ML17324B205
| ML17324B205 | |
| Person / Time | |
|---|---|
| Site: | Cook  |
| Issue date: | 01/09/1987 |
| From: | Alexich M INDIANA MICHIGAN POWER CO. (FORMERLY INDIANA & MICHIG |
| To: | Harold Denton Office of Nuclear Reactor Regulation |
| References | |
| AEP:NRC:0514R, AEP:NRC:514R, NUDOCS 8701200057 | |
| Download: ML17324B205 (56) | |
Text
REQULATORY FORMATION DISTRIBUTION SY N (RIDS)
ACCESSION NBR: 8701200057 DOC. DATE: 87/01/09 NOTARIZED:
NO DOCKET 0 FACIL: 50-315 Donald C.
Cook Nuclear Power Plant>
Unit i. Indiana 0
05000315 P
50-316 Donald C.
Cook Nuclear Power Planti Unit 2i Indiana 5
05000316
'AUTH. NANE AUTHOR AFFILIATION ALEXICH>N. P.
Indiana fc Nichigan Electric Co.
RECIP. NANE RECIPIENT AFFILIATION DENTONi H. R.
Office of Nuclear Reactor Regulation.
Director (post 851125
SUBJECT:
Forwards load drop analysis of main load block of auxiliary bldg crane. In event that load block should fall from max heighti load block vill strike top of spent fuel pool racks.
Fee paid.
DISTRIBUTION CODE:
A033D COPIES RECEIVED: LTR ENCL SIZE:
TITLE:
OR Submittal:
USI A-36 Control.of Heavg Load Near Spent Fuel-NUREQ-Oh NOTES:
RECIPIENT ID CODE/MANE PWR-A EB NRR SINQH A 01 WIQQINQTON, D PWR-A RSB INTERNAL: ADN/LFl'lB NRR NEIQHBORS09 NRR PW ADTS F
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5 5
1 1
1 1
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1 EXTERNAL: LPDR NSIC 03 06 2
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1 TOTAL NUBBER OF COPIES REQUIRED:
LTTR 25 ENCL '3
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INDIANA8 MICHIGAN ELECTRIC COMPANY P.O. BOX 16631 COLUMBUS, OHIO 43216 January 9,
1987 AEP NRC'0514R Donald CD Cook Nuclear Plant Unit Nos.
1 and 2
Docket Nos.
50-315 and 50-316 License Nos.
DPR-58 and DPR-74 AUXILIARYBUILDING CRANE TRAVEL LOAD BLOCK DROP ANALYSIS Mr. Harold R. Denton, Director Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Uashington, D.C.
20555
References:
- 1) Our Letter AEP:NRC:05140, dated February 14, 1986 2)
NRC Safety Evaluation Report (SER) dated February 27, 1986
- 3) NUREG-0612, "Control of Heavy Loads in Nuclear Power Plants,"
dated July 1980
Dear Mr. Denton:
This letter and its attachments transmit a load drop analysis of the main load block of the auxiliary building crane of the Donald C.
Cook Nuclear Plant.
As noted in Reference 1 and the NRC Safety Evaluation Report (Reference 2),
we were required to complete a load drop analysis for the main load block within one year of the date of that SER.
Attachment 1
contains the load drop analysis that was performed for us by Exxon Nuclear Company (ENC), the suppliers of our current spent fuel pool racks.
Attachment 2 is an independent evaluation of the mechanical analysis portions of the ENC report noted above by our consultant Dr. J.
D. Stevenson of Stevenson
& Associates (S&A).
The criticality and radiological consequences sections of Attachment 1 have been reviewed by AEP personnel.
The attached analysis concludes that in the unlikely event the load block should fall from its maximum height, it will strike the top of the spent fuel pool racks.
For the postulated
- accident, the load block itself will be unable to penetrate the upper grid portion of the racks and the kinetic energy of the block will be absorbed by crushing the upper structure of the fuel rack.
- However, the hook may cause of the rupture of the grid, with a subsequent penetration of the hook to a maximum depth of 29.5" into the active fuel region. If such penetration should occur, a
maximum of four fuel assemblies could be damaged.
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Hr. Harold R. Dento AEP:NRC:0514R The radiological consequences and the potential for criticality as a
result of this accident have been examined.
It was conservatively assumed that this will result in four times the release cited in the Updated FSAR for the consequences of a fuel-handling accident in the auxiliary building.
In this case, the potential two-hour doses that a person would receive at the site boundary could be as high as 7.2 rem to the thyroid and 2.12 rem whole body.
Although this is well below the 10 CFR 100 limits and the restrictions cited in Reference 3 (i.e.,
one quarter of the 10 CFR 100 limits), it requires that radioactive iodine be filtered through the spent fuel pool filter system.
A criticality analysis was performed assuming that the four damaged fuel assemblies were moved to their most reactive configuration.
Analysis of this configuration was performed using a 2 x 2 array of damaged fuel assemblies at the center of each of an infinite number of arrays of 10 x 10 undamaged fuel assemblies of infinite length.
The result of this calculation was a Keff of 0.94, at a one-sided 95% confidence level based on use of the KENO V computer code.
This is below the value of Keff of 0.95, which is suggested as an acceptable limit by Section 2.2 of Reference 3.
Based on the above, we believe that even though the actual weight of the load block is approximately 4.25 tons, it should not be considered a
heavy load for the purpose of compliance with Technical Specification (T/S) 3.9.7, provided the spent fuel pool ventilation is operable and the auxiliary building is under the negative pressure required by T/S 3.9.12.
In the event the above conditions of T/S 3.9.12 cannot be complied with, we will administratively require the main hoist to be deenergized and carry no load on the main hook when the load block is moved over the pool.
This latter condition is 'the same as the current requirement of T/S 3.9.7.
We believe that an analysis of the handling of heavy loads can take credit for the charcoal filters, as noted in Appendix A, Item 1 (4) of Reference 3,
provided that we meet the conditions of T/S 3.9.12 with respect to auxiliary building negative pressure.
A response is requested from the NRC staff by February 28, 1987, in order to ensure continued operation of the Donald C.
Cook Nuclear Plant.
The reason for this is that a footnote has been added to T/S 3.9.7 that expires on that date.
The purpose of that footnote, as stated in Reference 2,
was to allow sufficient time to complete an analysis of the consequences of a postulated drop of the main load block.
We believe that this letter and its attachments fulfillthat requirement.
A check in the amount of $150.00 has been enclosed for NRC processing of this submittal.
l Mr. Harold R. Dento AEP:NRC:0514R This document has been prepared following Corporate procedures which incorporate a reasonable set of controls to insure its accuracy and completeness prior to signature by the undersigned.
Very truly yours, cm P.
lexich Vice President
>(s(~
Attachments cc:
John E. Dolan W.
G. Smith, Jr.
- Bridgman R.
C. Callen G. Bruchmann G. Charnoff NRC Resident Inspector
- Bridgman J.
G. Keppler
- Region III
ATTACKKNT 1 TO AEP:NRC:0514R
EQON NUCLEAR COMPANY, INC.
2101 HORHRAPIDS ROAD. PO BOX 130. RICHLANDY/A99352 509 3~58100 TEI.EX 15 2828 Attachmentl AEP:NRC:0514R
~JAr~ 0 8
~gg<
January 6,
1986 AJM-87-001 Mr. R.
B. Bennett American Electric Power 1 Riverside Plaza P. 0.
Box 16631
- Columbus, Ohio 43216-6631
Dear Mr. Bennett:
Subject:
D.
C.
Cook Spent Fuel Pit Load Dro Analysis This is in response to your request for ENC to perform an analysis of the consequences of dropping the 4.25 ton hook/block assembly into the spent fuel pit from its full height.
It was concluded from the analysis that the accident will not cause a
criticality incident, and that the released radiation dose as a result of such an accident is less than one-fourth of the 10 CFR 100 limits, i.e.,
75 rem thyroid and 6.25 whole body.
Furthermore, the kinetic energy limit of 240,000 in-lbs recommended in NUREG-612 can be applicable to the Cook Plant.
The analysis is included as three appendices to this report.
Appendix A is the mechanical analysis which predicts the extent of the
- damage, Appendix.B shows the results of a confirmatory test, and Appendix C shows the inputs to the criticality analysis.
Mechanical Analysis The sketch on page A-1 illustrates the geometry of the problem.
It is assumed that the hook/block is dropped from a height of 39 feet above the surface of the pool.
It then travels an additional distance of 23.7 feet through the water before impacting the top of the fuel rack.
The velocity at the time of impact with the water surface is 50.1 ft/sec.
Oue to the drag and buoyancy of the water, the velocity of the hook/block increases only slightly as it drops through the water so that the velocity of impact with the top surface of the fuel storage rack is 51.6 ft/sec.
This is shown on page A-5.
The grid of bars near the top of the rack, together with the upper portions of the fuel storage cells, will absorb the impact-ing energy of the hook/block.
The sketch on page A-17 shows the position of the hook/block relative to one of the fuel storage cells when it finally comes to rest.
It can be seen that the hook will penetrate to a point 29.5 inches into the active fuel region.
The block itself will not penetrate the grid, but will crush the top 19 inches of fifteen fuel storage cells.
The sketch on page A-1 shows a top view.
The hook will damage approxi-mately four assemblies seriously, and will cause superficial damage to several surrounding assemblies.
1 I
Mr. R.
B. Bennett January 6,
1986 AJM-87-001 It is estimated on page A-18 that 360 kgs of uranium will be released in the form of pellets and fragments which will fall through the water.'ost of this debris will be trapped at the top of the first undamaged fuel assembly spacer.
The radiation. dose total release will be 7.2 rem thyroid and 2.12 rem whole body, as calculated on page A-18.
These mechanical calculations are based on the principle that the hook/block will continue to move through the structure until the kinetic
- energy, which was available upon impact, is dissipated by gross distortion and crushing of th'e upper structure of the fuel storage rack.
The lower portions of the rack and the fuel assemblies will experience some vibra-tion, but this will be restrained by the spacers and well damped by the water.
Furthermore, the length of time required to bring the hook/block to a stop from the time of impact with the fuel rack is only 124 msec.
Since the fuel assembly will have a lateral period of vibration of 500 to 1000
- msec, the energy will have been absorbed before the lower sections of the fuel assemblies receive any lateral dynamic forces.
The longitudinal resonant frequencies are hi@her, but only the elastic components of the longitudinal waves will be transmitted any appreciable distance.
These elastic waves wi 11 not produce any significant damage.
Criticality Analysis The maximum credible reactivity condition was conservatively modeled.
An infinite arr ay of infinite length rack modules (10x10 bundle array per module) was modeled.
Each module contained 96 undamaged bundles and four damaged bundles.
The four damaged bundles were conservatively assumed to be in a 2x2 array in the center of the module.
The nominal dimensions of the storage racks were used in all cases.
Nominal new fuel dimensions were used for the pellet and clad.
All pellets were 95K Theoretical Density U02 with an enrichment of 4.0X.
The undamaged bundles were modeled at the nominal rod pitch (0.496"), while the damaged bundles wer e modeled with the 0.5272" rod pitch.
Thus, the damaged bundle was expanded to fill the entire storage cell.
This is the most reactive configuration in that the bund'le moderation has been
- improved, the bundle size has been increased, and the water gap between the damaged bundle and its absorber plate has been decreased.
The absorber plate was modeled as B4C with a 8-10 loading of 0. 020 gm per square cm.
This is considerably lower than the minimum certified value of 0.0234 qm per square cm, and is therefore, conservative.
(See reference 6).
The system described was explicitly modeled using KENO-Va and 16 group cross sections with resonance self-shielding corrections by BONAMI.
Replicate calculations using the 27 group ENDF/B-IV cross section library prepared by NITAWL were also performed.
All codes and cross sections are part of the SCALE (reference
- 5) system which has been extensively bench-marked against data from critical experiments.
A listing of the KENO input is provided for details on the model.
A listing of the input to NITAWL is also provided for details of cross section preparation.
I A
Mr. R. B. Bennett January 6,
1986 AJM-87&01 Thc KENO k-eff for this vorst case model ie 0.934 +/- 0.0045 using 16 group cross sections, and 0.935: +/- 0.0040 using 27 group cross sections.
Therefore, there is no evidence of differences due to cross section sets.
Supplemeatary benchmarking of the methods employed werc performed using data from reference 7 ~ hll of the cases selected employed absorber plates betveen bundles; i.c., they are close to the conditions of this analysis.
Thc average aad standard deviation of the calculational biases for the nine cases analysed (16 group) vere 0.0021 and 0.0019, respectively.
Pooling the variances from KENO and the bias determination results in an overall standard deviation of 0'0049.
The k-eff from KENO vas calculated using 103 generations of 400 neutroas..
The one-sided 95% probability Student t vith 100 degrees of freedom is 1.66
'he corrcspoading one-sided 95Z confidence upper limit on the k-eff ie:
k-eff (95'L) ~ 0.934 - 0.0021 + 1.66~0.0049
~ 0.940 Therefore, the k-eff is lese than the limit of 0.95 with 95X confidence.
Very truly yours, A. J. Martenson Mechanical Design Consultant L. D. Gerrald Criticality Safety Specialist AJM:sh Attachments Mechanical Analysis Review
& Approval:
Criticality Analysis Review
& Approval:
Ro Go Hill Senior Engine r J
. Pieper rporate Licensing-Quality Assurance Reviever (Cri ica
)
p Date i~~~z Date Fuel Design Approval:
G. J.
- sselmaa, Maaager Puel Design Date xc:
Ch Brown GJ Busselman LD Gerrald (2)
RG Hill AJ Martenson (2)
JE Pieper RB Stout
Mr. R.
B. Bennett 4
January 6,
1986 AJM-87-001 REFERENCES 1.,
Handbook of Hydraulic Resistance, Coefficients of Local Resistance and of Friction, I.E. Idel 'chik, AEC-tr-6630, Published for the U.S.
Atomic Energy Commission and the National Science Foundation, Washington, D.C. by the Israel Program for Scientific Translations, 1960, Available from the U.S.
Department of Commerce-Clearinghouse for Federal Scientific and Technical Information, Springfield, Va.
22151.
2.
Formulas for Stress and Strain, Raymond J.
Roark, Third Edition, McGraw Hill Book Company, Inc.,
1954.
3.
Source Book on Stainless
- Steels, Compiled by The Periodical Publica-tion Department of the American Society for Metals, Copyright 1976.
4.
Stainless Steel Cold-Formed Structural Design Manual, 1974 Edition, American Iron and Steel Institute, 1000 16th Street, NW, Washington, D.C. 20036.
5.
"SCALE:
A Modular Code System for Performing Standardized Computer Analyses for Licensing Evaluation",
6.
"DC Cook Units 1
8 2, New and Spent Fuel Storage Array Criticality Safety Analyses",
XN-NF-81-47(P),
Rev.
2.
7.
Baldwin, M.N., et al, "Critical Experiments Supporting Close Proximity Water Storage of Power Reactor Fuel", BAW-1484-7.
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True stress. true strain tensile plot shows factors which are important in an analysis of press formability.
I 3
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I <ich occur during testing, while thc latter is pluttcrl using original specimen dimcnsin>>s.
The "true" test is therefore a morc accurat<< i>>di-cation of the'performance of a material dur-ing deformation.
The true stress-true strain tensile properties ivhich are significant in an analysis of press formability (shown in the graph,"
include:
Yield Stress (a) Thc stress at iv)rich a specimen shows deviation from strai ~ht linc proportionality of stress to strain.
~ Stress at lfaximum Load
(~) The
. stress at the highest load (in pounds) sus-tained by the specimen.
~ lfaximum Uniform Strain jr)~faxi-mum value of straining before uniform defor-mation ceases and localized deformation;used necking take place.
This is tltc strain at point of maximum load.
Xfodulus of Work Hardening-Slop<<of the plastic region of the true stress-true strain curve.
~fodulus indicates rate of cold ivork hardening.
~ Deformation IUork (A) Area under true stress-true strain curve to point of m;rxi-mum load.
This is a measure of the work (in inch.pounds) required to elongate thc tensile specimen through the region of uniform strain.
results of any value in a study of press rorm-abilty (see box above).
The Formability Factor Xfaximum uniform strain is the most im-portant factor in press formability.
A stain-less steel blank can be formed as long as every area is deforming uniformly.
As soon as the strain in any section of the steel sur-passes the maximum uniform strain, localized necking will occur at that point, leading to rupture.
Total stain (or elongation) in a ten-sile specimen is unsatisfactory as an indicator of press formability, since only an undeter-mined amount of total "stretch" is uniform.
However, factors other than uniform strain METAL PROGRESS
A-l+
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Standard Stainless and Heat-Resisting Steels Msnhannat Psoerncs ot anuatcd Matcnat
~t Room temperature horn nas P~oocrnas st ancutsd Mstcnas at Un fernale stare Aist Tyoe (0NSI Austcnit K (c 201 (SNLOO)
Typ'cat Con pca bon. 9(r (al 16 14 Cr, 3.5-5.S Ni, O.IS C. $>7.$ sta, LO Si, O,ON P, 0 OX 9, 025 N form (b)
Sheets Qnys Tubing Tens Ie Qrengps.
los psi I I5 IIS 115 rictd Qrength.
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($20200) 17 19 Cr. 44 ki 0.1$ C. 7 5 10 0 Ma, 1.0 Si, 0060 P, 0030 S. 025 rt Sheets Qnpl Tubing
>105 105 105 55 55 55 5$
55 cs Rb 90 Rb 90 Rb 90 100 145 200 220 5$
95 150 110 5$
38 15 5
205 (S20500) 16.5 lb Cr, I 1.15 Ni, 022025 C, ll 15.5 Mn, 1.0 Si, 0,060 t', 0.030 S, I 1.15 Mo, 0224.40 N
Plates 301
($30100) 302 IS30200) 16 18 Cr. 6 4 Ni O.IS C, Z.O Mn, I 0 Si, O.ol5 P, 0030 5 17 1$ Cr. 8 10 Ni O.LS C. 20 Stn, 1.0 Si, 0olS P, 0030 S
Ptatcs Sheets Stnys Tubing 8<<s Ptatcs Sheets Stnys Tvtung tcire 105 110 110 105 35 35 40 3S 35 55 60 fio 50 bhn 165 Rb 85 Rb 8$
Rb 85 Shn 150 RbN Rb 85 Rb 4S Rb 85 Rb 63
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'to tyPe 302 for tnidatvsn resistance 303
($30300) 303se LS30323) 17 19 Cr. 8 10 Ni, 0.15 C. 2.0 lln, !2) Si.
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ISZ 235 267 40 40 40 37 61 61 45 40 35 30 67 6S 62 60 52 31 8$
90 100 06 125 30l t530400) 304L (SS)403) 1$ 20 Cr. 8 1050 Ni 0.08 C, 2.0 Ma, 1.0 Si, 0.04S P. ONO S Ib 20 Cr. 412 Ni 0.03 C. ZO Ma. IO Q, 0045 P, Ol)X 4 Bars Ptates Sheets Stnys Tubing Wire Plates Sheets Qfiys Tvtvng 1$
bl bl 14 3$
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- N
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'Ifire 42 44 240( ~I Xs (SX500) 17 IS Cr. 10.50 13 Ni, 0,12 C. 2.0 Ma, 1.0 Si, 0045 P. OOX 4 Plates Sheets Stnys Tubing cfire Rb N Rb N Rb N Rb 17
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APPENDIX B
B-I DETERMINATION OF AVERAGE CRUSHING FORCE It is a simple matter to calculate the buckling load of the upper section of the fuel storage cell since it consists of a square duct.
The duct is made up of four plates which are welded on the corners.
When the buckling load is exceeded, two of the opposing plates will buckle inward while the other two plates buckle outward.
In this way, the intersection of the plates remains a right angle, and no moment is transferred'across the joint.
This mode of deflection would absorb the least energy.
Each of the four plates is then simply supported on al 1 edges.
Note on page A-8 (copied from Ref.
- 2) that the buckling load is independent of a/b for values of a/b greater than 0.8.
This shows that the buckling of simply supported plates (and therefore thin walled ducts) is fundamentally different from that of thick walled columns.
The thick walled column will collapse at the axial location of maximum bending
- moment, and thereafter form a plastic hinge.
The applied load will fall off quickly, and there-fore, very little energy will be absorbed.
This happens because all of the energy i s absorbed at the short plastic hinge.
On the other hand, each of the sides of the thin walled duct will fail independently (at a relatively low load),
but the corners provide sufficient restraint to avoid complete collapse.
As explained in Reference 4 (page 38 attached),
the plate elements continue to carry load after buckling.
A periodic wave is formed in the. thin walled column which means that many plastic hinges must occur in order to crush the column.
It follows that the crushing of this thin walled column will absorb much more energy than the buckling of a thick walled column.
This periodic deflection curve for the thin walled duct explains why the buckling load is nearly independent of the a/b ratio.
The length of the column does not matter since each characteristic section of the column has its own buckling load.
In order to demonstrate this effect, two crushing tests were performed.
The first was a 4.7 inch steel duct with a 25 mil wall.
Page B-4 illustrates the geometry of the steel duct and shows the experimental data of load vs.
deflection.
A calculation of the critical stress and buckling load is also included.
Notice that the measured load of 1.9.tons (or 3,800 lb.) is much hiqher than the calculated value of 1,430 lb.
This is probably due to the fact that the duct was fabricated from two sheets of metal which were bent to shape and joined in two corners.
The two joints consist of interlocking bends of metal and are, therefore, very stiff. Page B-5 shows the load-deflection-curve.
As the theory suggests, very high forces are maintained after buckling.
The force dropped off to one-half of the initial buckling
- load, and occasionally dropped to values as
.low as one-fourth of the buckling load.
It would be conservative to assume that the average crushino force is one-third of the buckling load.
0 gt ~
8-2 In fact, the initial buckling load was exaggerated by the two stiffened joints.
These joints probably contributed much more to the initial buckling strength than to the subsequent load carrying ability of the duct.
The load carrying ability after buckling depends on the stretching and folding of the plates.
If the two joints were not stiffened, the ratio of sustained crushing load to the initial buckling load would have been greater than I/3.
Page 8-6 shows two photographs of the partially crushed, steel duct.
Note how two plates buckle outward, while the other two buckle inward.
After this occurs, the load began to pickup again, and this will reoccur many times until the duct is completely crushed.
One of the stiffened joints can be seen in the picture.
Page 8-7 i llustr ates the geometry of the second buckling test.
This was done with a soft aluminum extrusion with a thicker wall.
The calcu'lated buckling stress of 51,920 psi is much greater than the yield stress of this material which is probably less than 10,000 psi.
The duct actually fai led at a stress of 12, 500 psi Failure was i nitiated by yielding of the material rather than the application of the critical stress.
The stress vs. deflection curve shown on. Page 8-8 shows an average crushing force of half the buckling force.
The crushing pattern is'uch like the previous test.
Note from the photographs on Page 8-9 that this soft material experiences a great deal of deformation before the load picks up for the second collapse.
As before, the duct will collapse many times until it is completely crushed.
This test again confirms the fact that it is very conservative to assume that the average crushing load is equal to one-third of the buckling load.
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I b/2 b/2 Rg. C.5 Local Buckling and poit Bucidlng Strength of Stiffened Compression Element Following the same approach used in Refer-ence 4, the unit for stress has been changed to ksi in the 1974 Specification instead of psi used in the previous edition.
2.3 Properties of Sect)ooe Unlike columns or shells, plate snd sheet ele-ments possess a large strength reserve after buck-ling, unless buckling occurs at stresses approach-ing the yield point for sharp yielding materials or at large inelastic strains for materials such as stainless steels which do not have s definite yield point. For example, Figure C.5(a) shows the buckled form of s stiffened compression element (a sheet which is supported along both unloaded edges by thin webs or edge stiffeners and can be regarded as simply supported), uniaxially loaded by a compression force. Although the element has buckled, snd out of plane waves have developed, it is still capable of sustaining additional load, snd the member of which the element is a part does not collapse. This behavior is a result of the mem-
/
brane stresses which are developed in the element transverse to the direction of loading. UnstifTened elements (sheets which are supported along one unloaded edge only, the other unloaded edge being unsupported) behave in a similar fashion, except that the strength reserve after buckling is rela-tively small because less membrane action is possible.
The general equation for the critical buckling stress of isotropic sheet elements is 12 (1-rAx) (w/t)'-'r,.l>
wherec= critical buckling stress E
initial modulus of elasticity
= Poisson's ratio in the elastic range q
plasticity reduction factor w
flat widt t
= thickness k
buckling coefficient Inspection of Eq. (C.l) reveals that the ratio of flat width to thickness of the sheet element is an important parameter; the critical stress decreases with increasing width-thickness ratio.
To keep the width-thickness ratio reasonably small, thus maintaining larger critical stresses, compression elements are frequently provided with intermediate longitudinal stiffeners between webs or between a web and an edge stiffener (Fig-ure C.3).
In practical design the effective width concept is widely used for taking the postbuckling strength of compression elements into account.
Figure C.5(b) indicates the post buckling stress distribu-tion in a stiffened compression element. The solid line is the actual stress distribution over the actual element width, w. The dashed line is the equival-ent uniform stress distribution, equal in intensity to the edge stress of the actual distribution but only applied over an effective width b. The total load carried by the element is the same for both distributions. Applications of the effective width concept are given in Section 2.3.1 of the Specifi-cation.
The effective width concept is used explicitly in computing the properties of sections which con-tain stiffened or multiple-stifTened compression elements. Because the effective width is a function of the element edge stress, it follows that the prop-erties of the section are also functions of the stress level. For this reason, when computing the effec-tive area, moment of inertia, and section modulus, proper recognition must be given to the effective width of stiffened and multipleatiffened,compres-sion elements as a function of the edge stress and the flat-width ratio. The applications of the pro-vision are included in Sections 2.4 and 3.6 of the Specification.
2.3.1.1 Stiffened Element Without fntermedlate Stiffenere The efTective width relations used in the previous edition of stainless steel specifi-
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APPENDIX C
C-1 NITAWL INPUT LISTING
'DC COOK HOOK DROP,4.0XENR, EXPLICIT MODEL 0$ $
6 7 8 ll 18 19 9 0 20 1$ $ 0 21 5RO 10 2RO -1 0
T FOR U-235,U-238:
XXXOY=UNDAMAGED, XXX2Y=DAMAGED Y=1 FOR INTERIOR RODS, Y=2 FOR EDGE RODS DAMAGED PITCH=0.52718",
UNDAMAGED=0.496" 2$ $ 92235
-92501
-92502
-92521
-92522 92238
-92801
-92802
-92821
-92822 40302 25055 8016 1001 24304 26304 28304 13027 5010 5011 6012 3** 92501 293.15 2 0.38481 0.2065 1773.7 9.40644-4 1
15.9994 185.225 1
238.051 196.683 1
1 92502 293.15 2 0.38481 0.1622 1773.7 9.40644-4 1
15.9994 185.225 1
238.051 196.683 1
1 92521 293.15 2.0 0.38481 0.1537 1773.7 9.40644-4 1.0 15.9994 185.23 1.0 238.051 196.68 1
1 92522 293.15 2.0 0.38481 0.1369 1773 F 7 9.40644-4 1.0 15.9994 185.23 1.0 238.051 196.68 1
1 92801 293.15 2 0.38481 0.2065 74.85 2.22902-2 1.0 15.9994 7.8165 1.0 235,044 0.4431 1.0 1.0 92802 293.15 2 0.38481 0.1622 74 '5 2.22902-2 1.0 15.9994 7.8165 1.0 235.044 0.4431 1 '
1.0 92821 293.15 2.0 0.38481 0.1537 74.9 2.22902-2 1.0 1 5.9994 7.8165 1.0 235.044 0.4431 1.0 1 '
92822 293.15 2.0 0.38481 0.1369 74.9 2.22902-2 1.0 15.9994 7.8165 1.0 235.044 0 '431 1.0 1.0 40302 293.15 1.0 0.0635 0.2125 191.39 4.25181-2 1.0 6R0.0 1.0 MN IN 304 SS,
- SLAB, AVG THK = 0.05",
DANC=O.O 25055 293.15 1.0 0.127 0.0 511.83 1.73644-3 1.0 55.847 387.309 1 '
58.71 79.3401 1.0 1.0 4** F293.15 T
~ (i(r~
C-2 KENO-Va INPUT LISTING DC COOK HOOK DROP, 4.0$ ENR READ PARAMETERS THE=50.0 GEN=103 NPG=400 LIB=41 TBA=2.0 FLX=YES FDN=YES XS1=YES NUB=YES PWT=YES END PARAMETERS READ HIXT SCT=1 INTERIOR ROD, UNDAMAGED BUNDLE MIX=1 92501 9.4064-4 92801 2.2290-2 8016 EDGE
- ROD, UNDAMAGED BUNDLE MIX=2 92502 9.4064-4 92802 2 '290-2 8016 INTERIOR ROD, DAMAGED BUNDLE HIX=3 92521 9.4064-4 92821 2.2290-2 8016 EDGE
- ROD, DAHAGED BUNDLE MIX=4 92522 9.4064-4 92822 2.2290-2 8016 Z IRCALLOY MIX=5 40302 4.2518-2 WATER MIX=6 1001 6.6740-2 8016 3.3370-2 304SS FOR CAN MIX=7 24304 1.7430-2 25055 1.7364-3 26304 ALUMINUM MIX=8 13027 6.0242-2 MIX=9 5010 6.6707-3 5011 2.7081-2 6012 8.
END HIXT READ GEOMETRY UNIT 1
COH='NTERIOR ROD, UNDAMAGED BUNDLE'YLI 1
1 0.38481 2P100.0 CYLI 0
1 0.3937 2P100.0 CYLI 5
1 0.4572 2P100.0 CUBO 6
1 4P0.62992 2P100.0 UNIT 2
COM='DGE
- ROD, UNDAMAGED BUNDLE'YLI 2
1 0.38481 2P100.0 CYLI 0
1 0.3937 2P100.0 CYLI 5
1 0.4572 2P100.0 CUBO 6
1 4P0.62992 2P100.0 UNIT 3
COM='UIDE TUBE, UNDAMAGED BUNDLE'YLI 6
1 0.3937 2P100.0 CYLI 5
1 0.4572 2P100.0 CUBO 6 I 4P0.62992 2P100.0 UNIT 4
COM='NTERIOR ROD, DAMAGED BUNDLE'YLI 3
1 0.38481 2P100.0
'CYLI 0
1 0.3937 2P100.0 CYLI 5
1 0.4572 2P100.0 CUBO 6
1 4P0.669514 2P100.0 UNIT 5
COM='DGE ROD, DAMAGED BUNDLE'.6462-2 4.6462-2 4.6462-2 4.6462-2 5.9359-2 28304 7.7182-3 4380-3
CYLI 4
1 0.38481 2P100.0 CYLI 0
1 0.3937 2P100.0 CYLI 5
1 0.4572 2P100.0 CUBO 6 I 4P0.669514 2P100.0 UNIT 6
COM='UIDE TUBE, DAMAGED BUNDLE'YLI 6
1 0.3937 2P100.0 CYLI 5
1 0.4572 2P100.0 CUBO 6
1 4P0.669514 2P100.0 UNIT 7 COM='NIT 7 IS UNDAMAGED BUNDLE'RRAY 1 2*-10.70864 -100.0 REPLICATE 6
1 4R0.6731 2RO.O I REPLICATE 7 I 4RO. 1905 2RO.O 1
REPLICATE 8
1 4R0.0254 2RO.O 1
REPLICATE 9
1, 4R0.18034 2RO.O 1
REPLICATE 8
1 4R0.0254-2RO.O 1
REPLICATE 7
1 4R0.0762 2RO.O 1
REPLICATE 6
1 4R1.45542 2RO.O 1
UNIT 8 COM='NIT 8 IS DAMAGED BUNDLE, PITCH=O. 5272"
'RRAY 2 2*-11. 38174 -100. 0 REPLICATE 7
1 4R0.1905 2RO.O 1
REPLICATE 8
1 4R0.0254 2RO.O 1
REPLICATE 9
1 4R0.18034 2RO.O 1
REPLICATE 8
1 4R0.0254 2RO.O 1
REPLICATE 7
1 4R0.0762 2RO.O 1
REPLICATE 6
1 4R1.45542 2R0.0 1
ARRAY 3 3*0.0 COM='RRAY 3 IS loxlo BUNDLE ARRAY'ND GEOMETRY READ ARRAY ARA=1 NUX=17 NUY=17 NUZ= 1 LOOP 2
1 17 1
1 17 1
1 1
1 1
2 16 1
2 16 1
1 1
1 3
6 12 3
3 15 12 1
1 1
3 4
14 10 4
14 10 1
1 1
3 3
15 3
6 12 3
1 1
1 END LOOP ARA=2 NUX=17 NUY=17 NUZ=l LOOP 5
1 17 1
1 17 1
1 1
1 4
2 16 1
2 16 1
1 1
1 6
6 12 3
3 15 12 1
1 1
6 4
14 10 4
14 10 1
1 1
6 3
15 3
6 12 3
1 1
1 END LOOP ARA=3 GBL=3 NUX=10 NUY=10 NUZ=1 LOOP 7
1 10 1
1 10 1
1 1
1
C-4 8 561 561 111 END LOOP END ARRAY READ START NST=S NBX=S END START READ BOUNDS ALL=SPECULAR END BOUNDS END DATA
ATTACE4ENT 2 TO AEP:NRC:0514R
Attachment 2
EP:NRC:0514R STEVENSON R, ASSOCIATES a structural-mechanical consulting engineering firm 9217 Midwest Avenue
~ Cleveland, Ohio 44125
~ (216) 587-3805
~ Telex: 980101 1608C 86cl438 January 5,
1987
,(a,/81 Mr. R.
B. Bennett American Electr1c Power Service Corporat1on l Rivers1de Plaza P.O.
Box 16631
- Columbus, OH 4321-6631
Dear Mr. Bennett:
Per your and Mr. Satyan Sharma's
- request, I have performed an 1ndependent review of the mechan1cal analysis of the D.C, Cook Spent Fuel Pit Load Drop Analysis transmitted to you by letter from A.J. Martenson of the Exxon Nuclear Company Inc. dated 7 November 1986 and revision 2 of Appendix A and a
new Appendix 8 to that Analys1s transmitted to you on or about 31 December 1986.
Based on this rev1ew, it is my opin1on that the depth of penetration of a free fall drop oF the crane book assembly into the spent fuel pool and impacting the spent fuel racks of 19 inches for the block plus 30.5 inches more for the hook is a reasonable est1mate of the expected depth of pentration.
Please adv1se if you requ1re any clar'ificat1on of this letter.
S1ncerely,
(
glQ i)
(
N l~
~
~
John D. Stevenson President CC:
Mr. Satayan Sharma