View source for A06968, Responds to NRC 871215 Request for Addl Info on SEP Topic III-6, Seismic Reevaluation & Clarifies Questions Raised During 880406 meeting.W/13 Oversize Drawings

{{Adams
| number = ML20197F415
| issue date = 06/03/1988
| title = Responds to NRC 871215 Request for Addl Info on SEP Topic III-6, Seismic Reevaluation & Clarifies Questions Raised During 880406 meeting.W/13 Oversize Drawings
| author name = Mroczka E
| author affiliation = CONNECTICUT YANKEE ATOMIC POWER CO., NORTHEAST NUCLEAR ENERGY CO., NORTHEAST UTILITIES
| addressee name = 
| addressee affiliation = NRC OFFICE OF ADMINISTRATION &amp; RESOURCES MANAGEMENT (ARM)
| docket = 05000213
| license number = 
| contact person = 
| case reference number = TASK-03-06, TASK-3-6, TASK-RR
| document report number = A06968, A6968, B12936, TAC-51937, NUDOCS 8806130208
| package number = ML20197F421
| document type = CORRESPONDENCE-LETTERS, INCOMING CORRESPONDENCE, UTILITY TO NRC
| page count = 182
| project = TAC:51937
| stage = Meeting
}}

=Text=
{{#Wiki_filter:d NORTHEAST t#TILITIES                                        o.nor.i Orric.. . seio.n sireet. seriin. connecticui H'ARTFORD. CONN ECTicUT 061410270 L    L T2 C Z;O'l1"d"""                                               (200 665-5000 June 3, 1988 Docket No. 50-213 A06968 B12936 Re:    SEP Topic III-6 ISAP Topic No. 1.05 U.S. Nuclear Regulatory Commission Attn: Document Control Desk Washington, D.C. 20555 Gentlemen:
Haddam Neck Plant Information on SEP Topic III-6 Seismic Re-evaluation (TAC #51937)                                                   t In a December 15,         1987 letter (I)     the NRC Staff requested aoditional information regarding SEP Topic III-6, Seismic Re-evaluation for the Haddam Neck Plant. The Connecticut Yankee Atomic Power Company (CYAPC0) met with the Staff on April 6,      1988 in the NRC's Rockville, Maryland offices to answer questions on this SEP topic and ISAP Topic No. 1.05.
In Attachment 1 CYAPC0 formally documents answers to the questions posed in your {{letter dated|date=December 15, 1987|text=December 15, 1987 letter}} and provides additional clarification for the questions raised during our meeting en April 6, 1988.
If you have other questions please contact us.
Very truly yours, CONNECTICUT YANKEE ATOMIC COMPANY
                                                                /Y&          A/
E. J. Jlf o' czka (f Senior Vice President cc:   W. T. Russell, Region I Administrator A. B. Wang, NRC Project Manager, Haddam Neck Plant J. T. Shedlosky, Senior Resident Inspector, Haddam Neck Plant (1) See the A. B. Wang letter to E. J. Mroczka, "Haddam Neck Plant, SEP                                     j Topics III-2, III-4 and III-6," dated December 15, 1987.
8806130208 880603 PDR ADOCK 05000213 r-a                  conn _

Docket No. 50-213 B12936 Haddam Neck Plant Attachment 1 Response to Questions Regarding SEP Topic III-6 June 1988

o HADDAM NECK NRC'S REQUEST FOR ADDITIONAL INFORMATION ON SEP TOPIC III-6, "SEISMIC REEVALUATION" QUESTION II-A Adequacy of the interface between the pipe gallery and the containment exterior shell (Task 1).
: 1. The licensee should clarify how the 0.02 inch differential end displacement of the pipe gallery was obtained. From the licensee submittals, it appears that the displacement was due only from the NS earthquake component with the contribution from the other components excluded.

===Response===
This response addresses the concern as to how the differential displace-monts at the ends cf the pipe gallery (PG) vere obtained, considering that the Containment Exterior Shell (CES) and the Primary Auxiliary Building (PAB) are structurally connected by the pipe gal: ry structure.
The method used to determine the differential displacement was by per-forming Time History analysis on a coupled three-dimensional dynamic lumped mass model of the PAB ond the CES, connected by the pipe gallery (Figure IIA-1). Thia was the same model used in the parametric study performed to determine the effect of dynamic interaction between struc-tures and the me+ hod of developing the model and the member properties used in the study can be found in Appendix B to Reference 1(b). In the model, element No. 10 and No. 11'and their connectivity nodes 10, 11, and 12 represents the pipe gallery structurc.      The cross-sectional properties used in the analysis for the connecting member are shown in Figure IIA-2 and are extracted from Reference 2.
The time history analysis was performed both in the E-V and in the N-S directions for the coupled model. The ground input motion was in the form of acceleration time history shovn in Figure IIA-3.
The results in the form of time histories of displacement responses of nodes 10 and 12 are as shown in Figures IIA-4 to 7. Table IIA-1 gives the maximum relative displacements extracted from the time history plots and vere used to calculate the interaction forces at the interface of the pipe gallery with the containment exterior shell. The maximum displacement of about 0.02 inch is in the E-V direction, which is transverse to the pipe gallery, as compared to the displacement of 0.018 inch in the N-S direction. Thus, the use of the maximum displacements in both directions to calculate the end forces are included in the strength evaluation.
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Haddam Neck.--NRC'O Requ:2t for                                        Pags 2 Additional Information on SEP Topic III-6, "Seismic Reevaluation" QUESTION II-A
: 2. The licensee should provide justification for the adequacy of deter-mining the end forces /monents of the pipe gallery considering only the differential end displacement and not the combination of the effects of differential displacements and the pipe gallery inertia load.

===Response===
The end forces / moments of the pipe gallery were determined for the maximum relative displacements of the pipe gallery in the E-V and in the N-S directions.
In the dynamic model, both the PAB and the CES vere assumed to have fixed boundaries at their foundation level; the pipe gallery connects both structures at close to the foundation level and is assumed structurally decoupled in the vertical direction.           Therefore,  the evaluation of differential displacement in the vertical direction has been neglected.
The interaction forces calculated for the pipe gallery / containment (Node
: 12) assumed that the end of the pipe gallery is being fixed.        The review of the concrete drawings, Attachment IID-lii, indicates that the con-struction of the pipe gallery is not an integral part of the containment vall. The construction joint at the interface of the PG/CES provides a significant amount of relaxation to the constraint forces.        As concluded in the general discussion of the pipe gallery interface, the loads on the containment exterior structure justify that the structural evaluation is very conservative.       However, the pipe gallery inertia forces in the vertical direction vere determined and combined with the pipe gallery interaction forces while evaluating the local effects in the containment exterior shell.
The end reactions have been determined from the known displacements along with the applicable stiffness terms used in the seismic analysis.           The pertinent terms are listed in Figure IIA-8.       The summary of the interface loads extracted from References 2 and 3 are listed in Table IIA-3. There-fore, response to parts 1 and 2 of this question, both directions of dis-placement and inertia effects vere utilized in the evaluations eventhough the resulting forces deemed insignificant.
QUESTION II-A
: 3. T!.e licensee should provide the procedures used to determine the local shell stress at the interface using the end forces and moments on the l        pipe gallery as applied loads.
(
l   , Response The end forces and moments generated at the ends of the pipe gallery were
!    imposed on the containment structure to evaluate the local ef fects. The l   method of analysis, the assumptions made, and the boundary conditions l   used, extracted from References 2 and 3, are summarized herein.
l l .
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l Haddam Neck--NRC's Requ;st far                                          Pagn 3 l Additional Information on SEP                                                  l Topic III-6, "Seismic Reevaluation" l
l Summary of The Analysis Procedure and Results                                  l In order to determine the local shell stresses at the containment inter-face using the pipe gallery's end forces and moments as applied loads, a finite element model of the containment exterior structure was developed.
The computer code SAP-4 vas employed to perform this analysis.
Quadrilateral plate /shel) elements were used in modeling the cylindrical structure.       Since the pipe gallery end reactions dissipate quickly, a short cylinder rather than the entire containment structure was used.
Also, in order to take the adva'tage of plane of symmetry of the cylinder, only one half of the cylinder was considered such that there is at least one axis of symmetry existing. Therefore, in this case, the plane of the symmetry selected cuts through the center of the cylinder and the center line of the pipe gallery.        Thus, the finite element model generated is shown in Figure IIA-9; and for clarity, the location of the pipe gallery interface has been indicated by cross-hatched section.
The pipe gallery's end forces and moments vere applied as joint loads at common nodes where the pipe gallery connects to the containment shell.
Since only one half shell was considered for the analysis, the loads have been categorized into symmetric and anti-symmetric depending upon the direction of the input loading.       Based on this, the axial load F(y), the vertical load F(z), and the bending moment M(x) were categorized as symmetric because they produce the symmetric shell behavior about the axis of symmetry.       Similarly, the shear load F(x) and the lateral bending moment M(z) vere categorized as anti-symmetric loads. The mode of appli-cation of these loads are diagrammatically indicated in Figure IIA-10 and
: 11. The breakdovn of the end forces into the containment's nodal loads from each component of the pipe gallery's end reaction and the summary of the loads applied at each node are tabulated in Figure IIA-12, 13 and 14.
The boundary conditions imposed on the one half shell model vere such that the degrees of freedom are fixed or released forcing the model to act the same way as the actual shell structure vould behave under applied loada (see Table IIA-2).
The results of the analysis as element forces are tabulated in Table IIA-4. These forces are local to the containment shell structure and have been added to the other forces during the structural evaluation of the containment structure for SSE loads.
QUESTION II-A
: 4. The licensee should address the adequacy of maintaining boundary at I           the interface of the Containment Exterior Shell and Pipe Gallery.
l l
l

===Response===
In the coupled dynamic model of the containment exterior shell and the PAB vith the pipe gallery used as a connecting member, one of the assumptions made in the modeling is that the pipe tunnel rigidly connce.ts with the I
containment exterior shell.
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Haddam Neck--NRC's Requ;st fer                                       P gn 4 I
Additional Information on SEP Topic III-6, "Seismic Reevaluation"                                               i Although the results of the seismic analysis indicates very small relative displacements at the ends because of its stiff nature, it has developed very high end reactions. This fixed end assumption is an upper bound and the forces are very conservative.      These forces are imposed on the con-tainment and evaluateo with other containment design loads, in evaluating the structural strength.
However, the review of the as-built drawings indicates that the pipe gallery valls connect to the containment with a 6-inch by 2 1/2-inch shear key along with No. 4 dovels spaced at 12 inch 0.C. This type of detail more realistically represents a hinge connection; and if it was assumed in the seismic analysis, the moment reactions vould have disappaared and in turn resulted in lover stress levels in the containment vall section.
Therefore, whether it was assumed fixed or hinged, the interface boundary would remain integral and sound. Moreover, the structural connectivity of the pipe gallery is limited to the concrete portion exterior to the Con-tainment Shell. Therefore, the local stresses developed in the vall have little or no influence on the containment pressure liner which is interior to the boundary to the con-containment vall. Thus, the liner, as a tainment, vould also maintain its functional integrity.
General Discussion of the Pipe , Gallery Interface Loads on the Containment Exterior Structure The forces evaluated at the end of the Pipe Gallery (PG) with the Containment Exterior Structure (CES) was due to the PG's end relative displacements along with its own Inertia forces in the vertical direction.
The rigid connectivity of the pipe gallery with the containment In vall, tnat reality was assumed by the URS/ John A. Blume Associates as monolithic.
the fixity cannot be achieved with a standard shear key type of construction joint with no reinforcement embedded.
Fig. IIA-15 is the plan view of the construction joint, with #4 devels pl_eed at 12 inches on centers. These dovels are not enough to resist an axial force of F(y) = 5100 kips (Table IIA-3) if not all of the        other Therefore, forces like moments and shears tabulated in the Table IIA-3.
the meaningful application of PG's end reactions as an applied load on the containment exterior shell vould dissipate due to the relaxation of the         '
construction joint before it engages the CES.
In addition, after review of the concrete drawings the pipe gallery's placerent of concrete was done after construction of the containment vall in a sequence, the base slab first, folloved by the vertical vall and the placement of concrete for 18 inch roof slab would be the last.
In a situation like this, the shrinkage of concrete at the construction joint    (CJ) would be significant and the magnitude can be fairly predictable by the approach outlined and reported by the ACI Committee 209, ACI 209R-82, "Prediction of Creep, Shrinkage and Tempetature Effects in Concresa Structures," Reference 7.
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Haddam Neck--NRC'c Requist fer                                              P gt 5 Additional Information on SEP Topic III-6, "Seismic Reevaluation" The Haddam Neck Plant was constructed prior to 1967, and tne concrete placement shrinkage has already attained its "Zero" level. By using this data and the "As-built" construction sequence, the amount of shrinkage can be estimated both in the horizontal as well as in the vertical directions in terms of strains (c).
From Reference 7.
sh (U sh u sh t Vhere:
t - time in days (c,3), - shrink.tae strain, inches / inches (e,3), - ultimate (in time) shrinkage strain
                                  - 530.4 x 10'' inch / inch Therefore:
( c,3 ) - 233.4 x 10-' inch per inch t - 28 days DIFFERENTIAL DISPLACEMENTS COMPARISON DIRECTION            SHRINKAGE            SEISMIC            REMARKS (INCHES)            (INCHES)
N -S DIRECTION                                             Length - 10 ft.
(xXIAL)               0.028               0.018         (half of the pipe gallery)
     ~E-V DIRECTION                                              Length - 23.5 ft.
(TRANSVERSb)     l      0.066               0.02          (half of the
            -                            ~~                        vidth of the PG.)
     ~ VERT.                                                     Length - 18 in.
0.0042              0.006         (depth of the top slab)
By observing the above differential displacements, the gap created at the construction joint due to the shrinkage of the concrete is greater than the differential displacements due to a seismic event.
Therefore, it can he concluded that the relaxation of the construction joint at the interface of the pipe gallery with the containment vall, not only decouples the structure's frequency response but also eliminate the pipe gallery's and reactions. The structural evaluations made by URS/ John A. Blume Associates of ti e containment vall using the pipe gallery's end forces as an applied load to the containment is very conservative and an upper bound value, eventhough they are not realistic.
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Haddas Neck--NRC's Requ;It for                                         Pags L Additional Information on SEP Topic III-6, "Seismic Reevaluation" OUESTION II-B Radial beams at che floor of containment internal structures (Task 2).
: 1. Inconsistent statements between the letter from V. Counsil of CYAPC0 to D. Crutchfield of NRC dated January 19, 1984 and the February 1983 "Seismic Reevaluation of Major Structures of the (CYAPCO) Pover Plant, Volume II--Containment Structure," performed by URS/ John A. Blume Associates, makes it difficult to determine whether the radial beams or the ring beams are overstressed.

===Response===
There are no upper steam generator supports installed either to the radial beams or to the ring beams at the operating floor of the internal struc-ture. If they have been installed, the ring beams (not the radial) would have been overstressed.
URS/J. A. Blume & Associates performed Seismic Analyses of the RCS without the upper steam generator supports;          the evaluation resulted in an acceptable RCS condition with a minimum number of modifications.           The beams at the operating floor level, subsequently reevaluated for the loading conditions without the inclusion of loadings from the upper steam generator supports, vere found to be acceptable.
QUESTION II-B
: 2. The licensee should address the following:
: a. Indicate whether the upper steam generator (S/G) supports have been installed; and if so, provide the as-built details,
: b. If the upper S/G supports have been installed, provide the docu-ment (s) which indicate only the ring beams are overstressed together with the proposed modifications.

===Response===
Since there are no upper steam generator supports installed, the response I   to the above questions is not required.
i   OUESTION II-B l
l   2. The licensee should address the following:
: c. If the upper S/G supports have not been installed, provide the document (s) which indicates that the radial beams were reevaluated and found acceptable as stated in the letter from V.      Counsil of l
CYAPC0 to D. Crutchfield of NRC dated January 19, 1984.
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                                                                            "J Baddas Neck--NRC'c RequIst far                                            Pagm 7 Additional Information on SEP Topic III-6, "Seismic Reevaluation"

===Response===
To determine the floor beam's actual member forces, they were evaluated by performing the Finite Element Method of Analysis.              The loads and the loading combination used as an evaluation criteria from Reference 1(a) is as follows:
U - D + L' + 0 + E Vhere U = Total load to be resisted D - Dead Load L' - Actual Live Load (including piping and equipment loads) 0 - Operating Loads, if any E = SSE Loads (the results of E-V, N-S, and vertical direction of motions were combined by the SRSS method)
The beam's location, for details see Attachment IID-ii, and their nominal sizes have been identified in Figure IIB-1. The capacity of the beams vere evaluated using the Strength Design Method according to the provi-sions of References 5 and 6, including appropriate capacity reduction factors. Table IIB-1 provides the comparison of the beam's capacities vs.
the actual forces and moments developed in each member for the expected loads in the loading combination.
The results indicate margin of safety exists in the radial beams, as well as in the ring beams, for the limiting loads during a seismic event.
QUESTION II-C Anchorages for the Refueling Vater Tank and the Demineralized Vater Storage Tank.      (TASK 4)
: 1. To date, information on this issue has not been received by the staff.
: 2. The licensee should provide information addressing this issue, or provide a schedule for doing so.

===Response===
for the Demineralized Vater Storage Tank (DVST)   was The analysis report provided to the staff as part of Attachment 1 to V. G. Counsil's letter to D. H. Crutchfield dated February 9,1983. The analyses reports for both the DVST and the Refueling Vater Storage Tank (RVST) vere included in the ten-volume "Final Report - Hechanical and Electrical Equipment - Seismic Reevaluation Program - Haddam Neck Nuclear Generating Station" prepared by Stevenson Associates, September 1983. This report was forwarded to the staff     by V. G. Counsil's letter to D. H. Crutchfield dated January 19, 1984.      The analyses of both the DVST and RVST (again provided herein as CIVIL 1.2

Haddam Neck--NRC'c Requrat fer                                         Pagn 8 Additional Information on SEP Topic III-6, "Seismic Reevaluation
* Attachments IIC-1 and IIC-2), indicated that modifications to the anchorages for these tanks vould be required for seismic adequacy. The attachment to         V. G. Counsil's letter to D. H. Crutchfield dated January 19, 1984 titled "Safety Related Equipment (Mechanical and Elec-trical)" stated that, "Modifications indicated necessary to the Demineral-ized Vater Storage Tanks and Refueling Water Storage Tank have been completed".
The modification for the DVST consisted of the bumper and tie rod system attached to the tornade vall as presented in the analysis report - see Attachment IIC-1. This modification was completed in November 1983.
The modification for the RUST consisted of adding eight (8) additional 1-1/4" anchor bolts and anchor bolt chairs as presented in the analysis report - see Attachment IIC-2.      This modification was completed in March 1983.
QUESTION II-D Q.II.D Structural Integrity of the Auxiliary Feedvater Pumphouse (AFVP).
(TASK 5)
Response Introduction The auxiliar" feedvater pumphouse is a light steel framed structure located to the vest of the containment exterior shell. Figure IID-1, shov the steel framing at EL. 65'-6" supports the asphalt roof and four main steam line spring hangers. The intermediate floors consist of very light stcel platforms and are typical of those in Figure IID-2. The struc'.ute is supported on one side by the containment and on the other side by steel columns on 3 feet thick concrete shield vall at G'-Line. The tornado vall, which supports the dead veight of the piping, is interconnected to the shield vall by 2 feet by 2 fee t concrete beams at EL.      39'-6",  along l   vith the steel bracing.
The enclosed drawings in Attachment IID (16103-22090, SH.8, 10, 20 and 34 l   and 16103-51048, SH.1 to 4) provides the framing details and the beams l   connecting details of the concrete valls.
The evaluation report provided to the staff as part of Attachment 1 to V.
G. Counsil's letter to D. H. Crutchfield dated May 2, 1984, entitled "Seismic Reevaluation of the Auxiliary Feedvater Pumphouse at the Haddam Neck Nuclear Pover Plant, Hay, 1984" vas due to the gravity load, the estimated seismically induced pipe reactive loads, in addition to the seismically induced boundary displacements at the containment interface.
l l
The primary load in this preliminary evaluation vere due to the main steam and the feedvater systems, conservatively estimated preliminary loads due to SSE. Based on the preliminary loads, modifications to two columns and three beams at G'-Line vere proposed pending a confirmatory analysis on the structure after systems verification analysis.
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l Haddam Neck--NRC's Requzst fer                                        Paga 9 Additional Information on SEP Topic III-6, "Seismic Reevaluation" l
At present the structure encloses the isolation valves of the main steam       l system and pipe supports for the main steam and feedvater systems. The roof framing supports four spring hangers for main steam piping. Under this loading the structure does not require any modifications, and vill maintain its structural integrity even in an SSE event.                        l Therefore, the proposed confirmatory structural analysis includes:
i)   Seismic Analysis li) Static Analysis The seismic analysis of the structure vill be performed by developing a lumped-mass model and the masses are connected by weightless elastic members. Figure IID-3, the cross-section of the AFVP looking north, show the relative locations of the steel platforms with respect to the shield and tornado valls. Figure IID-4 is the proposed seismic lumped mass model which represents the behavior of AFVP under seismic excitation.             In general the mass locations in the model are at the platforms level. The member connecting the masses are veightless elastic springs, as indicated in the seismic model,                                                    i The support of the springs are assumed fixed at the base, EL. 15'-6" and hinged condition of all the supports interface with the containment vall.
The Hodal analysis response spectra method vill be employed using ground response spectrum, Figure IID-5, as an input motion.             Seven percent structural damping vill be used for SSE to determine the maximum response of the structure. The assumed support input motion at the base of the structure, EL. 15'-6",      may produce a higher structural response as compared to the response spectrum of the containment structure. Since the lateral supports of the platforms are hinged at the containment interface, the response spectrums at different heights may not drive the structure.
However, before passing judgement on which one controls, a study vill be made in determining the responses both ways and a conservative results vill be used for the structural evaluation. The modal responses vill be obtained by exciting the model in the east-vest, north-south and in the vertical directions, one at a time. The results of the analysis vould be in the form of the acceleration and the displacement responses at the mass points.
0.II.D-1.      The licensee should provide justification for:
a) not considering soil-structure interaction;

===Response===
The foundations for both the chield vall and the tornado vall are indepen-dent footings of a 5 feet effective depth to EL. 15'~6". The top of the rock surface is at about EL. 10'-0".
Since the foundation bottom is about 5 feet 6 inches off of the rock surface, the effective of soil structure interaction vill be assumed to be negligible because of the following:
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Baddam Neck--NRC'c R;quist for                                         Pags 10 Additional Information on SEP Topic III-6, "Seismic Reevaluation"
: 1. The depth structural fill is shallow, confined between two hard boundaries and becomes very stiff, with a very high frequency content.
: 2. The superstructure is very light and very flexible, which would nor-mally produce a lov frequency response.
: 3. As the structure is light, the foundation footings vill be lightly loaded.
: 4. Since the primary loading on the structure vould be from the pipe reactions, a substantial margin of safety vill be maintained against the foundation's bearing pressure.
Therefore, in the dynamic model, while performing the seismic analysis, the base of the model vill be fixed at the foundation level.
0.II.D-1. The licensee should provide justification for:
b) decoupling the turbine / service building from the AFVP.

===Response===
The mathematical model representing the AFVP vill be decoupled from the turbine / service building because they are physically separated by more than 25 feet. The only connection between the two structures is by two-30 inch deep beams at EL. 48'-9",  from F-line at the service building to the steel columns on G'-line.
The beams vere notched intentionally by cutting off the top and the bottom flanges of the beam.       This type of connection vould make the beam very flexible and, in addition to the buildings separation, justifies the decoupling assumptions.
0.II.D-2. The licensee should prov.ide detailed drawings of the beam and column connections to the containment exterior shell,           the tornado vall and the shield vall.

===Response===
The details for the beam column connections are similar to the typical of those suggested by AISC, Manual of Steel Construction.             However, the enclosed drawings provide "as-builts" both for the steel as well as the concrete connection details at the containment structure's interface, including the shield vall and the tornado vall, Attachment IID-i.
l 0.II.D-3. Provide clarification as to how the shield vall and the tornado vall, as well as their connecting beams, were analyzed.
l I
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Haddam Neck--NRC's Requ at fcr                                         Paga 11 Additional Information on SEP Topic III-6, "Seismic Reevaluation"

===Response===
The limiting load on the shield vall and the tornado vall vould be from the pipe restraint's reactive loads on the structure due to SSE. At present there are no pipe restraints and the size and the shape of these restraints vill be finalized pending piping systems analysis.
The confirmatory static analysis is to be performed by using the following loading combination:
U = D + L' + 0 + E Vhere:
Ur Total Load to be Resisted D = Dead Veight L'= Actual Live Load 0 = Operating Pressure & Temperature Load E = SSE Loads (includes structure's inertia effect and the pipe support loads)
The shield vall and the tornado vall vill be evaluated as shearThe     valls in beam the N-S direction and as a frame action in the E-V direction.
connection to the vall vill be evaluated locally in addition to the beam strength and vill be compared to the allovables per ACI 349-80, code requirements for nuclear safety-related concrete structures and the AISC Rth Ed., Dec. 1980.
0.TI.0-4. Provide clarification as to how the steel floor plates were included in the seismic model.

===Response===
l The steel floor plate is a cover plate on the steel platform grid.         The mass of the steel plate vill be included in the dynamic lumped mass model.
0.II.D-5. Provide quantitative results justifying the assumption that     a static analysis of the AFVP was sufficient.

===Response===
The structural analyses performed on the AFVP vas based on upper bound loads imposed on the structure to evaluate its integrity.         The primary load on the structure was an estimated pipe reactive load from the Main Steam and the Feedvater Systems due to SSE.      Since the structure is very l
l     light and its supporting beams vere assumed to be rigidly connected to the l     containment. The containment's SSE relative displacements were used, in addition to the piping loads, as its boundary displacements.
1 cIv!L1.2 l

Haddam Neck--NRC's Requ2st fer                                          Paga 12 Additional Information on SEP Topic III-6, "Seismic Reevaluation" However, the present scope of work contains the seismic forces resulting from the proposed dynamic analysis vill be added on to the static analysis to arrivo at the total effect.      The expected completion of this ef fort will follow after the piping analysis.
0.II.D-6. Provide clarification as to how the containment displacements were input to account for both direc: tons of horizontal motion.
0.II.C-7. Provide justification that the equivalent static analysis of the AFWP in the vertical direction using 0.3g vertical acceleration is adequate to cover the response of the higher structural modes.

===Response===
The confirmatory structural analysis vill be based on the structure's inertia forces rather than the boundary displacements. The member force evaluation vould be the SRSS value of the inertia forces due to the N-S, the E-V and the vertical directions.
0.II.D-8. Provide details of the design modifications made to the AFVP.

===Response===
Currently the Main Steam and the Feedvater Piping Systems are in the preliminary stages of reanalysis.       At present, there are no additional pipe supports installed.
The proposed modifications to two structural columns and three beams was based on the preliminary loads from the pipe supports were not made pending reanalysis of the systems. Under the present configuration, the structures require no modifications and vill maintain its structural integrity during an SSE event.
The structural evaluation for the loading combination, as stated in response to question II.D-3 and along with the acceptance criteria, vould indicate modifications if it became .necessary.
0.II.D-9. Provide an evaluation of the soil bearing pressure at the AFVP foundation.

===Response===
The spread footings (foundation for the shield vall and the tornado vall) vill be strength evaluated as part of            the confirmatory structural evaluation.       The    foundation's  strength   verification    includes   the determination of the soil bearing pressure and the safety factor against sliding.
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Haddam Neck--NRC's Requ;st for                                        Pags 13 Additional Information on SEP Topic III-6, "Seismic Reevaluation" 0.II.D-10. Provide clarification as to how the piping loads were obtained and what spectrum was input to obtain such loads.

===Response===
The AFVP contains four (4) systems that require seismic reanalysis. The systems are main steam, main feedvater, auxiliary feedvater, and main steam drains (main steam supply to the auxiliary feedvater pumps), all of which have existing supports located in the AFVP.
Both the auxiliary feedvater and main steam drain piping have been reana-lyzed. The main steam and main feedvater piping is currently in the pre-liminary stages of reanalysis and is expected to progress consistent with the dates submitted under the Integrated Implementation Schedule for ISAP TOPICS 1.04 and 1.05. The Integrated Implementation Schedule was sub-mitted by letter, dated November 13, 1987.
Preliminary piping loads have been developed using the ADLPIPE (Version ID) piping analysis program. Each piping stress problems was modeled and analyzed in accordance with the following two documents:
: 1. Piping Stress Analysis Guidelines for Connecticut Yankee, and
: 2. Connecticut Yankee Safety Related Pipir; Seismic Qualification Program Criteria Document-Revision 2.
Both of the above documents vere submitted to the NRC for their review and approval for use in the SEP effort at Connecticut Yankee.
Each system was analyzed for dead weight, thermal, and dynamic seismic loadings including ZPA's (zero period acceleration) and SAD's (seismic anchor displacement). Vhere piping is routed through several buildings or various floor elevations, an enveloped response spectra was developed and applied to the entire piping geometry.      PVRC damping criteria (Code Lase N-411) is employed in accordance with the requirements stipulated in the CY Criteria Document.
The various floor acceleration response spectra used in the piping analyses vere obtained from the site specific Connecticut Yankee seismic analysis prepared by URS/ John A. Blume & Associates, Engineers, Volumes I through V, dated December 1982.
QUESTION II-E Adequacy of the percentages of the modal mass participating in the modal mass participating in the structural response (TASK 7).
: 1. The participating modal mass for the Turbine-Service Building is adequate.
: 2. For the Primary Auxiliary Building, Screenvell House, and the Con-tainment Structure, the use of residual response to account for the CIVIL 1.2

Haddam Neck--NRC's Requrat for                                          Pags 14 Additional Information on SEP Topic III-6, "Seismic Reevaluation" missing mass of the structures is adequate, provided that the follow-ing items are satisfactorily resolved:
: a. Whether residual stresses vere included in the stress evaluation of the containment exterior shell and interior structure to account for the inadequate modal mass;
: b. The method was used to obtain the residual responses;
: c. Whether the residual responses and building foundation mass were included in the sliding and overturning analyses;
: d. How the effect of the missing structural mass was accounted for in the generation of in-structure response spectra.
: 3. The licensee should provide information on the percentage of modal mass used in the vertical analysis of the Primary Auxiliary Building.
Response to Questions II-E.2, a, b, e and Question II-E.3 Dynamic analytical models of the Primary Auxiliary Building, Screenvell House and Containment Structure vere developed for the subj ect study.
Hodal superposition response spectrum analyses vere performed, assuming that modes with frequencies of vibration greater than 33 Hz. are rigid.
These analyses were performed using URS/Blume SAP IV program, which includes provisions to include the effects of residual response from rigid modes.
Attachment IIE-1 gives a detailed technical discussion of the procedure used to calculate these residual stresses.         In brief, URS/Blume's LAP IV program calculates the residual inertia forces computer per equation (19) and (20) of Attachment IIE-1.         These inertia forces are then app 12ed to the structure as static loads to determine residual stresses.             These residual stresses are then added to modal stresses on an SRSS basis.
The particular questions raised in Question IIE-2d does not apply because in-structure floor response spectra are developed on the basis of a model superposition time history analysis. In these analyses, the time history of the sum of the modal relative accelerations (contribution of modes with frequency less than 33 Hz.) is added to the ground acceleration time history (contribution of rigid modes) to produce the floor acceleration time history. This floor acceleration time history is then used to generate floor response spectra.        This is a rigorous solution which does not make assumptions regarding residual or "missing structural mass".
Respor.se to Question II-E.c Copies of the sliding and overturning analyses for the relevant structures are given in Attachment IIE-2. The dynamic responses used in these calcu-lations are from the URS/Blume SAP IV analysis results. These do consider the contribution of residual modes, as described above.
c!VIL1.2

H:ddas Neck--NRC's Requ;2t fer                                       Page 15 Additional Information cn SEP Topic III 6, "Seismic Reevaluation" Attachment IIE-2 also shows that the mass of the foundation mat was con-sidered in these calculations for the Screenvell House and the Containment Structure. The Primary Auxiliary Building is founded on valls and columns which are embedded in soil and resting on vall or spread footings. The factors of safety against sliding and overturning moments for this build-ing are primarily a function of the mass of the building above grade, the soil passive pressures and the capacity of the valls to resist the passive soil pressures.      The mass of footings is negligible and almost equal to the mass of soil they displace.
References
: 1. Seismic Reevaluation of Major Structures of the Connecticut Yankee Atomic Power Plant:
: a. Volume I--Seismic Analysis and Evaluation Criteria, prepared by URS/J. A. Blume & Associates, dated December 1982
: b. Volume II--Containment Structure, dated February 1983
: 2. Calculation No. JAB-7937-04-CF-01; Containment Exterior St':ucture (Analysis), Development of Mathematical Model and Response Spectrum Analysis, by URS/J. A. Blume & Associates
: 3. Calculation No. JAB-7937-54-CF-01; Containment Exterior Structure (Evaluation), Structural Evaluation for SSE Loads by URS/J. A. Blume &
Associates
: 4. Calculation No. JAB-7937-55-CF-01; Containment Internal Structure, Structural Evaluation for SSE Loads by URS/J. A. Blume & Associates
: 5. ACI 349-76; American Concrete Institute,        "Code Requirements for Nuclear Safety-Related Concrete Structures"
: 6. ACI 318-77; American Concrete Institute, 3 Building Code Requirements for Reinforced Concrete"
: 7. ACI 209R-82; Hanual of Concrete Practice, 1985 Part I, Materials &
General Properties of Concrete.
KLipdm 5/19/88 CIVIL 1.2

TABLE 00 A-1 DIFFERENTIAL DISPLACEMENTS OF THE PG WITH CES AND PAB DIFFERENTIAL EXCITATION             DISPLACEMENT E-W DIRECTION (BX )        0.02 INCH 0.018 INCH N-S DIRECTION (h)
NOTE:
PG - PIPE GALLERY CES - CONTAINMENT EXTERIOR SHELL PAB - PRIMARY BUILDING l
l L-

i TABLE AD A-2 BOUNDRY CONDITIONS AT NODES A &             B b                  8      8y    82 x   Dy   Dz        x SYMMETRIC           1    0     0      0       1      1 LDADS ANTI-SYMMETRIC        0     1    1       1      0     0 l          LDADS i
l NOTE:
  !             O - MEANS; DISPLACEMENT IN THAT DOF IS FREE.
1 - MEANS; DISPLACEMENT IN THAT DOF IS FIXED.

Y A3(LE 00 A-3 INTERFACE LOADS AT PG/CES O       '   #
PG/CES    FORCE DUE TO FORCES    TIMi. HISTORY D.WT    FT. K]PS )
F(x )
1830.                  1830.
TRANS. SHEAR F(y) 5100.                  5100.
AXIAL F(2) i                                244.0     244.0 VERT i
M(x )                    610.0     610.0 M(z )     37,600.                 37,600.
1 l
1                                                      _

                                                                     )
TABLE 00 A-4 FORCES        IN THE         -
CONTAINMENT EXTERIOR SHELL TANG.      VERT VERT     HOOP                        HOOP LOADING     FORCE    FORCE fhff     MOMENT MOMENT KIPS /IN KIPS /IN KIPS /IN FT KIPSNT FT RIPSET PIPE 0AL'.EY SYMMETRIC       3.01    5.91     4.82      324.      178.
LOADS P]PE GALLEY ANTI-SYMMETRIC   0.32     1.66     1.83      73.5      70.
LOADS TOTAL       3.33     7.57     6.65      397.5     248.
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Beam element
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(Rigid link)                                                          ' 22 O 2I n      Node number 20
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jl9 PG     Pipe Gallery .                                                        p gg CES    Centainment Exterior Shell 18 o l'8 b
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CROSS SECTION                    OF PIPE GALLERY PROPERTIES:
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L = 20 FT. ( LENGTH OF THE PG. )
CONCRETE:                                                             )
fe' = 3000 PSI                                         :
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V     = 0.17                                            '
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           .02     . . . . l....l....l....l....l....l....l....l....l...~                                                        -
        .      0.00        2 00        4.00         6.00     8.00     10.00     12.00      14.00      16.00     18.00     20.00 T I ME (SECOND)
COUPLED MODEL OF ECS + PAB FIG llA-6
[

RESPONSE TIME HISTORY OF NODE 10 IN Y DAMPING = 0.07                                                    -
02      , , ,,      ,,,,    ,,,,          ,,,,     ,,,,      ,,,,    ,,,,    ,,,,     ,,,,    ,,,,
  ,                  01  _
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a                -
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                   .02       , .   ..l....l....l....l....l....l....l....l....l....                                                  -
                .       0.00        2.00     4.00         6.00      8.00      10.00   12.00   14.00    16.00   18.00        20.00 T I ME (SECOND)
COUPLED MODEL OF ECS + PAB FIG li A-7 i

FDGURE 00 A-8 8
E
:                                                     N00E Q2) l N00E 01) 1 3
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l           Fy = AE   ( gy )
i                   L l

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FDGURE 00 A-t@
CONTAINMENT STRUCTURE INTERFACE          -
F y --*   F(z )
Fy     :
p(z )
l                 Z                AXIS OF                          Z O        : -Y- - -                                 O    = -Y- - - - -
2 F y --* si 2F(z )                          -SYMMET
                                                 - - - - - R YFy - o F(z ) --                B A
Fy --          F(z )       1r                                                1r X                                                  X i
l i
l ACTUAL STRUCTURE                                            $SHELLMODEL FOR WITH APPLIED FORCES                                               ANALYSIS j
l                SYMMETRIC LOADS F(y)= 5100 KIPS F(z ): 244 KIPS M(x) = 610 FT. KIPS l                                                                                             ,.

o F0GURE 00 A-11                                                             !
L CONTAINMENT STRUCTURE ANTI       -
SYMMETRIC LOAD F(y )              DUE TO M(z)
      *               '      " F(x )                          AXIS OF                            U F(x )
()2  :  y___  _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ . ()Z : y __    _.
SYMMETRY                        A                          B n                                                                 n sr F(x)        X                                                                 X ACTUAL STRUCTURE                                                 {SHELLMODEL WITH                                                          FOR APPLIED FORCES                                                     ANALYSIS NON - SYMMETRIC LOADS F(x ) : 1880 KIPS M(z ) = 37,600 FT. KIPS

FDGURE 00 A-12 LOAD ( DISTRIBUTION                                )FROM THE PIPE GALLERY fSYWWETRY
         .u    )                                    NODE 00                                NODE U10) p                                  <                              (      --
l 9         .s, & *Fx (TnAn.s.v.t:tst)                                           ,
    ,4                                                   .
                                                           /                               /
NODE 004)                        l NODE Ull)
PG END FORCE                                                                              NODAL LOAD NODAL FORCE DISTRIBUTON K1PS & FT. K1PS                                                                                   (KIPS )
                                                 *T                                   =   5 (10 3)     Y (104 )
* TY (110 )
* TY (111)        8           1638 E          AXI AL)
S o        M(x ) : 610
* 610 '          +15
                                                                                                          ~
h    (VERT. BENDING)                   (103)
* iY (104 )
* iY (110 )
* iY (111)        10x4         (wos 10)
'                                                                                     =    2 VERT)
(103) = f z (104 ) = f z (110 ) = 1z (333 )       8             1 31
'8        F(x)= 1830                                                                                    +229
* 1830          ~
lh         (L ATERAL)                 x (103)
* i* (104 )
* i* (110) * #* (111)             8         (wcs 235)
  .w m s @'   M(z ) = 37,600                                                                37600 ls l
S'  [L AT. BENDING)             Y (103)
* TY (104 )
44(2)
'e.
                                                                                                                   .m

FDGURE 00 A-13 l
AT CONTAINMENT, NODAL APPLIED LOADS.
103                104              210 NODES                                                                            111 (KIPS 1              (KIPS)          (KIPS)           (KIPS 1 FORCE COMPONENTS            Fx      Fy    F2      Fx    Fy  Fz   Fx     Fy  Fz   Fx      Fy     Fz F(y > = 5100                 638.                638.            638.             638.
M(x ) = 610                   15.                -15.             15.             -15.
F(z ) = 244                         31.               31.             31.                 31.
TOTAL DUE TO SYMM.                 653. 31.          623. 31.        653. 31.         623. 31.
F(x)= 1830            229.                 229.            229.            229.
M(z): 37600                  427.       !        427.           -427.            -427.
TOTAL DUE TO ANTI-SYMM. 229. 427.         229. 427.      229.  -427. 229.   -427.

FDGURE 00 A-14

==SUMMARY==
OF PG'S INTERFACE NODES                                               .
WITH CES h SHELL MODEL G SYMMETRY NODE 003)                                  NODE U10) a      Y''                                     {     --
              !?
S I
u      *                                             --
l NODE 004)                                ! NODE 0111 SYMMETRIC LOADS                           ANTI-SYMMECTRIC LOADS AT NODES (KIPS)                               AT NODES (KIPS) fy a 103 = 653 (was 648) fy @ 104 = 623 (was 628) fy a 110 = 653 (was 648)                             fy a 110 = 427
;          fy @ 111 = 623 (was 628)                             fy 9 111 = 427 fz : 31 KIPS                               fx = 229 K1PS           -
l             ( e 103,104,110,111)                          ( e 103,104,110,111) 1 i
                                                                                         ^*

FDGURE 00 A-15 I
       .                                   ;tt'. 6' k
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                                             ,vhygi.
d vit .x vi    e SHEAR KEY
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LINER co   -   -f5 '
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                     }E. 6p fkI7 h ki'k'.                      {
                                                     \
4'-6"
                      =                                      =
PLAN CONNECTION DETAILS PG/CES INTERFACE (DWG.16103-50086, SH.3)

FDGURE 00 A-14

==SUMMARY==
OF PG'S INTERFACE NODES                     .
WITH CES h SHELL MODEL 4 SYMMETRY NODE (103)             NODE U10) m     .                   .
o                           !
N l     _
NODE UO4)            l NODE 011)
SYMMETRIC LOADS      ANTI-SYMMECTRIC LOADS AT NODES (KIPS)          AT NDDES (KIPS) fy a 103 = B53 (was 648) fy a 104 = 623 (was 628) fy a 110 = 653 (was 648)        fy a 110 = 427 fy a 111 = 623 (was 628)        fy @ 111 = 427 fz = 31 KIPS            fx = 229 KIPS
( e 103,104,110,111)    ( e 103,104,110,111) l

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l i'   GeddOD Cesk--D2C80 Ragsett for                                                                                                i Additiemal Information en SBP l
Topic 113-6, ' seismic Reevaluation
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ATTACIDGNT IIC-1 i
i

==SUMMARY==
OF THE REPORT ON M-11, DENINERALIZED WATER STORAGE TANK c
ll
[
I CIVIL 1.2

A TTA CHiYfivf.27- C-/
5.11   H-11 DEMINERAllZE0 WATER STORAGE TANK e
l l
l l
i i
610

SYSTEM COMPONENT NAME Dr-l- . t0.6, rt,, 7tCOMPONENT N2                                     M - //
LOCATION       dufi/c.                                  ELEVATIDN                2. 3 '- 6 "
COMPONENT SAFETYFUNCTION:ACTIVEO PASSIVE l@ 2O S-LIST PAGE N               l 'lf METHOD OF ANALYSIS:_ One                        D ,l ...:,,- ,,l. K.; lL e. E le . ~ s
( 2 e y t % / 4l 4 .,,, c M fr ee/, m                    D ., ,, .. . , E    A.. le / , s/, , t,l.,y dkc & ,le - t br TTD- 7c 2 +
SPECTRAL CURVES USED:                      6 ru ,, d. 6.,.,,;c DAMPlNG VALUE ASSUMED: 7 9e / Law ls',vcl ; o.r Oc> ( I/, s I, l., ,, )
ACCEPTANCE BEHAVIOR CRITERIA USED:
C.n GS y .'a ld.
COMPUTER CODE USED: G,rr Ecuc~r uvtr,. N rwe REMARKS:
Cef: TID- 7o 2 4
.          " Al>o x y oa . ,l         Ve d a L TL lu .'' E J. ke -,,,); . Oc t . t e s t
           * % ,s       o f [e n ,,, a 'e. De s , . -  P,,,jo s,', ,u     f, toelded S& eel 0 L s 4, <. , e.      T. Lu . "      to, n ,'. L        L M, bel,e ll REV.N2 0           i NORTHEAST                      BY        :rna eno UTILITIES-                    DATE #4 Wi7ls
{3                 QALCULATION COVER                 HADDAM NECK                    CHK'D NV FUP SHEET                                               DATE sf4'sM           ,
bPPR. Mr \DS DATE      Iny ib/r3 611

81C152.1 0591A 5.11.1 INTRODUC110N The Demineralized Water Storage Tank, TK25-1A, is located in the yard, southwest of the containment, at grade elevation. An elevation of the DWS tank is shown in Figure 5.11-1. The tank is 25'-0 in diameter and 30'-0 from grade to top roof support angle. The tank is covered by a fixed dome roof of 25'-0 radius. The plate material for this tank is 5052-F aluminum. Anchorage consists of 8-l' diameter A 36 threaded rods embedded in the slab foundation and bolted to aluminum anchor bolt chairs as shown in Figure 5.11-2. The maximum liquid height is 29'-0 from the bottom of the tank. This tank is surrounded by a concrete octagonal tornado wall that is l'-0 thick and 17'-0 high.
The foundation consists of a O'-8' thick octagonal concrete slab which sits on a backfill of clean sand above rock. The distance from the bottom of slab to the rock surface is approximately 5.3 ft. However, the more conservative 6.0 ft dimension was used in the soil structure interaction analysis.
These calculations were made to evaluate the seismic design adequacy of the Demineralized Water Storage Tank for the seismic response spectra input defined in Figures 5.11-3 and 5.11-4.
5.11.2 

==SUMMARY==
OF RESULTS The results of this analysis show that the anchorage of the Demineralized Water Storage Tank is not adequate to resist the forces and moments produced by the prescribed seismic disturbance defined by figures 5.11-3 and 5.11-4. Analysis indicates the anchor bolts will fail in tension from flexible tank and fluid overturning effects. It should be noted that failure of the anchorage does not necessarily mean that this tank will not maintain structural and leak tight integrity, as defined in category P-1 of ' Allowable Stress Criteria & Damping Values for NUSCO Haddam Neck." However, the probability of the tank's leaking or the shc11 buckling is increased due to this anchorage failure. After anchor bolt failure, the tank must be treated as being unanchored and local upitft of the tank bottom from the overturning effects would occur.
Seismic tests and observed tank response as the result of earthquakes have shown that greatly increased compressive stresses appear in the tank shell on the sides opposite to the uplift. In addition the welds between the tank bottom plate and wall cylinder would be required to transfer very large bending moments not considered in design as the result of uplift.
The potential for soil-structure interaction is also considered in evaluating the tank using the procedure defined in Ref. 7.10. However, since the foundation bottom is within 6'-O' of the rock surface, the effect of soil-structure interaction is negligible, as can be seen when from the magnitude of the computed spring stif fness shown in Figure 5.11-6a is compared to the values for rigid rock, i
612

6'       Goose NEC.( VENI b/    SC'45 E N qDOME Roof                       '4/ R.S (DOU BLE LAP WELD)                                                        -N" SQ HlNGED MANh0LE i                                                              '
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18'    MANHOLE i   (o EdO'd                                                     ,
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25'- O~ L D.
MATERIAL PL ATES - Al. 5052 - F i                   STRUCTURAL - 6061 - TG l W EIGHT l                   EMPIY            11.067 lbs.
FULL - 8'9936;              lbs.
                   .,Figur,e 5,11-1 Demineralized Water Storage Tank Dimensional Properties 613 -.
m

12'- 9 %" R ADIUS 4 '/2                   _
5"         _\
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No.       DESCRIPTION         i DIMENSION
: 1.       5083 ALUMINUM R.               5"x 5' x 1".8 RE0'D.
: 2.      __505 2 ALUMINVM lt. SK t W' .10 REQ'O.
: 5.       l' @ ANCHOR BOLT 5, NUTS ( WASHEO.
Figure 5.11-2 Anchor Bolt Chair Assembly for D.W.S.T.
l l
614 L-                                                                                                                    ,.

        <s -
o.0 DAMPING o.5   -
7 Y.
g        _
z  o4 o
4 m &
hj o.s   -
w O        -
O o.2   -    .
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o.o         i'''l.''     ''il'i'iiii'l>           ''li''''''il>''''' l             ''l  ' ' ' ' '  I o.O        5.0         10.0         15.0      20.0       25.0          30.0    35.0         40 0 FREQUENCY       (Hz.)
Figure 5.11-3 Horizontal Ground Response Snectrun at. Rock Surface

DAMPING o.6              _
                                                                                                               ]of, o.
z    o.4         -
O r                _
4 tr e   $    o.s
                    -  w O                ~
O o.2         -
                                                                                                                      ~
o.1 o.o           .........l........il.       i>>>>>>l.........l      ....... l.........l.........l.........I o.o         s.o          io.o          is.o      20.0        25.0          30.0      35.o      40.o FREQUENCY         (Hz.)
Figure 5.11-4 Vertical Groun<f Response 59ectrm at Rock Surface a
_ - _ _ - _ _ _ _ _ _        -~ _ _--           -     -.           .                      .-          .   ,                               ..

The procedure used to detemine soil-spring constants is given in Appendix 5.11 A.                              Further, it should be noted that the dominant frequency for impulsive response of the tank is approximately 6 Hz which is very close to the peak of the applicable response spectra so any shift in frequency that might be caused by the soil-structure would not significantly effect maximum seismic response.
The tank was analyzed treating some of the fluid as impulsive and moving with the tank and some of the fluid as sloshing and oscillating at a much lower natural frequency.
The calculated natural frequencies for this tank are the following:
0.35 Hz 1st mode sloshing of fluid in tank.
6.24 Hz 1st bending mode of horizontal impulsive fluid and tank motion.
17.79 Hz 2nd bending mode of horizontal impulsive fluid and tank motion.
30.28 Hz 3rd bending mode of horizontal impulsive fluid and tank motion.
41.63 Hz 4th bending mode of horizontal impulsive fluid and tank motion.
Total overturning moment at the base of the tank is; 61725 k-in.
The total resisting moment, equal to the total weight of fluid and tank times the radius, at the base of the tank is; 119614 k-in.
No rigid body rotation of the tank will occur but there may be uplift if the anchor bolts were to fail.
The tank shell was checked for buckling caused by shear plus bending plus axial compression at a point 45 degrees from the maximum compression in the shell. Buckling caused by bending plus axial compression was checked at the point of maximum compression. The results of these analyses are shown in Tables 5.11-1 and 5.11-2 respectively. The factors of safety are for the critical buckling stress and computed by the method of Reference 7.9.
6V

Table 5.11-1      Sumary of Tank Shell Buckling Stress Including Shear, psi Com3uted Stress                          Critical Buckling Stress (l)
Axial (f a) Shear (fy) Bending (fb)               Axial (Fe) Shear (Fy) Bending (Fb1 39        1549         2052                    6000        3781       5251
_ Factor of Safety
* Combined Axlal, Shear & Bending 1.18 Table 5.11-2 Sumary of Tank Shell Buckling Stress, psi Computed Stress             Critical Buckling Stress (l) Factor of Safety #
Bending (fb) Axial (fa)          Bending (Fb) Axial (Fc)          Bending & Axiai 2902            39            5251     6000                         1.19 The anchor bolts and anchor bolt chairs were checked for overstressing.
The anchor bolts are A36 steel rods with an ultimate strength of 58 ksi.
The top plates of the anchor bolt chairs are 5052 aluminum plate with a yield stress equal to 10 ksi and an ultimate tensile stress equal to 25 ksi. The computed stress values shown in Table 5.11-3 for the anchor bolts are calculated tensile stresses assuming that no deformation of the anchor bolt chairs occurs. The computed stress values shown for the anchor bolt chair top plate are bending stresses caused by tension in the-anchor bolts. These stresses, shown in Table 5.11-3, are for the bolts and bolt chairs located in Figure 5.11-5.
In this analysis it is assumed that the shear on the base of the tank is transferred to the foundation by friction between the tank bottom and foundation.
Table 5.11-3       Sumary of Anchorage Stresses, psi Computed Stress            Allowable Stress Bolt   Chair-Top P1 ate     Bolt    Chair-Top P1 ate Bol t  Factor  of Safety Chair-Top  P1a_fe A     85870       55050         40600          10000          FAILS       FAILS B     77670       49800         40600          10000          FAILS       FAILS C     54630       35030         40600          10000          FAILS       FAILS D     21240       13620         40600          10000          1.91        FAILS (I) Critical buckling stresses are computed from buckling criteria in Ref 7.8
   *          .67                         i          .67 p
618

In calculating bolt loads the stabilizing effect of the weight of the fluid and the relative stiffness of the tank bottom plate has not been considered hence results shown are quite conservative. Any detailed redesign of the anchorage should will consider all applicable l'oads ar.d compatible structural response.
C 0
0                     0 0                    o Ao                                o o
o D E
Oc FIGURE 5.11 5 Anchorage Layout The total hydrodynamic and hydrostatic pressure en the tank shell was computed and hoop stresses were calculated at the bottom of each tank shell ring as shown in Table 5.11-4.
Table 5.11-4     Hydrodynamic and Hydrostatic Pressure and Resultant Stresses PLATE       DEPTH                                   ALLOWABLE THICKNESS    IN FLUID                     HOOP STRESS    STRESS INCHES       INCHES          PT , PSI      KSI        KSI (Tension)
RING 4     .1875         79.9                4.393      3.51 5        10 RING 3     .1875        169.25               7.936      6.348         10 l
RING 2     .22          258.6              11.355       7.742         10 i RING 1     .3           348                14.798       7.399         10 l
l               Hoop Stress = PT.r t
!                      2    2       2 PT = (P1 + P2 + Py) 1/2 + pSTATIC      (Ref. 7.7) i 619 m

TANK     N S-  ll
[                  387.OG" -- 22                                . ,,.
y                                359.75'                gg s
n 357.5 328.67        o  18 m = 0.01767 303.51 < rt 0.075G9 1
275 .11         to o.12499 246.32 o 15 O.15549 217.42 o u O.17253 188.4G      o  is o.161G5 178.75 -- 12 159.48 o it O.18643 130.49 o to 'O.188 *)2 (o
jota 9     79 o.19019 89.375 --8 4
72.5     "7 0,19085 43.5 o6 o.1911G 14 . 5" <   5 0.19129           _
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                                                                         '3 Kes 5 Ah10' K4 e2                                          K'' I 47 ' ids EII"'
Ko B.47 w10' K-Iw / RAD.
x    Kos ?.327s.lf Khu o1 E[                                 Kc= IO'o K 29/ RAD.
Figure 5.11-6a D.W.S.T. Dynamic Model 620 SCALE     '/0 l'-O" m-J

6t. . Un) 387.06 22-TANK      NE 11 - DYNAMIC, MoDEL It **
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357.5l ''                   699.94   699,94     699.94 15.81 x10' 7.9 < 10' 7.9 ( 10'                                 h 328.07 to o     m 30.01707 2 375.8  H87.9      M 879     54.35,16 27.2 e 10'                    27.24 #                h 235,62'   117,91    117 . 9 1 5.3x10'      2.46 < 10' 2.46x10'                             @
303.5117 o    0.07569 275.11 iG o@012499
                    @0.15549                                                                                         @
246.32 15 o 6
176.825 88.413     88.413    3.93x 10     1.9 9 . 10'-              1.99 x lo' 217.42 14 o
                    @ 0.1725 3                                                                                       @
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101.49 94                                                                                                             Q 89.375 a-72.5 7         0.19085                                                                                                  _
4 43.5 60      @o.19116                                                                                                 3 283.026  141, 513   141.613   c.33x 10 5 S.19x lo' 3.19 xf o' 14.5     eo  0.19129 4          o r,'          >
a a / rigure s.11-68 o.w.s.r. oynamic soeei 621

1 a.oa                                                                                                    l
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S81.0G as'                                                                                               i TANK      NS DY N Ahl t. MoDEL 360 u__                          A tu2                              I , 1"'          23         r~'
359.75 25 Au 2w 2     Av,ist It         L        gut '
357.5 2r 17                   699.94    699.94      699.94 15.81,10' 7.9 x 10' 7.9 x 10'             @
328.67 2i o      m 30.01767 2 375.a   lI87.9       1187 9       54.35<@ 27.2 e 10'     2 7.2 < m'  @
235.62'    117.91                   5.3xto' 2.66.10' 117 . 9 )                          2.66 x 10'   @
303.5120 o      0.07569                                                                             @
275.11 19 o   @0.12499                                                                              @
               @o.15s49                                                                              @
244.32 180 176.825  88.413      88.413       3.9st lo'   1.99 < 106 1.99 x to' 217.42 17 <Q o.1725 3
" Or)?                                                                                               @
6 3<
IL             g o.1816s 178.75 it 159.48 u          o.18s 43 130.49 to o O 0.18892                                                                             @
               @o. I9oi9     207.497 103.749        103.749       4.67 x (O' 2.34 < 10'  2.34 x i o' 101.49 9 <
89.375 a--
72.5     7<      o.19085 43.5 Go
             @o.1   9116 4
3 263.026    141. 513    141.513      6.38xto' 3.19x 10'      3.19 x fo''
14.5    so      0.19129 4            O "3
2                     d Y
           '[                 z 4    Figure 5.11-6c        0.W.S.T. Dynamic Model W/ Bumper
                     /                                        System 622

1 where:
P1     impulsive motion hydrodynamic pressure P2 = sloshing motion hydrodynamic pressure Py = vertical hydrodynamic pressure Some type of additional support for the Demineralf.ed Water Storage Tank will be necessary. The space between the tank shell and tornado wall is limited, so that it will be difficult to add anchor bolts to the foundation. A conceptual design of a bumper and tie rod system, attached to the tornado wall, that will reduce the seismic moments acting on the tank base by a factor of 7, is presented. This bumper system raises the fundamental frequency of the impulsive fluid and tank motion to 18.23 Hz, well away from the peak of the ground response. Conceptual drawings and details of the bumper-tie rod system are shown in Figures 5.11-7 through 5.11-10. By applying this reduction factor to the anchor bolt computed stresses in Table 5.11-3 it can be shown the existing anchorage system is adequate if the bumper system is vsed.
5.11.3 LOAD CRITERIA Failure of the tank anchorage was considered to be the area of greatest concern because of the analytical methods used in the original design, which in accordance with the TID 7024, considered the tank as rigid.
The ground response spectra was used in the analysis of the tank.                           The basic ground response spectra is considered to be representative of accelerations at the rock surface, which is 6 feet belcw the foundation of the tank. The effeet of intervening soil was included by the use of soil springs. Curves for seven percent damping were used with impulsive mode response.              One-half percent damping was used for the sloshing mode of response.
Nozzle loads on the tank were assumed to produce limiting stresses in the nozzles of values equal to Sa11/3. These loads were compared with the loads on the tank anchorage system from the earthquake. The largest nozzle induced load was .17% of the overturning movement on the base of l
the tank determined from the earthquake. Because of the small values determined, the effect of nozzle loads on tank supports were not combined with earthquake loads.
5.11.4 STRESS CRITERIA The stress criteria are as shown in Section 3 with these exceptions:
Allowable stress in bending for the tank shell is governed by buckling criteria, Sal'1 - 5251 psi.
Allowable stress in shear for the tank shell is governed by buckling criteria, Sall                3781 psi.
623

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4 8' x 4" x /2''             ROLL RADIUS = 12'-6"                                                                             '
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B            __                __        IB Figure 5.11-8 Typical Bumper with Tie Rods                        625
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L 4" x 4" x %" ( 9" LG. ( 2 REQ'D.)
R. %' x SK .(.2 REQ'D.)
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VIEW     B-B Figure 5.11-9 Typical Tie Rod Bracket 626                                  -
                                                                                                                                         -- a

L 8 x A''x '/2 L (o " x 4 ' x 3/g*                                       8OM RADIUS =
12 '- 6' ft. 5/4" x SK.            NEOPRENE PAD.
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e-SEC.TIO N A-A Figure 5.11-10 Typical Section of Bumper 627

5.11.5 METHOD OF ANALYSIS The SAP IV computer program was used for the analysis of the impulsive motion of the tank and fluid, and the sloshing motion was computed and combined with the computer analysis by hand calculations.
Since the tank is supported on compacted backfill, the effect of soil structure interaction was considered in the modeling of the tank.
Equivalent soil springs were computed for vertical, lateral and rocking motions of the tank. A torsional spring abcut a vertical axis was considered very stiff and the largest value for spring constants in the computer program was used.        The computed equivalent soil springs, because of the proximity of the rock to the surface, were also very stiff. These springs had a minimal effect on the tank's fundamental frequency of vibration compared to a rigid base modeling assumption. The fundamental frequency of the tank system model is approximately 6 Hz which is very near the peak of the response spectrum and produced the large loads on the anchorage. Any error in response frequency introduced by actual support conditions may tend to move the response away from the peak and, therefore, reduce resultant stresses and loads.
The boundary elements for the computer model were input with the approximate soil spring stif fness in each direction.      The tank was modelled as a cylindrical beam stick with fluid mass lumped at points along the length. Sketches of the computer stick model are shown in Figures 5.ll-6b and 5.11-6c along with the values for the section properties and lumped masses.
628                                    ,
i

5.11.6 CALCULATIONS e
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5.llA APPENDIX Soil - Structure Interaction For Demineralized Water Storage Tar' 4
e e i
I 647 i

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     .i. // A APPENDIX-Soil - Structure interaction was modelled for the tank by assuming a rigid circular disk for the foundation and applying equivalent soil springs to the boundary elements.
: 1.          Soil Constants Supposition (Assume)
Y       p            E              v                  G (ft/s)  (ib/in3)     (psi)                             (psi)
                                       *1       *1                              *1 Sand                          2000    0.061                       0.40               9.09x104 Rock                                                          *2              *2 (granite)                                          lix106         0.40               3.93x106
                                          *1 Assumption 2
pVs Gsand =                               *2; the average value of granite from g                             Values for Rock computed from compressibility measurements by Brace 0.061 (200x12)2                (1966) at confining stress of 3-5
              =                                Kilobars, Lambe, Whitman 376.4                       Soil Mechanics, 1969
              = 9.09x104 (psi)
E Gr ock =
2(1+v) llx106
              =              = 3.93x106 (psi) 2(1+0.40) where.
Vs = shear velocity p = density E = Young's Modulus y = Poisson's Ratio G = Shear Modulus The following calculations are from Reference 7.10 648
: 2. Rotational Soil Spring considered as semi-infinite elastic continuum.
In case of vertical leading the displacement w is 2
P      1    'Z W = 41G" T . R' + (I ~ # '.
and                                  '
2                         '
3W         P    '3 X 2             X 2
e=3x=4mg.                   g3        + F .c1- v>            ,
where P        = force on the surf 3ce (x,y,:) = coordinates of the point R         = length between the point and loading point R=      (x2 + y2 + z2) in case of rotational loading applied the above formula
                    'the stress      0z =My X /3 7
                                 -         X    3'.X       +       2 -(1 - V ) dX dY 0=4EG                                                       ,
2              2     '
                                                 ,3 Z22  p cosg         p2 cosg        rdbd6
                        , _ 1 _ , M_y 2x 3 4tG Jy oJo            dr2+ z2p4             cr2+z )" 2-(1-0) dGga;2zjj'g+<,2 z ,3 +2 ,,cs-g42'+/-2zl'   a                    ,
a        3                          1         I        z2 7      U=21   3z         /a2 + z2
* 5 la2 +z2)u )
0 (r2,7ayu 2
{\
F3 2z2 ,g 0 (F +z2)w d F = jg2 + 7     2 2z
                                                                                   /
a = radius z = depth 649 i

on the surf ace (i = 0) 0 _1 - V , My  -
a 0
                ~
2G      Jy So,Odividedby00, that is the amplification function F
                                                                            '
* 2    1 4
3z 2 +      2 F,(Z/a) i 2act-0) 2z Ja%2 2  (a +z 2   2  )* +2 (1-0a2+z2  )1)2 z + a -2z; 3d)2 +    (y I    -
                       = 2 (1 - D )
* 2-({) /l+([f (1 +(h)f +2(1-0)fjl+(h)2 In this model a =
60" a -;
      ,,,,,,,,,,,,           z1      =72"(Assume) 2 G,     zi                     e = 0,-( l - F,(4) +(GyG f=C /a))
                                  , = the disp. in the case of semi-infinite soil of G1 lG 2 F,(2Va) = Fr.(7%so) = .6 21 0=Oi>            3 '2 5 t- amplification factor equivalent shear velocity, eVs eVs = 2000/ 6 39 3 ) 1/2= ?i9o f t/sec equivalent shear modulus, eG eG = 9.09 x 104 /0.393 = 2 313          I E (F'i)
Rotational Soil Spring Ke y = 0.40 Poisson's Ratio 8 eG a3 Kg=                     (by Borowicka)
                  = 3,47 , i o ' t Ib in/ rad
(
650
                                                                                                  , e

t
: 3.      Horizontal Soil Spring Horizontal displacement Ux surface force; Txz dxdy 2                          2   ,
                   -                                            1'X                                                                 X Ux = ra dx dy4EG                                                               F +1       T    +(1-2 R+Z        0)R'CI R(R+Z O So, when force is unifonn stress distribution for circular base 24 a ,__
Ux =          Uxt dr de 0
2
                         = pra. '_L 3 +4 4 _g                                                                        1-2u (fg g 2 _ q) 2 gm 2G                     2 Ji +g2 On the surface ( 2 = 6                                   0)
(2-v)
Uxo = rNG                                                                                                                -
                                                                                                               -2 6 + 17s     c j , 4, _ g ).
2 j
Fa m)= (2 U x = bx[                                      l - Fn(ZVa) +(G/G Fn (2Va))                       2 Fa(zVa) = Fu (7Wo) = .s44 J. Ux = U xi > . 469 eVs = 2000//,469)''' = 292c #b G - 9.09 10 /,469                                                                4
: i . 9 3 3. , o rg:-;        .
es Horizontal Soil Spring KH 32 (1-v) eGHa 7-8v
: l,41 io   i                                            lb/in 651
: 4.           Vertical Soil Spring As mentioned in the paragraph 2., the vertical displacement is expressed
_P ._L,' 4 +2(1-0)
W = 4tG R                         R As the results, Fy is expressed as follows,
                                                                                           /a)
Fy (Z /a) = / l + (Fa /                    2 (i v) ); z f +(1-2'v).(2 Fv(72Asd = s% >
Z W = Wi f I- FvC /a) +(G/G) FvCZVa)
                              = wi .           .
                                                     ,2u equivalent shear velocity eVs for the vertical eVs = 2000/ Y.1241'                    = SG75 ft/sec equivalent shear modulus eG for the vertical (psi) eG = 9.09 x 104/. 1+2 : ' 7 3 2 . n
* Vertical Soil Spring, Ky 4G a Ky =
(by Timochenko & Goodier 1951)
                        = 7.3n io ' .lb/in 652

The torsional spring constant, K , is larger than 1010 K-in/ radian and not acceptable to SAP IV input. It was set equal to 1010 K-in/ radian.
l l
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653 1
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Cctico Cook--CRC'c Coquest for Adtitio=01 10fercatisn e= sEP Topic III-6, ' seismic meevaluation" ATTACHMENT IIC-2

==SUMMARY==
OF THE REPORT ON M-12, REFUELING VATER STORAGE TANK
\
L CIVIL 1.2 l
\
5*

AritcM i   HrorH-C-2 5.12  M-12 REFUELING WATER STORAGE TANK 654
_m

/
SYSTEM                n r u e c i a r, COMPONENT NAME C n,e, o hc u ra.xCOMPONENT N2                                          41 - / z '
LOCATION           v,ico . exara. .<.f a cm,.rELEVATION                           :s ' o %
COMPONENT SAFETYFUNCTION ACTIVE                                    PASSIVE I            20 S-LIST PAGE NS             4 /o METHOD OF ANALYSIS: Je<                      /,o.     . . ,1.. / /,o , /< < /< , ,,.f M is.to,..          -r      ,c i. wi ,, ,     p,,..,,,,..     %,J                9, ed e,,4   m , /, n Jl . . I     h .i   rio. , s w .       So ;t n ,ca t ,        1,, 4 ,u s   ce c ,e_.. l , i L ,,,,,,,,;
SPECTRAL CURVES USED:                       6a ou        O ./e , c a r w o o c t c c c.
vu :s..
DAMPINGVALUE ASSUMED: , 7-. /nm e, c                               .       r % Se om e" c.        e     , ~ <-
co~. , uc .
ACCEPTANCE BEHAVIOR CRITERIA USED:
                                                                                                         "' " "' r C-_//...,4/e      6       fy, //
COMPUTER CODE USED: S4# / v /'d v na->c )                                 A" ys (eMie }
PEMARKSs            A D O i T/ ouM Anca w mc                      i: orcDCD r-ors -w C w.<
fo     D A .~ '    TIJ E        M C (/7A rel t    C 1% , r EG r A
     . Fo rt fioric /teruco o*              Aung:or , a           T wo oiusa:r uist. finirc
               ~
L- L. E M GerT*       EH EL"      F10 0EL-         WA:        U fE D .
REV.N2 o             I   a NORTHEAST                       BY       su Jw nw UTILITIES-                      DATE ?*& V#st Wlv
[8                                 R-COVER               HADDAM NECK                    CHK'D pnp W Per DATE NacAW*W SHEET APPR. 17?T *Ar %A
                                      ~
I                                                                                       DATE 7 /to W t. d(wh 3 655

81 C152.1 01078 5.

==12.1 INTRODUCTION==

The Refueling Cavity Water Storage Tank, TK4-1A, is located in the yard, northeast of the containment, at grade elevation. An elevation of the.
RCWS tank is shown in Figure 5.12-1. The tank is 35'-0 in diameter and S6'-0 from grade to top roof support angle. The tank is covered by a fixed dome roof of 35'-0 radius. The plate material for this tank is 5052-F aluminum. Anchorage consists of 8-1" diameter A 36 threaded rods embedded in the slab foundation and bolted to aluminum anchor bolt chairs as shown in Figure 5.12-2. The maximum liquid height is 35'-0 from the bottom of the tank.
The foundation consists of a 4'-10" thick octagonal concrete slab which sits on a backfill of clean sand above rock. The distance from the bottom of the slab to the rock surface is approximately 1.8 ft. However, the more conservative 6.0 ft dimension was used in the analysis.
These calculations were made to evaluate the seismic design adequacy of the Refueling Cavity Water Storage Tank for the seismic r'esponse spectra input defined in Figures 5.12-3 and 5.12-4.
5.12.2 

==SUMMARY==
OF RESULTS Two analyses were performed for the RCWS tank. The first analysis, with the tank modeled as a stick with lumped fluid masses, was used to compute natural frequencies and an upper bound on bending roments and shears for the tank shell buckling analysis. The second analysis used a two dimensional finite element shell model which included lift-off capability and local bending stiffness of the tank ba'se. Hydrodynamic pressures and inertia forces were applied to this model to coripute anchor bolt loads.
The results of the second analysis show that the anchorage of the Refueling Cavity Water Storage Tank is not adequate to resist the forces and moments produced by the prescribed seismic disturbance defined by Figures 5.12-3 and 5.12-4 without some type of modification. The present system of anchor bolts will fail in tensior. from flexible tank and fluid overturnih; effects.
It should be noted that failure of the ar.chorage does not necessarily mean that this tank will not maintain structural and leak tight
, integrity, as defined in Category P-1 of "Allowable Stress Criteria &
Damping Values for NUSCO Haddam Neck." However, the probability of the tank leaking or the shell buckling is increased due to this anchorage l
failure. After anchor bolt failure the tank must be treated as being l unanchored and local uplift of the tank bottom from the overturning effects would occur. Seismic tests and observed tank response as the i result of earthquakes have shown that greatly increased compressive j stresses appear in the tank shell on the sides opposite to the uplift.
l I
656

                                                                  '6'     GOOSE NECK VEi<
DOME ROOF O.53" ft.'s
                                                                                          -24" SQ HINGED MANHOLE
                                                                                          /
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e                                                            /
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                                                                       ~       '     "
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                                                                                         ~
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             'y                                                                           -   - R.4 - 0. 202, TH K.
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                                                                                              - R. 3 - 0.303" THK.
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(3)                                                                            _
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                            '                                                               ~
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         .        1
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                                                                                              -- it.1 - 0.504"THK.
v      ,                                        g.*9.
F                                        .-                                   . EL. 2 3'- O h, T.-
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SEE DETAll                                     3 5 '- O" I . D.
 ' MATERIAL i                   PLATES - AL 5052 -F STRUCTURAL - 6061 - TG i WEIGHT                                                                   .
:                   EMPTY - 26,053 lbs.
li                    FULL - 2, i27,30 7 lbs.
4 Figure 5.12-1 Refueling Water Tank Dimensional Properties t
657

I 17'- 9" R ADIUS i
li'-C.V[
OUTSIDE RAD.
5"             _
5"          _
                                                                                                          \
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1                                                 I        5554) Y4 V J_ = 1q%G"                             l
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      =                                                     'o                                             I i                                             l I
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5554 fy     p                          l t
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                                                                                          ,         / - _ __ __y P2
_  9 = 3"           _
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                                                                                                              '/i x 45*
2"     ;
                                                                     '~
No.         DESCRIPTION           i DIMENSION                  _
: 1.         5083 ALUMINOM R., 5"x 5"x 1".8 RE0h
: 2.         505 2 At.UMINUM 11. SK t '/ " 2.16 REQ'D.
: 5.          F@ ANCHOR BOLT 5, NUTS ( WASHEE$.
Figure 5.12-2 Refueling Water Storage Tank Anchor Bolt Chair Assemblies 658    -                                                -   -

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: c. .t o.o        .....>. .l........il. iiiii    il.......    .l........ l.........l.........l.........I o.o          5.o        io.o          is.o        20.o        25.0         30.0       35.o     40.o FREQUENCY          (Hz.)
Figure 5.12-4 Vertical cro::nd P.csponse Spectren at Pac!: Surface i

In addition the welds betweer tne tank bottom plate and wall cylinder would be required to transfer very large bending moments not considered in design as the .esult of uplift.
The potential for soil-structure interaction is also considered in evaluating the tank using the procedure defined in Ref. 7.10. However, since the foundation bottom is within 6'-0" of the rock surface, (l'-10")
the effect of soil-structure interaction is negligible as can be seen from the value of the computed spring stiffness shown in Figure 5.12-6b.
The procedure used to determine soil-spring constants is given in Appendix 5.124. Further, it should be noted that the dominant frequency for impulsive response of the tank is approximately 6 Hz which is very close to the peak of the applicable response spectra.
The tank was analyzed treating some of the fluid as impulsive and moving with the tank and some of the fluid as sloshing and oscillating at a much lower natural frequency.
The calculated natural frequencies for this tank are the following:
0.29 Hz     ist mode loshing of fluid in tank.
5.88 Hz     1st bendhig mode of horizontal impulsive fluid and tank motion.
15.59 Hz     2nd bending mode of horizontal impulsive fluid and tank motion.
25.93 Hz     3rd bending mode of horizontal impulsive fluid and tank motion.
35.84 Hz     4th bending mode of horizontal impulsive fluid and tank motion.
The total overturning moment at the base of the tank is; 166767 k-in.
The total resisting moment, equal to the total weight of fluid and tank times the radius, at the base of the tank is; 396104 k-in.
No rigid body rotation of the tank will occur but there may be upiift after failure of the anchor bolts.
Using the first analysis, the tank shell was checked for buckling caused by shear plus bending plus axial compression at a point 45 degrees from the maximum compression in the shell. Buckling caused by bending plus axial compression was checked at the point of maximum compression. The results of these analyses are shown in Tables 5.12-1 and 5.12-2 respectively. The factors of safety are for the critical buckling stress and computed by the method of Reference 7.9.
661

Table 5.12-1         Sumary of Tank Shell Buckling Stress Including Shear, psi Computed Stress                      Critical Buckling Stress (l)
Axial (fa) Shear (fy) Bending (fb )                  Axial (F C) Shear (Fy) Bending (Fb) 44         1494               1663                 7200         6916       6295 Factor of Safety
* Combined Axial, Shear & Bending 2.11 Table 5.12-2 Summary of Tank Shell Buckling Stress, psi Computed Stress                 Critical Buckling Stress (l) Factor of Safety #
Bending (fb) Axial (fa)            Bending (Fb) Axial (FC )          Bending & Axial 2352              44            6295         7200                        1.76 The anchor bolt stresses and anchor bolt chair to tank weld stresses from the second analysis are shown in Table 5.12-3a. The locations of the bolts are shown in Figure 5.12-Sa. The 1" diameter anchor bolts are assumed to be 36 ksi yield stress steel with an ultimate tensile stress of 58 ksi. The weld metal, for the aluminum anchor bolt chair to tank connection, is 39.6 ksi ultimate tensile stress.
Table 5.12-3a Sumary of Anchorage Stresses, ksi Computed Stress             Alloweble Stress           Factor of Safety Diameter Bolt Chair-Top Plate Bolt Chair-Top Plate Bol t Chair-Top Plate 1"       A 91,05           36.1          40.6        18.0            FAILS      FAILS 1"       B 49.27           19.3          40.6        18.0            FAILS      FAILS (l ) Critical Buckling stresses are computed from buckling criteria in Ref. 7.8
     *             .67                          i        .61 FS =      +                                FS =    +
i
)
662
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To correct the potential anchor bolt failure the static shell analysis was redone with eight 1-1/4" diameter anchor bolts added to the present 1" diameter anchor bolt system. The 1-1/4" diameter anchor bolts are assumed to be equivalent to Maxi-Bolts manufactured by Drillco Devices, Ltd. The allowable tensile load for these bolts is 85,024 lbs. according to ACI-349 Appendix B. Table 5.12-3b shows the anchor bolts and chair-tank weld stresses for the anchor bolts located in Figure 5.12-5b.
Table 5.12-3b Sumary of Anchorage Stresses, ksi Computed Stress            Allowable Stress      Factor of Safety Diameter   Bol t     Chair-Top Plate Bol t Chair-Top Plate Bol t Chair-Top Plate 1"       A 40.6          15.93        40.6        18.0       1.0         1.13 1 -1/4"  B 27.54         16.41        85.05       18.0       3.10        1.09 1"       C 27.12         10,64        40.6        18.0       1.5         1.69 1-1/4"   D 10.124         6.04        85.05       18.0       8.4         2.98 1"       E 18.748         7.35        40.6        18.0       2.16        2.45 1 -1/4"  F 12.0           7.15        85.05       18.0       7.09        2.52 o
B o                   o Ao                                o B
O Figure 5.12-Sa    Anchorage Layout for Static Analysis (1st Run)
O    o       F 0          0 C
O o
8    /                      0 O(
A o                            o 60                          0
                                      .O                    O o          o go 0            F i           Figures 5.12-5b    Anchorage Layout for Static Analysis (2nd Run) l 663 l
[                                                                                  _

r The total hydrodynamic and hydrostatic pressure on the tank shell was computed and circumferential stresses were calculated at the bottom of each tank shell ring as shown in Table 5.12-4.
Table 5.12-4       Hydrodynamic and Hydrostatic Pressure and Resultant Stresses PLATE          DEPTH                                                                                                                                    ALLOWABLE THICKNESS        IN FLUID                                                                                                                      HOOP STRESS     STRESS INCHES           INCHES        PT , PSI                                                                                                         KS!        KSI (Tension)
RING S       .1875"          76.375'        4.5                                                                                                             5.04             10 RING 4       .202'          162.28125'      8.09                                                                                                            8.410            10 RING 3       .303'          248.1875"      11.419                                                                                                           7.913            10 RING 2       .403'          334.09375"     14.708                                                                                                           7.664            10 RING 1       .504'          420'           18.012                                                                                                           7.505            10 Circumferential Stress = Prer t
2        2    2 Py(x, e) = (P) +P2 + Py) I/2 + PSTATIC                                                                                                              (Ref 7 7) where:
Pi = impulsive motion hydrodynamic pressure Py = vertical hydrodynamic pressure P2 = sloshing motion hydrodynamic pressure 5.12.3 LOAD CRITERIA failure of the tank anchorage was considered to be the area of greatest concern because of the analytical methods used in the original design, which in accordance with the T107024, considered the tank as rigid.
The ground response spectra was used in the analysis of the tank. The basic ground response spectra is considered to be representative of accelerations at the rock surface, which is 6 feet below the bottom of the tank. The effect of intervening soil was included by the use of soil springs. Curves for seven percent damping were used with impulsive mode response. One-half percent damping was used for the sloshing mode of response.
Nozzle loads on the tank were assumed to produce limiting stresses in the nozzles of values equal to Sall/3. These loads were compared with the loads on the tank anchorage system from the earthquake. The largest nozzle induced load was .24% of the maximum overturning moment at the base of the tank determined from the earthquake.                                                                                                             Because of the small values determined, the effect of nozzle loads on tank supports were not combined with earthquake loads.
664
_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _                            ~

5.12.4 STRESS CRITERI A The stress criteria are as shown in Section 3 with these exceptions; Allowable stress in bending for the tank shell is governed by buckling criteria, Sall = 6295 psi.
Allowable stress in shear for the tank shell is governed by buckling criteria,     Sall    6916 psi.
5.12.5 METHOD 0F ANALYSIS The SAP IV computer program was used for the dyntmic analysis of the impulsive motion of the tank and fluid, and the sloshing motion was computed and combined with the computer analysis by hand calculations.
The tank was modeled as a cylindrical beam stick with fluid mass lumped at points along the length. Sketches of the computer stick model are shown in Figures 5.12-6a and 5.12-6b along with the values for the section properties, lumped masses, and spring stiffnesses.
Si' ice the tank is supported on compacted backfill, the effect of soil structure interaction was considered in the modeling of the tank for the dynamic analysis. Equivalent soil springs were computed for vertical, lateral and rocking motions of the tank. A torsional spring about a vertical axis was considered very stif f and the largest value for sping constants in the computer program was used. The computed equivalent soil springs, were also very stiff because of the proximity of the rock to the surface. These springs had a minimal effect on the tank's fundamental frequency of vibration compared to a rigid base modeling assumption. The fundamental frequency of the tank system model is approximately 6 Hz which is very near the peak of the response spectrum and produced the large loads on the anchorage. Any error in response frequency introduced by actual support conditions may tend to move the response away from the peak and, therefore, reduce resultant stresses and loads.
The ANSYS computer program was used for the static analysis of the tank and anchorage system. The tank was modeled as a shell using two-dimensional finite elements for the bottom and shell platn . The anchor bolts were modeled as beam elements.
An earthquake will produce three different hydrodynamic pressures on a tank shell: an impulsive pressure (Pj) arising from fluid moving in phase with the tank shell; a sloshing pressure (P )2 due to fluid sloshing (mainly a surface effect); and pressure from the vertical earthquake (Py) which is equivalent to an increase or decrease in the hydrostatic pressure. These hydrodynaic pressures were comput(d such that the combined pressure would produce a maximum overturning nKeent.
This condition is produced when the invulsive pressure (P1 ) and sloshing pressure (Pg) are at a maxingra w the vertiev) pressure (Py) is at a minimum. Since these pressur          unsid m d to ar.t independently, they are combined by SRSS and adda             to tu hydrosw tic pressure to produce a total pressure as                   the location of the point of investigation on the tank t'                    The comptne'ts of the total 665

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366.22 '9 0        0.15004 343.c,18-                   247.511    123.755 123.755 10.h 10' 5A6 x 106 5Acx fo' 333.0411o          0.26124                                                                      15 297.7(o'6 o@o,s3829                                                                            @
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           <j    y/                                                                                .
Figure 5.12 -6a Refueling Water Tank Dynamic Model 666

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Figure 5.12-6b Refueling Water Tank Dynamic Model SCALE 8/tr.
667 l
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pressure are shown applied to the tank shell and bottom in Figure 5.12-7.
A computer generated plot of the shell model is shown in figure 5.12-8.
P i ( x , 6 ) = W
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WI     weight of implusive fluid II     height to centroid of impulsive volume from bottom pSa) = pseudo-acceleration of ist bending mode impulsive motion Wy     total weight of fluid psap = pseudo-acceleration of 1st sloshing mode 0    diameter of tank H = height of fluid (ZPA)y = vertical zero period acceleration P=massdensityoffluid x = depth in fluid from free surface 6 = angle from point of maximum compression to point of investigation, measured counterclockwise positive 668

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l URS/BLUME 130 Jessie Street (at New Montgomery)                                                                                           )
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SUBJE CT $ C4Ul!;fIEf.      'T~       "l ?I . -              SHEET ! S            OF      .S A
                                    /Tc*t , s ..a . L C,2 < a i: ,r W D r !<                    soem __d/C/S; L C y,1 *" '-                                  ..-           or . dDu             OATE   1/G/D STEVENSON                       Art Li&L 7- I 77 dJ /5.                                      csmo ey fref. out _2 ./3 -12.
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  &. . ASS.O..C.I ATE.S g z , " < ' ' - 4 c-T., --- - - - -                                     nEv no                          -
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                     "7y    ,T My          =           2. '43     90    e  =      l-f 3          ,
NOT C/d M' Pt:A hi                ~
A%         tb I'/         .vo t5 c.E         4     '/M/ L 44                9/!d 'E O
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                                                                                                     .      C     / r /206_0 NCD d         7%       ,''1-       /s M,.=           19 40'2-         Ib- on M   ,'        l'7 JT'I lb ->^
M  ,-
M 4 9+        '
                                               .h-   -
W     =       t0 4     96          /s c
                   =
V,            to 4 ?o             Ib s l
i l         k' - -  '     . O'/2-      9a    /fr          It)1,   G N) Ft      C),fd     r l
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l 693 m

5.12 A APPENDIX Soil - Struc'.ure Interaction For Refueling Cavity Water Tank l
\
l l
l i
I 694

5.12. A APPENDIX Soil - Structure interaction was modelled for the tank by assuming a rigid circular disk for the foundation and applying equivalent soil springs to the boundary elements.
: 1.        Soil Constants Supposition (Assume)
Y          p            E           v         G (f t/s)    (1b/in3)     (psi)                  (psi)
                               *1           *1                     *1 Sand                    2000       0.061                    0.40      9.09x104 Rock                                                  *2         *?
(grani te)                                      11x106      0.40       3.93x106
                                     *1 Assumption pVs Gs a nd a ---                      *2; the average value of granite from 9
Values for Rock computed from compressibility measurements by Brace 0.061 (200x12)2                (1966) at confining stress of 3-5
         =                                Kilobars, Lambe, Whitman 376.4   ,
Soil Mechanics, 1969
          = 9.09x104 (psi)
E Gr ock =
2(1+v) 11x106
          =            = 3.93x106 (psi) 2(1+0.40) where, Vs = shear velocity p = density E = Young's Modulus y = Poisson's Ratio G = Shear Modulus The following calculations are from Reference 7.10 695
: 2. Rotational Soil Spring considered as semi-infinite elastic continuum.
In case of vertical loading the displacement w is                     ,
7                    '
P          'Z' 1
W = 41G ' Y , R' +
('~ '      s and 2
aW _       P     '3 X Z                                   '
O = 3X         4EG     -
p5
                                                         +X =2-(1 - v)
R3                   ,
where P     ' force on the surface (x,y,z) = coordinates of the point R     ' length between the point and loading point R=     (x2 + y2 + z2) in case of rotational loading applied the above formula the stress     Fz =My X /3 7 2
1     . My 3XZ          X 2 -(1 - V ) dX dY g=4EG Jy              X           6           3               ,
R          R 2         2    2
                              , My 2a el                                                 edhde2-(1-0
_    1
                                                '3  Zg2cos0 ,(r cos gr2,2z)*
                    - 4EG Jy a Jo               dr2+z2)sA 2            4                   2 2 1    .My ,g,                             Z 4KG Jy         '2 Z /a2 z23Z, (a2 1-z  2    )M +2-(1-0):t22 la2+za   ,
ta -2Z f'        (r2[z yw dr =f h-f2 a                             7 2  +h(g2,2)n)      2 2
N              2 Z + a*       -2Z l
(\     'O lI + 2 )
gp, /d +2 2                                        )
l a = radius z = depth I
l 696
__m

on the surface (Z = 0) g _1 - V . My   -
a 0        2G    Jy So, 9 divided by so, that is the amplification function F 2     o l
32 2 +       7 4
I 2 z*+ a -2z F (Z/a) = 2 a a-0i 22 /a2 +2   .
2 (a2 +22 )3 +2      i Ja2+z2 (1-0J             h
                                              . '2-(f)3d)((f)4 I
2 (1- 9) ,          .
                                                               /l +(pf {i +({)f +2(1-2))}l+(p)2 -2 In this model
               ,L-- a      -              a = 210"
                                             =72"(Assume)
                  ,,,,,,,,,            z1 Zi                   e = 0,-{ l -F,(4) +l GyG} F (2v'a ))
lG
             ..g i
9, = the disp. in the case of semi-infinite soil of G1
               'G 2 F,(ZVa) = Fa72C /2l0) - 0.779 G = 0, = 0.239 t- amplification factor equivalent shear velocity, eVs eVs = 2000 /           (0.239)1/2 = 4091 ft/sec equivalent shear modulus, eG eG = 9.09 x 104 /0.239 = 3.803 x 105 (psi)
Rotational Soil Spring Ke 8 eG a3                                       v = 0.40 Poisson's Ratio Ke=                     (by Borowicka)
                          = 1.5'7 x 1013 lb in/ rad i
697 l
l                                                                                                      m
: 3.       Horizontal Soil Spring Horizontal displacement Ux surface force; Txz dxdy 2                 2 L dxdy                              X Ux = 4rG 1Tl<PX +1 R+2-          + (1-2 0)R".I ~ R(R + Z ) I s
So, when force is unifonn stress distribution for circular base 1 2K(a __
Ux =) ] Ux r dr de
                              '             2                     '
                    = Tx7 a. ,L 3 +4 4 _g4,1 -2 u gi +42 _ q) 2G     2 ji +g2           2              -
g, On the surface ( 2 = 6 = 0 )
Uxo = r G       ( 2 - 1> )
Fg m = (2 v)-        /      2 26+l-2 U'+Y-5b Ux = Ux1      1 - Fs C Va ) +(G/GJ Fn C Va))
Fu (ZVa) = Fn (72/2l0) = 0.642
            . . Ux = Uxi = 0.3 73 eVs =2000//O.373 = 32 75 ft/s eGs = 9.09 104 /O 373 = 2.4 3 7 10' psi        .
Horizontal Soil Spring XH 32 (1-v) eGHa KH" 7-8v l                = 2.59 x 108 lb/in 698 J

l
: 4. Vertical Soil Spring As mentioned in the paragraph 2.,        the vertical displacerr.ent is expressed W = 4 x.G           ,     + 2 (1 - 0)]
As the results, Fv is expressed as follows, f v( 2 /a) = j l + (Fa )'     2(              2 7 (l _p py,(2/a)
Fv(72/2l0) = 0.34 3 2
        .'. W = Wi f I- FyC             /a) +(G/G)' FvC2Va))
                  = Wi  0.114 equivalent shear velocity eVs for the vertical eVs = 2000/(.114)l/2 = 5923 f t/sec equivalent shear modulus eG for the vertical eG   = 9.09 x 104 /0.114 = 7.974 x 105(psi)
Vertical Soil Spring, Ky 4G a Ky =                      (by Timochenko & Goodier 1951) l
          = 1.12 x 109 lb/in 1
699 w

The rotational spring constant, Kg , is larger than 1010 K-in/ radian                    j and not acceptable to SAP IV input. It was set equal to             10 10    K-in/ radian
                                                                           ~
A similar value was selected for rotional stiffness about the vertical axis.
I l
700

         ,   _                            _   .. . . - - _ .              ~.
CSCdSO 23ck--CCC's scc 800t for Additional Information on SEP Topic III-6, ' Seismic Reevaluatica" AITACHMEFf IIE-1 SAP 4 RESIDUAL RESPONSE MODIFICATION:
(1) AN IMPROVED PROCEDURE FOR MODAL ANALYSIS (2) COMPUTER COD 3 VERIFICATION CIv!L1.2 4

Q AN lMPROVED PROCEDURE FOR MODAL ANALYSIS by Dilip P. Jhaverl l INTRODUCTION in the dynamic seismic analysis of multi-degree-of-f reedom systems, as well as contintous systems, the method most frequently used is the modal anal-ysis procedure.        In this procedure fi rst the natural modes of the systems are computed in terms of their f requencies and mode shapes which uncouple the system equations. The response of the various modes, in terms of the uncoupled generalised co-ordinates, are then computed. The system response is then found by adding the modal responses.           The efficiency of the modal analysis procedure lies in the fact that it reduces the problem of integrating a large set of coupled differential equations to a set of uncoupled differential equations.          Further, only a few modes usually con-t ribute to most of the responses.        This reduces the cost even further.
There are no well established rules in determining the number of modes to be included in a particular analysis.           This depends very much on the nature of the dynamic load acting on the structure, and the structure's dynami c p rope rt i es . Such quantities as modal participation factors, which indicate in a gross manner the proportion of the dynamic load acting on a mode, are used to provide an indication of the importance of the mode. Also, the f requency range in which the dynamic load has significant powe r usually p rovides an indication as to the modes that will contribute I Vi ce-P res i den t , URS/ John A. Blune & Associates , Engineers, San Francisco, CA

  ~-
I i
significantly to the response. These considerations and the experience and Judgement of the analyst, usually determines the number of modes considered in a particular analysis.
Herein a procedure is developed to estimate the error involved in terms of proportion of load neglected in the analysis due to the truncation         l to a particular number of modes.
An approximate procedure involving static analysis for correction of response in the case of seismic load (l.c., support motion) is also pre-sented. Theory and examples of application of this procedure are given in the next two sections. The last two sections discuss the significance of these procedures and conclusions.
j i                                                                                    -

l THEORY The equations of motion of a multi degree-of-f reedom system may be written in the standard matrix form as mV + cv + kV         = fr(t)                                          (1) in which V,    9, V are respectively the relative displacement, velocity and             4 acceleration vectors of size Nxl where N is the number of dynamic degrees of freedom of the system.          m, k and C are, respectively the NxN mass, stiffness and damping matrices,          f is the vector of amplitudes of applied forces and r(t) represents the variation of the force vector with time.
The choice of components of displacement vector V is such that equations
* of motion (1) form a coupled set of equations.            In the nodal analysis procedure, advantage is taken of the fact that there exists a set of normal or generalized co-ordinates Y in which these corresponding undamped equations mV + kV     =    fr(t)                                                (2) are coupled.       The two sets of co-ordinates are related by v =      Cy                                                           (3) where 4 is the NxN mode shape matrix such than Ct m;   =   I                                                         (4) and         ,tm?    =   [' u,2 },                                                 (5) where I is an NxN Identity matrix, and ['u, s} is an NxN diagonal matrix consisting of of as the Ith diagonal element, og being the Ith frequency                      ;
of the system.
I I                                                                                                    r

th; cqu tions of motions (1) for th2 dnmpad systcm can clso bo uncoupled if the damping matrix c is orthogenalised by the same transformation, i.e.
T C c0   =
[*2A.4J t
(6) where ['2A,.u.s)
                .t is the diagonal matrix and A.tis the damping for the Ith mode. The conditions under which the damping matrix c is orthogenalised is discussed fully in Reference 1.
i Substituting transformation (3) into equations of motion (1), premulti-              ,
plying by C# and using relations (4), (5) and (6), we get the uncoupled set of N equations.
IE + [' 21.u.,]E + [' u ,]Y     =  Fer/t)                        (7) t t          t where T
F   =  cf                                                        (8)
F is the force vector in the normal co-ordinates. We can write the force vector in geometric co-ordinates, f , in terms of F using equations (8) and (4):
1 f   =
[:")~p    = m:F                                          (9)
Note that 0 is a NxN matrix and its {nverse can be shown to exist for s table sys tems.
In practice it is common to use only a few modes, say M, M<N, which gener-ally dominate the responses.         Equation (9) can be used to obtain an estimate of the residual forces not considered by restricting the number of modes used in the analysis.          First let us partition the mode shape mat rix C into two parts

t       [C A 'B]                                                  (10) where 4   is NxM matrix of the M modes considered and gt is Nx(N-M) matrix 4
of the N-M modes not conside. red.       Equation (9) can then be written as F
f  =
[M] [CA 'B3IF}   B (II) where F is partitioned into Mxl vector FA and (N-M)x1 vector FB following the partition of f in equation (10)
F         4             ' I F   =
{pb)    =  {A)fp     =   {ApI }                           (12)
B (B            'B expanding equation (11),
f   =             "
fA+ B          M# FA A + N# FBB                           II3) or f     "      -f A "            F                                  Il4)
B                    # - M*A A where Bf is the residual portion of the force vector f excluded by limiting the analysis to M modes.
Thus fg can be evaluated from (13) once the M modes of the system have been obtained.      Comparison of elements of fb and f can then provide an estimate and the portion of applied force vector being negleceed by restricting analysis to M rzdes, m

DISCUSSION AND APPLICATIONS For further discussion of the signi ficance of the above formulation, we will restrict our attention to the case of seismic forces i.e. , base support mot ion. The force vector f is given by f  =   MD (15) where D is the Nx1 vector of displacements caused by rigid body displace-ments of the base in a given direction and -P(t) la then the acceleration time-history of the base support motion. F in equation (8) then defines the modal participation factor (MPF).         Note that from equation (9) and (15) we get the relation D  =    CF (16) a relation that has been directly or indirectly recognized previously by some investigators.
Applying equation (15) to equation (13) and (14), and using equation (16) ,
one gets
                       *           =
f b
F M*B B      M(D - CAAF)                                  IU) and CF BB
                         =
(D - tAA F)                                            (10) where FA , FB are the rmdal participation factors of M modes considered and (N-M) modes not considered, respectively.
In seismic analysis it is generally found that the first few modes or generalised co-ordinates, (when they are arranged in ascending order of thei r f requencies , i .e. uy<U2*UJ.         . .<u <. . . < u ) generally dominate the response. This is usually reflected in the values of modal (note:    M and m used interchangeably as mass matrix)
                                                                                          ~

participation f actors which are generally smaller for the higher modes.
Also, recognizing that there is no amplification in the ' acceleration response spectra of recorded earthquake accelerograms beyond a certain frequency that frequency has been designated a "cut-off" frequency fe, i.e.
S (f, A) = liigl mar for f>f, where S (f, A) istheresponsespectrumvalueandlliglma Is the peak ground acceleration for frequency f and damping ratio A.       The modes with frequencies greater than f, may then be considered as rigid modes and the absolute acceleration response WBdue to those modes may be evaluated as follows.
U   "
B      #B I' #a'YB
                      =
eF'II BB d!###
                      =
(D-c  AAF N !* *                                         (I9)
Thus hB can be evaluated without the kncwledge of the 8 modes having frequencies greater than f, .      The inertia forces due to this rigid B modes are given by cF g   =  $  8                                                    (20) which may then be applied as a static force to obtain the residual displace-ments, Vg , and, more importantly stresses in the structure,       it may also be noted that all of the rigid modes will essentially have the same time history as the base support acceleration, a fact that may be used in a modal superpositiorat time-history analysis.

Rev. 7
                                                     ,                            D:te 10/1/79 DS-            ,
FORM 4.3-2 URS/BLUME VERIFICATION OF COMPUTER PROGRAM
* Program:     SAP 4 Version:     NGE 4.0 Principal Use: SAP 4.is a structural analysis program for static and dynamic response of linear structures and piping systems. It can be used for static analysis , eigenvalue extraction, mode superposition response history analysis, response spectrum analysis, and direct integration response history analysis.
Description of program modifications: The modifications. in this version consist df:
(1)   In response spectrum analysis.more options, in addition'to SRSS, can be used in combining modal response of structures if their natural modes are closely spaced.
(2) A new element, superelement, is added into the element library of SAP 4.
(3) In response spectrum analysis, the residual responses of the structure 9't -
due to cut-off of the high frequency modes are calculated and printed.
(4) Sequential file I/O is replaced by random access I/O to save computer time in 'I/0.
Location of Verification Documentation: NGE Division computer library.
The above program has been verified, checked, and approved in accordance with established procedures, and I certify to this as noted below:
Verified by                 _ _h. ,                     Date // 28 ,
Checked by "W' h                  Date lI (H- f Po Approved by          M           -        VA            Date      / //R      '
N Manager of Computer Operations or Project Manager g,.
* Revised in Revision 7
                                                                    ,e.  .m.w.       . m ee *=   ******.'-oe

    - _ . . . . .  ,     .   . . _ . . . _ . . . . . _ . - . = _ . . . . . . _ . . _ . , ._ . . - . . . . _ . . - . . . - _ . _ __ _ , _ _ _ . . .
t
                                                                                                                                                     +
SAP 4 Residual Response Modi fication I                     Contents-
: 1. Purpose' and Description
: 11. Input /Outpu t : Modi fi ca tions Ill. Re fe rences IV. Computer implementations V. Veri fication Problems 4
   .O i
I l
l l
i l
t i
f-l l
l l
h      .

I
: 1. Purpose and Description in the response spectrum analysis, the common practice is to extract only those natural f requencies below 33 Hz. Those unextracted modes are belonged to the rigid range and generally are unaccounted for in the analysis.
The purpose of this modification is to include the residual response due to        .i t runcated hi gher modes. Since all those modes above 33 Hz are in the rigid range, the forces contributed by them can be considered as static loads.
Therefore, the total response of a system can be obtained by adding the dy-               .
namic response and the residual static response. The de tailed theoretical derivation of the residual response is shown in Reference 1 and will not be repeated he re. The modi fication of the input / output of the original SAP 4    .
program is shown in Section 11.
In addi tion to residual responses, nodal translational and rotational accelerations are calculated and printed. However, only those accelerations-in the direction of earthquake ground motion should be used because these accelerations are absolute. The accuracy of accelerations in di rections Cs othe r than that of earthquake ground motion is questionable due to dif fi-culties in defining the relative spectral acceleration spectra.
Translational and rotational residual nodal forces and residual nodal accelerations are calculated in the program. Nevertheless, only transla-tional residual responses are printed.

                                                                                                  ~2 I I . - input / Output Hodi ficat ions A. Input All input to the ' SAP 4 program remains unchanged.
: 8. Output
: 1. The residual nodal forces due to 1 9 ground accelerations are printed af ter the nodal displacement printout.
: 2. The residual nodal accelerations due to Ig ground acceleration as printed.
3    Nodal translational and rotational accelerations are printed.
C
'l
  +
4 4
d f
1 I
              .._.._.___u.

1 111. Re fe ren ces,
 !          1. Jhaveri, Dilip P., "Evaluation of Residual Response in Truncated Ndal Analysis Procedure," Paper No. 12/6 presented at Civil Engineering and Nuclear Power ASCE Specialty Conference,' Knoxville, Tennessee, September 17, 1980.
i C
I 4
i

t IV. Computer Implementations A. Subroutines involved in Modificatico                                            ;
The subroutines involved in the modification are listed below under the category of overlay.
: 1. Overlay.0
: a. SAP 4
: b. TH REE D
: c. SHELL
: d. PRINTD
: e. STRSC
: 2. Overlay 5
: a. TPLATE 3    Overlay 15
: a. RESPEC
: b. SPECTR
: c. STRESR B. Modi fi cation Outline Brief outlines of' the modifications in the subroutines listed in A are given in the following:
: 1. SAP 4              TAPE 30 and TAPE 77 are added into the program s t atement
: 2. TH REE D           Element stresses are printed and saved on TAPE 30.
3    SHELL              Plate elenent stresses are printed and saved on TAP E 30,
: 4. PRINTD             Residual nodal forces , residual nodal acceleration responses are printed and saved on TAPE 30.
t

P
: 5. STRSC              Nodal forces of plate elements are calculated, printed, and saved on TAPE 30.
: 6. TPLATE             Plate element sti ffness matrices to element in-cidences are saved on TAPE 77 7   RESPEC             Hore core storage is allocated in this subroutine due to the modification in subroutines SPECTR and STRE S R.
: 8. SPECTR             Residual nodal forces and residual nodal accelera-     .
tions are calculated in this subroutine.
9   STRESR             Element nodal forces for each mode and total element nodal forces are calculated, printed and saved on TAPE 30.
C. Flow Charts Figures 4.1 to 4.8 provide flow-charts of the logic involved in the modification of residual response.
C I
m

URS/BLUME 130 Jeule Street (at New Montgomery)                                                                          [
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URS/BLUME 00 Jessie Street (O New Montgomery)
San Francoco, Caldorn'a 90W5                                                                      SHEET NO am                   a                                                     EF(HCfjM DATE CLIENT               SUBJEb7                        .
CWD N 0 b DATE 8+ N' A. rXNAu lc_- AM4LYs15 VO gesspuAL             Non4L      Fmas               -
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t> <                                       <am k
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pier                    paiarp
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* S D                           aspew er     5 T* t
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                                                                          +> +5=
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URS/BLUME 130 Jessie Street (3 New Montgomery)                                         ,
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Caddan Geck--CRC's Dequest for Additional Information on SEP Topic !!I-6, "seismic Reevaluation
* ATTACHMENT IIE-2 EVALUATION OF STRUCTURAL STABILITY:
(1) SCREENVELL HOUSE (2) CORTAINMENT STRUCTURE (3) PRIMARY AUXILIARY BUILDING CIVIL 1.2

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130 Jessie Street (at New Montgomery)                                                                                 )
San Francisco, CaMornia 94105                                                                  SHEET NO 4.4,05        '
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San Fe, CaMornia 94105                                                                                                                                   Stet NO 7 6,o /
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Ucddon Cock--ERC's Bcquost for                                              ,
Additional Information on SEP Topic III-6, ' seismic Reevaluation
* ATf.iCHMENT IID LIST OF NUSCO DRAVINGS (i)   AUXILIARY FEEDVATER PUMPHOUSE 16103-51048:
SH.1 - H.S. & F.V. Piping Layout SH.2 - Ficor Platform Beams & Connections SH.3 - Schedule of Steel Inserts in Concrete SH.4   -    Steel Columns Base Support Details 16103-22090:
SH.8 - Pinned Buttress Connection Between Shield Vall &
Tornado Vall SH.10 - Vall Bumpers SH.20 - Installation Layout Assembly of Pinned Buttress & Base Connection for Tornado Vall and Shield Vall SH.34 - Tornado to Shield Vall Base Connection 16103-56026:
SH.1&2- Concrete Details for Tornado Protection (ii)   REAcr0R CONTAINMENT INTERIOR STRUCTURE 16103-50096:
                        - Interior Concrete Outline EL. 48'-6" 16103-50098:
SH.1 - Interior Concrete Details EL. 48'-6" (iii)   REACTOR CONTAINMENT EXTERIOR STRUCTURE 16103-50086:
SH.3 - Pipe Gallery Interface Concrete Details CIv!L1.2 w _-

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OVERSIZE DOCUMENT PAGE PULLED SEE APERTURE CARDS
_.ER O, _e,ze _e m _ _ __ _e                             ls APERTURE CARD /HARD COPY AVAILABLE FROM RECORD SERVICES BRANCH FTS 492-8989 l
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