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Ty Lin & Associates Calculation, Capacity of Liner Under Combines B-Axial Compression & Shear.
	ML18102B027
Person / Time
Site:	Salem
Issue date:	11/02/1969
From:	; AFFILIATION NOT ASSIGNED
To:	;
Shared Package
ML18102B015	List: LR-N970230, Forwards Response to RAI Re Request for Change to TS Re Containment Design Temperature,Dtd 961011,970203 & 0321; ML18102B019; ML18102B020; ML18102B025; ML18102B027;
References
	NUDOCS 9705050081
	Download: ML18102B027 (129)
	v • d • e

{{#Wiki_filter:ATTACHMENT 5 T. Y. LIN & ASSOCIATES CALCULATION CAPACITY OF LINER UNDER COMBINED Bl-AXIAL COMPRESSION AND SHEAR

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ATTACHMENT 6 PSE&G RESPONSES TO NRC RAI ON SALEM PSAR QUESTIONS 10.3 b, 10.3 c & 10.4

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Additional Information Requested April 9, 1968 10.3 (b)

QUESTION 10.3 (b)

Describe: (b) The analytical procedures and techniques to be used in liner anchoraee design including sample calculations . ANSWER The lin'!r 'Was originally checked for anchor spacing of 30" x 30" in the dome and 15" vertically ....,ith 24" horizontally in the cylinder as shown in the PSAR Pages 5.1-24 to 5.1-26. The ans....,er to Qu~~tion 10.2 (d) shows a revise~ anchor spacing of 20" x 20" in the dome and 15" vertically ...,1th 16" hori-

        %On tally in the cylinder, which results in a much higher factor of safety against    bucklin~.
                                     ~
   *'  The anchor which is a 1/2         Nelson Stud must resist tensile and__ shearing loads.

It is common conservative practice to assign thP. laternl load needed to prevent coluom buckling as 2% of the buckling load. The maximum compressive stress 1 was o!lpproximately 25,000 psi. The total ~oad per plate would be 15" x 3/8" x 25,0QO.psi

_141,0000. Therefore; the tensile load per anchor is 141,0000 x

.02

2,820# which yields a stress of 2,8200/.2 sq. in.
14,150 psi.

There is also the possibility of a negathe. pressure of 3.5 psi, with the

      *maximum tensile load per anchor .. 3.5 psi x 20" x 20"

1~400#.
The shear load ~n the gnchor will be due to strain in the liner. Assuming the strain to approach yield of .1%, the anchor deflection "Would be 20 11 x .OC.l ..

02". The anchor ia more flexible than this. (See PSAR Page 5 .1-29 and ansver to Question 10.2 (d).)

i. AMENDMENT 6 I i. 1* ~, .*

      .:.. Additional Information Requested April       9, 1968                        10.3 (c)

(~)

 ~-..       QUESTION   l~.3  (c)

Describe: 0

          * (c)    The failure mode and failure prop agation characteristics of anchorages.

Discuss the extent to which these characteristics influence leak tightness integ:-ity. ANS\*rER The anchorages can fail by failure of the studs in shear or tension, by studs pulling out from the concrete or by studs tearing off from the liner plate.* The most likely mode of failure 1.3 by tearing .away from the plate. The anchors will be dc$igned so that failure o~curs in the anchor rather than the plate, thereby insuring that the leaktigh~ integrity of the containment liner will be maintained. Tests will be made to verify this. If failure sho~*d develope, it would _be a random stud failure due to poor workmanship during stud attachment. This would not impair thP. liner integrity nor would it cause progressive failure. The de~igr. load per anchor is low,and if an anchor should fail, the load it would have carried is easily distributed to the adjacent anchors. II* I I I I AMENDMENT 6

Additional Information Requested April 9, 1968 10.4 QUESTI~N 10.4 For the design of the anchors, elastic and inelastic buckling of the liner should be considered. The study should prove that no chain reaction can occur and that the possibility of ma~sive bucklinp, of the liner, and mass failure of anchors is excluded .. ANSWER We have sho'lffi under the answer to Question 10.2 (d) a sketch of a typical liner anchor. This design allows for differential movement between the plate and the concrete occuring during local buckling or for any other cause. The fact that the anchor will allow lateral movement ur.der severe loading conditions will prevent the anchors shearing off and the possibility of chain reaction anchor failure. i j i I

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AHENDHENT 6

ATTACHMENT 7 MODIFICATION SKETCHES FOR REACTOR COOLANT PUMP PLATFORMS AND CONTAINMENT SPRAY PIPING SUPPORT STRUCTURE

1EC-3450 Unit 1 Containment Spray Piping Support Structure MDs S1, 52, 53, 54 CDs 5503, 5504

Part 2 MODIFICATION AND TESTING MD NO.: Sf-~. FORM NC.DE-WB.ZZ-0001-15 CRANGE NO.: \ E.C. -"!>y :)C. SECTION 11.0 T-MOD NO.: _ __ MODIFICATION DOCUMENT COVER SHEET PACKAGE NO:~_\:;..__ __ CP REV. NO. : ..!!: 8 PAGE NO. REV. NO. 1 J1"1 2 0 3 0 4 0 5 JO/ JI' 7

                                                           *o      0 I

DOCUMENT CHANGED: DRAWING 223114 A 8989 NUMBER AND REV, THAT CHANGE IS Rev, 3 BASED ON) (PROVIDE TYPE, ACTUAL CHANGE TO BE MADE:

1. Modify structural connection "DETAIL M", Elevation 268'-2" (one location) by grinding out weld at gusset plate and replacing bolting material. S~e ,Pr:j 'Z ~F HD Ar l~c.._t::,.._q.

Slotted holes in horizontal 8WF member shall be elongated if necessary to satisfy gap requirements. Beam flanges shall be notched if necessary to satisfy gap requirements

2. Modify structural framing at Elevation 246' -2 11 (two locations) to f?eL,eve:ThG'<lllM~H. S'Ofs:s-s11rs e>u.E!."Th e.ua.v'...,f.o-ra-if.i!.tt.-i"11:r. Construction may choose OPTION 1 or OPTION 2 depen~ing upon accessibility.

See ~es .3 4 l/ 11>Pi:::11.5 MP r-r. /1>ca.:t: ** ,,'!:I. CONSTRUCTION OPTION 1

Modify structural connection "DETAIL K", {two locations} by grinding out weld at gusset plate and replacing bolting material ..

Slotted holes in horizontal SWF member shall be elongated if necessary to satisfy gap requirements. Beam flanges shall be notched if necessary to satisfy gap requirements CONSTRUCTION OPTION 2 See Mo S 2. for cut beam details. , Revise DETAIL P and Section 6-6 as shown on pages 7 ~ 6 respectively, of this MD (DETAIL P and Section 6-6 changes are for drafting only}

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BY & DATE & DATE MCRs DATE NO.

SUMMARY

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ATTACHMENT 8 SALEM PSAR PAGES 5.1-18 THRU 5.1-30
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Under an accident condition if one or more insulation panels are displaced from the liner, the liner will locally buckle at these places. Specifications for the attachment of the insulation and careful inspection of the installation will prevent to a great extent the possibility of dispiacement of the insulation. In any event, local buckling of the liner will not propagate failure in adjacent anchors because of the flexibility of the stud anchors which re-distribute the loads from the worst stressed point to lesser stressed areas.
Quality of both materials and construction of the containment vessel will be assured by a continuous program of quality control and inspection by Public Service Electric and Gas Company, and/or its field representatives, and Westinghouse Atomic Power Division, as described in Section 5.1.2.7. 5.1.2.4 Design Stress Criteria The design is based upon limiting load factors which are used as the ratio by which loads will be multiplied for design purposes to assure that the loading deformation behavior of the structure is one of elastic, tolerable strain behavior. The load factor approach is being used in this design as a means of making a rational evaluation of the isolated factors which must be considered in assuring an adequate safety margin for the structure. This approach permits the designer to place the greatest conservatism on those loads most subject to variation and which most directly control the overall safety of the structure. In the case of the containment structure, therefore, this approach places minimum emphasis on the fixed gravity loads and maximum emphasis on accident and earthquake or wind loads. The loads 5.1-18
utilized to determine the required limiting capacity of any structural element on the containment structure are computed as follows: (a) c l.OD +/- O.OSD + 1. SP + 1. 0 (T + TL) + LOB (b) c = 1. OD +/- O.OSD + l.25P + 1.0 (T' +TL') + l.25E + l.OB (c) c l.OD +/- O.OSD + l.OP + 1. 0 (T" + TL") + 1. OE I + 1. OB (d) c l.OD +/- O.OSD + 1. lOWt + LOB + 1. OP b Symbols used in these formulae are defined as follows: C: Required load capacity of section. D: Dead load of structure and equipment loads. P: Accident pressure load as shown on pressure-temperature transient curves. T: Load due to maximum temperature gradient through the concrete shell and mat based upon temperatures associated with 1.5 times accident pressure. TL: = Load exerted by the liner based upon temperatures associated with 1.5 times accident p~essure. T': = Load due to maximum temperature gradient through the concrete shell and mat based upon temperatures associated with 1.25 times accident pressure. TL':= Load exerted by the liner based upon temperatures associated with 1.25 times accident pressure. T": Load due to maximum temperature gradient through the concrete shell, and mat based upon temperatures associated with the accident pressure. TL": Load exerted by the liner based upon temperatures associated with
E: =
the accident pressure. Load resulting from assumed design earthquake or wind whichever is 5.1-19
greater. E': Load resulting from assumed hypothetical earthquake. B: Load resulting from buoyancy affect of ground water. W* Wind load due to tornado. t" Pb: = Bursting pressure loading associated with a tornado. Load condition (a) indicates that the containment will have the capacity to withstand loadings at least 50 percent greater than those calculated for the postulated loss-of-coolant accident alone. Results of the analysis using load condition (a) are shown in Figure 5.1-5. Load condition (b) indicates that the c-0ntainment will have the capacity to withstand loadings at least25 percent greater than those calculated for the postulated loss-of-coolant accident with a coincident design earthquake. Results of the analysis using load condition (b) are shown in Figure 5.1-6. Load condition (c) indicates that the containment will have the capacity to withstand loadings at least as great as those calculated for the post-ulated loss-of-coolant accident with a coincident assumed hypothetical earthquake with no loss of function. Results of the analysis using load condition (c) are shown in Figure 5.1-7. In order to appraise the relative influence of the individual loadings that form input for the design, the following plots are provided: Load plots for the dead, pressure, liner thermal, seismic and wind loads are shown on Figures No. 5.1-8 through 5.1-10. The buoyant load does not affect the walls nor the dome, and therefore no plot for this load is given. 5.1-20
The temperature gradient through the wall is essentially linear on both the insulated and uninsulated portions and is a function of the operating temperature internally and the average ~mbient temperature externally. Accident temperatures mainly affect the liner, rather than the concrete and reinforcing bars, due to the insulating properties of the concrete. By the time the temperature of the concrete within the interior of the concrete shell begins to rise significantly, the internal pressure and temperature in the containment shell due to the accident, will have been drastically reduced from their maxima. The temperature gradient between the outside and inside of containment during operating conditions induce stresses in the structure which are internal in nature; tension outside and compression in the inside of the shell. The resultant force is zero. Loading combinations concurrent with these temperature effects may cause local stresses in the outside horizontal and vertical bars to reach yield; however as local yielding is reached, any further load is transferred to the unyielded elements~ At the full yield condition, the magnitude of final load resisted across a horizontal and vertical section remains identical to that which would be carried if the temperature affects were not considered. Thus the overall carrying capacity of the structure and the factor of safety of the structural elements are not affected. In the attached Figure No. 5.1-11, sketch 1 shows a section through the wall which may be either a vertical or a horizontal section; sketch 2 shows the initial stress distribution set up by a thermal gradient; sketches 3, 4 and 5 show how adding increments of axial tension causes the com-pression on the concrete to vanish and the tension in the steel to increase. When any of the rows of steel reaches yield, the additional load is trans-ferred to the other rows which are below the yield point. This stress 5.1-21
redistribution continues until in the condition shown in Sketch 5, the
magnitude of final load resisted will be identical to that which would be carried if the temperature gradients were not considered.
This approach will not*affect the overall carrying capacity of the structure and the factor of safety of the structural elements will be consistent. Ultimate strength design justifies this argument because a basic assumption of ultimate strength design approach is that of trans-ferring stresses through local redistribution and equalization. The mat will also be analyzed utilizing load conditions (a), (b), (c) and (d) with the inclusion of buoyant forces where they result in more severe stress conditions. It will also be analyzed for loads occurring only at operating conditions. If the loads resulting from wind on any portion of the structure exceed those resulting from earthquake, the wind load "W" will be used in lieu of "E" in the appropriate load condition. Although no hypothetical wind load will be assumed for load condition (c), a check will be made to determine the maximum wind pressure that is tolerable in combination with pressure in the containment. All structural components will be designed to have a capacity required by the most severe loading combination. The loads resulting from the use of these equations will hereafter be termed "factored loads." The load factors utilized in these equations are based upon the load factor concept employed in Part IV-B, "Structural Analysis and Proport-ioning of Members Ultimate Strength Design" of ACI 318-63. Because of the refinement of the analysis and the restrictions on construction procedures, 5.1-22
the load factors in the design primarily provide for a safety margin on the load assumptions. The design will include the consideration of both primary and secondary stresses. The design limit for tension members (i.e., the capacity required for the design load) will be based upon the yield stress of the reinforcing steel. No steel reinforcement will experience average strains beyond the yield point at the factored load. The load capacity so determined will be reduced by a capacity reduction factor "0" which will provide for the possibility that small adverse variations in material strengths, workman-ship, dimensions, and control, while individually within required toler-ances and the limits of good practice, occasionally may combine to result in under capacity. For tension members, the factor "0" will be established as 0.95. The factor "0" will be 0.90 for flexure and 0.85 for diagonal tension, bond and anchorage. For the liner steel the factor "0" will be 0. 9 5 for tension, 0. 90 for compression and shear. The liner will be designed to assure that no strains greater than the strain at the guaranteed yield point will occur at the factored loads. Sufficient anchorage will be provided to assure elastic stability of the liner. The liner for the dome will be 1/2-inch thick. This liner will be re-inforced with structural tees which shall be welded to the liner at about 60-in. centers in meridional and circumferential directions. In addition
to the tees, 1/2-inch* diameter hooked studs will be welded to the liner 5.1-23
at the center of each grid formed by the tees. Therefore, the maximum unsupported span for the dome liner will be 30 inches in both directions. The lowest three courses of plates (approximately 24 feet of height) of the liner for the cylinder will be 1/2-inch thick. The transition knuckle at the bottom of the cylinder will be 3/4-inches thick. The balance of the liner in the cylindrical section of the containment will be 3/8-inch thick. Where there are an excessive number of penetrations in one area, the thickness of the plate will be increased from 3/8-inch to 3/4-inch for reinforcement.
'The anchors will be 1/2-inch diameter hooked studs welded to the liner plate at a maximum spacing of 15 inches vertically and 24 inches horizontally. Around penetrations the spacing will be reduced to ob~ain more effective anchoring of the liner at these openings. The studs for -
both the cylinder and dome will be hooked around the main reinforcing bars. The bottom liner will be 1/4-inch thick plate spliced at about 8 feet on centers. Each splice will be backed by a continuous structural tee section which will be embedded in and anchored to the concrete. The anchors will be 3/8-inch in diameter with a maximum spacing about 18-in. on centers. To insure transfer of loadings through the bottom liner without breach of liner leakage integrity, the equipment loads will be supported as shown in Figure 5.1-12. Each liner plate splice in the dome, cylinder and base mat will be covered by a pressurization channel. The buckling of the liner plates will be analyzed by formulas taken from Timoshenko and Gere, "Theory of Elastic Stability". The dome liner is a 5.1-24
continuous plate reinforced with welded structural tees. It is therefore reasonable to consider the case of a rectangular plate with clamped edges subjected to pressure in two perpendicular directions. If we designate G><. as the horizontal or hoop stress and G j as the vertical or meridional stress, then for a square plate: s.~3 rr 1 E h2 Where E = modulus of elasticity l2a. 2 (l-'ZS 1 )
~ Poisson's ratio a length of each side of square plate h thickness of the plate The cylinder liner is a continuous plate supported by a rectangular grid of hooked studs. It is conservative to consider the plate as a simple supported rectangular plate subjected to edge loads in two perpendicular directions. The equation for critical stress for this condition is: ~>C.-+" GJ(~
o.Z
)= Ge (I +-a..'J. )'2 b'2.
and Ge =i rriE.h-z l'Za.1. (1- ~ i) where a = length of plate in horizontal or hoop direction. b = length of plate in vertical direction. The maximum stress condition for the dome liner occurs under the (LOP + 1. OTL" + 1. OE') loading combination. Maximum G:J equals -9. 4 ksi and maximum ~)(.equals -9.4 ksi, or the sum equals -18.8 ksi. In compression (G)I.. +- G'J) critical = -39.0 ksi. Therefore, the factor of safety for the dome liner is 2.08. It should be noted that the method of determining factor of safety in the manner used above is extremely conservative, since it takes
no cognizance that we are dealing with factored loads and a hypothetical 5.1-25
earthquake which has an inherent factor of safety itself. The maximum stress condition for the cylinder liner occurs under the same loading combination as for the dome liner. Maximum~ equals -22.9 ksi and maximum c;;"~ equals -5.S ksi. Therefore:
= (04 KSI In Comparison:
The factor of safety is therefore 1.25 for this case. Again this is a conservative method for determining the factor of safety. The liner plate can be out of roundness either due to fabrication or because of bending under construction loads, specifically pouring of wet concrete against the liner. It can be out of round because of a more pronounced curvature toward the concrete wall, or a curvature away from the concrete wall. Additional curvature toward the wall will not be harmful as the concrete will resist buckling, since any buckling failure must be in the opposite direction. Fabrication out of roundness, away from the wall would be small since it must overcome initial curvature in the opposite direction, and also would be noticed in the visual inspection of the liner before the concrete is poured. In any case if it occurred, it would be very local in character and the surrounding area would take the overload from this spot. Springing of the plate out of roundness due to construction loads would be resisted by the natural curvature of the plate and the bracing of the liner
- by struts and wales when the concrete is being poured.
5.1-26
The liner plate thickness is subject to a permissible underrun of only 0.01 inches to comply with the ASTM Standards. This will in the worst case increase the stress 2-1/2 percent, and still have a satisfactory factor of safety. The possible variation in yield point should have no bearing on failure by buckling. The specifications call for a minimum yield stress of 32,000 psi. Manufacturer's guarantee of this minimum yield point will be obtained. The only time there would be a danger of failure is when the yield stress falls below the actual stress in either of the principal directions. The highest compressive stress in any direction is 22.9 ksi. This means that there is a.margin of 24% to reach the specified yield stress. The type, character and magnitude of stresses to which the liner will be subjected during normal conditions, proof test and under accident loads are shown in plots on attached Figures No. 5.1-13 and 5.1-14. The cyclic nature of the loadings are at the maximum as follows: Proof Test -- no more than 3 or 4 times in lifetime of structure. Accident loading -- expected to never happen. Operating to shutdown -- no more than 100 times in lifetime of structure. The change in temperature from operating to shutdown condition is from 120°F to 60°F and the number of times given above is based on this occurring more than twice in a year for a lifetime operation of 40 years. Since refueling is only once a year, this figure is very conservative. The margin of capability of the liner plate is as given in the table.below:
5.1-27

Condition Normal Operating Temp. and Pressure
.o. T = 60°F p = 0 G Hoop -15.7 ksi
(;D Vert.
-19.7 ksi Margin of Capability 1.22 Proof Test AT= 0°F +25.0 ksi +19.2 ksi 1. 28 p = 54 psi Accident l.5P AT= 246°F +30.5 ksi - 9.0 ksi 1.05 p = 70.5 psi Accident l.25P AT= 225°F +27.7 ksi -21. 8 ksi 1.16 and Earthquake p = 59 psi EQ.= l.25E Accident 1.0P ~T = 189°F - 5.5 ksi -22.9 ksi 1. 25 and Earthquake p = 47 psi EQ.= l.OE' The margin of capability under bi-axial compressive stresses is conservatively based on simply supported plates. In actual condition the liner plate is continuous at its edges and therefore will tolerate higher stresses before buckling. The margin of capability is based on the specified minimum yield stress where tension stresses govern.
The liner will be attached to the concrete by means of welded stud anchors. Tests were made on this type of connection at the University of Illinois under the sponsorship of Gregory Industries Incorporated, the makers of Nelson stud anchors. The results of these tests are reported in a paper in Highway Research Record No. 76 "Fatigue Tests of Plates and Beams with Stud Shear Connectors" by J. E. Stallmeyer and W. H. Munse of University of Illinois and E. A. Selby of University of Toronto. The part of the tests conducted, that is of interest in the design of the containment liner, is the
study of the shear capacity and load transfer distribution of the connectors 5.1-28
and the fatigue life of the connectors and plates. Of particular interest are the tests made on beams in flexure while a severe shearing load was externally applied to the connectors. This would approximate the accident condition where the studs would be subjected to shear due to possible localized buckling of the plate, or a slippage of the plate past the surface of the concrete as the temperature rises sharply in the plate. In the tests the stress cycle used was approximately from 14.0 ksi tension to 28.0 ksi tension and failure was produced in approximately 2,000,000 cycles. For operating conditions, a series of tests on flat plate specimens are of interest. Here the plates were stressed cyclicly between compression and tension. The worst stressed condition which is that between 20 ksi com-pression and 20 ksi tension produced failure after 180,000 cycles. This
testifies to the structural integrity of the plate on which lugs have been welded. The presence of studs did not in any way influence the yield point or ultimate strength of the base material.
If a single anchor or a group of adjacent anchors fail due to poor work-manship or material, then they are not effective in pr.oviding lateral support to that section of the plate. The liner will tend to buckle locally. In Manual No. 21 of the Gregory Industries Incorporated, makers of Nelson Studs, results of tests made on different size and shapes of studs subjected to torsion or shear loading is published. When subjected to shear loading the anchor exerts pressure on the surrounding concrete and deflects. For a 1/2-in. diameter anchor the deflection reaches 0.10 inches at approximately 92.5% of the ultimate shear strength of the material. This flexibility of the anchors will insure that anchors surrounding the 5.1-29
ones that have failed will transfer the load to other anchors and progressive failure will not take place. The axial force in the liner at the location where buckling occurs is equal to the critical load. In the adjacent area the forces are much greater due to the fact that the effective anchor spacing is much closer. The smaller the anchor spacing, the greater the capacity of the plate to resist buckling, and hence the larger the load. The inequality of the axial forces in adjacent areas places the effective anchors under a shear load. However, the flexibility of the anchors insures a redistribution of loads to the surrounding anchors and plates so that shear failure is avoided. In analyzing the liner around penetrations, it was considered that the
~I liner is basically not a load carrying member, but because of its integral relationship with the concrete it is subjected to the strains which the reinforced concrete imposes on it. The liner will be reinforced at each penetration according to the rules set forth in the A.S.M.E. Unfired Pressure Vessel Code Section VIII UG~36.
The weldments of liner to penetration sleeve will be of sufficient strength to accommodate the stress raisers around the openings. These welds will adhere strictly to ASME Section VIII requirements for both type and strength. In addition, each weld will have a pressurization channel placed over it which will add strength and stiffness to the welded area and assist in reducing stress in the weld and liner plate. 5.1.2.5 Foundation Design Criteria A field and laboratory investigation of subsurface conditions at the site
was carried out by Dames & Moore, Consulting Engineers in the Applied 5.1-30}}

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