ML13311A394

Rev 1 to Impact of Pipe Support Loads on Structures (IPSLS),Safety-Related

ML13311A394
Person / Time
Site:	05000000, San Onofre
Issue date:	10/10/1984
From:	BECHTEL GROUP, INC.
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ML13311A392	List: ML13311A391 ML13311A393 ML13311A394 ML13311A395 ML13311A397 ML13324A643
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M-37458, NUDOCS 8505310252
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Text

• E*

RETURN TO SERVICE DESIGN CRITERIA FOR SAN ONOFRE NUCLEAR GENERATING STATION, UNIT I IMPACT OF PIPE SUPPORT LOADS ON STRUCTstS (IPSLS)

SAFETY RELATED 14000-430; Joe No. 15691-430/471 BECHTEL POWER CORPORATION NORWALK. CALIFORNIA SCE WDRK PACKAGE 82-281 82-2"

-~553 10252 650514 PDR ADOCIK 0 5 0 20 Incorporated minor changes

& CE comments, 1

c/8 sued to SCE for -pprov 1

Issued SCZ or Approval

~ /984 Incorporated G! Goments and Issued to 0

8 r-F Ear Apprn59l D

3/5/84 Issue For so eview O

C 2/174 Revised for RTS Scope, Issued for BPC Review

& Comments sc.

AT

-aei~

-a ~

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SAN QNQFNE NUCLEAR esNERATIN STATION UNIT I TITLE IMPACT OF PIPE SUPPORT IDADS 14000-430; ON STRUCTURES (IPSLS)

O.15691-430/471 Ne 1

PROJECT DESIGN CRITERIA A2 A2 k a 1.0 if.

simple span A

k a 0.54 if the adjacent span is elastic.

k w 0.8 if the adjacent span is yielded.

References

1.

"Steel Structures Design and Behavior" Salmon, Johnson, 1971.

2.

"Bertero, V. Private Communications, 1984.

SAN ONOFRE NUCLEAN GENaRATING STATION UNIT 1 TITLE IMPACT OF PIPE SUPPORT LOADS 14000-430; ON STRUCTURES (IPSLS) ae.15691-430/471g,,1 PROJECT DESIGN CRITERIA sAs i

CONTENTS

1.

GENERAL DESIGN CRITERIA 1

1.1 General 1

1.2 Scope 1

1.3 Design Criteria 1

1.3.1 Seismic 1

1.4 Classification of Structures, Systems and Components 1

1.4.1 Quality Class 3

1.4.2 Safety Class 3

1.4.3 Quality Group 3

1.4.4 Seismic Category 3

1.4.5 Design Codes and Standards 3

1.5 Environmental Design Criteria 3

1.5.1 Normal 3

1.5.2 Post-Accident 3

1.6 Design Life 4

1.7 NRC Regulatory Documents 4

1.7.1 Code of Federal Regulations 4

1.7.2 Regulatory Guides 4

1.8 State and Local Regulations 4

1.8.1 Occupational Safety and Realth Administration (OSHA)

Regulations 4

1.8.2 California Administrative Code 4

1.8.3 SONGS 1 Documentation 4

1.9 References 5

2.

MECHANICAL DESIGN CRITERIA 5

3.

ELECTRICAL DESIGN CRITERIA 5

4.

CIVIL/STRUCTURAL DESIGN CRITERIA 5

4.1 General 5

4.2 Design Loads and Load Combinations 6

4.2.1 Dead Loads (D) 6 4.2.2 Live Loads CL) 6 4.2.3 Loading Combinations 6

4.3 Evaluation Procedure 7

4.3.1 Acceptance Criteria for Evaluation of Members 8

4.3.1.1 Structural Steel Intermediate Members 8

4.3.1.2 Acceptance Criteria for Structural Columns 11 4.3.2 Acceptance Criteria for Structural Steel Connections 12 4.3.2.1 Establishment of Connection Yield Strength 13 4.3.2.2 Material Properties 13 4.3.3 Allowable Stresses for Connection Evaluation 14 4.3.3.1 Shear on Bolts 14

SAN ONOFE NUCLEAR 6ggNaATING STATION UNIT I TITLE IMPACT OF PIPE SUPPORT LOADS 1400-430*

ON STRUCTURS (IPSLS) 0e.

m91-430671 1v.

PROJECT DESIGN CRITERIA sesii ii CONTENTS CContinued)

Page 4.3.3.2 Bolts Subjected to Tension 15 4.3.3.3 Bolts Subjected to Shear and Tension, Interactions 15 4.3.3.4 Bearing 15 4.3.3.5 Shear 15 4.3.3.6 Weld Stresses 15 4.3.3.7 Moment Connections 16 4.3.4 Rilti vik Expansion Anchors in Concrete 16 4.3.5 Rock Bolt Expansion Anchors 16 4.4 Design of Structural Modifications 17 4.4.1 Codes, Standards and Regulations 17 4.4.2 Load Combinations and Acceptance Criteria 17 4.4.2.1 Modifications to Structural Members 17 4.4.2.2 Design of Connection Modifications 20 4.4.2.3 Design of ilti Kvik Expansion Anchors 20 4.4.2.4 Design of Rock Bolt Expansion Anchors 21

5.

CONTROL SYSTEM DESIGN CRITERIA 21

6.

NUCLEAR DESIGN CRITERIA 21

7.

PLIANT DESIGN CRITERIA 21 APPENDIX A Critical Unbraced Length 1A

SAN QNQFtr NUCLEAR SENERATINg STATION UNIT I TITLE IMPACT OF PIPE SUPPORT LOADS 14000-430 ON STRUCTURES (IPSLS)

NO. 15691-430 v71.

PROJECT DESIGN CRITERIA paes

-iii or TABLES Page 1.4-1 Classification of Structures, Systems and Components 2

4-1 Allowable Member Ductilities 9

4-2 Allowable Design Loads for Concrete Expansion Anchors 18 4-3 Allowable Design Loads for Rock Bolt Expansion Anchors 19

AN QuoNt6 NUCLEAR 6SNECRATINB STATsON UNIT I TITLE IMPACT OF PIPE SUPPORT LOADS 14000-430 S1DN STRUCTURES (IPSLS)

NO.15691-430/471 1

PROJECT DESIGN CRITERIA 1

Ao 21

1.

GENERAL DESIGN CRITERIA 1.1 GENERAL Due to the pipe support reaction loads of safe-shutdown RTS piping the secondary structural elements need to be evaluated and upgraded, if required, to satisfy their seismic withstand capability.

This work is described as Impact of Pipe Support Loads on Structures (IPSLS).

1.2 SCOPE The design criteria address the analysis of existing structural members and their connections for the IPSLS condition and subsequent design and construction of modifications.

Modifications to existing structural steel elements shall be made by adding cover plates, stiffeners, providing additional members to support the loads coming from pipe supports or replacing existing members with larger size members.

Modifications to the structural connections involve replacing existing bolts with high strength bolts, increasing sizes of existing welds, installing new connectors in deck slabs and providing new seat angles or brackets to existing beans.

All structural elements or supports shall be designed to withstand predefined seismic loads without causing a failure which could impact a safety related structure, system or component. Included are all the plant areas in this scope of work.

1.3 DESIGN CRITERIA 1.3.1 SEISHIC All reevaluations and/or modifications for IPSLS shall meet the requirements of the Seismic Upgrade General Design Criteria Manual (Reference 1) unless stated otherwise in these criteria. The Design response spectra for DBE are defined in Appendix 2A of reference 1.

1L.4 CLASSIFICATION OF STRUCTURES. SYSTEMS AND COMPONENTS Quality class and seismic category assigned to project structures, systems and components is included in Table 1.4-1.

&AN QNQFRE NUCLEAR SENERATING STATIQN UNIT I TITLE MPACT OF PIPE SUPPORT LOADS 14000-430; ON STRUCTURES (IPSLS)

NO. 15691-430/471a.

PROJECT DESIGN CRITERIA 3

21 1.6.1 QUALITY CLASS quality Class is as defined in Section 1.4.1 of the Seismic Upgrade General Design Criteria (Reference 1). 'Table 1.4-1 defines the applicable quality class of the components associated with this modification.

1.4.2 SAFETY CLASS.

Safety class is as defined in Section 1.4.2 of the Seismic Upgrade General Design Criteria (Reference 1).

Table 1.4-1 defines the applicable safety class of the components associated with this modification.

1.4.3 QUALITY GROUP Quality group is as defined in Section 1.4.3 of the Seismic Upgrade General Design Criteria (Reference 1). Table 1.4-1 defines the applicable quality group of the components associated with this modification.

1.4.4 SEISMIC CATEGORY' Seismic category is as defined in Section 1.4.4 of the Seismic Upgrade General Design Criteria (Reference 1). Table 1.4-1 defines the applicable seismic category of the components associated with this modification.

1.4.5 DESIGN CODES AND STANDARDS Design codes and standards are identified in Section 4.4.

1.5 ENVIRONMENTAL DESIGN CRITERIA 1.5.1 MORMAL Normal environment is defined in Section 1.5.1 (Reference 1).

1.5.2 POST-ACCIDENT LOCA and MSB are not an evaluation and design requirement.

SAN ONQFWE NUCLEAR SENERATING STATION UNIT I TITLE IMPACT OF PIPE SUPPORT LOADS 14000-430; ON STRUCTURES (IPSLS)

.15691-430/471 1

PROJECT DESIGN CRITERIA 5

.-Op 21

1.9 REFERENCES

1)

"Seismic Upgrade General Design Criteria", San Onofre Unit 1 Project Design Criteria Manual, Rev. 2 dated July 12, 1984.

2)

"Balance of Plant Structures Seismic Reevaluation Criteria",

San Onofre Unit 1, dated February 17, 1981, reissued January 20, 1983.

3) "Manual of Steel Construction", Eighth Edition, American Institute of Steel Construction, Inc.

4)

"Guide to Stability Design Criteria for Metal Structures", Third Edition, Bruce G. Johnson Editor, John Wiley and Sons.

5)

Canadian Standards Association Code, "Steel Structures for Buildings Limit States Design, 1974.

6)

Newark et, al Airforce Design Manual, AFSWC-DR-62-138.

7) Kato and Akiyama, ASCE Spring Convention 1981, New York, Reprint 100,

'Ductility of Hembers and Frames',

Subject:

Buckling.

2.

MECRANICAL DESIGN CRITERIA Design, if necessary, will follow Chapter 2.0 of Reference 1.

3.

ELECTRICAL DESIGN CRITERIA Design, if necessary, will follow Chapter 3.0 of Reference 1.

4.

CIVIL/STRUCTURAL DESIGN CRITERIA 4.1 GENERAL mthe following design criteria apply to the analysis of existing structural

embers and their connections and subsequent design and construction of fmodifications required for the Impact of Pipe Support Loads on Structures (IPSLS).

For the analysis of the existing structural members and their connections, the provisions of Sections 4.2 and 4.3 in conjunction with the "Balance of Plant Structures (BOPS) Seismic Reevaluation Criteria, SONGS 1", Reference 2 will be applicable.

For the IPSLS evaluation, however, the requirements of these criteria will govern. The design of Structural modifications, when required, shall be done in accordance with Section 4.4.

BAN ONOFRE NUCLEAR 0ENERATING STATION UNIT I TITLE DMPACT OF PIPE SUPPORT LOADS 14000-430; ON STRUCTURES (IPSLS)

Ne.15691-430/471s.

1 PROJECT DESIGN CRITERIA 7

21 IPAGE Or SAM a Seismic Anchor Movement.

Z' member a Loads in the member due to DBE, other than pipe reactions (i.e., the inertial response of the structure).

D a Dead loads or their related internal moments and forces L = Applicable live loads or their related internal moments and forces T

= Thermal effects and loads during normal operating conditions based on the steady-state condition R a Mazimum pipe and equipment reactions during normal operating conditions based on the steady-state condition, if not included in the above loads Notes:

(a)

For the load combinations where D or L reduce the effects of other loads, the corresponding coefficients shall be taken as 0.90 for D and zero for L.

(b)

T will not be considered when it can be shown that the load is secondary and self-limiting in nature.

(c)

If more than 2 pipe supports from the same pipe are located on a single structural member, the member stresses from these pipe reactions will be combined in a statically consistent manner prior to combining this result, by SRSS, with the member stresses from all other pipe reactions on the same member.

4.3 EVALUATION PROCEDURE The purpose of the analysis is to determine the stress level of selected structural members which are subjected to pipe reaction loads, their connections and to establish their adequacy for the Return-to-Service (RTS) condition.

A.

The affected structural members and their connections will be evaluated with consideration of the pipe loads and loading combination described in Section 4.2.3.

SAN QN@QFR NUCLEAR SENERAimS STATIg on UNIT I TITLE IMPACT OF PIPE SUPPORT LOADS 14000-4301 ON STRUCTURES (IPSLS)

"e.v15691-430/471 PROJECT DESIGN CRITERIA VA0s 9

21 A.

The computed ductility of each structural member is determined by relating the associated forces.

The ductility will be computed as follows:

(2p

-1/2

%here a a Computed oment in member from elastic analysis Np a Plastic Moment Capacity of the section (Z F or Z F ) or reduced plastic moment capacity due to effeft of jrIload p

a Calculated ductility demand in the member B.

For biaxial bending, the ductility will be calculated by summing the moment ratio of each axis or:

Mt M

(2p -

1)1/2

+ M px py C.

If an axial load is present, ductility will be calculated as follows:

M H

(2p - 1)1/2 P

P y

px py D.

Structural member will be acceptable if p ma Acceptable ductility demand in the member, given in Table 4-1.

am p

a Caleulsted ductility demand in the member z

&AN QNOFEa NUCLEAR SENERATINe STATION UNIT I TITLE IMPACT OF PIPE SUPPORT LOADS 14000-430; ON STRUCTURS (IPSLS) 1..15691-430/471nev.

PROJECT DESIGN CRITERIA

_]11 21 b)

The torsional shear stresses in the connections will be less than the criteria limits.

c) Otherwise modifications will be implemented to eliminate torsion.

7.3 Shear a)

The plastic moment capacity of the beams will be reduced where the effects of shear are considered to be significant.

F.4 Lateral-Torsional Buckling a) Major axis bending only: unbraced length requirements corresponding to the demand rotation capacity will be satisfied (e.g. see Appendix A) b)

Minor axis bending only:

there will be no limit on the unbraced length for beams bent about the minor axis bending when b/t requirements are satisfied.

c)

Biazial bending:

There will be no limit on the unbraced length when the major axis bending moment is less than the critical moment and one flange of the beam is not totally plastic under both major and minor axis bending.

4.3.1.2 Acceptance Criteria for Structural Columns The AISC specification (Reference 3) does not provide for the ultimate strength-of columns in biaxial bending, the Code provisions are for single axis bending only.

In this evaluation both the major axis and the minor axis moment will be taken into account along with the axial load in determining the ultimate capacity of the steel columns.

The interraction equation used will be:

+

< 1. 0 (1)

Py 1.18 psz 1.67 py P

a Applied axial load, kips Py a Yield stress a section area (7 x A)

Mz a Applied moment, major axis kipyinches My a Applied moment, minor axis kip-inches Mpx z Plastic moment capacity, major axis, kip-inch (Z x

)

Hpy a Plastic moment capacity, minor axis, kip-inch (ZR x Fy) y y

SAN OnOFES NUCLEAR GENSE ATINO STATION UNIT I 14000-430; TITLE IMPACT OF PIPE SUPR 4N 1

4 7 1 1

a0 STRUCTURES (IPSLS)

No.15691-430 1

PROJECT DESIGN CRITERIA wAss 13

-4, 21 4.3.2.1 Establishment of Connection Capacity be connection Capacity, y, will be established by evaluating the following individual coonent as deemed necessary, using the allovables as described in Section 4.3.3.

Bolted connections:

1.

bolt shear

2. Dolt bearing on connecting material

3.

Weld between clip angle or plate and the supporting member

4.

Shear on the net area of the clip angles or connection plate

5.

Edge distance of the bolt

6.

Web tear out (block shear)

7.

Combined shear and tension on the bolts Moment Connections between Beam and Columns:

1.

Column or beam web panel-shear

2.

Column or beam web stiffeners will be checked against buckling.

3.

Various welds at the beam and column interface will be checked.

4.3.2.2 Material Properties

1. Bolts Type Yield Stress Ultimate Strength A307 60 ksi ainimum A325 92 ksi 120 ksi minimum (1/2 in. to 1 in.)

A325 81 ksi 105 ksi minimum (1-1/8 in. to 1-1/2 in.)

2.

A36 Steel 36 ksi 58ksi iniUlin aL

__________________)

-v%'~4~?

SAN ONOFRS NUCLEAR QENERATING STATsQN UNIT I TITLE IPACT OF PIPE SUPPORT LOADS 14000-430; ON STRUCTURES (IPSLS) ne.15691-430/471.v.

1 PROJECT DESIGN CRITERIA 15 21 III) Coefficient of Friction k -w 0.4 may be used if the slip-loaeof the connection with blast cleaned surfaces, is to be determined.

4.3.3.2 Bolts Subjected to Tension For high strength bolts A-325 or A-307 bolts, in tension Ft = 0.60 Fu where Ft is calculated on the gross area 4.3.3.3 Bolts Subjected to Shear and Tension, Interactions For the combined shear and tension forces applied on the bolts, the bolts will be acceptable if they meet the following:

(fv/0.6Fu) 2 +

/

2 0.75 Where: f a Applied shear load t = Applied tensile load 4.3.3.4 Bearing (On projected area of bolts)

F = 1.5 F p

n On projected area of bolts in shear connections; where F is the specified minimum tensile strength of the weaker of the connected ;arts in ksi.

Edge distance of the bolts will also be checked to assure that plate and tear out cannot occur.

4.3.3.5 Shear (Net area of clip angles or connection plate)

Thru plate material, such as web of beam or connection clip angles:

7 u0.6 F 4.3.3.6 Weld Stresses A.

Butt welds and full penetration welds will be assumed capable of developing the tensile yield stress of the base metal on their minimum throat section.

S"N ON@FmKg NUCL&AII GENCRATINGS TATIONi UNITI TITLE IMPACT OF PIPE SUPPORT LOADS 14000'430; ON STRUCTURES (IPSLS) e.15691-430/471, 1

PROJECT DESIGN CRITERIA PAss 21 6.4 DESIGN QF STRUCTURAL MODIFICATIONS 4nbe following design criteria apply to the design Ad construction of new modifications deemed necessary as a result of the evaluation of IPSLS a described in sections 4.2 and 4.3.

When structural modifications are required they will be made only to those specific components of the structural framing that do not meet the acceptance criteria.

The-following listed data shall be utilized for these design criteria.

4.4.1 CODES, STANDARDS AND REGULATIONS Refer to Section 1.7 of the ISeisuic 'Upgrade General Design Criteria (Reference I). SpecificallYs the following codes and standards will be used along with the provisions of this criteria for the design of the structural modifications.

AISC 1978 Specification for the Design, Fabrication and Erection of the Structural Steel for Buildings.

AWS D1.1-80 Structural Welding Code ACI 315-74 Manual of Standard Practice for Detailing Reinforced Concrete Structures ACI 318-77 Building Code Requirements for Reinforced Concrete ACI 349-80 Code Requirements for Nuclear Safety-Related Structures ASTH A36-77 Specification for Structural Steel ASTH A307-80 Specification for Carbon Steel Externally and Internally Threaded Standard Fasteners ASTM A325-80 Specification for Nigh-Strength Bolts for Structural Steel Joints, Including Suitable Nuts and Plane Eardened Washers.

4.4.2 LOAD CONSINATIONS AND ACCEPTANCE CRITERIA Loads and loading combinations will be same as Section 4.2.3.

4.4.2.1 Modifications to Structural Members The response of the modified structural member in the inelastic range will be acceptable if it meets the following requirements.

TABLE 4-3 ALUABLE DESIGN LOADS FOR ROCK BOLT EXPANSION ANCRORS

-4.

Preteasiofling

-!9)

I r

DeAow bd(2)

Torque Torque at Tension Ein His

.e Aesighora (1t-Lbs) to Installation Test Embed-in. C/C EdgeL Diameter Tenion Shear Expand (Ft-Lbs) Threads Load meat Spacing Distance o

(Inceter Tis)o (Kips)

Shell Not Lubricated (Kips) (Inches) (Inches)

(Inches)

(inches)

(Kips)(Kp)t 1

25(5)33(6) 8 12(8 200-250 300 30 14 10 6

1-3/8 50(5)66(6) 16 24(8 700-750 1100(3 60 18 14 8

(I )(6) 87950 (4)

(3) 2 0a 2

100(133 334 950-1000) 3300 120 24 20 10 (1) For 4000 psi (fc) or higher concrete.

51 do 4.9.5.

and cobndineatonpr (2) Subject to reduction per paragraphs 4.9.5.5 and 4.9.5.6 and combined interaction Per paragraph 4.9.5.7 of Reference I.

5 (3)

Preferred method is to pretension to specified test load using calibrated holloare

)1 hydraulic jack or calibrated stud tensioner.

(4) May be increased to 1500 ft. lbs. if required to prevent slight pullout of bolt which MY be experienced upon application of 3300 ft-lb torque for pretensionifg prior to grouting.

(5 Nanufacturer's recommended design load at 2:1 safety factor.

(6) These increased allowable loads are applicable only for "Abnorsal/Etree Environmental" (Design basis Earthquake) or "Faulted" loading combinations.

They are based on 0.9 times

i.

Hanufacturer's maximum working load to elastic limit.

) Preferred design load based on AISC limits using manufacturer's ultimate strength values.

(8) Design loads increased by 1.5 applicable only for "Abnormal/Extreme Environmental" (DBE) or "Faulted" loading combinations.

(9) The minimum embedment excludes the head assembly.

SAN QMQFEE NUCLEAR GENER ATIMS STATION UNIT I TITLE IMPACT OF PIPE SUPPORT OADS 14000-430; ON STRUCTURES (IPSLS) 15691-430/471 1

PROJECT DESIGN CRITERIA 21 2

.4.4.2.4 Desj&n of Rock Bolt Epansion Anchors acceptance Criteria same as Section 4.3.5.

5.

CONTROL SYSTEMS DESIGN CRITERIA Design, if necessary, will follow Chapter 5.0 of Reference 1.

6.

NUCLEAR DESIGN CRITERIA Design, if necessary, will follow Chapter 6.0 of Reference 1.

7. PLANT DESIGN CRITERIA Design, if necessary, will follow criteria for 15691-300 and 15691-339 for pipe support design.

A012 501 October 2, 1984 NRC-SEP Director, Office of Nuclear Reactor Regulation Attention: W. A. Paulson, Acting Chief Operating Reactors Branch No. 5 Division of Licensing U. S. Nuclear Regulatory Commission Washington, D.C. 20555 Gentlemen:

Subject:

Docket No. 50-206 SEP Topic 111-6 Seismic Design Considerations San Onofre Nuclear Generating Station Unit 1 The purpose of this letter is to submit the enclosed report entitled, Electrical Raceway Support Implementation Plan for Return to Service, San Onofre Nuclear Generating Station, Unit 1," dated September 1984. The report describes the electrical raceway support seismic reevaluation and upgrade plan which is being implemented as part of our return to service efforts at San Onofre Unit 1. The report also provides the basis for concluding that with the implementation of the identified modifications to raceway supports, the raceway system will have a seismic withstand capability of.67g without impairment of the overall integrity of the system and without impairment of the capability to achieve and maintain hot standby conditions.

If you have any questions or desire additional information regarding this subject, please contact me.

Very truly yours, M. 0. Medford Manager, Nuclear Licensing RO:2485F Enclosure bcc: (See attached sheet)

ELECTRICAL RACEWAY SUPPORT IMPLEMENTATION PLAN FOR RETURN TO SERVICE San Onofre Nuclear Generating Station Un it 1 September 1984

TABLE OF CONTENTS Section Page

1. Introduction 1

2. Scope of Raceway Support Modifications for Return-to-Service 1

3. Justification for Return-to-Service 2

4. Conclusion 7

5. References 9

6. Tables 10

1.

INTRODUCTION The purpose of this report is to describe the electrical raceway support implementation plan for return-to-service. Justification for the plan is also presented.

2. SCOPE OF RACEWAY SUPPORT MODIFICATIONS FOR RETURN-TO-SERVICE Re-evaluation of the electrical raceway supports is complete. The criteria which was employed for evaluations is stated in Reference 1. The final design of all supports requiring upgrade is substantially complete and most of the supports which did not meet the re-evaluation criteria have been upgraded in the past two years to withstand the 0.67g Housner DBE.

The raceway support modifications have been subgrouped into the following categories:

a.

Cable tray support modifications

b.

Cable tray tie-down modifications

c.

Conduit support modifications

d.

Replacement of masonry wall expansion anchors in ungrouted cells with through bolts.

The cable tray support modifications include the addition of longitudinal and/or transverse supports, upgrading of support members and connections, etc. (exclusive of tie-down modifications).

Table 2.1 identifies the approximate number of raceway support modifications which will be implemented prior to return-to-service (RTS) for each of these categories.

Table 2.1 also provides a comparison of this RTS scope with the total number of support modifications which have been identified as not meeting the reevaluation criteria.

As shown in Table 2.1, the support modifications which will be installed prior to return-to-service include:

a.

A significant portion (over 80 %) of the total tray and conduit support modifications which have been identified as not meeting the re-evaluation criteria.

b.

A significant portion (over 80 %) of the tray tie-down modifications, including all tray tie-down modifications for cantilever supports.

c.

All masonry wall expansion anchor replacement modifications.

Table 2.2 shows the locations of required modifications. As shown in this table the RTS scope for tray and conduit modifications are distributed throughout the plant.

Table 2.2 includes both new supports (which were added to reduce the existing support spacings and provide additional longitudinal brace points) and the modifications of existing supports. As shown in Table 2.3, relatively few new supports were identified as being necessary to meet the re-evaluation criteria compared with the number of existing supports installed at San Onofre Unit 1 (approximately 1,200 cabletray supports and 7,300 conduit supports.) Additionally, as shown in Table 2.3, most of the new supports which were identified to meet the re-evaluation criteria are included in the return-to-service scope.

These tables show that a very significant portion of the raceway supports will be modified prior to return-to-service.

Because of this upgrade, the raceway systems will have a significant increase in margins even though there is a technical basis, as discussed in the following sections, for not upgrading the raceway support systems.

3. JUSTIFICATION FOR RETURN-TO-SERVICE Justification for the return-to-service scope identified in Section 2 consists of:

a.

Conservatisms associated with the evaluation and final design process.

b.

The evidence of high seismic capacity of raceway systems as observed in the extensive testing programs described in References 2 and 3.

3.1 Conservatisms Associated with the Evaluation and Final Design Process The evaluation and final design process employed for SONGS Unit 1 utilized conservative parameters for seismic input and damping as well as a conservative methodology. Each of these considerations is addressed in the following discussions.

Seismic Input.

The basic seismic input (instructure response spectra) used in the re-evaluation and the design of the modifications of the support systems was described in Reference 4. Subsequently, in November 1982, in response to a request expressed by the NRC staff, the conservative nature of instructure spectra was demonstrated. A sunmary of the conservatisms is provided in Table 3.1.

Damping.

The damping values utilized for the re-evaluation of the cable tray support system are conservatively based upon the results of the testing program performed by ANCO Engineers, Incorporated (See Reference 2).

-3 The damping measurements that were documented in Reference 2 were conducted on trapeze and braced cantilever type supports. The cable tray damping data obtained from the ANCO testing were grouped into sets according to the direction of input, the type of tray support, the spacing of bracing and the amount of cable loading. The trend of the data suggests that in most cases a bilinear relationship exists between the damping ratio and the acceleration level.

A typical acceleration vs.

damping relationship is shown in Figure 1.

Extensive evaluation of the test data as prese-ted in Reference 2 was also performed as discussed in Reference 7. Various combinations of systems braced in the same direction were combined and the mean curves and the 15%

non-exceedance probability curves (corresponding approximately to mean minus one sigma) were computed utilizing statistical procedures. Table 3.2 provides the critical damping values from the curves in Reference 7 corresponding to 0.67g, which is the ground level acceleration (lowest acceleration level) for evaluation and design in the SONGS 1 seismic reevaluation program. Using the aceleration vs. damping relationships presented in Reference 7, it is noted that the damping will be higher at acceleration levels above 0.67g.

The conclusions of the testing program concerning damping apply to cable tray support systems in general, as the largest portion of the system damping was the result of the amount of energy dissipated between the adjacent moving cables and through friction between cables. The type of tray and the type of tray support system being utilized (trapeze, cantilever, etc.) was not a significant factor in determining the overall system damping. Three tests (11-118, 11C, and 11D) were performed on cantilever raceway support systems. The configuration tested consisted of a five support, three tier cantilever system with trays supported every 8 feet. A review of the test data indicated that the magnitude of damping observed was somewhat higher than the comparable trapeze system and the resonant frequency was about equal.

Therefore, damping values corresponding to those in Table 3.2 are appropriate for SONGS 1 cable tray support systems. However, to be conservative, a critical damping value of 15% was selected. Table 3.3 shows the damping values used in the re-evaluation and final design of the raceway systems.

Methodology.

The basic concept used in the re-evaluation process is an equivalent static analysis with consideration given to the dynamic character of the seismic loadings. The methodology uses manual calculations and engineering judgement to predict the behavior of a continuous complex system by simplified models, and no credit is taken for load sharing between adjacent supports or the continuous nature of the raceway systems. This process does not reflect the following considerations which are attributable to the actual behavior of raceway support systems as observed in the test programs:

a.

Inelastic action of the support system

b.

Actual damping within the system

-4

c.

Internal load redistribution of a continuous system

d.

Actual material strength versus the allowable design values used for material properties

e.

High level of reserve capacity inherent in steel structures due to material ductility.

Because the evaluation and design process does not reflect the above considerations, the results are considered to be conservative estimates of the stress levels. This, in turn, leads to a conservative amount of modifications necessary to restore design margins. This is illustrated in Table 3.4, which shows the comparison between the test and re-evaluation results for a cable tray support system which was successfully.tested in the Reference 2 test program. As described in Table 3.4, the application of the re-evaluation criteria and methodology to the actual test specimen configuration concluded that modification of the tested support system would be necessary in order to meet the re-evaluation criteria. It should be emphasized that the acceleration values recorded during the testing were greater than the values used in the calculations, and even at the higher test values no damage was observed.

It is concluded that the overall methodology, criteria and procedures used in the evaluation and design process are conservative and that the number of modifications identified would have been significantly reduced if more sophisticated methods which account for the factors a. through e. above had been utilized.

3.2 High Seismic Capacity of Tested Raceway Support Systems Over the past two decades, many earthquakes have occurred within the United States. Of these, several were of sufficient magnitude to cause structural damage to industrial facilities. Following such strong earthquakes, inspection of power generation and distribution facilities has offered valuable information as to the overall performance of engineered structures. The 1971 San Fernando earthquake has been of particular interest in this regard. It was one of the most severe earthquakes Southern California has experienced in recent history. A survey of structural damage to the Sylmar Converter Station, located within a few miles of the epicenter, provided data relative to the behavior of electrical distribution equipment and electrical raceway systems when excited by strong ground motion. Of special interest was the fact that simple unbraced raceway hanger systems were able to survive the earthquake without major structural damage.

Another finding was that even at locations where a minor amount of structural distressing occurred, the cables within the tray systems did not lose their functional integrity.

The fact that the converter station's unbraced support system survived the San Fernando earthquake generated interest regarding the practicality of using similar systems in nuclear power plants.

-5 In the years following the San Fernando earthquake, an increasing effort has been put into the design of earthquake resistant structures. Included in the list of structures are nuclear power plants. As early as October 1971, design guidelines were developed by Bechtel that outlined methodologies for the engineering of raceway supports. In March 1974, Bechtel issued a design standard by which most seismic raceway supports have been designed. This standard closely followed the guidelines set forth in USNRC regulatory guides and standard -eview plans, which were also being developed during the same period of time. Designs based upon these criteria have tended to require substantial amounts of bracing. By contrast, the Sylmar Station support systems were essentially unbraced.

Consequently, it appeared that either the design methods or the design criteria, or possibly both, were unnecessarily conservative.

In 1976, a plan was initiated to test electrical raceway systems. The goal of the testing was to establish the best possible approach to create an economical, yet adequate, support system for electrical cabling within nuclear plants. By the first part of 1977, a clearly defined program that outlined the types and sizes of raceway systems that would be tested was established. In the last months of 1977, testing was begun by Anco Engineers, Incorporated. Full scale installations of both cable tray and conduit raceway systems were tested. By the end of 1978, over 2000 individual dynamic tests had been performed, generating over 50 volumes of raw data.

The details of each phase of each task will not be explained in this report. This information is detailed at length in Reference 2. Instead, the overall philosophy is discussed. In addition, some specific examples are included.

In general, testing was performed by starting with the simplest test setup as possible. Initally, this involved testing of cable tray or conduit on rigid supports independent of their trapeze type hangers. These tests provided information that was useful in the next set of tests in which the cable tray or conduit was mounted on trapeze hangers that were totally unbraced.

This step allowed the collection of meaningful data related to a flexible hanger system.

Next, bracing was added to the hangers to restrict certain modes of vibration and attempt to begin the simulation of an in-situ seismic restraint.

Again the data developed in previous tests was valuable in understanding the behavior of the more complex system.

To augment the dynamic testing performed on the shake table, several static and quasi-dynamic tests were performed.

The static tests were performed on cable trays.

There were five types of cable tray used in the test program.

It was the goal of this test sequence to develop a better understanding of tray section properties and to establish an upper limit as to static load carrying capacities of the trays. The quasi-dynamic tests were performed to establish low-cycle fatigue characteristics of standard strut type connections. The data.collected from both types of testing was used to establish the design criteria for the tray and connection components.

The trays and conduit were loaded with miscellaneous sizes of electrical cable. The design load for the 24-inch wide tray was 50 lb/ft. Tests were run with cable weight varying from 0 to 50 lb/ft. Conduit loadings were also varied depending upon the size of conduit. Most of the conduit testing was performed using maximum cable loads. Conduit not filled to the maximum was loaded only to 50% of the maximum weight. All cable used as fill was typical power plant material.

The earthquake time history used to formulate the majority of shake table input motions was the Bechtel Horizontal Synthatic Time History - H1.

This record was selected due to its conformance with USNRC Regulatory Guide 1.60. In addition, a select group of four historical earthquake records was used during a limited group of tests. However, the actual input motion to the shake table was not the input motion corresponding to any one of the records mentioned. Rather, a modification to each record was made to account for effects of building amplification for the purpose of creating a "worst case" shake table input motion.

The strut supported systems that were tested survived all testing without loss of function. The type of damage that was observed in a few cases consisted mostly of fracturing of strut type angle fittings. This damage was due to low cycle fatigue resulting from significant ductile-plastic deformation that occurs at connections during large amplitude loading. Of the four angle fittings that are used to attach the hanger to the overhead steel (i.e., two fittings per vertical element, two vertical elements per hanger), never did more than one fitting of the four fracture during any one specific large amplitude test. Most of the systems were tested at input levels corresponding to 1.0 to 3.0g's maximum acceleration. These input levels were demonstrated to be equivalent to ground motion levels of 0.25 to 0.75g free-field acceleration. Never in the course of some 2000 dynamic tests did a total structural collapse of a strut-supported raceway occur. Nor was there any loss of function in the electrical circuits that were monitored.

The test results indicate that for conduit and tray supports similar to those in SONGS 1, seismic acceleration levels about the same or higher than expected for SONGS 1 can be achieved without impairment of the structural integrity of the raceway support systems.

The test results on combined tray and conduit support systems similar to those at SONGS 1 indicate that no loss of circuit continuity is expected when these support systems undergo maximum displacements of about 3-4 inches. The effects of seismic induced displacements on the structural integrity of the raceway support systems were evaluated based on the results of these tests. For cases where raceway systems would be subjected to differential displacements greater than 3-4 inches, the elements of structures were modified to assure that the resulting differential displacements would be less than 3-4 inches.

-7 The seismic capacity of the raceway systems was also observed in more recent seismic testing (Reference 3).

Although the design input levels for these tests were less than the SONGS 1 input levels, Reference 3 concluded that the seismic capacity of the raceway systems is attributable mainly to the high level of damping and the high level of reserve capacity inherent in steel structures due to material ductility.

The results of some of the specific tests conducted in the Reference 2 test program which are applicable to the type cand configuration of raceway supports at San Onofre Unit 1 are discussed in the following sections.

Testing of Rigidly Supported Conduit. For these tests conduits were attached by clamps to a strut mounted rigidly to the vertical testing surface. The test setup simulated conduits attached directly to a structural wall (Fig. 3a). The testing was conducted to determine the ultimate capacity of the conduit clamps. The supports were spaced at eight foot intervals, typical at SONGS 1, with 3/4" and 2" diameter rigid steel conduits attached. The conduit clamps utilized in the test program are equivalent (similar in design characteristics) to those used at San Onofre Unit 1.

Several tests were run on each setup. These tests included both uniaxial and biaxial dynamic loading. A sinusolidal input motion was used. In the biaxial tests, the input motion in the vertical direction was one half the input motion in the horizontal direction which was directed parallel to the conduit axis. Vertical slippage of the clamps was considered to have occurred when a displacement of 0.1" or greater occurred at the clamp locations.

For the 3/4" diameter conduits, the fragility levels were in excess of the shake table capacity of 13 to 15 g's. For the 2" diameter conduits, Table 3.5 shows the maximum input acceleration levels obtained during testing.

These tests were conducted with sinusoidal input motions, and input frequency to conduit frequency ratios of 0.83 to 0.90. The average acceleration input achieved with no slippage in these tests (excluding the two nonrepresentative tests results shown in Table'3.5) was 11.40g's.

The 2" diameter rigid steel conduits were also tested with various clamp types for sixteen feet support spacing. For the B-2013 clamps (equivalent to the P1117 clamp used in SONGS 1), vertical slippage was observed at a 1.10g sinusoidal input at the conduit resonance frequency.

Rigid steel 4" diameter conduits were also tested with various clamp types for twenty feet support spacings. For the 8-2013 clamps, vertical slippage was observed at a 1.14g sinusoidal input at the conduit resonance frequency.

Testing of Flexible Supported Conduits. For these tests conduits were attached to a horizontal strut which was connected to two vertical struts which in turn was attached to the testing facility (Figure 3b). These tests simulated conduits supported by trapeze or cantilever hangers. The tests were conducted to determine the adequacy of the clamps attaching the conduit to its support.

-8 Rigid steel 2" diameter conduits with supports at 8 foot intervals were tested with the B-2013 clamp type. Table 3.6 gives the input acceleration levels obtained during the biaxial sinusoidal tests. For these cases, clamp slippage did not occur. The acceleration amplitudes listed represent the limits of the shake table, therefore, the fragility input levels for these specimens are in excess of the amplitude given.

Both 2" and 4" diameter rigid steel conduits on trapeze type supports were also tested for ten feet support spacing. Tests were conducted with and without lateral bracing at the middle support. The objective of the testing was to determine the dynamic characerisics of conduit runs supported by standard trapeze raceways. Peak accelerations as high as 3.4 g's were recorded at the hangers during the test. Slippage of the conduit hold down devices were not detected and there were no failures associated with the conduit or its coupling devices.

Testing of Combined Tray and Conduit Support System. This test was conducted to determine whether fracturing of conduit fitting, pullout of a conduit from a panel, or a large deformation of the cable tray side rails could induce an interruption of electrical signals or jeopardize the quality of interlocking materials. Both 2" and 4" diameter conduits were assembled as shown in Figure 4 and tested.

Table 3.7 shows the results of the testing. It should be noted that large horizontal and vertical displacements were achieved between the base attachment point and the cable tray attachment points. During the testing there was no loss of circuit continuity nor was there any change in the integrity of the insulation. No failure of the conduit fittings or electrical boxes occurred other than the loosening of conduit fittings.

Minor loosening of the conduit to tray clamp occurred at approximately 80%

of the input acceleration values given in Table 3.7. Large distortions of the cable tray side rail occurred at maximum input, however, the tray was still able to adequately support the installed cables. After retightening the clamp, the input acceleration was advanced to the capacity of the shake table and no further evidence of clamp loosening was observed.

4.

CONCLUSION It is concluded that this implementation plan for return-to-service will provide for a seismic withstand capability of 0.67g without impairment of the overall integrity of electrical raceway support systems, and without impairment of the plant capability to achieve a hot standby condition (Mode 3).

This conclusion is based upon:

a.

Due consideration to the scope of raceway modifications to be installed for return-to-service as addressed in Section 2.

b.

Recognition of the conservatisms associated with the evaluation and final design processes which were utilized in the identification of raceway support modifications, as discussed in Section 3.

-9

c. The observed high seismic capacity of raceway support systems which are similar to the San Dnofre Unit 1 support systems, as discussed in Section 3.

d. Consideration that the tray and conduit raceway support systems are redundantly supported.

5. REFERENCES

1. Enclosure to letter from K. P. Baskin (SCE) to D. M. Crutchfield (NRC), dated August 17, 1982; enclosure entitled "Electrical Raceway Supports, Seismic Reevaluation Criteria" dated August 12, 1982.

2. "Cable tray and Conduit Raceway Seismic Test Program," 1053-21.1-4 Volumes I through IV, prepared for and in collaboration with Bechtel Power Corporaton, Los Angeles Power Division, Norwalk, California, 1978.

3. "Shaking Table Testing for Seismic Evaluation of Electrical Raceway Systems". The SEP Owners Group Under the Direction of KMC, Incorporated by URS/John A. Blume and Associates, Engineers, April 1983.

4. Enclosure to letter from K. P. Baskin (SCE) to D. M. Crutchfield (NRC), dated July 9, 1982.

5.

"Final Progress Report for the San Onofre Nuclear Generating Station Unit 1, Auxiliary Feedwater System Project" Draft for Comments, Seismic Safety Margins Research Program, June 18, 1982.

6.

"Soil Backfill Conditions, San Onofre Nuclear Generating Station Unit 1", August 12, 1982, and its revisions and addenda dated April 18, 1983, September 1, 1983, September 20, 1983, and November 28, 1983.

7. "Report on Cable Tray Support System Damping Values" Vogtle Electric Generating Plant, March 3, 1983. Enclosure to letter from D. Hutton to D. G. Eisenhut, dated March 5, 1982.

02490

-10 IABLE 2.1 N

SCOPE OF RACEWAY SUPPORT MODIFICATIONS FOR RETURM-TO-SERVICE (RTS)

Approx. Number Number of Support of Support Modifications Percent of Support Modifications Necessary to Meet Modifications in RTS Scope Evaluation Criteria in RTS Scope Cable tray support 459 617 modifications Cable tray tie-down 871 1,020 84%

modifications Conduit Support Modifications 1,316 1,596 83%

Replacement of masonry wall expansion anchors 231 231 100%

TABLE 2.2 LOCATIONS OF SUPPORT MODIFICATIONS TO BE INSTALLED PRIOR TO RTS APPROX. NLMBER OF AREA BUILDING/LOCATION MODIFICATIONS IN RTS SCOPE Tray Tray Conduit Tie-Downs 1

Containment-25 329 117 2

North Turbine Extension 43 49 60 3

480V Switchgear Room 53 121 109 5

East Turbine Extension 68 126 80 6

West Turbine Extension 60 106 86 7

South Turbine Extension 16 69 61 8

Auxiliary/Radwaste building 27 79 58 9

Intake Structure 19 92 10 Control Building 36 157 15 10 4160V Switchgear Room 87 158 4

11 Transformer Yard 14 12 Tank Area 4

1 4

14 Outside Area 36 22 42 16 Diesel Generator Building 1

43 3

17 Diesel Generator Building 3

..3 Totals 459.

1316 v 871

TABLE 2.3 NEW SUPPORT LOCATIONS APPROX. NUMBER OF NEW SUPPORTS TO BE INSTALLED NECESSARY TO MEET AREA BUILDING/LOCATION' PRIOR TO RTS RE-EVALUATION CRITERIA Tray Conduit Tray Conduit 1

Containment 1

30 1

45 2

North. Turbine Extension 13 9

25 12 3

480V Switchgear Room 6

19 9

19 S

East Turbine Extension 11 73 17 85 6

West Turbine Extension 3

4 3

7 7

South Turbine Extension 4

9 9

18 8

Auxiliary/Radwaste Bldg. --

17 2

21 9

Intake Structure 12 12 10 Control Building 3

16 3

25 10 4160V Switchgear Roon 3

67 9

82 11 Transformer Yard 12 Tank Area 14 Outside Area 5

6 6

7 16 Diesel Generator Bldg.

6 7

17 Diesel Generator Bldg.

9 99 Totals 49 273 84 349

TABLE 3.1

SUMMARY

OF INSTRUCTURE RESPONSE SPECTRA CONSERVATISMS Structures Conclusions Containment Sphere, and The spectra are based on soil-structure Reactor Building Inter-action damping values limited to 10%

in the horizontal directions and 17% in the vertical direction which is conservative, and the seismic input motion is applied at the foundation level without reduction in amplitude due to embedment effects. In addition, it has been shown that the spectra currently used in SONGS 1 for evaluation and design, envelope by a considerable margin the spectra in Reference 5.

Diesel Generator Building, The spectra arebased on time-histories that and Sphere Enclosure Building envelope the San Onofre Unit 2 and 3 ground design spectra which is more conservative than the Housner ground design spectra (the applicable seismic input for the seismic re-evaluation program).

Control and Administration The spectra at each elevation is an envelope Building of the responses of different locations at that elevation and thus are a conservative representation of the response levels expected.

Circulating Water Intake The spectra used are conservatively taken as Structure, and the Reactor the ground design response spectra without Auxiliary Building any reduction in amplitude due to embedment effects and without change in frequency content.

Fuel Storage Building, The spectra are an envelope of responses due Ventilation Equipment Building, to soil stiffness parameters corresponding and the Turbine Building to 95 percent compacted native backfill conditions and backfill as characterized in Reference 6.

-13 TABLE 3.2 PERCENT CRITICAL DAMPING VALUES BASED ON CABLE TRAY TESTS

% Damping Direction of Mean minus one Type of Bracing Input Motion Mean Sigma 8', 16', 32' Transverse 35 28 Transverse and Vertical One or Two Longitudinal 28 21 Longitudinal and Vertical TABLE 3.3 DAMPING VALUES FOR SEISMIC RE-EVALUATION Uamping Item Percent of Critical Conduit Supports 7

Cable Tray Supports 15 Combined Conduit and Cable Tray Supports 15 (same support)

-14 TABLE 3.4 COMPARISON OF TEST AND EVALUATION RESULTS The single-tier trapeze tray support system shown in Figure 2 was used and consisted of five supports spaced eight feet apart. The depth for the upper anchors to the bottom of the tier was 4'6". Unistrut P1001 was used for the vertical and horizontal members. The cable loading was 50 lb/ft.

Test Result Test Table Peak Observation ZPA Recorded Case*

Response

Evaluation Result 1

(No transverse 3.2g 1.94g No damage Both configurations were bracing)

(at middle was observed evaluated using the re support) when tested evaluation criteria and with combined methodology in Reference 1.

horizontal and Calculations were performed vertical earth-for 1g horizontal and 1g quake input.

vertical seismic accelera tions. The equivalent static method with a 1.5 factor was used for the seismic load computations.

Neither case satisfied the re-evaluation criteria and it was determined that the necessary modifications 2

would include the addition (With a trans-2.8g 2.19g No damage of at least two transverse verse brace at (at end was observed braces.

the middle support) when tested support) with combined horizontal and vertical earth quake input.

• Cases 1 and 2 represented amplified input motions which corresponded to equivalent free-field ground accelerations of 0.54g and 0.60g. respectively.

-15 TABLE 3.5 MAXIMUM INPUTS FOR 2" CONDUIT, 8' SUPPORT SPACING BIAXIAL TESTING, RIGIDLY SUPPORTED Ratio of input frequency Maximum input achieved to conduit frequency with no slippage (gas) 0.90 6.36*

0.85 12.39 0.90 4.95*

0.90 8.13 0.90 12.03 0.83 12.03 0.83 13.09

• These test series were conducted with coarse step size increases in acceleration input. Therefore these values, while conservative, are not representative of the test fragility levels.

TABLE 3.6 MAXIMUM INPUTS FOR 2" CONDUIT, 8' SUPPORT SPACING BIAXIAL TESTING, FLEXIBLY SUPPORTED Ratio of input frequency Maximum input achieved to conduit frequency with no slippage (gas) 1.19 12.75*

1.00 10.97*

The input level was limited by the shake table capacity.

-16 TABLE 3.7 COMBINED TRAY AND CONDUIT TESTING RESULTS Acceleration at Relative Displacement Gross Displacement Conduit Attachment Between Conduit Between Conduit Point Attachment Point Attachment Point Direction lop Second and Support Point and Ground Tier Tier a) 2" Rigid Steel Conduit H

1.25 1.50 2.10" 4.22" V

0.50 2.53" b) 4" Rigid Steel Conduit H

1.11 1.44 1.75" 3.87" V

0.42 2.34" tMKnarr:0249D

48 44 00 oo 0

"J 10 ma0us 0

0

• 0 0

• 0 00.

• 2

- o R

-011

-2)

FU 4.pr 0

us-

-ASE C,/) /O %. EXCEEDAWCE CURVE~

(e) A LOWE'R 8O//HD DE/C/V C(/RVC FIGURE I TYPICAL DAMPING TRENDS FOR CABLE TRAY TESTS

CABLE Lal&#446 P/~W~?.

RA YCONFIGUiRA T/OM 7E57TE

F/GURE 34. RIGIDLY 5UPPORT-D CONDUITS P.5FL E X4/7y SUPPORTED CONDUITS Ie

Oto COAIPSITFRACIWAV o

+

AOA7O 7'///4.B A a 1; Ftl L

f5T/

A

ADl 2 NRC-SEP December 23, 1983 Director, Office of Nuclear Reactor Regulation Attention: 0. M. Crutchfield, Chief Operating Reactors Branch No. 5 Division of Licensing U. S. Nuclear Regulatory Commission Washington, D.C.

20555 Gentlemen:

Subject:

Docket No. 50-206 Return to Service Plan Seismic Reevaluation Program San Onofre Nuclear Generating Station Unit 1 Over the last several months, we have met with NRC management and NRC staff personnel to discuss a return to service plan for the seismic reevaluation program. As discussed in these meetings, the basic premise of this ptan is that all San Onofre Unit 1 structures and systems whose failure coul& cause an accident and/or whose function is required to get to and maintain a safe shutdown (hot standby, Mode 3) will be available following a 0.67g earthquake.

The details of the return to service plan are provided in the enclosed report. Based on the implementation of this plan, San Onofre Unit 1 can r4turn to power without undue risk to the health and safety of the public even tonsidering the possibility of a major earthquake at the plant site.

As discussed in our previous meetings, it is necessary that SCE obtain agreement-with the NRC staff on the criteria to be applied to the seismic reevaluation program for return to service by mid-January 1984. As such, it is imperative that any questions you may have regarding our intentions in the enclosed plan be resolved as soon as possible. If there is anything that we can do to facilitate the NRC staff review, we will be pleased to provide additional information and/or meet with you as appropriate.

If you have any questions or comments on the enclosed plan, please contact me as soon as possible.

Very truly yours, M. 0. Medford Manager, Nuclear Licensing JLR:0541F bcc: (See attached sheet)

SEISMIC REEVALUATION PROGRAM RETURN TO SERVICE PLAN SAN ONOFRE UNIT 1 December 23, 1983

SAN ONOFRE NUCLA GENERATING STATION UNIT I FUNCTIONALITY CRITERIA FOR PIPING SYSTEMS IN RESPONSE TO THE OBE EVENT Prepared for:

Southern California Edison Prepared by:

EDS Nuclear Inc.

December 1983 EDS Report No. 04-0310-0063 Revision 2

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page i TABLE OF CONTENTS Page

1.0 INTRODUCTION

1 2.0 DEVELOPMENT OF THE FUNCTIONALITY CRITERIA 2

3.0 ANALYSIS PROGRAM APPROACH 6

4.0 SELECTION OF REPRESENTATIVE PIPING SYSTEMS 7

5.0 PIPING SYSTEM FUNCTIONALITY ANALYSIS 8

5.1 Elastic Analyses 8

5.2 Time History Generation 10 5.3 Nonlinear Piping Component Correlation 11 5.4 Nonlinear Analyses 12 5.4.1 AC-19 Nonlinear Analysis 13 5.4.2 MW-01 Nonlinear Analysis 15

6.0 CONCLUSION

S 17

7.0 REFERENCES

18

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page ii TABLES 5.1 AC-19 System Frequencies 5.2 MW-01 System Frequencies 5.3 AC-19 Elastic Stress Results 5.4 MW-01 Elastic Stress Results 5.5 AC-19 Linear vs. Nonlinear Analysis Results - Moments 5.6 AC-19 Nonlinear Analysis Results - Strains 5.7 AC-19 Nonlinear Analysis Results - Collapse Moment Comparison 5.8 AC-19 Linear vs. Nonlinear Analysis Results - Support Loads 5.9 MW-01 Linear vs. Nonlinear Analysis Results - Moments 5.10 MW-01 Nonlinear Analysis Results - Strains 5.11 MW-01 Linear vs. Nonlinear Analysis Results - Support Loads 5.12 MW-01 Linear vs. Nonlinear Analysis Results - Accelerations

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page iii FIGURES 4.1 AC-19 Mathematical Model -

Linear Analysis 4.2 MW-01 Mathematical Model -

Linear Analysis 5.1 AC-19 Mathematical Model -

Nonlinear Analysis 5.2 MW-01 Mathematical Model - Nonlinear Analysis 5.3 AC-19 Design Time History 5.4 MW-01 Design Time History 5.5 AC-19 Enveloped Response Spectra 5.6 MW-01 Enveloped Response Spectra 5.7 6-Inch Schedule 40 Carbon Steel Elbow Moment-Strain Curve 5.8 6-Inch Schedule 40 Carbon Steel Elbow Moment-Strain Curve 5.9 5x4 Tee Moment-Deflection Curve 5.10 5x4 Tee Moment-Strain Curve 5.11 MW-01 Elbow 4 In-Plane Bending Moment Response History 5.12 MW-01 8-inch Elbow Collapse Moment

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 1

1.0 INTRODUCTION

Southern California Edison (SCE) has been participating in the Systematic Evaluation Program (SEP) for San Onofre Nuclear Generating Station, Unit 1 (SONGS-1) for several years.

As part of this program, SCE is demonstrating that all safety-related piping systems at SONGS-1 can withstand a design basis earthquake (DBE) at the site. The Nuclear Regulatory Commission (NRC) and SCE have agreed upon criteria for piping which, when satisfied, will assure the integrity and functionality of the piping systems during and after a DBE.

These criteria are based on the ASME Boiler and Pressure Vessel Code,Section III (referred to as the Code) requirements, which are widely accepted as appro priately conservative design bases.

However, it has been recently proposed that the Code requirements are overly conservative when applied to cases involving short-term dynamic loads, such as the DBE. Diverse nuclear industry groups, such as the ASME code committees, Electric Power Research Institute (EPRI),

the

NRC, and concerns in foreign countries (e.g.

Japan and West Germany),

have for the past few years been investigating this issue (References 1 through 8).

The final resolution of the issue and changes to the Code are still several years away, however the trend in the results is clear -

piping systems subjected to short-term dynamic loads can maintain integrity and functionality at stress levels well above Code allowables with out a decrease in safety margins.

At SONGS-1, SCE is presently upgrading all piping systems to meet the SEP criteria. Due to the large amount of construction work involved in completing the upgrade program, the schedule for return to power may be extended. To allow a more expeditious schedule, SCE may instead complete partial modifica tions of the piping systems.

The success of this approach depends upon revising the criteria for piping integrity and functionality to take credit for the observed higher margins against failure under short-term loads.

The new criteria, which we have called the functionality criteria, is shown to be applicable for SONGS-1

piping, and includes sufficient conservatism to cover any uncertainty in the seismic analysis procedures.

EDS has developed such a functionality criteria in response to SCE's intention to return to power prior to completion of the seismic upgrade program for piping.

Section 2 of this report describes the functionality criteria and its basis.

Justification of the criteria requires demonstration of applicability to SONGS-1 piping systems.

The analysis program used for this justification is outlined in Section 3.

The program involves selecting representative piping systems at SONGS-1 (Section 4) and demonstrating through nonlinear analysis that integrity and functionality are maintained at the criteria stress limits (Section 5).

The conclusions of the study are summarized in Section 6.

The functionality criteria is justified for SONGS-1 without regard for its intended application.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 2

However, the arguments in support of the criteria may differ if they are applied in low probability situations (e.g. occurrence of a DBE over a short time period) or to systems with less safety significance (e.g. accident mitiga tion systems as opposed to safe shutdown systems).

2.0 DEVELOPMENT OF THE FUNCTIONALITY CRITERIA The objective of the seismic upgrade program being performed at SONGS-1 is to demonstrate the ability of the unit to successfully shutdown following a 2/3 g level earthquake. The design requirements to assure safe shutdown are those requirements imposed by the NRC under the SEP.

This design basis originates from the ASME Code requirements for piping systems.

However, operability and functionality (i.e. the capability of a system to function immediately after an earthquake until safe shutdown is achieved) can often be established using less restrictive criteria.

.The criteria requires that a system have the capacity to function during and immediately after an earthquake. This level of system performance is consis tent with less restrictive load limits than those specified by the NRC for the faulted condition -- limits that allow permanent deformations of a finite nature. The bases of these limits are general functionality and plastic limit analysis considerations.

For piping systems, the criteria allows an increase in the primary stress allowable to twice yield for carbon steel components and 2.2 times yield for stainless steel components.

These piping stress allowables, which are compared to stresses calculated from a linear elastic analysis, reflect the added capacity of a piping system beyond Code limits when subjected to short-term seismic loading.

These allowables are justified through the discussion and analyses that follow.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 3 The SONGS-1 functionality criteria were developed to include significant iden tifiable conservatisms inherent in ASME Code piping analysis procedures.

These conservatisms include:

• Damping values range between 2 and 5 percent for DBE loading as specified by Reg. Guide 1.61 [9].

Observed damping values for piping systems at high stress levels are much higher due to effects such as gaps in supports and flexible boundary conditions.

Current Pressure Vessel Research Council (PVRC) task group activities are investigating redefining damping to higher, more reasonable values [10].

• Strain rate effects are neglected in the criteria. These effects can sig nificantly increase the yield stress in dynamic loading cases.

• Stress intensification and flexibility factors considered are extremely conservative as defined in the linear elastic analyses.

These factors result in greater susceptibility to yielding under smaller loads in compo nents such as elbows and tees; however, there is no consideration for load redistribution to other components following initial yield.

• Pressure effects which increase the ultimate load capacity of components are not taken credit for in the criteria, although pressure stresses are included in the evaluations.

Component thicknesses are normally greater than the nominal dimensions specified.

This increase in thickness can have a significant effect on component capacity.

Actual material strengths are generally at least 10 percent greater than Code specified minimums.

Current Code allowables for dynamic loading are also recognized as extremely conservative, especially for seismic motion.

For the elastically-calculated stress levels of 2.0 Sy for carbon steel -piping and 2.2 S for stainless steel piping, actual yielding of the piping systems are expe ted to be of a limited local nature.

This prediction is based on the characteristics of seismic motion as well as the nonlinear behavior of piping systems:

The energy in any seismic motion is finite.

As a piping system yields locally, much of the input energy is absorbed as strain energy, and the kinetic energy of the system is reduced.

Nonlinear damping effects significantly decrease the response of a system after some amount of yielding.

.* SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 4

• The inertial effects of a typical piping system limit the deformations and hence the extent of local yielding.

• The system redundancy allows yielding at multiple locations.

In this

manner, system collapse due to formation of a mechanism is highly unlikely, and loading will be redistributed to different components such that excessive yielding will not occur in any one component.

Additional qualitative insight into the dynamic behavior of piping can be obtained from operating plants which have experienced strong ground motions.

The El Centro Steam Plant [1l], Lawrence Livermore Laboratory, and the Hamaoka Units in Japan have all been subjected to earthquake motion without disruption of operation. SRV discharge piping systems in BWR plants have also been sub jected to dynamic loads without damage, where conventional analysis indicates dynamic stresses well above current Code allowables.

Recently, the PVRC Task Group on Dynamic Loading has undertaken a program to develop more rational criteria for the evaluation of piping systems under transient loading [12]. This group recognizes the conservatism in current ASME Code practice, and is sponsoring research into the behavior of typical piping systems under dynamic loads to direct Code considerations towards the actual response and failure modes of those systems.

Several experimental programs to investigate the yielding of piping systems have recently been completed or are currently underway.

ANCO Laboratories has performed two sets of dynamic tests on Kraftwerk Union piping systems in West Germany [1,2,3].

One set utilized nine typical small-bore piping configurations of up to 300 feet in length with a variety of components and standard piping support systems.

These systems were subjected to both low and high frequency loads of various amplitudes corresponding to seismic and aircraft impact loads, respectively.

The maximum low frequency excitations with a maximum peak acceleration of 12 g were applied for durations of ten seconds. The maximum high frequency excitations with a maximum peak accelera tion of 24 g over the 20 to 40 Hz frequency range were applied for durations of approximately one second. Peak acceleration response of 50 g, peak dis placements of 50 cm, and plastic strains in excess of 0.6 percent were reported. Linear elastic analysis predicted dynamic stresses over four times ASME Code allowables.

Even for these extreme loads, there was no observed failure due to plastic collapse,

leakage, or loss of pressure-retention capability. This program was presented to West German licensing agencies to justify existing installations without backfitting for dynamic loaas, and to provide licensing support for the elimination of primary stress requirements for these loads on small bore (less than 2-inch diameter) piping.

. SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 5 High-excitation testing to benchmark dynamic nonlinear analysis methods for piping [4] is currently being conducted for EPRI. One test has been completed on a 4-inch Schedule 40 ferritic steel piping system.

This system has a length of 20 feet and consists of two elbows and three runs of piping.

The system was designed to ASME Class 2 rules.

The system was pressurized to design allowables and subjected to various dynamic excitation levels corres ponding to seismic events.

The primary purpose of this initial test was to demonstrate the feasibility of dynamically exciting piping systems to levels far in excess of current Code allowables.

The maximum dynamic excitation level corresponded to seven to eleven times a typical SSE spectra for a plant in a low to moderate seismic region.

This excitation level results in stresses which exceed Level D Code allowable stress limits by a factor greater than three.

Permanent and visible deformations were observed, but there was no plastic collapse or loss of structural integrity in the pressur ized piping.

Input accelerations were greater than 14 g, and response accel erations were greater than 21 g in one elbow. Plastic strains greater than 1.5 times the yield strains were recorded.

A limited amount of dynamic component testing has also been conducted [5,6,7, 8]. Straight pipe test data on fixed and pin-ended spans were developed in a joint Lawrence Livermore National Laboratories/Sargent and Lundy study.

Strain levels with corresponding stresses up to 130% of yield were observed.

A Japanese experimental study tested carbon and stainless steel elbows and tees well into the plastic range with harmonic excitation.

No failure or structural instability was observed in any of these tests.

These dynamic tests on piping systems indicate that typical piping systems can withstand extreme seismic loading conditions without plastic collapse.

Therefore, it is justifiable to develop functionality criteria which allows reasonable deformation of a piping system but still ensures that a safe shut down can be achieved following a 0BE.

Justification of the functionality criteria for piping subjected to a DBE is therefore provided by:

Inherent conservatisms in standard piping system properties and design techni.ques.

Demonstrated functionality of typical nuclear plant piping systems sub jected to seismic events and high-excitation dynamic testing.

Extremely low probability of occurrence of a DBE with the plant in the present design condition.

0)

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 6 Final justification of the functionality criteria limits for SONGS-1 is achieved through a nonlinear analysis program on representative systems from the plant. This provides program measurable evidence of the adequacy of the criteria.

The analysis program approach is outlined in the following section.

It is noted that a similar approach was used to successfully license the 2.0 Sy stress limit for seismic response of piping at Commonwealth Edison's Dresden and Quad Cities plants as part of their IE Bulletin 79-14 program [13].

3.0 ANALYSIS PROGRAM APPROACH The purpose of the nonlinear analysis program is to show that typical piping systems at SONGS-1 remain functional at elastically-calculated functionality criteria stress limits.

The load combination considered in the criteria is Gravity + Pressure + DBE.

.Thermal expansion was not considered as part of the criteria.

Two representative piping systems were selected for the functionality study.

Numerous piping systems were reviewed to choose these two systems.

It was desirable to choose systems typical of most of the piping at SONGS-1, and to provide a variety of material and component parameters.

Elastic analyses were then performed on the two systems.

Gravity, pressure, eigenvalue, and seismic analyses (both response spectrum and time history methods) were performed. These elastic analyses provided the following information:

Gravity, eigenvalue, and time history analysis results were used to provide correlation of results with the nonlinear analyses.

This insured proper development and accuracy of the nonlinear analysis models.

Gravity and pressure analyses were performed to assess the magnitude of those stresses compared to the total elastic stress levels required for the functionality study.

The gravity and pressure stresses were negligible and were excluded in the -nonlinear functionality analyses.

This assumption is discussed in detail in the analysis section.

Response spectrum analyses were performed to identify the critical direction of seismic input motion for the nonlinear analyses.

The results of the reponse spectrum analyses were also used to determine the scale factor on the input motion needed to produce maximum stresses at the required functionality limits.

Seismic time histories were developed for each piping system which enveloped the required SONGS-1 design response spectra.

The base motion used was 10

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 7 seconds of an El Centro 1940 acceleration time history record.

An iterative process was used to adjust the response across the frequency range of

interest, such that the final response spectra generated from the time histories closely matched the SONGS-1 design spectra.

Nonlinear analysis models were developed such that the components which comprise the model were correlated to experimental behavior.

Static loading analyses were performed on those components, and their material properties were adjusted such that their global response closely matched that of a similar experimental component.

This modeling technique provided increased accuracy in the piping responses predicted by the nonlinear analysis.

After extensive modeling checks were performed to verify the accuracy of the nonlinear analysis models, direct time integration analyses were performed with the scaled design time histories.

Response time histories of the critical components were obtained, and maximum moments and strains were reviewed to assure the functionality of the systems.

Finally, the results were used to make conclusions regarding the adequacy of the SONGS-1 functionality criteria for DBE loading.

4.0 CHOICE OF REPRESENTATIVE PIPING SYSTEMS Approximately twenty piping systems at SONGS-1 were selected for initial review. These twenty systems were those having stresses reported in excess of the SEP allowable stress level in the as-built configuration.

The April 1982 submittal by SCE to the NRC [14] was used to obtain the stress levels to select these twenty systems. Available support installation status informa tion was then reviewed to determine the number of supports requiring installation or modification for the final design and the number of those supports that had been installed at the time.

The selection of systems was based on the following considerations:

Location and magnitude of overstress Support requirements for final design condition System function

.Material properties System geometry, variety of components Based upon the review, two piping systems were chosen for the functionality study. These systems are designated problem numbers AC-19 and MW-01.

AC-19 consists of 2-1/2-inch and 1-inch lines which carry water to cool the primary shield wall.

All AC-19 piping is carbon steel A-53 Type B.

MW-01 consists

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 8 of 8-inch and 6-inch lines anchored at recirculation heat exchangers.

All MW-01 piping is A312 Type 304 stainless steel.

Figures 4.1 and 4.2 show the portions of AC-19 and MW-01 piping included in the analyses.

Problem AC-19 was selected because it requires the completion of much of the support work to meet SEP criteria.

The final design condition requires the addition of three lateral supports and one vertical support. The piping is small diameter carbon steel piping with relatively low diameter-to-thickness ratios.

Therefore, the components are fairly stiff, with low stress intensification factors. The geometry of the system is complex, with a great variety of different components such as 2-1/2-inch long-radius elbows, 5-0 bends, and 2-1/2x1 tees. A relatively even distribution of high stresses was anticipated for these lines.

Problem MW-01 was chosen to complement the system parameters investigated in the AC-19 analysis.

MW-01 piping is made of stainless steel with a high diameter-to-thickness ratio.

The 8-inch piping is Schedule 105; therefore, the components are flexible and have high stress intensification factors.

The system also has a variety of different component types.

High stresses were expected at a few local areas.

The two problems selected from the variety of systems at SONGS-1 provide a good representation of the various piping component,

material, and system types present in the plant.

Both carbon and stainless steel materials are represented, as well as piping components of different size and flexibility.

The systems both have typical run configurations with a mix of various component types. Although the seismic stress levels in the systems were not at the functionality stress limits, the input motions were increased to obtain the desired maximum elastic stress.

5.0 PIPING SYSTEM FUNCTIONALITY ANALYSIS This section describes the analysis methods used to demonstrate component and system functionality at the maximum elastic stress limits specified by the functionality criteria.

Preliminary elastic analyses are first discussed.

Input time history generation is. described.

Nonlinear analysis methods, assumptions, and results are then presented.

Overall conclusions are discussed in Section 6.0.

5.1 Elastic Analysis Mathematical models of each piping system were first developed.

These models include standard ASME flexibility factors and stress intensifi cation factors for the components. Material properties were obtained from the ASME Code Appendix I [151 for the design temperature of each system.

. SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 9 All elastic analyses were performed with the EDS computer program SUPERPIPE [16].

Gravity analyses for each system were first performed.

Low stresses were observed, as these systems are well supported in the vertical direction.

The functionality study considers seismic loading in one horizontal direction, and although the load combination Gravity +

Pressure + Seismic is addressed by the study, gravity stresses were omitted for the following reasons:

• Gravity stresses represented only 5% of the total stress in critical components when the maximum system stresses were increased to 2.0 Sy for carbon steel or 2.2 Sy for stainless steel.

• Since the earthquake load is scaled such that the maximum stresses equal the functionality stress limit, omitting gravity stress causes this scale factor to be greater.

During horizontally-applied seismic motion, piping components are stressed in different locations around the pipe circumference than when gravity loading is applied.

By omitting the gravity loading, the effects of the seismic loading are maximized, producing conservative strain data in the nonlinear analysis.

Pressure stresses were also calculated and found to be insignificant.

Pressure loading was not included in the functionality analysis because low to moderate levels of pressurization have a beneficial effect on piping response, in that it stiffens the piping system and increases the bending resistance of the components.

By neglecting pressure effects, strains are slightly overpredicted.

Unpressurized piping is also more susceptible to ovalization of its cross-section and reduction of flow area.

Eigenvalue analyses were performed to determine the fundamental frequencies of system response. Tables 5.1 and 5.2 list the modes and frequencies below 33 Hz for AC-19 and MW-01.

Seismic response spectrum analyses were then performed using the SONGS-1 design spectra for each global axis direction to determine the critical direction of seismic input for the nonlinear analyses.

For problem AC-19, the global X-direction response spectrum analysis produced an even distribution of high stresses in many components.

Thus, when the seismic motion in the X-direction is scaled such that the maximum stress level is 2.0 S,

extensive yielding of the system should result. For problem MW-0IY results for the X-direction response spectrum analysis also predicted overstress in more than one location.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 10 Therefore, the seismic load was applied in the global X-direction for the nonlinear analysis of both lines.

Tables 5.3 and 5.4 summarize the stresses in critical components from the X-direction response spectrum analyses for problems AC-19 and MW-01, respectively. These tables also show the final factored stress levels for the systems.

These factored stresses and their related bending

moments, system support loads, and accelerations are used to compare results with the nonlinear analyses later in this section.

The X-direction seismic time history and response spectrum scale factors are 2.68 for AC-19 and 7.85 for MW-01.

These scale factors were determined such that stresses in the critical elbow elements were at the functionality limits.

Thus, stresses in a few other components exceeded the functionality limits.

From the results of the elastic analyses, it was observed that there are areas of low stress in the piping systems.

To minimize the cost of the nonlinear analyses, it was desirable to eliminate as many piping degrees of freedom as possible.

Runs of pipe were removed from both systems in areas remote from the critically stressed piping.

The removed piping was modeled in the reduced system by specifying lumped masses and stiffnesses at the cutoff points.

The reduced models are shown in Figures 5.1 and 5.2 with the node numbering scheme used in the nonlinear analysis.

All elastic analyses previously discussed were performed on the reduced models with excellent correlation.

Critical frequencies were maintained in the reduced models, which assured an accurate and cost-efficient model for use in the nonlinear analyses.

5.2 Time History Generation To perform nonlinear seismic analyses of the two piping systems, it was necessary to obtain input time histories to meet the SONGS-1 seismic design requirements.

An iterative process was used to adjust the response at different structural frequencies such that the SONGS-1 design response spectra were properly enveloped by the response spectrum generated from the time histories.

A ten-second record from the El Centro 1940 earthquake motion was-used as the base motion.

One time history was generated for each analysis problem.

The SONGS-1 design response spectra used to match the time history response were envelopes of the two horizontal design spectra (N-S and E-W) which were used for the SEP analyses of AC-19 and MW-01.

The Fourier components of the El Centro motion were scaled such that the final time history produced an acceleration response spectrum close to that used in the SONGS-1 design.

The EDS computer programs FREAK [17]

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 11 and RESPEC [18] were used in the iterative process to obtain the design time history.

When the time history-gene)ated response spectra and the design response spectra were matched, the resulting displacement time histories were baseline-corrected to remove the drift in the motion.

The final acceleration time histories are plotted in Figures 5.3 and 5.4. The resulting response spectra are compared to the original design spectra in Figures 5.5 and 5.6.

To verify the adequacy of the generated time histories, a linear time history analysis was performed on problem AC-19 and compared with the results of the response spectrum analysis.

The time history analysis produced stresses slightly higher than those calculated in the response spectrum analyses.

This step showed that the time history generated conservatively predicted the system response.

5.3 Nonlinear Piping Component Correlation The computer program ANSYS

[19] was used to perform the nonlinear analyses.

The models were composed of elastic and plastic straight pipe elements and plastic elbow elements.

To maintain functionality, the

elbow, tee, and straight pipe elements must not distort excessively during the OBE event.

To assure that the system models accurately predict the piping behavior in the field, the ANSYS elbow elements were correlated with measured response in experimental studies.

Also, since ANSYS does not have a specific tee or branch connection element, an equivalent component was developed by adjusting properties of the four straight pipe elements used to model the branch connection.

These tees were also correlated with experimental data.

In finite element analysis, certain geometric and material property rela tionships are idealized.

In the ANSYS analyses, only a bilinear stress-strain relationship can be used for the non-proportional loading encountered in seismic analysis. This bilinear relationship is adjusted so the behavior of the elbow and tee elements closely matches the exper imental results. In the element correlation task, it was found that it was-not possible to obtain a good match-for both the momentdeflection data and the moment-strain data for. a particular element.

In an elbow,.

this is attributed to additional ovalization modes not included in the ANSYS model.

However, by matching the moment-deflection curves closely, the proper global response is assured.

Additionally, by matching the moment-deflection curves, a conservative moment-strain relationship is produced.

Thus the ANSYS-calculated strains can be considered an upper bound response of the component under the seismic load.

Figures 5.7 and 5.8 show the moment-deflection and moment-strain curves for a carbon steel elbow.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 12 To develop the ANSYS elbow models, the ORNL/NUREG-24 elbow study [20]

was used.

This study loaded 6-inch (nom'inal) commercial carbon and stainless steel elbows to produce predominately plastic response.

One stress-strain curve for elbows was developed for each problem, since AC-19 is a carbon steel system and MW-01 is a stainless steel system.

After these stress-strain relationships were determined, they were used directly to establish the elbow materials in each problem.

The Karman flexibility factor was used to allow for changes in elbow size and cross-section.

To develop the ANSYS tee models, results of the study by Ellyin [21]

were used. This study loaded tees of various run and branch sizes with in-plane and out-of-plane couples.

Loading was applied to produce plastic distortion of the tees.

To model the tee with ANSYS elements, standard pipe components were used for the run pipe.

For AC-19 tees, the run pipe was predicted to remain elastic, and elastic pipe elements were used in the model.

For MW-01, plastic pipe elements were used for the run pipe.

Two plastic pipe elements were used for the branch pipe in each problem.

The first was a relatively stiff element extending from the run pipe axis to the surface of the run pipe.

The other element was relatively flexible.

Deflections at the notch of the tee and at a point farther up the branch pipe were matched with experimental curves.

Again, this produces an extremely conservative moment-strain relationship.

Figures 5.9 and 5.10 show the moment-deflection and moment-strain relationships for the correlated carbon steel tee.

5.4 Nonlinear Analysis The mathematical model for the nonlinear analysis was developed with elbow and tee components which closely match experimental behavior.

Other straight-pipe components were modeled using the standard ANSYS pipe elements with ASME Code material properties at the design temperature.

Damping for the DBE seismic event was taken to be 2 percent from the fundamental frequency of the system to 50 Hz.

Alpha-beta damping using the current stiffness matrix was used.

Although the nonlinear analysis model was developed to closely predict actual behavior of the piping systems, they still contained inherent conservatisms.

Actual material strengths are greater than Code-specified minimums.

Code-specified minimums were used in the analysis.

Component thicknesses are normally greater than nominal values.

This increases the strength and moment-carrying capacity of the components.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 13

• The boundary conditions specified for each problem are conservative.

• Actual system damping is higher than code-specified values.

Two percent damping is used.

• Strain rate effects which enhance yield strength are conservatively neglected.

• Pressure effects increase collapse moments of components.

These effects were conservatively neglected in the analysis.

The nonlinear analysis models created with ANSYS were verified by comparison with SUPERPIPE linear analyses.

The previously created elastic SUPERPIPE models were modified to have the same material properties, and flexibilities as the nonlinear model.

Gravity, eigenvalue, and seismic time history analyses were performed. Both non linear ANSYS models showed excellent correlation with the linear SUPER PIPE models. The linear time history analyses were used to predict the time that each system-would begin to yield and the time when system response would be maximized.

The predicted time of maximum response using the linear analyses gives an upper bound limit to time of significant response in the nonlinear model.

Because of yielding and increased daMping and energy absorption in the nonlinear systems, actual maximum response occurs earlier than in the elastic system. This was observed in the analyses of both AC-19 and MW-01.

Thus the nonlinear analyses were not carried out to the end of the seismic time history.

Instead, analyses were performed to a time just beyond the time of maximum reponse predicted by the elastic analyses.

5.4.1 AC-19 Nonlinear Analysis The linear time history analysis of AC-19 predicted that first yielding would occur at 2.0 seconds at Elbow 3 and that maximum response would occur at 5.5 seconds.

In the nonlinear analysis, first yield -occurred in Elbow 3 (Refer to Figure 5.1 for designation of components).at about 2.24 seconds, slightly later than predicted.

Soon after the elbow experiences yielding, the piping near the support at Node 16 yields, followed by Elbow 1.

These components accumulate strain until strong motion starts at about 5 seconds into the earthquake.

At this time, additional straight pipe segments yield (at Nodes 13, 14, etc.),

and the maximum response is reached at 5.0 seconds.

The analysis was run to 6.0 seconds, and it was seen that response was significantly decreased in that final second.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 14 In the nonlinear analysis, moments in the critically stressed components were significantly rcduced.

Table 5.5 compares bending moments from the linear and nonlinear analyses for the more highly stressed components.

Significant moments were reduced a minimum of 16%.

The reduction in moment was mainly due to the "detuning" of the system as it yielded.

The frequency of the system decreased as yielding occurred.

For the AC-19 system, this caused the response to move away from the spectral peak, thus a decrease in response for the entire system was expected.

The variation of moment reduction throughout the system is due to the redistribution of total load to the yielding components.

Strain data from the nonlinear analysis of AC-19 is reported in Table 5.6. Very low strains were calculated for the AC-19 piping system, with a maximum strain of 0.74 percent reported in Elbow 3.

Maximum strain for a straight pipe section was 0.41 percent at Node 16.

The response of the piping in the area of the tees produced displacement-induced loads on the tees.

The nonlinear analysis predicted artificially high moments in Tee 1 because of the stiff model used, which did not allow the required deflection of the piping.

Strain energy methods were used to predict a maximum moment of 1.13 k-in, which is in the elastic range of behavior.

Thus, the tees in AC-19 were not expected to yield under the applied loading conditions.

Functionality of the AC-19 piping system was assessed by comparing the maximum moments in each type of component (elbow, tee, and straight pipe) to ASME collapse moments and by comparing the calculated strains to measured strains in experimental studies.

Table 5.7 compares theoretical collapse moments with calculated moments.

All moments in the AC-19 system were below the collapse moments except the moments in Elbow 3 and the straight pipe adjacent to the support at Node 16.

Moments at these two locations exceeded the collapse moment by 6 to 7 percent; however, due to the conservatism of the collapse moment determination and the low strain levels in the piping system, these moments were considered acceptable.

Strains for AC-19 were compared to the strains reported in the ORNL/NUREG-24 study used for the elbow correlations.

For all carbon steel elbows tested in the study, the elbow strains calculated in the AC-19 analysis were in the range of measurement. Maximum ovality in the experiments was 6.5 percent.

This ovality corresponds to a flow area reduction of about 0.3 percent, which is insignificant.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 15 The piping of AC-19 was considered functional for the following reasons:

• The system and component models were conservatively developed as previously discussed, therefore response was overpredicted.

• The ASME collapse moments are extemely conservative.

They correspond to a ductility of 2.

Component test data show that piping is functional at moments in excess of the ASME collapse moment.

• The strains reported in the nonlinear analysis were conservative because of the material law used to match the global response.

These conservatively calculated strains were well within the allowable strains reported in experimental studies and resulted in an insignificant flow area reduction.

The impact on support loads was also investigated.

Table 5.8 compares support loads for the elastic and nonlinear analyses.

Because of the frequency shift previously discussed, the loads on all supports were reduced in the nonlinear analysis.

5.4.2 MW-01 Nonlinear Analysis The linear time history analysis of MW-01 predicted that first yielding would occur at 0.6 seconds at Elbow 4 and in the straight pipe at Node 19 (Refer to Figure 5.2 for designation of components).

Maximum response was predicted to occur at 1.9 seconds.

In the nonlinear analysis, very slight yielding of Elbow 4 occurred at 0.075 seconds; however, significant yielding of the system did not begin until 0.55 seconds into the seismic motion.

At this time, Elbow 6 also yielded, followed by Elbows 2 and 5 in the next 0.2 seconds. The maximum moments were observed at 0.9 seconds, however maximum strains occurred at about 1.5 seconds. At this time, yielding was observed in Elbow 3 and the straight pipe at Node 19.

The analysis was run to 4.0 seconds.

Results showed-no significant response after the maximum response at 1.5 seconds.

Tee 1, which was very highly stressed in the linear response spectrum analysis due a high stress intensifi cation factor, did not yield. This was due to the inherent flexi bility of the branch connection which is not included in the ASME Code provisions for component modeling.

In the nonlinear analysis, critical moments in the highly stressed components were significantly reduced.

Other moments of smaller magnitude increased due to load redistribution following yielding.

Unlike AC-19, which has a very low fundamental frequency, MW-01

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 16 became "tuned" as the system yielded.

The response moved closer to the spectral peak as yielding occurred. Table 5.9 compares the bending moments from the linear and non-linear analyses for the critical components.

Figure 5.11 compares the linear and nonlinear moment response at Elbow 4 for the 4 seconds of applied seismic motion.

This figure shows the reduction in moment and a slight frequency shift between analyses after the significant yielding occurs at 1.5 seconds.

Strain data from the nonlinear analysis is reported in Table 5.10.

The highest strain in an elbow was 2.0 percent at Elbow 4.

Maximum strain in a straight pipe was.07 percent at Node 19.

Functionality of the MW-01 piping system was assessed by comparing the maximum moments in each type of component experiencing plastic deformation (elbow and straight pipe) with ASME-defined and theoretical collapse moments and by comparing the calculated strains to strains reported in experimental studies. Figure 5.12 calculates the ASME collapse moment at twice the deflection at yield for the 8-inch elbow in the MW-01 system.

The maximum resultant bending moment in Elbow 4 was 87 percent of the ASME collapse moment.

The maximum resultant moment in the straight pipe at Node 19 was 75 percent of the theoretical collapse moment.

Strains for the elbows in the MW-01 system were compared with strains reported in the study of thin-walled elbows by Imazu, et al. [22].

In this study, for elbow strains of 2.0 percent, flow area reduction of 5 percent was

reported, which is not significant.

The piping of MW-01 was considered functional for the following reasons:

The system and component models were conservative as previously discussed, therefore response was overpredicted.

All calculated moments were well below the theoretical collapse moments, which allow a ductility limit of 2.

Strains in the critical elbow were conservatively calculated, yet resulted in a predicted flow are reduction of only 5 percent.

This flow area reduction was considered to be insignificant.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 17 The impact of nonlinear piping on support loads and accelerations was also investigated.

Table 5.11 compares the support loads calculated for the linear and nonlinear analyses.

Table 5.12 compares accelerations for selected nodes for the two analyses.

Because of the frequency shift of the system and nonlinear load redistibution previously discussed, some suport loads and accelerations increased while others decreased.

6.0 CONCLUSION

S The nonlinear analysis program supports the functionality criteria and shows that typical SONGS-1 piping systems remain functional at elastic stress levels of 2.0 Sy for carbon steel and 2.2 Sy for stainless steel.

Although the nonlinear analyses made conservative assumptions in modeling and load definition, moment and strain levels in both systems were within experimentally verified functionality limits.

Critical moments in both systems were significantly reduced in the nonlinear analyses. AC-19 was a "detuned" system, such that the response was reduced after yielding occurred.

For AC-19, all moments and support loads were reduced.

This suggests that for lightly supported, flexible systems, functionality criteria which allow component yielding provide a rational method of evaluation. MW-01 was a "tuned" system, such that the reponse was increased after yielding occurred.

Despite the increase in response, the yielding allowed a redistibution of load in the system such that functionality was maintained.

This shows that for both "detuned" and "tuned"

systems, redistibution of load following yielding provides the required load reduction to insure system functionality.

Therefore, the elastic piping stress limits of 2.0 Sy for carbon steel and 2.2 S for stainless steel specified in the SONGS-1 functionality criteria provi e assurance that the piping systems are capable of withstanding DBE loads without loss of function.

These criteria allow local yielding in components such that load redistribution reduces maximum moments and stresses, yet provides limits on the extent of yielding such that functionality of the system is maintained.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 18

7.0 REFERENCES

1. Sand,

Lochau, Schoor, and Haas, "Experimental Study of Dynamic Behavior of Piping Systems Under Maximum Load Conditions Analysis", Kraftwerk Union, Federal Republic of Germany, ASME 1982 Orlando Conference, 1982.

2.

Ibanez, P.,

Keowen, R. S.,

and Renty, P. E.,

"Experimental Study of Dynamic Behavior of Piping Systems Under Maximum Load Conditions - Testing",

ANCO Engineers, Culver City, California, ASME 1982 Orlando Conference, 1982.

3.

"Quick Look Report:

Dynamic Testing of a Pressurized Piping System Beyond the Elastic Limit",

Preliminary Report, ANCO Engineers, Culver

City, California, prepared for EPRI, October 1981.

4.

Shibata, H.,

et al., "Test and Evaluation about Damping Characteristics of Hanger Supports for Nuclear Power Plant Piping Systems (Seismic Damping Ratio Evaluation Program)", Paper K6/4, Sixth Structural Mechanics in Reactor Technology Conference, Paris, France, August 1981.

5.

Moscone, "Damping Values of Nuclear Power Plant Components",

Westinghouse Corporation, Nuclear Engineering and Design 26 (1974).

6.

Ishiki, Nishizawa, et. al.,

"Nonlinear Seismic Analysis and Test",

1979, U.S. - Japan Seminar on HTGR Safety Technology.

7. Teidoguchi, H.,

"Experimental Study on Limit Design for Nuclear Power Facilities During Earthquakes", 1975.

8. Campbell, R. D.,

Kennedy, R. P.,

and Thrasher, R. D., "Development of Dynamic Stress Criteria for Design of Nuclear Piping Systems, Structural Mechanics Associates, Inc., Report SMA 17401.01, November 1982.

9.

U.S.

Nuclear Regulatory Guide 1.61, "Damping Values for Seismic Design of Nuclear Power Plants," October 1973.

10.

PVRC Technical Comitee -on Piping Systems, Task Group on Damping Values, Meeting Minutes, September 27, 1982

11.

"Equipment Response at the El Centro Steam Plant During the October 15, 1979 Imperial Valley Earthquake," NUREG/CR-1665, October 1980.

12.

PVRC Task Force on Dynamic Stress Criteria letter of November 24, 1981 re:

Request for Proposal on Dynamic Stress Criteria.

13.

"Quad Cities Unit 1 Functionality Study of Piping Systems in Response to the SSE Event," EDS Report No. 01-0590-1135, Revision 0, December 1980.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 19

14.

"Balance of Plant Mechanical Equipment and Piping Seismic Reevaluation Program, San Onofre Nuclear Generating Station, :nit 1," April 1982

15. Appendix I, ASME Boiler and Pressure Vessel Code,Section III, Division 1, "Design Stress Intensity Values, Allowable Stresses, Material Properties and Design Fatigue Curves,"

American Society of Mechanical Engineers.

16. SUPERPIPE, V. 15c, 6/28/82, EDS Nuclear Inc., San Francisco, CA

17.

FREAK, V. 12/1/76, EDS Nuclear Inc., San Francisco, CA

18.

RESPEC, V. 10/06/75, EDS Nuclear Inc., San Francisco, CA

19.

ANSYS, Rev. 3 Update

67K, June 1, 1979, Swanson Analysis Systems Inc.,

Houston, PA

20. Greenstreet, W.L.,

"Experimental Study of Plastic Responses of Pipe Elbows, "ORNL/NUREG 24, February, 1978.

21.

Ellyin, F.,

"An Experimental Study of Elasto-Plastic Response of Branch Pipe Tee Connections Subjected to Internal Pressure, External Couples and Combined Loadings," Welding Research Council Bulletin 230, September 1977.

22.

Imazu, Sakakibara, Nagata, and Hashimoto, uPlastic Instability Test of Elbows Under In-Plane and Out-of-Plane Bending,"

Paper E 6/5, Sixth Structural Mechanics in Reactor Technology Conference, Paris, France, August 1981.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 20 TABLE 5.1 AC-19 SYSTEM FREQUENCIES Mode No.

Predominant Direction Frequency, Hz Period, Sec 1

X 2.0 0.513 2

X 3.7 0.267 3

Y 6.8 0.148 4

X 7.2 0.139 5

Y 9.0 0.111 6

Z 11.1 0.090 7

X 13.9 0.072 8

X 18.7 0.053 9

X 21.6 0.046 10 Y

22.9 0.044 11 X

25.8 0.039 12 Z

29.2 0.034 13 Y

30.5 0.033 14 Y

31.9 0.031

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 21 TABLE 5.2 MW-01 SYSTEM FREQUENLIES Mode No.

Predominant Direction Frequency, Hz Period, Sec 1

X 7.7

.129 2

Z 9.4

.106 3

Y 12.9

.078 4

X 14.7

.068 5

Z 16.5

..061 6

Y 23.8

.042 7

Z 25.5

.039 8

X 27.0

.037 9

Z 27.5

.036

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 22 TABLE 5.3 AC-19 ELASTIC STRESS RESULTS Elastic Stress(2),ksi Location(1)

Design Spectrum 2.68 x Design Spectrum Elbow 1 18.6 49.7 Elbow 2 14.8 39.5 Elbow 3, 25.9 69.4 (2.OS (3))

Pipe @ Node 11 15.2 40.7 Pipe @ Node 12 12.8 34.1 Pipe @ Node 13 20.0 53.5 Pipe @ Node 14 20.1 53.7 Pipe @ Node 16 28.9 77.4 (2.2 Sy (3))

Tee 1 32.3 86.6 (2.5 S (3))

Tee 2 16.3 43.5 NOTES:

(1) From model in Figure 5.1 (2) Elastic Stress = 0.75iM/Z, 0.75i > 1.0 where i =.ASME Class 2/3 Stress Intensification Factor M = Resultant-of two bending and torsional moments Z = pipe section modulus (3) For A-53 B Carbon Steel, Sy = 34.7 ksi at 110*F

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 23 TABLE 5.4 MW-01 ELASTIC STRESS RESULTS Elastic Stress(2),ksi Location(L)

Design Spectrum 7.85 x Design Spectrum Elbow 2 2.75 21.6 Elbow 3 2.94 23.1 Tee 1 11.57 90.8 (3.6 Sy (3))

Elbow 4 7.01 55.0 (2.2 Sy (3))

Elbow 5 4.05 31.8 Elbow 6 3.71 29.1 NOTES: (1) From model in Figure 5.2 (2) Elastic Stress = 0.75i M/Z, O.75i >

1.0 where i = ASME Class 2/3 Stress Intensification Factor M = Resultant of two bending and torsional moments.

Z = Pipe Section Modulus (3) For A312 TP304 stainless steel, Sy = 25.0 ksi at 200OF 01

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 24 Table 5-5 AC-19 LINEAR VS. NONLINEAR ANALYSIS RESULTS - MOMENTS Bending Moment, k-in Location (See Figure 5.1)

Linear Nonlinear Percent Change Elbow I In-Plane 37.9 18.5

-51 Elbow 2 In-Plane 33.8 25.8

-24 Elbow 3 In-Plane 69.4 53.2

-23 Pipe @ Node 14 About Vertical Axis 54.7 45.7

-16 Pipe @ Node 16 About Vertical Axis 82.4 53.9

-35 Tee 1 (1)

Out-of-Plane 12.0 7.0 (2)

-42 NOTES:

(1) These moments are reported at the centroid of the tee element.

Actual moments in the tee are somewhat lower.

(2) This moment is an upper-bound moment based on a stiff tee model.

Actual moment is lower.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 25 Table 5.6 AC-19 NONLINEAR ANALYSIS RESULTS -

STRAINS Linear Analysis Nonlinear Analysis Location (See Figure 5.1)

Stress, ksi Maximum-Strain,-Percent Elbow 1 @ Node 2 49.7 0.49 Elbow 2 @ Node 7 39.6 Remained Elastic Elbow 3 @ Node 8 69.4 (2.0 Sy (1))

0.74 Pipe @ Node 14 53.8 0.21 Pipe @ Node 16 77.5 (2.2 Sy (1))

0.41 Tee 1 86.6 (2.5 S

)

Remained Elastic(2)

Notes:

(1)

S =34.7 ksi (2) See text for discussion

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 26 Table 5.7 AC-19 NONLINEAR ANALYSIS RESULTS - COLLAPSE MOMENT COMPARISON Nonlinear Analysis Collapse Percent of Location (See Figure 5.1)

Resultant Moment (1), k-in Moment, k-in Collapse Moment Elbow 1 (Std. long radius) 25.5 26.4(2) 97 Elbow 3 (5-D bend) 53.6 50.4(2) 106 Pipe @ Node 16 53.9 50.4(2) 107 Tee 1 1.1(3) 2.8(4) 41 NOTES:

(1) Resultant moment is taken as the SRSS of the two maximum bending

)

moments.

(2) Theoretical collapse moment =

(D3 - D0)

(3) Based on strain energy considerations for actual tee behavior.

See text for discussion.

(4) ASME collapse moment at twice deflection at yield.

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 27 Table 5.8 AC-19 LINEAR VS. NONLINEAR ANALYSIS RESULTS - SUPPORT LOADS Support Load, k Node(l)

Direction( 1 )

Linear Nonlinear Percent Change 11 Y

5.98 4.82

-19 14 Y

0.50 0.46

-10 15 Y

0.28 0.25

-11 16 X

1.88 1.31

-30 Y

0.13 0.11

-15 20 Lateral 2.83 2.19

-23 Y

0.23 0.20

-13 28 Y

0.84 0.61

-27 Z

2.04 1.69

-17 Notes: (1) See Figure 5.1

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 28 Table 5-9 MW-01 LINEAR VS. NONLINEAR ANALYSIS RESULTS - MOMENTS Bending Moment, k-in Location (See Figure 5.2)

Linear Nonlinear Percent Change Elbow 3 In-Plane 57.7 69.1

+20 Out-of-Plane 4.6 7.0

+52 Elbow 4 In-Plane 118.3 96.2

-19

@ Node 14 Out-of-Plane 80.2 65.6

-18 Elbow 4 In-Plane 30.3 28.1

-10

@ Node 16 Out-of-Plane 6.3 5.7

-7 Pipe About Vertical Axis 306.9 198.1

-35

@ Node 19 About Horizontal 53.1 39.0

-27 Axis Tee 1 In-Plane 24.8 21.8

-12 Out-of-Plane 5.5 6.6

+20

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 29 Table 5.10 MW-01 NONLINEAR ANALYSIS RESULTS - STRAINS Linear Analysis Nonlinear Analysis Location (See Figure 5.2)

Stress, ksi Maximum Strain, Percent Elbow 3 @ Node 8 23.1 0.10 Elbow 4 @ Node 14 55.0 (2.2 Sy (1))

2.0 Elbow 4 @ Node 16 31.8 0.42 Pipe @ Node 19 38.9 0.07 Tee 1 90.8 (3.6 Sy (1))

Remained Elastic Notes:

(1) S = 25.0 ksi

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 30 Table 5.11 MW-01 LINEAR VS. NONLINEAR ANALYSIS RESULTS - SUPPORT LOADS Support Load, k Node(l)

Direction(1 )

Linear Nonlinear Percent Change 11 Y

2.00 3.37

+69 Z

0.89 0.99

+11 19 X

11.30 8.31

-26 Y

1.64 1.25

-24 20 x

4.64 2.59

-44 Z

1.50 1.16

-23 24 Y

0.41 0.91

+125 Z

2.09 2.47

+18 28 X

3.82 4.60

+20 Z

1.38 1.40

+1 Notes: (1) See Figure 5.2

SOUTHERN CALIFORNIA EDISON Report No. 04-0310-0063 Revision 2 Page 31 Table 5.12 MW-01 LINEAR VS. NONLINEAR ANALYSIS RESULTS - ACCELERATIONS Acceleration, g Node(l)

Direction(1)

Linear Nonlinear Percent Change 10 x

8.12 6.52

-20 Y

0.05 0.09

+67 Z

0.18 0.17

-1 12 X

8.12 6.51

-20 Y

0.31 0.21

-31 Z

0.96 0.53

-45 14 X

8.08 6.47

-20 Y

1.62 1.73

+7 Z

2.68 2.80

+5 21 X

5.21 5.71

+10 Y

0.55 0.79

+42 Z

1.74 1.72

-1 Notes: (1) See Figure 5.2

tv

.4.1 1

2 6*6

.10 Arr 00

-4 0,46.~

Ib C

/dl ~ tA~

FIGUR 4

A M

h i

M l

i r

y

'I t~J* 4 st So pt FIGURE ~

~

~

~

~

~

~

~

j 4.1 Ito aheaia odl LnarAayi

f goo fi*t i

• e'

• I'i Lostd m-li

~~K~::i4 4

FIGURE 4.2 MW-01 Mathematical Model

-Linear Analysis

X z

0

~~130 lali 9

1 a

FIGURE 5.1 AC-19 Mathematical Model - Nonlinear Analysis

So 3S 3

3 3O 3 1 104O l

0 1,0

>1 41 4

197 Ic FIGURE 5.2 MW-01 Mathematical Model

-Nonlinear Analysis

1.2 C)

H 0.0~

-1.2 0.0 2.0 4.0 6.0 8.0 10.0 TIME, SEC FIGURE 5.3 AC-19 Design Time History

H E-4 0.0

-1.

2 CC A

C.L..S&

.4I 1

I.

L JL1 L

-J LJ IC L

J A

0.0 2.0 4.0 6.0 8.0 10.0

TIME, SEC FIGURE 5.4 MW-01 Design Time History

6.0 5.0-t I

11 Envelope of SONGS-1 4.0 N-S & E-W Design Spectra Generated From Time 3.q History 2.0

______j_

1.0 I

I I

1.0 5.0

10.

50.

FREQUENCY, HZ FIGURE 5.5 AC-19 Enveloped Response Spectra

2.0 Envelope of SONGS-1 N-S & E-W Design Spectra Z 1.5 H

Generated From Time EA History 1.0

1.

O-

^

tI 0.5 I

I I

1.0 5.0

10.

50.

FREQUENCY, HZ FIGURE 5.6 MW-01 Envelope Response Spectra

350 ANSYS 3

EXPERIMENTAL 300' 250 LOAD APPLIED PERPENDICULAR TO PLANE OF ELBOW 200 6" SCH 40 L.R.

ELBOW CARBON STEEL 150 100 50 0.25 0.50 0.75 1.00 1.25 DEFLECTION, IN FIGURE 5.7 6-Inch Schedule 40 Carbon Steel Elbow Moment-Strain Curve

350 EXPERIMENTAL 300 250

.-00 ANSYS Z200.

150 LOAD APPLIED PERPENDICULAR TO PLANE OF ELBOW 6" SCH 40 L.R. ELBOW 100

/CARBON STEEL 0.1 0.2 0.3 0.4 0.5

STRAIN, PERCENT FIGURE 5.8 6-Inch Schedule 40 Carbon Steel Elbow Moment-Strain Curve

100 1'ANSYS EXPERIMENTAL 80 MOMENT 60 DEFLECTION 40 4x5 TEE CARBON STEEL 20

.02

.04

.06

.08

.10

.12 DEFLECTION, IN FIGURE 5.9 5x4 Tee Homent-Deflection Curve

100 EXPERIMENTAL 80 SANSYS 60 07, MOMENT 7

6STRAIN MEASUREMENT 20-

/4x5 TEE 7/

CARBON STEEL I

I I

I I

as 0.1 0.2 0.3 0.4 0.5 0.6

STRAIN, PERCENT FIGURE 5.10 5x4 Tee Moment-Strain Curve

100

,-----Linear 80 Nonlinear of it if I

60

20.

O~

z

!I i

It

- 20 40 It It Ii

-so

-80j

-1001

-120

.11 II I

.50 1.00 1.50 2.00 2.50 3.00 3.50 4.00 4.50 5.00 TIME, SEC FT(UTRE 5.11 MW-(01 Elbow 4 Tn-Plane Bending Moment Response History

COLLAPSE MOMENT MAXIMUM RESULTANT MOMENT 125 100I YIELD MOMENT 75 P

M 50 DEFLECTION,5 25 0.2 0.4 0.6 0.8 1.0 1.2 1.4 DEFLECTION, IN FIGURE 5.12 MW-01 8-Inch Elbow Collapse Moment