CMC 88306 satisfactory?
3.0 Response
( ' o. The purpose of Observation CTS-00-05 was to investigate the baseplate. This is illustrated by the following reprint from the Observation:
"1.0 Description The anchor bolts, baseplate / angle and channel of contilever support Detail "E" were originally designed as two-way restroints to resist oxial loads on the channel and moments about its major axis. In order to use Detail "E" on a cable troy riser, where it must act as a three-way restraint, the channel section was modified to resist moments about its weak exis. The ability of this l
! configuration to function as intended, i.e., to also resist moments about the weak oxis, could not be guoronteed since the anchor bolts and the baseplate / angle were not evoluoted for such a load."
The channel was correctly analyzed by Gibbs & Hill in Calculation SCS-146C, sets 4 and 8. .,
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Comanche Peak ASLB Hearings '
(..- Response to CASE Questions Question No.: Walsh #6 .
Page 2
- b. CMC-88306, Rev. 4, was originated and opproved by the some person. This is acceptable for the following reasons:
. There is a controlled list of people authorized to approve CMC's for construc-tion prior to design review. In the cose of CMC 88306, the approver was on .
that authorized list.
e Project procedures do not prohibit someone on the authorized opproval list from also being on originator.
. The subject CMC is on interim release for construction purposes. Each CMC receives a subsequent design review by the original desigra organization in accordance with Gibbs & Hill Procedure DC-7.
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Curanche Peak ASLB Hearings
\ Recense tc, CASE Questions Ques:Sn No.: Walsh #7 ExhiL1 No.: None 1.0 CASE Question CTS-00 006 What is the "significant design margin" as shown in the resolution?
2.0 Cygno Interpretotion Observation CTS-00-06 states that "... further analyses by Gibbs & Hill (see Cygno Technical File i1.2.1.50, pp. 31-69), incorporating Cygna's comments, revealed that sufficient design margin existed to compensate for the increased stress levels." The
" increased stress levels" refer to the potential increase in stress levels due to the items noted in the observation.
Please quantify the design margins.
3.0 Response To demonstrate the adequacy of a judgement made in their qualification of standard
- details A, B, C, and D by similarity to standard detail D;, Gibbs & Hill performed on
( analysis using the NASTRAN code. For the purposes of this analysis, the C6 x 8.2 section was oriented to match details A, B, C, and D. The results of this analysis are contained in calculation SCS-10f4 C, Set #1, where it is shown that the member interaction ratio for the C6 x 8.2 section is 0.94 (maximum). This ratio is based on on analysis using troy weights of 35 lb/ft 2 and which included troy support self-weight excitation. The "significant margins" are due to the fact that the interaction alone was G% below allowable and the troy loads were assumed to be 22% larger than the octual loads.
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l Comanche Peak ASLB Hearings -
{. Response to CASE Questions Question No.: Walsh #8 Exhibit No.: None 1.0 CASE Question 1
CTS-00-07. The analysis that included the beam element did not consider prying action and the flexibility of the baseplate to determine the center of compression.
2.0 Cygno Interpretation N/A.
3.0 Response Gibbs & Hill performed a refined analysis of the frame and baseplate to resolve Observation CTS-00-07. Cygno reviewed the results of this onelysis and judged the frame, baseplate and anchor bolt design to be adequate.
In order to quantify the adequacy of that engineering judgement, relative to the anchor bolt design, Cygno performed on analysis of the frame / baseplate system using fixed boundaries at the honger-to-baseplate connections. The fixed-end loads developed at
{. these boundary points were then applied to a baseplate model. Cygno's program PSDS (Pipe Support Design System) was utilized for the analysis and design check. PSDS includes a standard baseplate /onchor bolt routine that considers rnechanisms, such as prying action and baseplate flexibility.
The results of this analysis show the following design margins:
Tensile Load Shear Load Design interaction Bolt No. (Ibs) (Ibs) Ratio
- I 500 1540 .10 2 4;i0 1830 .75 3 3040 1890 .45 4 2970 ,. 1820 .45 5 4210 1530 .65 -
- Design Interaction Rotio = (tensile load /ollow.)S/3 + (shear load / allow.)S/3 < l.0
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Comanche Peak ASL3 Hearings
(' Response to CASE Questions Ouestion No.: Walsh #8 Page 2 The design interaction ratio equation, using on exponent of 5/3, was originally contained in Revision 0 to Cygna's review criteria for Comanche Peak. In Revisior I of the Comanche Peak review criteria, the exponent was reduced to 1.0 to be consistent with the equation octually used by Gibbs & Hill.
Further justification for the 5/3 exponent is provided in the response to Walsh Ovestion
- 5. It is also important to note that these results contain the following conservatisms:
lumped troy masses, enveloped response spectro, higher than octual tray weighrts (35 psf vs. 28 psf).
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ITD4 DESCRIPTION ATTACHMENT W9-1 STRESS PROBLEM 6 p y__ C 3 3 _ ccq __J 4 3 ( (Page 1 of 1) l _ 6 ff[g /4 _
d SRP A/A REV. COMPL. ADDIT 10NAL TNFORt'ATION
_j . All dimensions and elevations verified //A
- 2. Valve orientations verified (
- 3. Orientation verified for line mounted equip, pg .
BRHL/GHH N/A . REY. I//////////
- 1. Support mark nu ,bers verified /JA
- 2. Support locations verified hB BRH 4FX- /) :5%-oo 7-F43R. REV. '2 I/((((((// .
- 1. Direction of support verified ,/
- 2. Type of support verified /
- 3. General conficuration verified V 4: Clearances {where applicable) -
/
- 5. Location verified and appropriate red hkN
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[63 marks entered on
& GHH SF-Y-F5- 028 REV. /
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'T ~Ud W- y-j ,_g INSPECTION DISPOSITION The above listed documents were reviewed, no -
change.to location or function has occurred Inspected by: MSN M//fev t//,/f3 sherefore inspection is not required .
(As-Built Inspector)
// 2/2Ahr TShBC DATE QAABC DATE
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- Response to CASE Questions l
Question No.: Walsh #10 .
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Exhibit No.: N/A l J
1.0 CASE Question WD 07-02 What document did Cygno see that showed the temperature indicator would l be installed at a later dote?
i 2.0 Cygno Interpretation What was the basis for closing Cygno Observation WD-07-02? What documentation was ,
reviewed?
3.0 Response Based on a conversation with Texas Utilities personnel, Cygno learned that temperature elements are normally installed offer all other work in on area is completed. This is done in order to avoid demoge to the instrument during construction. When Cygno performed the Spent Fuel Pool Cooling System walkdown, pointing activities were still underway.
Further review otso showed that local indicators, such as this one, are not safety-related C- devices.
The key documents reviewed by Cygno relevant to closing Observation WD-07-02 are discussed below: ,
- l. Instrument installation Checklist (Form No. 2-81)
Form 2-81 is required to be completed by Comanche Peak procedure 35-1195-ICP4. In this case, it indicated that the device was not installed and that the "discreponey" was " turned over to Brown & Root completion and TUGCO".
- 2. The G-list was checked to ensure that the device was non-sofety.
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1 Comanche Peak ASLB Hearings
(' Response to CASE Questions .j Guestion No.: Walsh #11 Exhibit No.: None l
1.0 CASE Question Pipe stress checklist, note 3, item a: !
, 1) What is the basis for considering that the ef fects were negligible?.'
- 2) What pipe stress run did Cygna look at, since the inclined load was used in the design of support RH-t-010-003-S22R?
2.0 Cygno interpretation Pipe stress checklist (PI-02), note 3, states the following:
- 3. The following' supports were modeled along the coordinate oxiss rather than inclined. The impact is negligible.
- a. RH-1-010-003-S22R at data point 1253 (8.6 degrees)
- b. SI-I-042-001-S22R at dato point 793 (7.S degrees)
- c. What was the basis for concluding that support RH-1-010-003-S22R was adequate?
- b. What pipe stress run was evoluoted?
3.0 Response
- o. Support RH-l-010-003-522R is a simple restroint, inclined 8.6 degrees from a line drown perpendicular to the pipe. Cygna judged that this small inclined angle would not significantly offect the support design or the piping analysis. ,An important element of this judgement is that the 8.6 degree, os-built alignmemt is only 3.6 degrees beyond the construction tolerance of 5.0 degrees.
In order to verify the adequacy of this judgement, Cygna requestedi that Gibbs &
Hill reonalyze piping segment AB-I-70. For this reonalysis, the pipiing model was revised to include ~the following:
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( Response to CASE Questions Question No.: Walsh #1l Exhibit No.: None j Page 2 o Supports RH-l-010-003-S22R and SI-t-042-001-522R were modeled with skew angles of 8.6 and 7.S degrees, respectively.
o Support RH-l-010-003-S22R was modeled as two trunnions with snubbers located 7 inches from the pipe centerline.
o Support RH-1-064-010-S22R was modeled l'-4" west of the elbow.
The results of this reonalysis are contained in Attachrnent Dil-l (Gibbs & Hill Calculation), D11-2 (computer output without modifications), and D11-3 (Computer output with modifications). These results are summarized below:
Maximum System Stress (psi)
ASME Equation Old New Allowable 8 9,039 9,039 18,480(I) 9 (upset) 21,094 21,103 22,180((2) 9 (emergency) 24,4SI 24,463 3) 33,260(4) 10 22,883 22,883 27,600(S) 1I 27,881 27,881 46,080 Notes:
(1) 1.0 Sh e per ASME B&PV Code, Section til, Porograph NC-36S2.1
. (2) 1.2 Sh e Per ASME B&PV Code, Section lit, Paragraph NC-3652.2 (3) 1.8 Sh , per ASME B&PV Code, Section til, Paragraph NC-3611.3c (4) Sa = f(1.2S S c + 0.2S Sh ) where f = 1.0, for no more than 7,000 thermal cycles, per ASME B&PV Code, Section 111, Paragraph 36S2.30.
(S) So+ S h,-jier ASME B&PV Code, Section lit, Porograph 36S2.3b.
where S[= 18480 for material SA-312, TP 304 at 280 F
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Se= 18800 psi Per ASME B&PV Code, Section ill, Appendix I
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( Comanche Peak ASLB Hearings Response to CASE Questions Question No.: Walsh #11 Exhibit No.: None Page 3 Support Loads (Ibs)
Normal Upset Emergency Old New Old New Old New Allow.*
RH-l-010-003-S22R 1705 1459 3534 4519 3967 5189 15700 105 164 -1724 -2894 -2756 -3565 -15700
- Per NPSI Load Data Capacity Sheet, dated 6/81, for an SRS No.14 strut.
Regarding the following line excerpted from Attachments Dil-2 ano Dil-3, the allowable Equation (9) stress for emergency conditions is 1.8 Sh p>er ASME B&PV
(-_ Code, Section lil, Paragraph NC-3611.3c. The comparison to 1.2 S h ni ADLPIPE is a built-in precaution, not a poss/ fail test.
Stress Summary (Equation 9 Emergency and Foulted Conditio ns)
SEC MEM SEO POS EON 9 Additional in: formation 20 52 896 BEG 13016 20 S 897 END 24451 Equation 9 ex:ceeds 1.2 Sh Nozzle Loads (Ibs)
Old Nev.
Load 3084 254!
Allowable 3 3120 3120 Ratio 0.98 0.81 In summary, the reonalysis showed no change in the pipe stresses, a decrease in
- nozzle loads, and support loads well below the allowable. This verifies the original engineering judgement, b
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( Response to CASE Questions Question No.: Walsh #11 Exhibit No.: None Page 4
- b. Gibbs & Hill pipe stress run AB-I-70, Rev. 0,' was evoluoted by Cygno os noted on several Observations, including PI-00-01.
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