ML20084Q930
| ML20084Q930 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 05/18/1984 |
| From: | Finneran J, Iotti R EBASCO SERVICES, INC., TEXAS UTILITIES ELECTRIC CO. (TU ELECTRIC) |
| To: | |
| Shared Package | |
| ML20084Q918 | List: |
| References | |
| NUDOCS 8405210590 | |
| Download: ML20084Q930 (43) | |
Text
{{#Wiki_filter:. _ _ _ _ . ___ __. ._. UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION BEFORE THE ATOMIC SAFETY AND LICENSING BOARD 1 In the Matter of )
) Docket Nos. 50-445 and TEXAS UTILITIES ELECTRIC ) 50-446 COMPANY, ET AL. )
) (Application for (Comanche Peak Steam Electric ) Operating Licenses)
Station, Units 1 and 2) ) t AFFIDAVIT OF ROBERT C. IOTTI AND JOHN C. FINNERAN, JR. REGARDING THE EFFECTS OF GAPS ON STRUCTURAL BEHAVIOR UNDER SEISMIC LOADING CONDITIONS I, Robert C. Iotti, having been first duly sworn hereby depose and state, as followst I am Chief Engineer of Applied Physics for Ebasco Services, Inc. In this position I am responsible for directing analytical work in diverse technical areas, including analyses of the response of piping and support systems to dynamic events, including earthquakes. I have been , retained by Texas Utilities Generating Company to coordinate and oversee the technical activities performed to respond to the Licensing Board's December 28, 1983, Memorandum and Order (Quality Assurance for Design) . A statement of my educational i and professional qualifications was transmitted with Applicants' letter of May 16, 1984, to the Licensing Board in this proceeding. 0
i I, John C. Finneran, Jr. , hereby depose and state as followst I am employed by Texas Utilities conorating Company as Project Pipe Support Engineer for the Comanche Peak Steam Electric Station. In this position I oversee the pipe support f design activities of each organization performing pipe support design work for Comanche Peak. A statement of my educational and i l professional qualifications is in evidence as Applicants' Exhibit 1428. O. What is the purpose of this affidavit? A. In this affidavit we address CASE's allegations regarding i the ef fect of gaps (a.o. , bolt hole tolerances) on the i g structural behavior of pipe supports under seismic loading , conditions. This affidavit is in partial response to Item 9 of Applicants' Plan to respond to the Board's December 28, { 1983, Memorandum and order (Quality Assurance for Design). Q. What is CASE's allegation regarding these effects? A. CASE argues that bearing type connections are inappropriate j for use as mechanisms for supporting structures during
- seismic events.
Q. What are the bases for this assertion? A. There are two based for CASE's argument. First, CASE argues that in a bearing type joint it is impossible to predict how 1 many bolts are involved in the transfer of shear from the support to the wall.1 second, CASE argues that the presence 1 See CASE Proposed Findings of Fact and Conclusions of Law, Sections VII, XXI. J
of gaps in the joints under dynamic conditions can be
" disastrous."2 In both instanceu, thoro la an underlying ,
s concern by CASE that the Lolt holes in support base plates are " oversized," thus creating " gaps" that must be specially analysed. We will address each of the issues separately. We will also put in perspective the question of " oversized" bolt holes. Q. Before addressing these matters, please explain the l dif ference between bearing and friction connections and the factors which seem to be of concern to CASR. , , j A. Bearing type connections are connections where shear forces , between the two joined components (in most instances at 4 Comanche Peak the base plates and the concrete), are reacted l by bearing of the bolt surface on the surface of the bolt ; j hole. Friction type connections are those where the same shear forces are expected to be fully reacted by friction ); forces created by proloading the cornection bolts. For a baseplate, the friction forces exist between the baseplate and concrete surface and between the baseplate and bolt washer surface. It must be recognised that friction i ' connections will ultimately revert to bearing connectiong when the shear load exceeds the friction developed at the interface plane. In the ultimate condition, all joints are bearing joints. 2 cAgg gnhibit 763 at 3t Tr. 6605-6624. t t
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O. Do you agree with CASE's assertions concerning the distribution of shear in multiplo bolt, bearing type connections?3 A. No. CASE's position is promised on an orroneous assessment of industry practice and bolt interaction theory. CASE argues that "
. . . the usual proceduro (industry practice) is to assume that two holes react the load regardless of the number of bolts in the pattern (for patterns of 4 bolts or more)". Not only is this not industry practice but it is contrary to sound engineering principles. This is readily apparent if one considers a pattern with, for example, twenty bolts. CASE would have one believe that it is in(ustry practice to assume that only two bolta can react imposed shear loads. Obviously, if this were the case, and no more than two holta could be counted on to react shear loads, every bolt in a multiple bolt pattern would have to be significantly overdesigned. If this were done, every connection designed for shear might as well be designed with only two bolts to begin with. This obviously is an illogical result, but one which flows directly from CASR's position.
CASR's argument is also contrary to sound engineering principles, recognised in authoritative texts concerning the deJign of bolted joints. " Plastic Design of Steel Framen," 3 cAgg Proposed findings, section VII at 10,11, and Section XXI. l l
I by Lynn 3. Meedle (Attachment A), at pages 7-9, discusses the inherent plastic action that exists in alantic denton. l Therein it is noted (at 8) that there are "a number of examples . . . in which the ductility of steel has been 1 counted upon in elastic design knowingly or not -- but certainly not through direct application of plastic design procedures." The authob states (at 7) that "perhaps the outstanding example of this variance between elastic design assumptions and the actual truth is to be found in the ordinary riveted or bolted joint." The author goes on to explain that inelastic (plastic) action occurs to assure that all bolts will eventually participate in reacting the shear load. Another text which addresses this matter is " Structural Design Guide to AISC Specifications for Buildings," authored by Paul F. Rice and Edward S. Hoffman (Attachment 5). Rice ! and Hof fman explain (at 268) that in designing bolted connections for loading in shear "the use of an average capacity for each of the several connector elements sharing the totti load is justified by allowing self-limiting
- localized stresses determined by an elastic joint analysis to exceed the yield point and create inelastic localised deformations of the connector materials, or by inelastic l
1 deformations of the connection elements (1.15.4)." 'To illustrate this principle, Rice and Hof fman employ a two bolt example, designed to account for inelastic behavior
i i ,. resulting from the cumulative effects of out-of-round holes, holes exceeding the bolt diameter by 1/16", and hole alignment. They conclude that both bolts share the load. In summary reliance on inelastic action in bolted connections to distribute shear actively to all bolts in the connection is a well recognized and valid assumption in elastic design. Q. Do these excerpts agree with the positions taken by the NRC i Staff in this proceeding? A. Yes. As Dr. Chen stated (at Tr. 6884) "it is usual practice to assume that all the bolts will react equally to a shear i load: the rationale being that even though initially one bolt or possibly two, let's say, out of a pattern of four will be engaged, some yielding of the first or second bolt will lead to equal load sharing eventually." Q. Are the tolerances Applicants employ for bolt holes those normally characterized in the industry as " oversized" bolt holes? A. No. The' term " oversized" holes has a generally accepted meaning in the construction trade. The 8th Edition of the AISC sanual of Steel Construction is quite instructive on
'- ',, s this point. At page 5-58 of the Manual, Paragraph 1.23.4.3
( Attachment C) states " Oversized holes may be used in any or alk', vplies'of' friction-typeconnections, 4 but they shall not
, he ,used in bearing-type connections. " Table 1.23.4 on that 1
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" ove rs i z e'd " holes. This table establishes tolerances for standard and oversized hole diameters.
Q. Do Applicants use " oversized" hole tolerances noted in the AISC Manual in the design of pipe support anchor bolts? A. No. Applicants do not utilize the oversize hole tolerances for anchor bolts. Applicants.use instead more stringent tolerances. Apolicants' specified hole sizes meet the following requirements: holes for bolts up to 1" are to be d
+ 1/16" (where d is the diameter of the bolt), and holes for bolts 1" and over are to be d + 1/8". AISC defines
" oversized" as d + 3/16" for bolts up to and including 7/8" diameter, d + 1/4" for 1" bolts, and d + 5/16" for bolts greater than or equal to 1 1-1/8" diameter. Although Applicants' criterion for bolts with diameters equal to or greater than 1" is 1,16" larger than the " standard" by AISC, it is 3/16" smaller than the AISC " oversized". Thus, Applicants' specifications for bolt hole tolerances can definitely not be called " oversized," as that term is generally used in the construction industry.
Q. What is the purpose of these bolt hole tolerances? A. These types of tolerances are absolutely necessary to facilitate construction. As demonstrated in this affidavit, a reasoned consideration of the principles related to the distribution of loads in these connections demonstrates that such an approach is appropriate and acceptable.
Q. What test data are available regarding the capacity of anchor bolts to withstand the type of displacements which could occur in the transfer of shear loads? A. There are two sets of data already in the record regarding bolt capacities in shear. Both Applicants' Exhibit 142D and the Cygna response to Doyle Question 16 contain data which demonstrate that b'olts 1" and greater have more than enough capability to deflect a worst case 1/8" and still carry their full rated load capacity. (See Attachments B and C of Applicants' Exhibit 142D, and Enclosure D16-1 to Board Exhibit April 1984 No. 1 (Cygna testimony).) Attachment "B" of Applicants' Exhibit 142D indicates that the 1 1/4," super kwick Hilti bolt did not fail until approximately .7 inches displacement. Therefore, the inherent safety factor to failure for this bolt with a maximum displacement of .125" (1/8") would be .7 divided by .125, or 5.6. Enclosure D16-1 indicates failure of a 1" Hilt' bolt at about .57" displacement. The safety factor inherent in this bolt would be .57 divided by .125, or 4.56. Further, Attachment "C" of Applicants' Exhibit 142D indicates a lower limit for slippage of a 1 1/4" Richmond Insert of about .4 inches. Thus, the safety factor in this case would be .4 divided by
.125 = 3.2. This test data is indicative of the displacement capacities of anchor bolts used at Comanche Peak. The data demonstrates that Applicants' bolts withstand the worst case slip of 1/8" that would be necessary to distribute shear forces equally to all bolts in the connection.
Q. Is'there anything else you would like to, add regarding CASE's assertions concerning the distribution of shear loads in multiple bolt, bearing connections? A. Yes. During cross-examination of Cygna in the April 1984 hearings , Mr. Doyle introduced a paper by James M. Fisher.4 M r. Doyle used a portion of that paper (page 87, Shear and Anchor Bolts") in support of his position. Mr. Doyle argues , that Fisher concurs with CASE's position that "not all bolts in a cluster" may be used to transfer shear. However, Mr. Doyle apparently overlooked the fact that Mr. Fisher is talking about anchor bolts used at the base of columns. As shown on p. 4-126 of the AISC Manual (Attachment D), anchor bolts used at the base of columns which have 1-inch to 2-inch diameter are permitted to have 1/2 inch oversize holes. In addition, Mr. Fisher recommends on (p. 87) that bolt hole sizes for anchor bolts should be 1.33 times the bolt diameter. These tolerances are much greater than those employed on. pipe supports at CPSES. It is clear that the condition described by Fisher is not the same condition 4 CASE Exhibit 1001: Tr. a t 6605-6624.
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appliable to designing joints for supports. This is best illustrated by Figure 1. Figure 1 shows ultimate deflections in shear and tension for different size Hilti bolts. Also shown in Figure 1 are the deflections that would be permitted if the bolts were loaded to their allowable values. The true interaction curve is approximated by a straight line between the ultimate displacement in tension and in shear. Also shown is the actual curve derived from combined shear-tension tests. i Similar curves are available in the Teledyne Report Generic Response to U.S. NRC I&E Bulletin 79-02 Base Plate / Concrete Expansion Anchor Bolts." Although the linear approximation of the interaction curve is not conservative for high values of tension displacements (which correspond to near ultimate pull-out loads), it is more than adequate for shear displacements, which are the concern here. To illustrate l what these test data indicate, let us take a 1" Hilti embedded 10 1/2" deep, and compute the margin of safety when loaded to its allowable values in tension or shear. In - either case there is a margin of safety of 5. Let us now assume that only one bolt out of many in the connection is , loaded in shear, as CASE argues should be done. Let us further assume that it will deflect through a bolt hole gap. of 1/16" before the other bolts begin to take some of the shear. This bolt alignment is the worst case condition for imposing shear stresses in the bolt. (See Illustration 1.) l l I w.+ ~ e - -e--4, - -~- -g - - - - -,- - p4.v-v,e --++wew s- ~ - , - e
Let us further calculate the deflection of the Hilti bolt, which would result from thermal expansion of the tubular frame. A worst case of .0485 inches deflection is computed by the method of Applicants' Exhibit 142D for a 1" bolt, embedded to 10 1/2", (see Figure 2 of Applicants' Exhibit 142D). eG D,oscwoWw LoAo y
*GV Sitt ShkWN D>$6KRMT3D ILLOSTRATION 1 If the direction of the load is as shown in Illustration 1, bolt number 1 will take all the shear until it deflects through the gap. In this instance, the gap is .0625 inches (1/16"). If bolt number 1 is. assumed to have already been loaded to its allowable value (7000 lbs) it would have deflected approximately .085 inches. Thus, the next bolt would have crossed the gap and started sharing the load, and the thermal expansion load would have already been totally relieved.
For the sake of being ultra-conservative, however, let us further assume a totally unrealistic scenario, i.e., the first bolt displaces as a result of its normal load, plus the thermal expansion load (which would have.been relieved and thus is not truly additive to the other displacements),
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. . .o plus the gap, before the second bolt engages. This would correspond to a gap for the second bolt of 0.085 + .0485 +
.0625 = 0.196 inches. From Figure 1, even under this incredible condition, the margin of safety of the first bolt would be .55/.196 = 2.8. In reality, the margin of safety resulting from the 1/16" bolt hole gap is 5 since the second bolt begins to share the load before the first bolt reaches the deflection at which the load would equal the allowable value. Even if the bolt holes are 1/8 larger than the bolt, the real factor of safety would still be 4.4. And if you were to play the same imaginary game of adding allowable displacement, thermal displacement and gap one would obtain a margin of safety of 2.1 (.55/.258).
To illustrate why the Fisher paper made the recommendations it did, let us compare this situation to that examined in 't the paper. The gap condition examined by Fisher may be as high as 1/2 inch. The real margin of safety of the first bolt engaged would only be .55/.5, or 1.1, for a 1" bolt embedded 10 1/2 inches. This apparently is why Fisher recommends consideration of shear distribution on a limited number of bolts, rather-than all bolts, where oversized holes such as he was examining may exist. Q. What is your response to CASE's concern 5 with the dynamic nature of the loadings? 5 CASE Exhibit 763 at 3 Tr. at 6605-6624.
.. 4. 4 l l A. CASE contends that the presence of gaps in joints under seismic conditions can be " disastrous." To the contrary, in a seismic event the first quarter cycle loading would cause preferentially loaded bolts to deflect in shear until the other bolts engage. Once the bolts have deflected, the gaps are uniform for all bolts. This is an over simplified explanation of what is a very complex phenomenon. To illustrate this point, refer to the above sketch. Bolt #1 will deflect until the second bolt engages (if the load is high enough to cause such deflection). As the load reverses, the locations of both bolts in the holes are now the same and both will take shear and deflect a like amount, and continue to do so, for the remainder of the dynamic event. Therefore, only during the first quarter cycle can there be preferentially loaded bolts. Thereafter, for subsequent cycles the load will be reacted by all bolts and there is no genuine concern for the capac'ity of the bolts to accept these loads. O. Are there any other matters that should be considered with respect to CASE's concern regarding the potentially adverse influence of gaps on a system's seismic response?6 A. Yes. It is important to recognize that the effect of gaps and other nonlinearities on the seismic response of systems cannot be defined in absolute terms. The effect is dependent on many factors, including the nature of the 6 Transcript at 13,705-13,717. l l l l
r* n excitation (magnitude and distribution of frequencies), and the size, orientation and number of gaps. The situation is further complicated by the fact that nonlinearities introduce impacts and hence impact damping, which is known to be related to the coef ficient of restitution (C ) via the formula:7
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C =e-R For steel on steel, average restitution coefficients of 0.6 would lead to impact damping f actors (8) of 8% , lower restitution coefficients would lead to progressively higher values of impact damping, and even with a very high restitution coefficient (0.8) the impact damping would be higher (3.5%) than the damping values recommended by the NRC in Regulatory Guide 1.61. This would mean that to account for just one of the effects of gaps, one would have to employ damping factors for portions of the system that exceed those specified by the regulations. Clearly, consideration of such effects would require complex analyses which depart from accepted practices. Another fact that should be recognized is that while the gap is being transversed, little or no seismic' input acceleration is being experienced. Also, depending on the nature of the gap (whether it is a gap in a bolt hole or the deadband in a mechanical snubber, or the play in a strut i 7 HEDL-SA-1769, D.A. Barta, " Analyses of Piping Systems with Nonlinear Supports Subject to Seismic Loading", ASME P & PV Third National Congress, June 1979 ( Attachment E) .
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assembly), a fraction of the seismic input may be introduced via friction (if permitted by the vertical excitation for a horizontal gap or vice versa). Thus, while transversing the gap, material damping takes place without a corresponding feed of energy from the seismic event. Obviously, the combination of intermittent energy input while damping continues produces a beneficial effect on the system response. All of the above-described effects cannot be accounted for in the typical linear response spectrum analyses which are used to design the systems at CPSES, and are only accounted for with difficulty by performing nonlinear time history,
! analyses. Thus, absolute generalizations as CASE contends should be made (CASE Proposed Findings at VII_-ll,12) simply 2
are not possible. Q. Is there a way to compare the results that would be obtained from a nonlinear analysis which considers the presence of gaps with those of the response spectrum analysis Applicants employ, in order to assess the effect of the gaps on seismic response? A. In some respects, yes. However, in trying to compare results of nonlinear time history analyses with response spectrum analyses it becomes very difficult to distinguish between the ef fects of having conducted a full time history analyses (whether linear or nonlinear) versus a response spectrum analysis, and the effects of the gaps by
n i themselves. In other words, these two analytical approaches are suf ficiently dissimilar that one may not discern whether particular results are attributable to differences in individual variables or assumptions (e.g., gaps) or the analytical techniques themselves. In view of these uncertainties, it may be concluded that a more viable approach to assess,the effect of gaps in designs which have been accomplished using response spectrum analyses (as in the case at CPSES), is to compare the support loads and ;;ipe stresses predicted by the response spectrum (without gaps) with those which would be predicted by nonlinear time history analyses of each system. To respond fully to this question, however, would be a virtually never ending and enormously expensive task. Fortunately, a number.of comparisons have been made between results obtained by response spectra analyses and nonlinear time history analyses which simulate the actual gaps in the system which provide reasonabale evidence of the effect of gaps. Two general conclusions can be derived from these comparative studies and tests which simulate seismic conditions in actual piping systems which are apt to have some gaps in the supports, as would be present in any piping support system. The conclusions which can be reached are that:
is 46 a) The seismic response spectrum method, which ignores the nonlinearities, is more conservative than the non-linear time domain method (which includes gaps), and b) The effect of gaps on reduction of response frequency is negligible due to the transient nature of the seismic acceleration loading. . Studies bearing out these conclusions are discussed below. The study prepared by Barta (footnote 7) compares results of the response spectrum analyses of an FFTF piping system to those of the non-linear analyses which modelled the deadbands (gaps) of mechanical snubbers. The gap used was 0.120 which is comparable to or in excess of that existing in supports. Table 2 in that study compares the snubber loads obtained by both analyses With the exception of two snubbers, all other support loads were higher in the response spectrum analysis, with the average b91ng about 1.45 times larger. A second example, taken from a study prepred by Badrian8, compares steam line stresses computed by response spectrum and non-linear time history methods. Table 3 in that reference provides the results. Here again, the response 8 Badrian, "A Seismic Analysis Review," Ebasco Services, Inc. ( Februa ry , 1977) (Attachment F). l
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ca ,, spectrum method is more conservative by approximately a factor of 2.3, although one point did show a higher stress for the non-linear model. Finally a third study 9 compared results of response spectra analyses performed on 4 in., 16 in, and 28 in. pipelines against those obtained by nonlinear time history analyses of the same systems with varying gaps. Tables 3, 4 and 5 in this study illustrate the difference in support loads and piping stresses that result from several modelling assumptions. Once again, the results of the response spectrum analyses are generally conservative with respect to non-linear analyses. In sum, existing studies and analyses indicate that the effect of gaps are adequately bounded by analyses performed using the response spectrum methodology (such as is employed at Comanche Peak), and further analysis of these effects using nonlinear time history analysis would be unwarranted. l l i ! 9 Barta, Huang and Severud,'" Seismic Analysis of Piping With Nonlinear Supports," Plant Analysis, FFTF Project,
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dW g;.g would require careful study prior to the application of plastic dem.n ,
to them. . r,.,,gt**n.. .m 3 . ',,c .. . 1 s 1.4 TACIT ACCEPTANCE OF DUCTILE S!HAV!OR M. W
. 6O'.
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.W@g.,. ..L All but the most recent texts and speciscations in the field cf structural 4...t .:TV.y;gd e .
steel design have required that the fiber stresses should nowhere exceed f the yield point of the material when a specided overload is applied. {yAQ y p,. .f As shown in Art.1.1 this has been.an appropriate criterion for simple ' beams because deformations start increasing rapidly at yield-point stress. W ii eh.Mi" But if it were argued that yielding could not be permitted in any part N[M~ .f' of a s:meture, then much of the past and present practice would be .M.W M.. ',. ; completely ruled out. hfNE.. . . ' y 27~ Both in buildinsi and in bridges, specifications allow the designer QMPi.7 g,h$$
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to use average stresses due to bending, shear, and bearing that result in - actual local yielding. Such cases occur in pins and rivets and at local hjrdygFfMU i . ' f,12(0ilLA. j' y.9c., points. This local yielding results from stress concentrations th:t are Qge... . ' neglected in the simple design focaulas. Plasti: action is thus depended E.%#.$ 3 l upon to insure the safety of steel structures, and experience has shown iNd;J ..
.. ~
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that average or nominct mwmum stresses form a satisfactory basis for $;$,4Ws
,.N .
. kg.g design.. . . ; .i . '. . ~ -
Perhaps the outstanding e.xample of this variance between elastic design assumptions and the actual tmth is to be found in the ordmary riveted or bolted joint. The' assumption commonly is ma.de ilmt each p/g%g
'*/
ir+-*y v.~ fastener carries the same shear force. This is true for the case of two WMx g$j.n 6 .U - fasteners in a line. Wnen more are added (Fig.1.5), then as long as h.%hh ; ig~ j;. . Y tha joint remams e!astic, the outer fasteners must carry the greater i cJ preportion of the load. For the example with four rivets, if each rivet N.,.Wp .'s. ,a4 %, transmitted the same load, then between rivets C and D one plate would gfk.. .- ..U k;b 4 ! ' ~ d ,*N % carrf perhaps three times the force in the other. Therefore it would .
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stretch chree times as much and would necessarily force the outer rivet 1 ' ; (D) to carry more load than was assumed. The actual forces would look something like those shown under the hmMng " Elastic" in Fig.1.5. hdif [g - YQM
. .W;.'Or T8dI(.$@ dN What eventually happens is that the outer rivets yield, redistributing ,F.
s k < 'Q. 3 ,.C.
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, r m.w.9s f:rces to the inner rivets until all forces are about equal ca shown. he *7 '~ ,
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Therefore the basis for design of a riveted joint is really its ultimste iL load.and not the attamment of 6Irt yield. . g'[.f.,.,# , q -.j]Q d. M @ Kj To a greater extent than we may reali:e, the mwmum, strength of a pfpih g$g%]j'.-M p;ay% structur. has always been the dommnt hen eriterion.- when the W g.jp QG. p F M os
@f~ 's g%yyy3q q2ey usual permissible working stress has led to designs that were condstently too conservative, then that stress has been changed. Present design T, f;MML g' u.}yg. 4 .
.y.~g g a-tt:a.qd.
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this is a safe procedsre. Thus, the stresses that are edeulated for elastic.
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design purpcses often are not true mxdmum stre ses at all; they simply 'N 8-if~ y provide an indcx for structural design. A number of examples will now be given in which the ductility of
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. F. . . %.Md steel has been countai upon .m ehstic dedgn-knowin;Iy or not-but Gu:{ . )f.1:%<--f -f .' l J2E.57.E
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certainly not through direct application of phstic design procedures. The following listing is in two categcries: (1) facters that are neglectai i ' .. .h .4U,'C"Jed@W.;hf/
'" becauce of the compensating effect of ductility, and (2) instances in l ' Q.9 .~ .
which the werking stresses have been ravised becaua the " normal" gv
~$.D.
.y tj. $..] value was too conser stive. ,r o. , low.mgt.ce listing, several examples in e2ch category will be discussed in further detail.
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'WL 'I. Factors that are neglected:
.W-1 (1) Residual stresses (in the case of flexure) due to cooling after
. % 0. rollirg.
(2) Residual stresses resultin; from the cambering of beams. Y. $
%,%s -
(3) Erection stresses. M;I.{ (4) Foundation settlements. d.;h (5) Over-stress at points of stress-concentration (holes. etc.). (6) Bending stresses in angles connected in tension by one leg only. n-@ (7) Over-stress at points of bearing. MM9 (8) Nonunifoma stress-distribution in spliecs. leading to desi;;u of. Wl4 connections on the assumptten of a um. form c..isen.bution ot i;: n . a., stresses among the rivets, bolts, or welds. (Discussed above.)
.W4 (9) Difference in stress-distribution arising from the "canti!crer" as compared with the " portal" method of wind stress analysis.
@Ud@ y II. Revisions in wo-king stress due to reserve plastic strength:
E : Ah.4.3 q (10) Bendinp: stres of 30 ksiin round pins. (11) Bearing stress of 40 kai in pins in double shear. W~%;. (12) Ba&g stress of 24 ksiin framed structures at points ofinterior 2.$. . '-a support.
. .. .( Consider item (1) for example. All rolled members contain residual
.d stresses that are tormei due to cooling after rolling or due to cold-1~.j r em i @ ~ ir e Figure 1.6 shows a typical W shape (sketch c) with a p.%
charactenstic residual stress pattern (sketch b). When load-carrying y@N bending stresses are applied (sketch c), the resulting strams are additive
' to the residual strains already present. As a result, the " final stress" T. ': t -
(sketch d) could easily involve yielding at working load. In the example VD t of Fig.1.6 such yielding hr.s occurred both at the compression flange
.h ! tips and at the center of the tension flange (sketch e). Thus it is seen
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tnat cooling residual stresses (whose influence is neglected and yet which c .j& are present in all rolled bes=s) may cause yielding in a beam even below Qp,. g5,: the working load. - ns.p.f: Ld'- 3. y;. ~ . e, e., er I-) I g(-) ; (-) I (-) l SM:ey i.~rs '.- l t i Residual Apphed ! Mnal seess lPN
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w ( k [d E4 Van Nsstrand Reinho!d Company Regional Offices: New York Cincinnati Atlanta Dallas San Franc:sco V. t p.N Van Nostrand Reinhold Company International Off!ces: pyt London Toronto Melbourne F.3 2 1, ~.1 Copynant 1976 by Litton Educational Publishing.Inc. 3 Library nf Congress Catalog Card Number: 75-4 0491 {@1
+. ISBN: 0 42 26904-8 V,s .a M A!! rights rese:ved. No part of this work covered by the copyr'.ght hereon may d
.rq be reproduced or und in any form or by any means-graph 2c. electronic, or mechanical, including photocopying, recording. taping, or information storage
;7; t and retrieval systems-without permisson of the pubiisher.
h g Manufacturedin.the United States of America y.J f
'd Published by Van Nostrand Reinhold Company 450 West 33rd Street. New York.N.Y.10001 to.!
UM
%. Published simultaneously irt Canada by Van Nostrand Retnheid Ltd.
m . [.j 15 14 13 12 11 10 9 8 7 6 5 4 3 2 1
.H
- s. s N{h V;:
7.] Library of Congress Catalogmg in Publication Data
- ry
$l Rice. Pau! F 1921-Structurd design guide to AISC specifications for buildings.
h Includes bibliographical references and index.
- 1. Structural design-Handbooks. manuals. etc.
i 2. Building-Contracts and specifications. I. Hoff. t3 man. Edward S.1920- joint author. II. Title.
$ TA658.3.R52 690 75-40491
!( ISBN 0-442 26948-8 r.
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- ' 2ss: . STRucTtJRAl. CESIGN GUCE TO AtSC SPEC FICATIONS'FOft SUILDINGS l . ,
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m}. : .. ; . . . . . . Wp y.:":.w.. ..: . - 3 9h q. , . joint 'ar well as the sh' ear,1.15.5.),cr trmi.rfgid(trdnsmitting a pre d;termined fraction .
,"fi, Q.S the. full nioment capacity as a rigid joint and furth:r ! cads in ' shear as a t!exible joint w corresponding angle change to supply rota: ion for the additionalloads,1.15.5; 1.2).
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r N 9. Q, . 6MG w -~.i.L h Flexible Conneerions. " Flexible" connections are designed to transmit shear withed exceeding allowable unit stres:es on the connectors as a group or the connection as T . M :sg14 (ss-R.;-7 f.Cijiff jj}oq whole. The use of an average enpacity for each of several connector elements sharing th
"' N'd~.fM;M. {, :.Q total load is justified by allowing self-limi:ing localised sites:es determined by an clast Pg joint analysis to exceed the yield point and create inelastic localized ceformations of th h ~ ~~4 ~ g
, .h connector materials. or by inelastic deformatiens of the connection elements (1.15.{
*l:-T T The simplest examples of to '"d deformatien occur in the assembly of bearing.typ Q:f2( @' bolted eennect: ens where the cumulative toie ances permi::ed exist on (1) out-of rouni in the bolts. (2) oversize holes @ ) and i31 cen:er to.cen:er location of the holes in th g'A3 different elements -onnected. The extreme degree of such inelastic action occurs withi
.$ two-bolt bearing type connec:fon where one belt is loosely fitted and one is very tight Q Until the material of the connec:ed elemen: surrounding :he loaded belt or the bol: yield and deforms (+h"), the load is not shared and a 50 percent adjustment will be develope
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as the load increases. For larger (and :hus more impcrtanti members. more bolts or rive: f. (hj -g will be required and the degree of adjustment required on each will be le:s. Lesser adj
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s ATTACHMENT E i
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/ HEOL.SA.1769 I
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AliALYSIS CF P!P!!!G SYSTEMS WITH NCNL!! TEAR Sf1PPORTS SUS.:ECTED TO SE!.WC LCA0!NG l
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i'. 'esl-- ((3 , Maren 1979 {'. .,G ** ' s / I. /
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ASME PSPY THIRO NAT!0ftAL CDiGAESS [ June 1979 - San Franctsco. Ca. [ a'd o MM (MINttRM OMWMtM 'M.MMM I' 0, i.4 = w e.a.no.iereCni.m.eme e,en - Mapesse Bestas Carpesense, ed. the Deparament of I, ' Isseg Camerset Na, ET.7541621M emntowf uctr.st rtoTiet M"4 7 9 ' '.*.~.*.* 3*, '"lll7. ".=O IlutfittNt*THlM OF THl3 IHW'L'MFNT IS IIMlJMITIU d . . te L Q, . w 4 e
dynamic degrees of freedom. These interfaces are described by specific forms 6 of forcing functions, for the linear analysis. Since the dynamic matrix is invertad.but once, option 2.5 is far more economical than the former, and ,i it was used.throughout;this investigation. ANALYSIS RESULTS ANO O!SCUSSIO.1 Single Degree of Freedom Results The response of the simplified spring / mass model shown in Figure 1 to the FFTF horizontal acceleration was detamined with damping asst. sed to be 2". of critical. Gap sizes were varied from zero to 1.0 inch (2.54 cm) and the oscil-lator frequency. calculated by disregarding gap effects, was varied from 1.0 to 30. Hz. The results, which are shown in Figure 7. reveal relatively s=all load magnification for all gas sizes at frequencies lower than the 2.5 Hz. seismic spectrum peak. At frequencies higher than 5. H2. Icad tagnifications due to impact against hard structure become increasingly larger with increases in both natural frequency and gap size. The response results shown in Figure 8 illustrate the effect of 20% of critical impact dssping on response attanuation. The nonlinear responses are less than the 21 of critical damped tir. ear response spectrum at all frequencies lower than 5. Hz. Furtnermore. the nonlinear responses remain lower than the peak of the response spectrum up to a frecuency of 30. Hz. The value of 20% critical damping for impact was chosen aroitrarily for this study. However, sosie fuel assembly impact tests (3) show maximum recounds of less than 301. The impact damping coefficient was shown in (3} to be related l to the coefficient of restitution by the following equation: 2 g, (-2:t//1-6) Where. 8 = Coefficient of Impact Damping C = Coefficient of Restitution R i. Maxieum rebounds of less than 30% approximate a 20% of critical impact damping i coefficient. Simulated Mechanical Snubber / Piping Results A simulation of the combined civil structure, snubber ar.d piping stiffness, shaus in Figure 2. was used to assess the sets. ate response characteristics over a range of snubber stiffness and damping values which were determined by ' snub 6er characterization tests. The assumed piping mass and stiffness corressenAto a 1.0 inch (2.54 cm) nominal diameter pim with insulation , f and filledt with sodium, and with a 141.5 inch (360. cm) span simply supported ' at each and to ground. A mechanical snubber with variable stiffness and - damping values wee assumed to be attached at one end to the piping at m4 span and at the other and to ground through civil support structure wi J variable stiffness. The snubher gap was assumed to be .030 inch (.0iG cm) : as measured by test. A range of snuther stiffness and damping values were usW which correspond to a small snutter of the type that are used to . support ses11 piping. The seismic loading was applied through large base sesses representing ground. Forces applied through the ground masses were - - scaled to produce base accelerations identical to the FFTF horizontal seismic motion. The civil support structure stiffness values were chosee , L to tune the systes natoral vibraties frequencies to 2.5. 5.0 and g.0 Hs.
- - - - - - - - - _ L_- , _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ __ _ _ _ _ _
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. TABLE 2 COMPARIION OF LINEAR SPECTRA AND NONLINEAR AttALYSIS FFTF PIPING SYSTEM S? LUBBER LOACS SEISMIC SNUSBER LOADS s LSF* l /5 Me SUPPORT LINEAR NONLINEAR h#a SPECTPA - AN/4.YSIS //~ /-<n f!O.
4140 / I-1 - X 5047 1423 Al 1 - Y 3289 3152 0 /6
' 2 - Z 5071 A 13 2552 1315 3 - X 1776 N
3 - Y 2391 i ' #' 4422 4376 j 4 - X 1327 486 171
'{ 4 5
Y Y 1819 2876 p.f i 5866 5514 f, o) 5 - I 2G92 2708 <70 7 5A - Y 5273 /, ' 8 5A - I 5826 6 - Z 10817 4716 L Ef 9369 8536 /, / o
, 7 - I 4852 3760 /,,79
~7 - Y ~
7A - X 9835 8994 t, of 4021 /,57 . 7A - Y 6G86 3670 2238 /. 4 v 9 - X 4227 3516 / 48 - 9 - Z . 6888 3418 AS/ 10 - Z _, a
'*1.0 L8F = 4.448 N.
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ATTACHMENT F APTR 12 3-E-2 {. ' 11.06.23 O o A SEISMIC ANALYSIS REVIEW -
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2 ORRIS BADRIAN ( EBRUAE*1977 S
*EBASCO SERVICES INCORPORATED .
NEW YORK
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TABLE 3 lb MAIN STEAM LINE (2MS-32-2SB) OBE STRESSES , 1 ( Unif'rm o Response
- Node Number Spectrum Method Time History Method Ratio of Spectrum (EQ) (PS) (PSI) (PSI) (SEC.) To History Methods 8 712 4,686 2,124 2.88 2.2 34 2 2,084 1,508 -2.97 ,
1.4 1,153 835 2.97 1.4 35 4 1,234 683 2.95 1.8 2,229 1,225 . 2.95 . 1.8 . 36 5 1,601 503 2.83 3.2 , 886 287 2.83 3.1 37 6 4,609 1,976 2.95 2.3 1,985 2.95 38 8 1,461 1,74'4 3.12 .84 1,744 3.12 39 850 1,388 691 3.62 2.0 692 3.62 40 9 1,635 421 6.01 3.9 2,956 747 6.01 3.9 41 10 3,291 1,075 5.78 1 3.1 1,821 607 5.78 . 3'. 0 i 42 12 3,324 2,403 9.85 1.4 l l
- Does not include relative anchor displacement effects
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- 1. Uniform Response Spectrum Method - Pipestress 2010 Calculation Number 1157 (July 6,1976)
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ATTACHMENT G { + i SEISMIC ANALYSIS OF PIPING WITH NONLINEAR SUPPORTS
- 0. A. Beren. 8. N. Hussig and L. K. Sevensd Plant Analyss. FFTF Proiect Westinghouse Henford Comseny l RicNens. Mngton J l
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l i ABSTRACT The modeling and results of nonlinear time-history seismic analyses for three sizes of pipelines restrained by mechanical snubbers are presented. Ranerous parametric analyses were conducted to obtain sensitivity information which identifies relative importance of the model and analysis ingredients. Special considerations for "W the pipe clamps and the mechanical snubbers based on experimental characterization data are discussed. Comparisons are also given of seinic responses, loads and pipe stresses predicted by standard response spectra methods and the nan 14naar time-history methods.
. IETEODUCTION In the high temperature and seismic envirotment of a nuclear power plant, i
such as the Fast Fluz Test Facility (FFTF), mechanical type snubbing devices are used to restrain piping actions during seismic events. The snubber is
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connected at one end to the building support structure and at the other and to piping thrauste a clamp assembly. Daring normal plant cperating conditions, these snubbers offer no significant resistance to thermal growth of the piping. thder seismic conditions the snubbers lock up at very low acceleration levels. ! Dynamic characterization tests were performed at HEIL on mechanical snubbers j to determine their stiffness, dynamic loss of motion due to free play in the support linkages, and conditions of lockup. The tests also disclosed that the r snubbers dissipate a large amount of energy per cycle which could provide a 4 significant enount of piping systes damping that may exceed that commonly Jaed for seismic response spectra analyses of nuclear piping. A [1]k (s described in U. S. Nuclear Regulatory Ccussission Guides, such as in2], two meth spectra or the time-history method. Due to theoretical limitations and computer implementation difficulties, the response spectra method can not be used to w , evaluate scae of the essential support characteristics such as free play and v local dampir.4 at snubber locations. Thus, the response spectra method is limited to linear analysis. On the other hand, the time-history method is able to inelnde support nonlinearities and local desping in the analysis but with h e in parentheses indicate refersones listed at the end of the paper.
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TAa12 3. 4 1NCN FIFELINg NANGEn IDAD % lB. (1 13F = %.bb8 5) NAN 083 auN 12 auN 13 RUN 7 RUN 8 BUN 9 pun 10 RUN 11 RUN 2 DUN 3 BUN b BUN 5 BUN 6
- 30. RUN 1 39 38 39 28 39 38 39 39 N - 9T 9h 134 107 39 39 o o 72 109 0 16 0 o o o N 8K 62 168 105 51 53 b4 '
51 51 50 51 51 50 N - 77 77 192 14% 51 65 79 k80 518 107 89 151 70 70 N - 6K 529 946 bli 70 30 33 3h 43 33 33 33 26 39 N - ST 105 125 91 33 0 o o 198 199 0 62 23 33 0 N - 51 b39 648 200 22 22 25 35 22 22 22 19 25 N bu 67 152 113 23 o o o 128 214 0 13 o o o N b5 191 269 156 36 29 31 31 36 31 31 31 31 25 M - 3T 73 175 115 0 0 o o 97 96 0 22 9 13 N - 31 52 176 99 38 55 b8 h1 55 43 211 48 68 48 68 N - 2T 36 270 o o o o 84 113 142 0 13 o o 241 N - 22 79 183 15 23 lb 26 210 180 20 60 40 65 21 45 t M - 11 142 359 22 18 25 22 21 82 22 22 22 M - IT Sb 131 e IDCATION MAXIMM PIPING STRESS % PSI (1 EBI = 6.89 les) 3.3 NODE 1 12.1 i NOT13 1. PIPE 3D analysis with rigi4 supporte and using enveloping design seismic spectra of Figures 10 and 11. 2. PIPESD analyals with support and clamp flexibility and using enveloping seismic spectra. 3. b. ANDYS model with support and clamp flexibility same as in RUN 2 but using calcula enubber stiffness and demping characteristics. Snubber gape = .030 in. (.076 cm).
- 5. same as aUN h but with snubber gape = .005 in. (.013 cm).
- 6. Same as Run b but with snubber gape = .015 in. (.038 cm).
- 7. Some sa DUN % but with enabber demping reduced 505,
- 8. Some se RUN b but with elemp stiffness decreased 305, 9
Same se DUN b but with clamp stiffnees increased 305 .
- 10. Some se SUN b but with time history compressed 10#.
- 31. Beme se RUN h but with time history espanded 105.
. 12.
13 Sene as BUM b but with onubber desping scaled in proportion to the energy across an4 spatina responses combined accordind to Regulatory Guide 1.92. Article 2.2.
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gm' ) 352 326 M8 3% 396 f a.bT 254 % 0 0 0 0 962 5 5Y 2883 o 2262 2178 2262 2172 2190 6 3 51 12566 2139 482 un 3 5 o 0 12 6 15 69e6 5 i~ I . Sar 5 = 5A1 9357 3237 3262 3236 2989 2999 3262 416o 3996 3719 LIST h150 E.63 % 525 3896 I 1526 1613 1hoS 1707 1672 h378 E.7I 77:12 2189 2241 2232 3818 lokal 1983 2219 E 7T 3212 3209 5068
.ffccted h, E = TAE 85026 3275 330h 2187 3158 2153 2203 2209 kkT2 5 . TAI ik2ho 22ho 6251 6272 5945 6204 6168 6167 h I . 105 31668 3 the FIFD0 #538
- FBI (1 FSI e 6.89 epa)
FIFDS tg3 3L30W 30. Lysis dh32 5738 30h3 5767 5718 1 16185 3267 ho98 3902 3573 3739 3716 6 1584 3567 15h51 1 % 66 8 % 563 13526 15353 1h076 15662 23%h 21310 22131 22182 23075 9 36989 23o31 8196 7779 7965 8007 9189 , 11 63906 7831 a 3h223 35357 35372 36058 e 13 57h04 35037 35737 .
- BUIn: M 1. FIFW emetrate with emppers fleentaity, ustes 25 eseyes DEE esteste eyestra l and ICS nothat aseet s - a h . l ME 2. ===H=. stas history emelrete with enkbar stifftene and Wing ther. .J esternettee, sushkar eso = .o30 La (.076 ao). -~
asernes eless esitrasse. seruel time hintery. . M5 3 . sese as Its 2 hs vita time histery tempressed 105. j' Em 4. emme es its 2 tot vita time history espamest lo$. NN 5 . sees as mis 3 Det vita mLeamos elamp esitrones. EN 6. Same es Em 3 but with assima slag ettftsene.wtan someisse vita roepueses EE 7 useimme respeesee tras puss 2 4 ese tw s. esse h es eaemie ame P.eate leadings eseerting to Deselesary Guide 1.92. i Jrslake 2.2. t t i 21
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5-21 339bb 16041 5826 5854 8259 5816 5831 5502 6108 17387 11584 10266 < u-2Y 9606 9fb9 4593 4667 6119 4583 b602 % 31 6 6889 15152 8336 1042 3 5 - 2A1 2h016 13271 50be 5001 5927 5079 5021 5039 5133 lb563 10TT3 11076 5 - 2Af 8599 1324T 3439 3696 Th86 3438 34%c 3368 3502 10419 6776 6596 8-31 11100 30095 54T5 56bi 13221 5bi5 5476 5217 5666 32725 23603 2bo62 s-5Y 6949 8976 2657 2688 kb91 2665 2666 2b53 2801 8602 5945 5676 s-6* 1919 21628 5534 5665 11143 5538 5530 5266 5783 29380 22729 22216 5-1I 5596 1T521 2371 2376 hTT2 2266 2419 21?.2 2465 36055 5962 8870 1770 3152 166T 1828 158b 1821 5849 kn60 4281 3-7Y T588 7496 1739 1604 1811 15917 5-9I 4654 2037% 1722 1846 2860 1719 1 Tit 5815 7924 5-9Y 6501 266b6 1523 1674 7804 1525 1520 lbO2 1697 18368 9033 8b92 5 - 11E b876 lb215 2891 2993 6318 2090 2092 2669 3034 14155 11180 8866 B - 121 5580 20670 3585 36T1 6973 3579 3607 3bl6 3756 15809 12332 1153b Fir 18G ELBOW SThESS % FSI (1 P01 = 6.89 epa) g glJ0W 30. 1 7092 20598 7513 1500 _ 1510 1515 Tib8 Tb51 28320 15163 16T12 5 16214 11521 861T e564 _,,, 8630 8608 $ lib 8774 21053 'lb120 10631 6 InTb5 16779 6316 6318 _ 8362 6380 5892 6565 29304 16305 15338 7 8T45 26512 5328 5315 _,. 5335 5320 5698 5162 25104 1338a in861 10 13068 28703 6865 7038 _,. 6066 6861 16b2 Tbb2 2b536 1591b !?bO3 12 8854 19308 5037 4982 ,,,,,, 5033 503T 4895 521b 16959 12354 10648 BOTES: BUE 1 - FIPSSD llamar analyete with reactor veneet ledge flealbility but with riste pipe supporte. RUB 2 - FIP3so Itnear analysts with reactor vessel ledge floalbility and with flesible pipe supporte. PDA 3 - ANDTS maalinear analyste with reactor weseel ledee flealbility, average clamp ettffnese, enutber test date of attffases ese desping characteristice, saubber sape = 0.030 in (0.076 cm). RUE % - Sene.se sue 3 but with enutter gape = 0.015 la (0.038 cm). SUB 5 - Same se stb 3 but wlth stb 2 support ettsfassees sad saubber empinge e 200 It> sec/la (350 E-eec/ca). BUR 6 - Some as 915 3 but alth slalanes blastal cleap ettffneeses at 5-7 sad E-9. Sus 7 - Same se puu 3 test with maata= blastal cleap ettffnessee et E-T and 5-9. NW 8 - Same as Suu 3 but with time history compressed 105. Em 9 - Some se DUE 3 but with time history espande4105. aus to- same as sus 3 but with zero seaplag and 1.8 times the bortsontal acceleretton/ttee history. RUE 11. Some se 305 to but with emubber deeples asaled down is pesportlos to the energy across the emutAer. BUR 12- Some ao RUS 11 but with spatial cyante of eetente motion opptted separately ama opettal re-sponsee combined accoretag to Segulatory thatae 192 Artlete 2.2. _j}}