ML20204H123

Rev 1 to Design Rept C-PENG-DR-006, Addendum to Cenc Rept 1444 Analytical Rept for Waterford Unit 3 Piping

ML20204H123
Person / Time
Site:	Waterford
Issue date:	03/22/1999
From:	Doney R, Mendrala C, Nadgor B ABB COMBUSTION ENGINEERING NUCLEAR FUEL (FORMERLY
To:
Shared Package
ML20204H122	List: ML20204H118 ML20204H123 ML20204H140 ML20204H146 ML20204H171 ML20204H182 ML20204H214 ML20204H245 ML20204H251 ML20204H259
References
C-PENG-DR-006, C-PENG-DR-006-R01, C-PENG-DR-6, C-PENG-DR-6-R1, NUDOCS 9903290017
Download: ML20204H123 (220)

Text

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Report: 11 oanes l Appendices: 244 canes Other Attachments: none 1 l

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DESIGN REPORT NO. C-PENG-DR-006 Rev. 01 Addendum to CENC Report No.1444 l j Analytical Report for i Waterford Unit No. 3 Piping <

j ABB COMBUSTION ENGINEERING NUCLEAR POWER 1

L CO*iBUSTION ENGINEERING, INC l

WINDSOR, CONNECTICUT I i i 1

j This document is the property of Combustion Engineering, Inc. (CE) Windsor, Connecticut,

and is to be used only for the purposes of the agreement with

CE pursuant o which it is furnished.

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2 VERIFICATION STATUS: COMPLETE i

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Authored by: ~

Date: dla22/99

C.[. Mendrala, Cognizant Engmeer '

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) . Approved by: b. @F Date: O M /O89 B. Nadgor, Inde' pendent Reviewer i f Approved by:  % Date: 34-9 R.O. Doney, Manag#r j

) It is hereby certdied this report has been properly and

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completely reconciled with the requirements of the Certidad by: _ . OJkMrX ASME Boiler and Pressure Vessel Code,Section III, Registration No. /Mi0 0 1989 Edmuon, no AMmria Staas of e-- --

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Dess S/M,/6 -

j 9903290027 990323 PDR 9 AD d.~-...OCK05000382d PDR f l

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t-C-PENG-DR-006 Rev. 01 -

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Page 2 of 1I l A.BSTRACT l This design repon is an addendum to CENC Report No.1444, Reference 3.3. Three locations on

) . the Hot Leg are evaluated for the installation of a Mechanical Nozzle Seal Assembly (MNSA); one  !

RTD nozzle, one PDT nozzle, and one Sampling nozzle. The components of the MNSA replace the pressure boundary previously assumed to be the J-weld connecting the nozzle to the pipe. Four holes are required in the Hot Leg for installation of each MNSA assembly. The impact of these holes on the design basis of the RCS Hot Leg is also examined in this design report. Other than the evaluation of the three above locations on the Hot Leg, there is no imput on the existing analysis of the Hot Les piping due to the installation of the three MNSAs.

l l These components are analyzed in accordance with:

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1. ABB Specifications 09270-PE-140, Revision 7

2. C-NOME-SP-0067, Revision 01

3. - ASME Code Section III,1989 Edition, no Addenda All requirements are satisfied.

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1 C-PENG-DR-006 Rev. 01 l Page 3 of 11

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RECORD OF REVISION '

Rev Date Pages Involved Author Reviewer Approval 1

! 00 3/19/99 All- Original C. Mendrala B. Nadgor K. Haslinger j 01 3/22/99 p.1, 2, 3, 4,10, Cl, DI C. Mendrala B. Nadgor R.O. Doney l

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C-PENG-DR-006 Rev. 01 Page 4 of 11 TABLE OF CONTENTS SECTION PAGE ABSTRACT..............................................................................................................2 1.0 INTRO D UCTIO N ........ .... . ..................... ................. .. . .............................. 5 2.0 SIGNIFICANT RES ULTS ........................... ... ........... .. .. ............ ....... ....... 6 2.1 SAMPLmo No7mI MNSA... .... . ... . . . . . .... ... . . . . . . ...7 2.2 PDT No7mR MNSA ... . .. ... .... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ......8 2.3 ' RTD No7m E MNSA.. . . .. .. .. .. ... . . . . . . . . . . . . . . . . . . . . . . .9 2.4 HoTLEG PIPNG... . ...

...................... . . . . . . . . . . . . . . ... . .. 10 3.0 RE FE RE NCES ............. ........ .... ....... ............ ..

.. ... ... ....... ............... .... 1 1 LIST OF ATTACHMENTS A ASME Code Year Reconciliation B

Calculatbn C-PENG-CALC-018, Revision 00, " Analysis of Waterford Unit 3 Hot Leg Sampling MNSA" C

Calculation C-PENG-CALC-0' ), Revisions 00 and 01, " Analysis of Waterford 3 Hot Leg PDT MNSA" D

Calculation C-PENG-CALC-020, Revisions 00 and 01, " Analysis of Waterford Unit 3 Hot Leg RTD MNSA" E

Calculation A-WATERFD-9449-1213, Revision 01, " Evaluation of Attachment Locations for Mechanical Nozzle Seal Assembly on Waterford Unit 3 Partial Penetration Welded Nozzles in the Hot Leg Piping" l

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, 1 C-PENG-DR-006 Rev. 01 Page 5 of 1I t 1.0 INTRODUC110N l

This design report summarizes the analyses performed for the installation ofMechanical Nozzle Seal Assemblies (MNSA) on three Hot Leg nozzles. The MNSA is a mechanical device that acts as a complete replacement of the "J" weld between an Inconel 600 instrument nozzle and the Hot Leg pipe. Its function is to prevent leakage and restrain the nozzle from ejecting in the event of a through-wall crack or weld failure of a nozzle. The potential for these events exists due to Primary Water Stress Corrosion Cracking.

i The components of the MNSA are analyzed, as well as the impact of the installation on the Hot Leg piping, in accordance with the ASME Code, Reference 3.4 l-l l.

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7 C-PENG-DR-006 Rev. 01 Page 6 of 11 2.0 SIGNIFICANT RESULTS The only Weterford 3 piping design report affected by this addendum is CENC-1444, Reference 1. The list of design report addenda below remain applicable to the RCS Piping and are not modified by this addendum.

" Analytical Evaluation of FAR #9270-74 for Louisiana Power and Light - Waterford Unit No. 3 Piping", Report No. CENC-1460, Febmary 1981.

" Addendum to the Piping Analytical Report for Entergy Operations, Inc. Waterford Unit 3", Report No. C-MECH-DR-001, Revision 00, December 1993.

" Addendum to the Piping Analytical Report for Entergy Operations, Inc. Waterford Station Unit 3", Report No. C-MECH-DR-004, Revision 00, December 1993.

The results of this design report are in accordance with the ASME Boiler and Pressure Vessel Code,Section III allowables (Reference 3.4).

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i C-PENG-DR-006 Rev. 01 Page 6 of 11 2.0 SIGNIFICANT RESULTS i

l The only Waterford 3 piping design report affected by this addendum is CENC-1444, Reference 1. The list of design report addenda below remain applicable to the RCS P

! and are not modified by this addendur:u

" Analytical Evaluation ofFAR #9270-74 for Louisiana Pov.er and Light - Waterford Unit No. 3 Piping", Report No. CENC-1460, February 1981.

"' Addendum to the Piping Analytical Report for Entergy Operations, Inc. Waterford Unit 3", Report No. C-MECH-DR-001, Revision 00, December 1993.

" Addendum to the Piping Analytical Report for Entergy Operations, Inc. Waterford Station Unit 3", Report No. C-MECH-DR-004, Revision 00, December 1993.

The results of this design report are in accordance with the ASME Boiler and Pressure Vessel Code,Section III allowables (Reference 3.4).

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L C-PENG-DR-006 Rev. 01

! Page 7 of 11 2.1 Sampilng Nozzle MNSA L

j' Attachment B contains the detailed analysis of the Hot Leg Sampling Nozzle MNSA L

' components. The following table summarizes the results for the Hot Leg Sampling Nozzle  !

MNSA components. The existing design analysis for the Hot Leg Sampling Nozzle is not l affected by the addition of the MNSA. l l Component Stress Calculated Stress /

Allowable Stress /  !

Category Usage factor Usage Factor l

(stress in ksi) (stress in ksi)

Tie Rod Design 6.45 26.9 Average 32.90 53.8 Maximum 41.39 80.7 Thread Shear 8.85 16.14 Usage Factor 0.40 1.00 Hex Head Bolt Design 22.51 28.3

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Average 38.34 53.8 Maximum 44.36 80.7 Thread Shear 12.05 16.14 Usage Factor 0.50 1.00 Top Plate Shear 0.91 9.72 Bending 7.30 24.3 .

Compression Collar Shear 9.12 9.72 Bearing 15.09 17.9

_Uoper Flange Thread Shear 6.16 9.72 Shear 6.98 9.72 Clamp Bolt Design 4.17 26.9 Average 30.71 53.8 Maximum 37.96 80.7 Thread Shear 6.14 16.14 Clamp Half Thread Shear 4.54 9.72 Lengths of engagement used in analysis:

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Tie Rod - Upper Flange: 0 5it (0.317 in. minimum)

. Hex Head Bolt - Hot Leg Pipe: 0.5 in. (based upon bolt thread shear; 0.373 in.

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minimum) i e

j Hex Head Bolt - Clamp Half: 1.00 in. (0.38 in. minimum) i i

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s C-PENG-DR-006 Rev. 01 Page 8 of 11 2.2 PDTNozzle ntNSA +

Attachment C contains the detailed analysis of the Hot Leg PDT Nozzle MNSA components.

' The following table summarizes the results for the Hot Leg PDT Nozzle MNSA -

components The existing design analysis for the Hot Leg PDT nozzle is not affected by the addition of - )MNSA.

Component Stress Calculated Stress / Allowable Stress /

Category Usage factor Usage Factor  !

(stress in ksi) (stress in ksi) ,

Tie Rod Design 8.58 26.9  !

Average 30.0 53.8 Maximum 37.4 80.7 Thread Shear 8.1 16.14 Usage Factor 0.424 1.00 Hex Head Bolt Design 22.51 28.3 Average 35.9 53.8 Maximum 42.1 80.7 Thread Shear 11.0 16.14 j

Usage Factor 0.50 1.00 Top Plate Shear 2.64 9.72 Bending 22.4 24.3 i Compression Collar Shear 9.26 9.72 Bearing 15.1- 17.9 I Upper Flange Thread Shear 5.60 9.72 Shear 6.98 9.72 Lengths of engagement used in analysis:

Tie Rod - Upper Flange: 0,5 in, (0.29 in. minimum)

(. Hex Head Bolt - Hot Leg Pipe: 0.5 in. (based upon bolt thread shear; 0.34 in.

minimum) t i

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C-PENG-DR-006 Rev. 01 l f Page 9 of 1I 2.3 RTD Nozzle MNSA Attachment D contains the detailed analysis of the Hot Leg RTD Nozzle MNSA components. The following table summarizes the results for the Hot Leg RTD Nozzle MNSA components. The existing design analysis for the Hot Leg RTD nozzle is not affected by the addition of the MNSA.-

Component Stress CrJeulated Stress / Allowable Stress /

Category Usage factor Usage Factor (stress in ksi) (stress in ksi)

Tie Rod Design 6.45 26.9 Average 35.48 53.8 Maximum 39.33 80.7 l Thread Shear 9.55 16.14 i

Usage Factor 0.359 1.00 j Hex Head Bolt Demien 22.51 28.3 Average 43.6 53.8 Maximum 48.98 80.7 Thread Shear - 14.15 16.14 l Usage Factor 0.50 1.00 l Top Plate Shear 3.00 9.72 Bending 23.88 24.3 Compression Collar Shear 7.02 9.72 Bearing 15.08 17.9 -

Upper Flange Shear 4.16

' 9.72 Thread Shear 6.64 9.72 Lengths of engagement used in analysis:

Tie Rod - Upper Flange: 0,5 in. (0.34 in. minimum)

Hex Head Bolt - Hot Leg Pipe: 0.5 in. (based upon bolt thread shear; 0.44 in.

minimum) t b

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C-PENG-DR-006 Rev. 01 l Page 10 of 11  !

L I 2.4 IHot Leg Piping -  !

The following issues were addressed relative to the attachment of the MNSAs in the Hot Leg Pi ping.

1. Maximum allowable bolt load

2. Area ofReinforcement Requirements for Nozzle Penetration

3. Primary plus Secondary Stress Intensity Range

4. Fatigue Usage Factor The results shown below are taken from the detailed calculation provided in Attachment E.

l Other than as discussed below, there is no impact on the existing analysis of the Hot Leg .

piping due to the installation of the three MNSAs.

2.4.1' Maximum A' llowable Bolt Load l . The maximum permissible bolt load based on the strength of the Hot Leg piping is 6.27 kips.

1 2.4.2 Area of Reinforcement Requirements for Nozzle Penetration l

For the standard MNSA design with minimum pipe thickness of 3.75 in.

Area Available = 3.939 in' > Area Required = 3.458 in' for the geometry plane at 45' to longitudinal ,

Area Available = 4.242 in' > Area Required = 4.091 in' for the longitudinal axis geometry For the sampling nozzle MNSA design with actual pipe thickness of 3.85 in.

Area Available = 4.730 in* > Area Required = 4.262 in' for the geometry plane at 20' to longitudinal Area Available = 4.942 in' > Area Required = 4.091 in' for the longitudinal axis geometry 2.4.3 Primary plus Secondary Stress Intensity Range For minimum pipe thickness of 3.75 in. Simx = 21.02 ksi < Allowable 3Sm = 55.2 ksi.

This value is less than the 36.33 kni calculated in the previous analysis.

2.4A Natigue Usage Factor j: For minimum pipe thickness of 3.75 in.IUmx = 0.382 < Allowable = 1.0 l- The fatigue usage factor has increased from the value of 0.0123 in the previous

, analysis.

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C-PENG-DR-006 Rev. 01 i Page 11 of 11

3.0 REFERENCES

3.1 Specification No. 09270-PE-140, Revision 7, " Project Specification for Reactor Coolant Pipe and Fittings for Entergy Operations, Inc., Waterford Unit 3", December 1993.

3.2 Specification No. C-NOME-SP-0067, Revision 01, " Design Speciscation for Mechanical Nozzle Seal Assembly (MNSA) Waterford Unit 3", March 1999.

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Report CENC-1444, " Analytical Report for Waterford Unit No. 3 Piping", May 1981.

3.4 ASME Boiler and Pressure Vessel Code,Section III 1989 Edition, no Addenda.

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F-110 T-053 P-001 f%R 23

• 99 14:57 ABB March 23,1999 NOME-99-C-0134 Mr. Bruce N. Proctor Entergy Operations, Inc.

Waterford Steam Blectric Station Unit 3 F.O. Box B Klilona, LA 70066

SUBJECT:

Mechanical Nozzle Seal Assemblies (MNSA)

Entergy Contract No. NWC00385 Proprietary Information

Dear Mr. Pmctor:

The following documentation has been submitted to Entergy for the subject contract.

1. A WATERFD-9449-1213 Rev 01, Evaluation of Attachment Locations for Mechanical Nozzle Seal Assembly on Waterford 3 Partial Penetration Welded Noulesin the Hot Leg Piping.

2. C-PENG-CALC-021 Rev 00, Determination of Waterford 3 Hot Leg Seismic Response Spectra and Accelerations for Usein Analysis ofMNSAs

3. Design Report No. C-PENG-DR-006 Rev 01 Addendum to Stress Report No. CENC-1444 Analytical Report for Waterford Unit No. 3 Piping Item I is an Attachment to Item 3. Pages 2 through 25 ofitem 1 are marked "ABB -

CENP Proprietary". Picase dinegard this proprictary designation. ABB does not consider this proprietary.

Item 2 includes an Appendix A which is stamped " Proprietary Information". Please disregard this proprietary designation. ABB no longer considers this information proprietary.

If there are any questions on the above. please call me.

Sincerely, y chn T. Mc Project Mana'ggr xc: G. Bundick ABB Combustion Engineering Nuclear Operations

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Cohsuontrunccino n PO Bnn dW) Twwwa s' (800) GF413 ti M D14v WIf%f I p (tN) W Ppi Winda GT 000%350f) tww 'rM7 CUMBEN WSM l

O CALCULATION CHANGE COVER SHEET B13.05 (Original R-Type or R-Type from Attachment 7.7)

Calculation: EC-M94-012 1 2 Number Revision Change Number O Markup Calculation Change @ Finalized Format Calculation Change O Addendum Initiating Document- No.: ER-W3-99-0198-00 Rev.: 1 Calculation Perfo.rmed Under:

l l Waterford-3 Procedures l X l Supplier Approved Procedures Computer Software Used: Ansys 5.3 N/A Code Version Disk l I

Deceription of Change: This design report is an addendum to CENC report # 1444. Three locations on i the HOT leg are evaluated for the installation of a Mechanical Seal Assembly (MNSA), one on RTD Nozzle. one PDT nozzle, and one Sampling nozzle. The impact of these holes on the design basis of  !

I the RCS Hot leg is also examined in this Design report. Other than the evaluation of the three MNSA i locations on the Hot Leg, there is no impact on the existing analyses of the Hot leg piping due to the installation of the three MNSAs.

This Calculation Change is considered NOT to be installed / implemented until the engineer initials and dates below. If no installation / implementation is required then initial and date to indicate an as-built condition. This Calculation Change may be canceled by initiating and dating below:

Initial Date Completely Installed / Implemented:

Partially Installed / Implemented ~

l Change Canceled: i 1

(list loads partially installed below)

(use additional pages as needed)

I have reviewed this Calculation Change in conjunction with any previous changes (see second caution in Section 5.5.1.3.1). The impact of these changes on the subject calculation do, ot result in exceeding its design basis / design margin.

ABB/CE,3/22/99 Mm.4v Dips k Dasgupta/ 3 3/E9 k D J.P. Burke / 3/23/99 Prepared By/Date Verified / Reviewed By/Date Approved By/Cate NOECP-011 Rev. 5 Attachment 7.8 1-

L l C-PENG-DR-006 Rev. 01 l

Page Al of A9 ATTACHMENT A ASME Code Year Reconciliation i

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Page A2 of A9 Construction Code Reconcm-tion for Waterford Unit 3 M~h==ical Nas Seal As e-blies (MNSA)

' The purpose of this reconciliation is to demonstrate fulfillment of the requirements for use of a

-later edition of the Construction Code for a Mechanical Nozzle Seal Assembly to be used at Entergy Operations' Waterford Unit 3 (WSES3). This is intended to allow the use of the ABB Combustion Engineering Mechanical Nozzle Seal Assembly (MNSA), which is built to an Edition of the ASME Code which is later than the Edition of the Code of Constmction to which the Main l Loop Piping was designed and analyzed.

In accordance with Entergy Operations, Inc. Project Specification No. 9270-PE-140, Rev. 07 the Construction Code for the Waterford Unit 3 Reactor Coolant Pipe and Fittings is the ASME 1971 Edition through Winter 1971 Addenda (hereinafter referred to as the Constmetion Code). l The Construction Code associated with the Installation is the same as for Design and l Procurement. Additionally, the Waterford:3 FSAR also incorporated the material fracture toughness requirements from the 1972 Summer Addenda into the plant Design Requirements.

j The ASME Section XI program at WSES3 is in accordance with the 1992 Edition, with ponions l of the 1993 Addenda (hereinafter referred to as the Section XI Code). The Constmetion Code  !

used for the Mechanical Nozzle Seal Assembly project is the 1989 Edition, No Addenda, of the

)

ASME Code,Section III (hereinafter referred to as the Replacement Code).  !

The WSES3 Mechanical Nozzle Seal Assembly Project involves both Repair and Replacement activities in accordance with the Section XI Code. Subparagraph IWA-4170(b) states that  :

" Repairs and installation of replacement items shall be performed in accordance with the Owner's Design Specification and the original Construction Code of the component or system. Later Editions and Addenda of the Construction Code or of Section III, either in their entirety or portions thereof, and Code Cases may be used." The Replacement Code (E89) is therefore acceptable for the Repair activities, which includes Installation.

The Constmetion Code for Design and Procurement is the 1971 Edition through Winter 1971 Addenda of the ASME Code as noted above. The Section XI Code (Subparagraph IWA-4170 (c)) specifies that " Items to be used for replacement shall meet the following requirements, unless ,

the alternative ofIWA-4170 (d) is adopted:" Subparagraph IWA-4170(d) specifies:

" Alternatively, an item to be used for replacement may meet all or portions of the requirements oflater Editions and Addenda of the Construction Code or Section III when the Constmetion Code was not Section III, provided that the following requirements are met.

C-PENG.DR-006 Rev. 01 Page A3 of A9 (1) The requirements affecting the design, fabrication, and examination of the item to be used for replacement are reconciled with the Owner's Specification through the Stress Analysis Report, Design Report, or other suitable method that demonstrates the item is satisfactory for the specified design and operating conditions.

(2) Mechanicalinterfaces, fits and tolerances that provide satisfactory performance are compatible with the system and component requirements.

(3) Materials are compatible with the installation and system requirements."

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t C-PENG-DR 006 Rev, 01 Page A4 of A9 These three requirements are addressed individually in below:

Reanirement (1)

"The requirements affecting the design, fabrication, and examination of the item to be used for replacement are reconciled with the Owner's Specification through the Stress Analysis Report, Design Report, or other suitable method that demonstrates the item is satisfactory for the specified design and operating conditions."

Discussion i

The Construction Code for the Reactor Coolant Pipe and Fittings is the 1971 Edition,  ;

i through Winter 1971. Addenda, of the ASME Boiler and Pressure Vessel Code. ABB

. Combustion Engineering Nuclear Power, has prepared a design specification for the Mechanical Nozzle Seal Assembly C-NOME-SP-0067, Rev. 01 specifying the Construction Code for the Mechanical Nozzle Seal Assembly to be the 1989 Edition of i

the ASME B&PV Code with no Addenda. As noted above, this is referred to as the  !

Replacement Code throughout the remainder of this document.

The Design requirements in the Reactor Coolant Pipe and Fittings Specification (9270-PE-140, Rev. 07) for ASME Code Class I components, are per Article NB-3000 of the Construction Code. The fabrication and Installation requirements are per Article NB- 1 4000 of the Construction Code. The Examination requirements are per Article NB-5000 of the Construction Code. Similar Articles specify the Design, Fabrication and Installation and Examination requirements of the Replacement Code. The corresponding Articles for Design, Fabrication and Installation, and Examination requirements of the Replacement Code are Articles NB-3000, NB-4000, and NB-5000, respectively.

- A comparison of each of the requirements of Design, Fabrication and Examination (called Inspection in the original Construction Code) for the Construction Code (71EW71 A) and the Replacement Code is provided below:

Design l

The MNSA uses the General Design Criteria (NB-3100) and Design by Analysis (NB-3200) for its design rules. Thus, reconciliation of Suh-Articles 3300,3400 and 3500 is l unnecessary.

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C-PENG-DR-006 Rev. 01 Page A5 of A9 There have been two primary changes in NB-3110 between the Construction Code and the 'teplacement Code. In the Construction Code, Design Conditions (NB-3112) included  !

a definition of Pressure, Temperature, and MechaNeal Loads and Operating Conditions l (NB-3113) included a definition of Normal, Upset, Emergency, Faulted at i Test Conditions. In the Replacement Code, Paragraph NB-3112 has been renamed " Design Loadings" and still includes Pressure, Temperature, and Mechanical Loads but directs the l

user to NCA-2142 for how to establish those loads. NCA-2142 in the Replacement Code l remains consistent with the rules in NB-3112 of the Constmetion Code. Also, in the j Replacement Code, Paragraph NB-3113 has been renamed " Service Conditions" and directs the user to NCA-2142 for classification of Service Conditions. Because th i

has final authority over what an event is classified after discussions with the Owner / Licensee, it was determined that ASME was at best. duplicating and at worst conflicting with the Regulatory Authority and the Code was revised to simply include the I analyticallimits for each condition (which were also renamed A, B, C, D, and Test). The I classification of events within each category is now left to be agreed to by the NRC and the Owner / Licensee. The MNSA Design Specification classi6es the events in +he same fashion as the Piping Specification which was agree to by the NRC and LPL and is considered reconciled. NB-3300 has added nomenclature and acronyms over the years and has become more refined and prescriptive. However, this Sub-article is concerned with components under external pressure, spherical and cylindrical shells, etc. which do not effect the design of the MNSA.

Stresses in the various parts of the MNSA are determined using common strength of materials clastic methods. These resultant stresses are then categorized and combined as

! required in NB-3200 and compare to specified allowables. This method is consistent between the Constmetion Code and the Replacement Code. The bolting items are

! designed and analyzed to NB-3230. Except for nomenclature differences, the L

Construction Code and the Replacement Code are identical iri this area. As such, they require no reconciliation.

l Finally, a Design Report has been prepared and demonstrates that the modified design is

! satisfactory for use for the design and operatir.g conditions specified in C-NOME-SP-0067,Rev.01 Fabrication and Installation The preponderance of rules in NB-4000 for Fabrication and Installation are concerned i

with welding, weld preparation, repair by welding, etc. There is no welding involved in the fabrication or installation of the MNSA; therefore, the welding requirements need no reconciliation.

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l C-PENG-DR-006 Rev. 01 Page A6 of A9 The Subarticles related to the MNSA are limited to NB-4100 and NB-4700. Within these, only Sub-subarticles NB-4110, 4120, 4130 and 4710 and Paragraph NB-4211 actually apply.

i NB-4110 is simply the Introduction to the Article and is essential!y identical between p

Code Editions. NB-4130, " Elimination and Repair of Defects" was renamed " Repair of j

Material" in the Replacement Code but other than that is essent.ially the same as the 2

Construction Code. NB-4211 " Cutting" is identical in both Codes. NB-4710 is specific

• to the installation of bolts and studs and is identical between Code Editions. Thus, none j

of these Sub-subarticles or paragraphs requires reconciliation.

4 The applicable portions of NB-4120 are slightly different between Code Editions. The 1989 Code has been editorially changed to refer to " Certificate Owner (s)" versus

" Manufacturer (s)"in the 1971 Code. Also, references to NA-8000 have been changed to NCA-8000. These changes were made to comply with other Code changes and are

) considered editorial in nature requiring no reconciliation. The technical requirement in the i

Replacement Code which does not appear in the Construction Code is found in the addition of Subparagraph NB-4121.3 which requires that a new surface exam of the material be performed if the material is machined in the course of fabrication or i installation.

This is an improvement over the Construction Code as it requires an p

additional examination intended to detect subsurface indications that are brought to the surface when the previously examined surface is removed. This requirement is not found in the Construction Code and is more conservative than the Construction ABB Code.

CENP does perform this examination after machining and the results are forwarded in the ~

i data package for the MNSA. With this, the Replacement Code is considered reconciled with the Construction Code.

i Examin.uon i

It should be noted that the Examination requirements in NB-5000 of the Construction '

Code are for welds and welded fabrication. This does not apply to the MNSA. There are no welds or welded parts used in the assembly. Thus, no specific reconciliation between NB-5000 of the Construction Code and NB-5000 of the Replacement Code is necessary.

However, the machined surfaces of the parts undergo a surface (liquid penetrant) examination after machining as required by the Replacement Code (NB-4121.3). This requirement is also found in NB-2547(c) of the Construction Code and NB-2547(c) and (d) of the Replacement Code.

The Section XI Code has separate examination and testing requirements for the installed i

replacement (IWA-4820, IWA-5000, IWB-2200 and IWA-2000) to assure proper j installation and operation of the replacement item.

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Reauirement (2)

" Mechanical interfaces, fits, and tolerances that provide satisfacto y performance are compatible with the system and component requirements."

RipfussiOA The Mechanical Nozzle Seal Assembly serves to replace the pressure boundary provi by the nozzle to hot leg pipe J-weld. As such, the system and component requirements be performed are the same as the J-weld: (1) maintain pressure boundary integrity leakage), and (2) prevent nozzle ejection.

The relevant interfaces, fits, and tolerances to maintain pressure boundary integrity (

leakage) are associated with the seal between the Split Packing (Grafoil) of the As!

and the Mechanical Nozzle. The leaktight integrity of the seal is demonstrated in the Leak l

Test Report provided with the Assembly.

The relevant interfaces, fits, and tolerances to prevent nozzle ejection are associated with the structural integrity of the Assembly under all design conditions assuming the internal J weld has failed or the nozzle has failed outboard of the J-weld. The stmetural integrity l

the Assembly under the specified loading conditions is proven in the Mechanical Nozzle Seal Assembly Design Report and in the Leak Tests noted above. The tests and De '

Report provide assurance that the mechanical interfaces, fit, and tolerances that prov satisfactory performance are compatible with the system and component requirements.

l t

L i

i f

i

. .-. _ - . - - . . . . _ - . - - - - . . - _ . _ . - . - . - - . . . - . . . ~ . - -

C-PENG-DR-006 Rev. 01 Page A8 of A9 Requia :;;t(3)

" Materials are compatible with the installation and system requirements."

Discussion The materials used to fabricate the Replacement item appear in an ASME published and NRC endorsed Edition of the Code and, therefore, may be considered acceptable for use in the Owner's Specification.

The WSES3 Mechanical Nozzle Seal Assembly is fabricated from SA-479 Type 304 austenitic stainless steel, and SA-453 Grade 660 high alloy, high temperature bolting material. Both of these materials are similar in general corrosion resistance to the Hot Leg RTD Nozzle material of SB-166 (Inconel) and are compatible with me Hot Leg base material SA-516 Gr 70. Differences in coefficient of thermal expansion have been considered and specific installation gaps have been integrated into the design to maintain Assembly integrity from cold to hot conditions. The Construction Code has no material specification for SA-479 Type 304 austenitic stainless steel or SA-453 Grade 660 high alloy, high temperature bolting material, therefore no comparison can be made between the Construction Code and the Replacement Code. The MNSA materials are exempted from the fracture toughness requirements found in the Summer 1972 Addenda.

Material properties such as coefficient of thermal expansion, Yield Stress, and allowable design stresses for the MNSA material have been shown to be compatible with the installation and system requirements by acceptable results in the MNSA stress analysis and Design Report. Analysis had determined that the stress and fatigue usage factor in all of the parts of the MNSA are below ASME Code allowables, thus being compatible with system requirements.

Relative to examination of original bar stock, the rules of NB-2000 apply. In the Construction Code, NB-2540 stipulated the requirements for " Examinations and Repairs of Forgings and Bar". The same requirements are found in NB-2540 of the Replacement Code including the Acceptance Standards in NB-2545.3 and NB-2546.3. A slight t

difference between the two Codes is the adtiition of Subparagraph NB-2541(c) which specifically directs the user to NB-2580 for the requirements for bar material to be used L ~ for bolting. Although Paragraph NB-2580 appeared in the Constmetion Code, it was

! ' assumed that the user understood its use. Later Editions of the Code replaced that

! assumption with clear direction as to where to go for these requirements. Paragraph NB-

. 2580 is identical in both Codes and was required in the Construction Code and is, 4

therefore, considered reconciled.

p * , ,w- - - . , ,

C-PENG-DR-006 Rev. 01 Page A9 of A9 lt is therefore concluded that, with respect to Material, the Replacement Code is reconciled to the Construction Code and the Owner's Specification.

Conclusion Based on the arguments in the preceding paragraphs, it is concluded that:

(1) - The requirements affecting the design, fabrication, and examination of the item to be used for replacement ue reconciled with the Owner's Specification through the Stress Analysis Report, Design Report, or other suitable method that demonstrates the item is satisfactory for the specified design and operating conditions. and (2) Mechanical interfaces, fits, and tolerances that provide satisfactory performance are compatible with the system and component requirements, and (3) Meterials are compatible with the installation and system requirements.

Therefore, it is concluded that the requirements of the Construction Code have been satisfactorily reconciled with the requirement of the Replacement Code for installation of the ABB CENP MNSA at WSES3.

C-PENG-DR-006 Rev. 01 Page B1 of B65 ATTACHMENT B 1

C-PENG-CALC-018, Revision 00,

" Analysis of Waterford Unit 3 Hot Leg Sampling MNSA" (65 pages including cover)

E

Design A.1alysis Title Prge

Title:

Analysis of Waterford Unit a dot Leg Satupling MNSA Document Number: C-PENG-CALC-018 Revi. n Number: 00

1. Vedncanon Statri:

- @ Complete - 0 NotRequired 0 CompletcwithContingencies/ Assumptions

2. Approvalof CosapletcJ Analysis This Design Analysis is complete and venSed. Management authorizes the use ofits results and attests to the quah of the Caani=at Engmeer(s), Mentor and Independent Reviewer (s).

Printed Name Signature Date Cognizant Engineer (s) B. Nadgor 6 k/ M 4 Q3//f/93 1 4 Mentor @ None a

" ' ;: ' ^ Reviewer (s) j k S. T. Slowik

_ ]. 3/ 7 3

M -- --  : Appel KR Hadnpr g[hffgjfgp3//j/jj i

3.

Package Contents (this secnon may be completed aAer M==d- -- u ^ approval):

Total page count, ielt body, WW attachments, etc. 64 List associated CD-ROM disk Volume Numbers and path names: @ None CD-ROM Volume Numbers Path Names (to lomst directory which uniquely WJ- to this h-t)

Other attachments (specify): @ Nome

4. Distdbution QA(2) BevBoya l . K.H. Haslinger C.L. Mendrala S.T. Slowik i

ABB Combustion Engineering Ni clear Power r-

M IDIF C-PENG-CALC-018 Rev. 00 Page 2 of 54 RECORD OF REVISIONS Rev Date Pages Changed Prepared By Reviewed By Approved By 00 03/18/99 Origin B. Nadgor S. T. Slowik K.H. Haslinger

~

ABB Combustion Engineering Nuclear Power

A It11 M IDIF C-PENG-CALC-018, Rev. 00 Page 3 Of 54 TABLE OF CONTENTS Section Page NO.

1 INTRODUCTION .. . .. . . . .. . . . ..... 5 1.1 OBJECTWE .

.5 1.2 ASSESSMENT OF SIGNIMCANT DESIGN CHANGES .

.5 2 METHODOLOGY . . . . .. . . .. 6 2.1 GAP AT NORM AL OPERAUNO CONDmONS..

2.2 DETERMINATION OFIMPACT FORCE., 6 )

2.2.1 .8 NetEjection Force F, 2.2.2 .8 Deflection ofCompesents Due to Nonle Ejection .

2.2.3 Impact Force..

.9

.10 2.3 STRESS EVALUATION OF THE MNS A COMPONENTS..

2.3.1 Tie rods.. .10  ;

2.3.2 .10 Hex HeadBolts (into Hot Leg)..

2.3.3 Top Plate..

.10 2.3.4 Compression Collar.. .11 2.3.3 UpperFlange.. . .I1

.I2 2.3.6 Clamp.. \

.I2 3 BASIC DATA AND ASSUMPTIONS . . . 13 3.1 SELECTION OF DESIGN INIUTS..

3.1.1 .13 Design and Operating Pressures and Temperatures...

3.1.2 MNSA Materials.. .13 3.1.3 No=le and Hot LegMaterials..

.13 3.1. 4 MaterialProperties.. . 13 3.1.3 MNSA ComponentDimensions.. .14 j 3.1. 6 .. ) 6 Nonle andNonle Component Dimensions.. i

.17

\

3.2 ASSUMPT10NS.. ..

3.2.1 Loadmg Conditions..

.I8 i 3.2.2

.18 \

Consideration ofSeismic Loads..

3.2.3 Friction Force.. .]8 l

.I8 t 3.2.4 Sealingpresare. ..

\ 3.2.3 Preload...

.I9

! 3.2.6 Dimensions.. . .

.I9

.19 4 SIGNIFICANT RESULTS - ..

..................20 5 iNuvS1S . . . .

.. . . . . 2 i 5.1 MNSA DESCRITDON., .

. 21 5.2 CONSIDERATION OF IMPACT LOAD..

\

5.2.1

. . 23 Relative Displacements Due to ThermalExpansion..

5.22

. 23 Cold Gap Setting vs. CalculatedDisplacements.. .

. 26

5. 2.3 Determ; nation offmpact Force.. . . ,

3.2. 4 linpact Force.. . 27 l

.30 \

5.3 STRUCTURAL STRESS ANALYSIS OViHE MNSA COMPONENTS . . 31 5.3. ) Tie Rod..

. 31 5.3.2 Hex HeadBolt (into Hot Leg).. .

.36 i

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I A Ik k I M INIP C-PENG-CALC-018. Rev. 00 Page 4 of 54 5.3.3 Top Plate..

5.3. 4 . 43 Compression Collar..

. 44 3.3.5 Upper Flange.

3.3.6 . 44 Clamp..

.44 )

5.4 FATIGUE ANALYSIS..

. 48 3.4.1 Normal Operating Presswe Force.. l

.49 j

3. 4. 2 Tie Rods.. .

5. 4.3

. .49 )

Hex Head Bolt..

. 51 \

5.5 CONSIDERATION OF HYDROSTATIC TEST PRESSURE CONDmONS .. . 52 5.6 CONSIDERATION OF FAU1.TED CONDmONS ..

. 52 6 REFERENCES .. . _ .

.. 53 LIST OF FIGURES FIGURE DESCRIPTION PAGE 1 Hot Leg Sampling MNSA. ..

. . 22 LIST OF APPENDICES APPENDIX A: ASSEMBLY DRAWING . . . .. ..A-1 APPENDIX B:

CALCULATION OF THE TIE ROD AVERAGE TEMPERATURE .B-1 APPENDIX C: QUALITY ASSURANCE FORMS . . . . . . . . .C-1 s

l ABB Combustion Engineering Nuclear Power l

M IDIF C-PENG-CALC-018 Rev. 00 Page 5 of 54 1 INTRODUCTION 1.1 Objective The objective of this design report is to present the rewits of the evaluation of the Mechanical Nozzle Seal Assembly (MNSA) to be installed on the Hot Leg Sampling nozzle at the Waterford Unit 3. l i

The MNSA is a mechanical device that acts as a complete replacement of the "J" weld between an Inconel 600 instrument nozzle and the Hot Leg pipe. Its function is to prevent leakage and restrain the nozzle from ejecting in the event of a through-wall crack or weld failure of a nozzle. The potential for these events exists due to Primary Water Stress Corrosion Cracking.

1.2 Assessment ofSignificant Design Changes This report presents the detailed structural and thermal analyses required to substantiate the adequacy of the design of the Waterford Unit 3 Mechanical Not.zle Seal Assembly as a replacement of the nozzle "J" weld. This analytical work encompasses the requirements set forth in Reference 6.1 and is perfonned in accordance with the requirements of the ABB CENP Quality Procedures Manual QPM-101 (Reference 6.2). i i

Addenda to the original Piping Design Report (Reference 6.3) were reviewed and it was determined that their results have no impact on the current analy:is and also that the current analysis does not impact their results. These Addenda Reports include:

CENC-1460 (2/81)

C-hECH-DR-001, Rev. 00 (12/93)

C-hECH-DR-004, Rev. 00 (12/93) l ABB Combustion Engineering Nuclear Power

A ItIt M IDIF C-PENG-CALC-018, Rev. 00 Page 6 of 54 2 METHODOLOGY The objective of this calculation is to analyze the Mechanical Nozzle Seal Assembly (MNSA) to be installed on the Hot Leg Sampling nozzle at the Waterford Unit 3. The methodology used in this calculation is based on the method developed in Reference 6.19, for a similar design.

2.1 Gap atNormal Operating Conditions A cold gap between the clamp and the top plate has to be established to account for the relative thermal expansions of the components (see representative drawing below).

l

\

I  ! j g -__ op l

(

l

,- m

/ f  %

m' The magnitude of the impact of the nozzle against the top plate is dependent upon this gap - the larger the gap, the greater the work done by the internal pressure, the greater the deflection of the components, and the greater the load on the components.

In order to determine the load impacting the MNSA components, if the nozzle ejects froni the hot leg, the gap between the clamp and top plate with the components at normal operating temperatur shall be determined. The thermal expansion displacements ofeach of the relevant components sha i first be determined and then added to or subtracted from the cold gap setting to determine the final operating conditions gap.

The linear thermal expansion displacements of each of the relevant components is calculated usin the following equation from Reference 6.25 (p. 53):

S = ot L AT where:

ABB Combustion Engineering Nuclear Power

A ItIk 2"E EFlF C-PENG-CALC-018, Rev. 00 Page 7 of 54 6= the displacement (deformation) oia component caused by linear thermal expansion et = the coefficient oflinear thermal expansion L= the length of the component AT = the temperature change from a reference temperature of 70*F to the applic4ble operating temperature The relative displacement,6,is determined by adding or subtracting the displacement of each of the individual components, as follows:

6, = Soom + Son a +6 5 ,..n, + Sa.., - So. w - Sn.. *

(* Sn ,. in this analysis is taken to mean the combined thermal displacement of the lower and upper flanges)

For determining the maximum relative displacement, it is assumed that the temperature of the tie rod and flanges increases from a reference temperature of 70 F to the ambient temperature of 120 F, and that the nozzle and clamp reach the normal operating temperature of 611 F. Because of respective

{

component lengths and coefficients of thermal expansion, these conditions produce the maximum '

relative displacement (6,.) between the nozzle and the MNSA top plate, such that the overall nozzle / clamp displacement exceeds the displacement of the top plate by a maximum amount. If there was no cold gap, the clamp would (theoretically) extend beyond the inboard surface of the top plate i by a distance of 6,..

The value of 6,. sets an upper limit on the cold gap setting, though the extreme temperature differences evaluated above would not beseen during plant operations.

For determining the normal operations relative displacement, a more reasonable temperature distribution is selected, based on more realistic and appropriately conservative conditions. These conditions produce the normal operations relative displacement (6, ,) between the end of the clamp and the MNSA top plate after heatup of the plant. If there were no gap between the end of the nozzle and the top plate at cold conditions, the clamp would (theoretically) extend beyond the inboard surface of the top plate by a distance of 6, op..

A final gap for the normal operating coaditions will be determined by subtracting the value of normal operations relative displacement (Sr op.) from ' he maximum cold gap setting.

ABB Combustion Engineering Nuclear Power

A ItIt M I H ip C-PENG-CALC-018, Rev. 00 Page 8 of 54 2.2 Determination ofImpact Force At the moment at which an instantaneous break occurs, the internal fluid pressure in the hot leg will eject the nozzle outward, with the nozzle impacting the top plate. In order to determine the stress effects of this impact on the top plate, it will be necessary to first determine the net ejection force acting on the nozzle. Once this net force is known, a relation can be defined betv een the work performed by the ejection force and potential energy stored in the deflection of trae affected components (at the point of maximum deflection). Once the deflection of the components is known, the impact force can be calculated and then used to determine stress effects.

2,2.1 Net Ejection Force, F, r

The nozzle will be forced out of the hot leg by the internal fluid pressure; this outward motion will be opposed by the friction force which the Grafoil Seal exerts on the external surface of the nozzle. The net ejection force acting on the nozzle, F., is the difference between the " pressure force", F,, and th seal friction force, Fr:

F. = F, - Fr 2.2.1.1 Force Due to InternalPressure 1

1, Motion of the nozzle at the moment at which there is an instantaneous break is due to the created by internal pressure pushing against the entire cross section of the nozzle. This force, F,, is determined as follows:

F, = p A where:

p= design pressure A= pressure area 2.2.1.2 Friction Force i

The determination of the friction force (Fr) provided by the Grafoil seal is made based upon the t

coefficient of friction for the seal against the nozzle and the radial load provided by the seal against the nozzle (produced by the compression of the seal).

Fr = P p A ABB Combustion Engineering Nuclear Power

A ItIt M IDIN C-PENG-CALC-018, Rev. 00 Page 9 of 54 where:

P= radial sealload (pressure)

=

coefficient of friction A= i surface area of the seal in contact with the nozzle surface  !

2.2.2 Deflection of Components Due to Nonle Ejection  !

1 The total deflection of the impacted components due to the ejection of the nozzle will be determined based upon the conservative understanding that all of the work put into the system by the net ejection force is converted completely into the poteadal energy of the deflected components (i.e., l there are no losses). The base equation for evaluating the total deflection is derived from Equa (a) on page 471 of Reference 6.25 and is presented, as follows:

l F, s = 1 K, &

• where:

F. = the net ejection force s= total distance traveled by nozzle '

l Ax = total deflection of MNSA tie rods and top plate, and K y= the equivalent stiffness of MNSA tie rods and top plate The total distance travekd by the nozzle, s, is equal to the distance of the gap at-temperature

(" Gap") plus the total deflection of the tie rods and top plate (Ax), or s = Gap +Ax.

The base equation may be re-written rkilows:

F, (Gap + &) = 1 K,, & 2 2

1 o - K,, &' - F, & - F, Gap = 0 2

In order to determine Ax, the stiffnesses of the tie rods and of the top plate are first calculated.

since the tie rods and top plate stiffness act in series against the impact load, the equivalent stiffne of tie rod-top plate systen;is calculated based upon a series stiffness equation from Reference 6.23 (p. 702).

K,, =

+

Ka K,, n ABB Combustion Engineering Nuclear Power

l l A IB C-PENG CALC-018. Rev. 00 Page 10 of 54 i

i The base equation developed previously is used to determine Ax:

r l

j Ax = [F. (F.2 + 2L, F. Gap)"] / K.,

L 1 2.2.3 Impact Force l

The impact force, F p , on the top plate and tie rods is then calculated:

Q =K yAr l

l 2.3 Stress Evaluation ofthe MNSA Components The stresses and fatigue in the MNSA components are examined considering, pressure, thermal loads and the impact force defined above.

2.3.1 Tie rods The Design tie rod load stresses are considered in accordance with NB-3231 and Appendix E.

The maximum tie rod load is determined by comparing the load created by preload and thermal expansion with the impact load. Whichever is greater is used throughout the remaining calculation as the tie rod loading.

The average and maximum stress in the tie rod and shear stresses in the threads are then calculated and compared to the corresponding ASME Code allowables.

The fatigue analysis of the tie rod is then performed, considering loads which may exist after weld or nozzle failure has occurred. The calculated usage factor is compared to the ASME Code allowable of1.0.

1 2.3.2 tiex Head Bolts (into Hot Leg)

The Design bolt load stre'sses are considered in accordance with NB-3231 and Appendix E.

Maximum Bolt Load Due to the flexibility in the design of the flanged connection between the MNSA and the Hot I

Leg, the impact from ejection of the nozzle will increase the load on the bolts. The stiffness of the flange relative to the stiffness of the bolts will determine what percentage of the impact loa will be effectively transmitted to the bolts.

l l

The total load on the bolt can be expressed by the following equation derived from Reference 6.24 (p. 579):

ABB Combustion Engineering Nuclear Power l

F1IfIF C-PENG CALC-018, Rev. 00 Page11of54 1

< 8 y F = Preload +

y'd' F ,,  ;

h, a, + Kw  ;

in t! e above expression, Kn.... is considered to be the equivalent stiffness of the components which are put in compression due to the torquing / tightening of the hex head bolts; these components include the upper flange (top and bottom pieces which are considered to act in parallel with each other), and the compression collar, which acts in series with the upper flange.

(The consideration of relationships between the top and bottom pieces of the upper flange represents a condition between two extremes that could be assumed for the components. The first, is that they act together as one solid piece. This would result in an unrealistically high stiffness. The second, is that they are both simply supported rings which act in series. Since the top ring is supported across the entire bottom surface, this would result in an unrealistically low stiffness. The assumption that the two rings act in parallel provides a stiffness that is between these two extremes and is concluded to be reasonable).

It may be concluded (from the above equation) that the greater the stiffness of the bolts as compared to the stiffness of the flange components, the greater the increase in load on the bolts from the impact (i.e., as Kn..,/Kwi -+ 0, the multiplier for Fi,,,,a 41).

The stiffness of each component is considered in the detailed analysis to calculate the maximum hex head bolt load.

The average and maximum stress in the hex head bolt and shear stresses in the threads are then calculated and compared to the corresponding ASME Code allowables.

The fatigue analysis of the hex head bolt is then performed, considering loads which may exist after weld or nozzle failure has occurred. The calculated usage factor is compared to the ASME Code allowable of 1.0.

2.3.3 Top Plate The shear and bending stresses in the top plate are calculated due to the impact load and compared to the corresponding ASME Code allowables.

2.3.4 Compression Collar The shear stress and bearing stress due to the preload of the hex head bolts are calculated and l compared to tia corresponding ASME Code allowables.

l ABB Combustion Engineering Nuclear Power

k l MIfir C-PENG-CALC-018, Rev. 00 1

l Page 12 of 54 l 2.3.5 Upper Flange The shear stress due to the preload of the hex head bolts is calculated and compared to the corresponding ASME Code allowables.

l 2.3.6 Clamp The average and maximum str%s in the clamp hex head bolt, and the shear stresses in the threa to the fullload acting on a nozzle are then calculated and compared to the corresponding ASM '

Code allowables.

i l ABB Combustion Engineering Nuclear Power

A IkIt 7"EBRIP C-PENG-CALC-018, Rev. 00 Page 13 of 54 l 3 BASIC DATA AND ASSUMPTIONS 3.1 Selection ofDesign Inputs 3.1.1 Design and Operating Pressures and Temperatures i

The Mechanical Nozzle Seal Assembly is considered a pressure-retaining component. The D Pressure is 2500 psia and Design Temperature is 650 F. Operating pressure and temperature are 2250 psia and 611 F, respectively(Reference 6.20). Ambient design temperature is 120 F (Reference 6.7).

3.1.2 MNSA Materials l

MNSA materials are taken from Reference 6.8.

Item Material Compression Collar SA-479, Type 304 Lower Flange SA-479, Type 304 Upper Flange SA-479, Type 304 Top Plate SA-479, Type 304 l Clamp SA-479, Type 304 Hex Bolt SA-453, Grade 660 Hex Nut SA-453, Grade 660 Tie Rod SA-453, Grade 660 Hex Bolt - Clamp SA-453, Grade 660 3.1.3 Nozzle and Hot Leg Materials l

Hot Leg Sampling nozzle and fitting matedals are taken from References 6.4,6.5, 6.6.

hem Material Referenc.s Nozzle Neck SD-166 6.5, 6.6 Safe End SA-182, Type 316 6.5, 6.6 Pipe SA-376, Type 316 6.4 Elbow SA-182, Type 316 6.4 i

ABB Combustion Engineering Nuclear Power

A ItIt M IFIF C-PENG-CALC-018, Rev. 00 Pa.r,e 14 of 54 3.1.4 Material Properties Material properties used in this analysis include coefficients of thermal expansion (at), moduli of elasticity (E), design stress intensity values (S.) and Yield Strength y Values IS ). These proper presented below and are found in the Appendices of Reference 6.9.

3.1.4.1 Coefficient ofLinear ThermalErpansion, a The following table presents the temperature-dependent coefficients oflinear thermal expansion various materials:

temperature SB-166 SA-479 Type 304 316 SS SA-453, Grade 660

(*F) (Alloy 600) (304 SS) (Alloy 660) 100 6.90 8.55 8.54 8.24 200 7.20 8.79 8.76 8.39 300 7.40 9.00 8.97 8.54 400 7.57 9.19 9.21 8.69 500 7.70 9.37 9.42 8.82 600 7.82 9.53 9.60 8.94 611 7.83* 9.55* 9.62* 8.95*

650 7.88 9.61 9.69 9.00

• by interpolation All coefficients are Coefficient B values from Table I-5.0, where Coefficient B is the mean coefficient 4

of thermal expansion X 10 in.fm/F in going from 70*F to the indicated temperature.

3.1.4.2 Modulus ofElasticity, E The following table presents the temperature-dependent moduli of elasticity for SA-479 Type 30 and SA-453, Grade 660:

l 1

0 ABB Combustion Engineering Nuclear Power l

)

i ABB C-PENG-CALC-018 Rev. 00 Page 15 of 54 temperature E

( F) 70 28.3 200 27.6 300 27.0 400 26.5 500 25.8 600 25.3 650 25.0*

709 24.8

• by interpolation All moduli of elasticity values are from Table I-6.0, where E = value given X 10' psi.

3.1.4.3 Design StressIntensity Value, S,,, 1 The following table presents the temperature-dependent design stress intensity values for various materials:

temperature SA-479 304 SA-453, Grade 660

(*F) S. S.

100 20.0 28.3 200 20.0 27.6 300 20.0 27.3 400 18.7 27.2 500 17.5 27.1 600 16.4 27.0 650 16.2 26.9*

700 16.0 26.8

• by interpolation The design stress intensity values for SA-479 Type 304 are from Table I-1.2; and the design stress intensity values for SA-453, Grade 660 are from Table I-l.3. All S. values are given in ksi.

i l

l ABB Combustion Engineering Nuclear Power

ABB C-PENG-CALC-018. Rev. 00 Page 16 of 54 3.1.4.4 YieldStrength Value, Sy The following table presents the temperature-dependent yield strength values for SA-479 Type 304:

temperature SA-479 304

( F) Sy 100 30.0 200 25.0 300 22.5 400 20.7 500 19.4 600 18.2 650 17.9 700 17.7 The yield strength values for SA-479 Type 304 are from Table 1-2.2. All Sy values are given in ksi.

3.1.5 MNSA Component Dimensions The bolts and tie rods have the following dimensions (References 6.8, and 6.18):

Bolts [0.500-20 UNF-2A] Tie Rods [0.375-16 UNC-2A]

Basic major diameter 0.5000 in 0.3750 in 4

Basic minor diameter 0.4374 in 0.297 in Basic pitch diameter 0.4675 in 0.3344 in 2

Tensile stress area 0.1599 in 0.0775 in z

Kn max (max minor 0.457 in 0.321 in diam. ofinternal thread)

Es min (min pitch diam. 0.4619 in 0.3287 in of external thread)

En max (max pitch diam. 0.4731 in 0.3401 in ofinternal thread)

Ds min (min major diam. 0.4906 in 0.3643 in ofextemal thread)

ABB Combustion Engineering Nuclear Power

71IfIF C-PENG-CALC-018, Rev. 00 Page 17 of 54 3.1.6 Nozzle and Nozzle Component Dimensions Various component dimensions are indicated below.

Sampling Nozzle Ref.

Pressure Diameter 1.00 in* 6.13 Length of Safe End 4.00 in 6.5,6.6 Length ofNozzle Neck 5.00 in 6.5,6.6 Length of Pipe + Elbow 9.535 in 6.4,6.8 2

Pressure Area = (n r ) 0.785 in 2 ,

• the value is conservatively rounded from 0.999 inches.

l l

k ABB Combustion Engineering Nuclear Power

A ItIt M IfIF C-PENG-CALC-018, Rev. 00 i

Page 18 of 54 i i

3.2 Assumptions 3.2.1 Loading Conditions '

If no crack is present, it is assumed that, except for preload and thermal expansion, the MNSA components are not loaded during normal operating conditions. An impact load may be experienced if there is a complete and instantaneous failure in the J-weld or a 360* circumferential crack in the nozzle, such that the nozzle would be forced outward against the top plate, closing any gap b the two components. After this event occurs, a normal operating load, without impact, would ex with the internal pressure holding the nozzle up against the top plate; this load would be cycli l

from essentially zero at Cold Shutdown to a maximum at normal operating conditions.

For the purposes of this analysis, it is assumed that there is a complete aa.d instantaneous failure ofl the J-weld (or a 360' circumferential crack in the nozzle) such that the nozzle is ejected outward impacts against the top plate, which will also then load the tie rods and other components. Th i

impact of the nozzle against the top plate conservatively represents the maximum load that the '

restraining components would experience.

3.2.2 Consideration of Seismic Loads Because of the nature of the accelerations from seismic events, only the tie rods will be evaluated the stress effects of the seismic event. The remaining MNSA components will not be sign affected. Separate seismic tests on similar MNSA configuration were performed to demonstrate anI adequate seal performance (see Reference 6.16).

3.2.3 Friction Force The effects of any impact of the nozzle against the top plate are dependent upon cenain as regarding the determination of the ejection force acting on the nozzle.

In an " ideal" (and worst case) break scenario, the crack would be complete, instantaneous and oriented such that no base or weld metal could interfere with the motion of the nozzle. In this ca the only resistance offered to the nozzle motion would be provided by the attached piping a l Grafoil seal.

In reality, the crack characteristics necessary for the " ideal" scenario would not exist, and, instea there would be potentially significant resistance offered to the motion of the nozzle by the crack surfaces and by integral material, if motion would be allowed at all.

1 l

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i

1 A It R

1. '11919 C-PENG-CALC-018, Rev. 00 Page 19 of 54 l In this analysis, a scenario which is somewhere "between" the " ideal" scenario and the "real" scenario will be evaluated: it will be conservatively assumed that motion will be allowed but in the presenc an opposing force provided by the crack metal and by the Grafoil seal. This opposing force will be l

accounted for by applying a coefficient of friction for the Grafoil-to-nozzle contact, as described below:

l A coefficient of friction (p) of 0.30 for Grafoil-to-nozzle contact will be used to detemtine the force which opposes motion of the nozzle. This value of 0.30 used for the (kinetic) motion of the nozzle ejection is higher than the values provided by the Grafoil seal manufacturer in Reference 6.21, which  ;

lists (static) coefficients in the range of 0.05 to 0.20 (see Table III of Reference 6.21). However, the l

application of a friction force based upon the coefficient value of 0.3 will be maintained on the basis '

that the actual force which would tend to limit or prevent motion in the "real" scenario would be i higher.

3.2.4 Scaling pressure Compression of the Grafoil creates a radial pressure against the nozzle surface of at least 3,100 for a preload of 30 ft-lb, based upon Reference 6.10. (This value will be used to determine a friction force on the nozzle from the Grafoil seal.)

3.2.5 Preload Nominal values of tie rod / bolt preload are used in this analysis since maximum values of preload not significantly increase corresponding preload stresses. (A check of the results indicate that use of the maximum preload values will result in stresses which will remain below, or will be on the order i

of, their respective allowables. Therefore, use of the nominal values is acceptatle).

3.2.6 Dimensions Referenced overall length of the assembly (15.30 inches) from Reference 6.8.1 was used in this report. Nominal design dimensions of the parts were used for the calculations of the relative displacement and for the stress analysis, except when noted. The use of the as-measured dimensions from Reference 6.13 produces slightly different values but does not affect the results of the current analysis.

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M IDIF C-PENG-CALC-018, Rev. 00 l

Page 20 of 54 i 4 SIGNIFICANT RESULTS The results presented below were determined using the assumptions defined and justified in Section l

3.2. There are no additional contingencies or assumptions that are applicable to these results. l All stresses are satisfactory and meet the appropriate allowable limits set forth in Section III of the ASME BoUer and Pressure Vessel Code (Reference 6.9).

Component Stress Calculated Stress / Allowable Stress /

Category Usage factor Usage Factor (stress in ksi) (stress in ksi)

Tie Rod Design 6.45 26.9 Average 32.90 53.8 Maximum 41.39 80.7 Thread Shear 8.85 16.14 Usage Factor 0.4 1.00 Hex Head Bolt Design 22.51 28.3 Average 38.34 53.8 Maximum 44.36 80.7 Thread Shear 12.05 16.14 Usage Factor 0.5 1.00 Top Plate Shear 0.91 9.72 Bending 7.30 24.3 Compression Collar Shear 9.12 9.72 Bearing 15.09 17.9 Upper Flange Thread Shear 6.16 9.72 Shear 6.98 9.72 Clamp Bolt Design 4.17 26.9 Average 30.71 53.8 Maximum 37.96 80.7 Thread Shear 6.14 16.14 Clamp Half Thread Shear 4.54 9.72 Lengths of engagement used in analysis:

Tie Rod - Upper Flange: 0.5 in. (0.317 in. minimum) e Hex Head Bolt - Hot leg Pipe: 0.5 in. (based upon bolt thread shear; 0.373 in. minimum) e Hex Head Bolt - Clamp Half: 1.00 in. (0.38 in. minimum)

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7%IfIF C-PENG-CALC-018. Rev. 00 Page 21 of 54 5 ANALYSIS S.1 MNSA Description The MNSA is a mechanical device that acts as a complete replacement of the "J" weld between an Inconel 600 instrument nozzle and the hot leg pipe. It replaces the sealing function of the weld using a Grafoil seal compressed at the nozzle outside diameter to the outer hot leg surface. The MNSA also replaces the weld structurally by means of threaded fasteners engaged in tapped holes in the outer hot leg surface, and a restraining plate held in place by threaded tie rods. This feature prevents the nozzle from ejecting from the hot leg, should the "J" weld fail or the nozzle develop a circumferential crack.

Drawing (Reference 6.8.1) for the Hot Leg Sampling MNSA is presented in Figure 1.

l l

l I

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i Allin C-PENG-CALC-018. Rev 00 Page 22 of 54 FIGURE 1 Hot Leg Sampling MNSA 4 .

i e, n j ,. n e

A ,,

A i

( '

/ _ _ 1 if f

, hy)/ :n =

e

i e t

I i

MMI 8

l 8

i 41k .

l ), i

N 3 i 0 y '

., D

/

~

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A ItIt M IfIF C-PENG-CALC-018 Rev. 00 Page 23 of 54 5.2 Consideration ofImpact Load 5.2.1 Relative Displacements Due to Thermal Expansion According to Section 2.1, the determination of the maximum relative displacement is based  !

j assumption that the temperature of the tie rod and flanges increases from a reference te 70 F to the ambient temperature of 120 F, and that the nozzle and clamp reach t temperature of 61l'F. Because of respective component lengths and coefficients of thermal expansion, these conditions produce the maximum relative displacement (6,.mx) between the n and the MNSA top plate, such that the overall nozzle / clamp displacement exceeds thej of the top plate by a maximum amount. If there were no cold gap, the nozzle would (

extend beyond the inboard surface of the top plate by a distance of 6, .

Hot Lee Samoline MNSA Design Temocrature Relative Disolacement I i

component temperature l ct L AT 6 l 4

(*F) (10 inlinl*F) (in.)

(*F) (in.) I nozzle neck 611 7.83 1.11* 541 0.0047 safe end 611 9.62 4.00 541 0.0208 pipe / elbow 611 i 9.62 9.535 541 0.0496 clamp 611 {

9.55 0.655** 541 0.0034 tie rod (lower)* *

• 120 8.27 5.11 50 (-)0.0021 tie rod (upper)"* 120 8.27 ]

8.19 50 (-)0.0034 flange j 120 8.60 2.00 50

(-)0.0009 6, x = 0.0721 l

• nozzle neck length = 15.3"(overall length from the drawing from Ref 6.8.1) - 0.655" (cl

. 1

+ elbow) . 4.00" (safe end) = 1.11" i

• • clamp length = 3.41" - 2.22" - 1.070/2" = 0.655" (Reference 6.8.2)

• the tie rod length is divided into two lengths - one up to the weld between the safe end and )

another equal to the length of the pipe + elbow and the part of the clamp above the elbow. The o1 '

length is based upon the overall assembly length less the thickness of the lower flange and the tI the two upper flange pieces.

i i

The value of 6,.x sets an upper limit on the cold gap setting, though the design temperature differences evaluated above would not be seen during plant operations.

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A ItIt M IFIF C-PENG-CALC-018, Rev. 00 Page 24 of 54 For determining the normal operations relative displacement, a more reasonable temperature distribution is selected, based on more realistic and appropriately conservative conditions. In this case, the following temperatures are assumed:

component temperature (T) nozzle neck / safe end 611 pipe / elbow 300 clamp 300 tie rod (inside) 500 i

tie rod (outside) 300 i

flange 550 The temperatures used in the determination of the normal operations relative displacement a average component values and are based upon the following:

• )

Nozzle Neck / Safe End - 611*F: due to the nozzle location in the hot leg pipe, the av temperature of the nozzle (external to the pipe) would be at or very near the hot leg temperatureI (611T).

Pipe / Elbow- 3007: this component would receiu some heat conducted through the nozzl Some heat would be lost during conduction but, overall, the pipe / elbow average templ '

would be elevated to a relatively high temperature, given its direct attachment to these hotter components.

Upper Flange - 5507: this component would receive some heat conducted through the ;

retainer and lower flange, as well as heat by convection / radiation through gaps.

Some heat would be lost to the insulation amhent though, overall, the flange average temperature would b elevated to a relatively high tempe<ature, given its proximity to these hotter compone temperature for this component is conservative as it will increase the impact gap.

Tie rod - inside - 5007: thic component would receive some heat conducted through the cen1 i

portion of the upper flange, with some heat lost to insulation ambient. A high temperature for I this component is conservative as it will increase the impact gap.

Tie rod - outside - 2507: this component would receive some heat conducted from the po l of the tie rod inside of the insulation. The average temperature of 2507 is based upon a he transfer evaluation which conservatively assumes a 500T heat source temperature and heat lost l ABB Cornbustion Engineering Nuclear Power

A IkIt D'1lElE C-PENG CALC-018. Rev. 00 Page 25 of 54 l

l to ambient at 120*F (see Appendix B). A high temperature for this component is conservative as it willincrease the impact gap.

l e Clamp - 300*F: this component would receive heat conducted directly from the pipe / elbow and nozzle fluid.

Hot Lee Sampline MNSA NOP Relative Disolacement component temperature a L AT S 4

(*F) (10 in./in./*F) (in.) (*F) (in.)

nozzle neck 611 7.83 1.11 541 0.0047 safe end 611 9.62 4.00 541 0.0208 pipe / elbow 300 8.97 9.535 230 0.0197 clamp 300 9.00 0.655 180 0.0014 tie rod (lower) 500 8.82 5.11 430 (-)0.0194 tie rod (upper) 250 8.47 8.19 230 (-)0.0125 flange 550 9.45 2.00 480 (-)0.0091 l 6,. = 0.0056 These conditions produce the normal operations relative displacement (6,,) between the end of the clamp and the MNSA top plate after heatup of the plant.

However, in accordance with Reference 6.22, chemistry samples are taken each week after flushing the sampling line for one half-hour. Consequently, the sampling line is exposed during short period of time to the Hot Leg operating temperature. In order to accommodate such conditions, the extreme NOP relative displaccinent has to be calculated based on assigning the temperature of 611 F to the pipe / elbow and clamp.

l ABB Combustion Engineering Nuclear Power

i Ak IkIt M IDIF C-PENG-CALC-018, Rev. 00 Page 26 of 54 Hot Lee Samoline MNSA Extreme NOP Relative Disclacement component temperature a L i

AT S 4

( F) (10 inlin/F) (in.) (*F) (in.)

nozzle neck 611 7.83 1. I 1 541 0.0047 safe end 611 9.62 4.00 541 0.0208 pipe / elbow 611 9.62 9.535 541 0.0496 clamp 611 9.55 0.655 541 0.0034 tie rod (lower) 500 8.82 5.11 430 (-)0.0194 l tie rod (upper) 250 8.4 8.19 230 (-)0.0125 flange 550 9.45 2.00 480 (-)0.0091 6,_ = 0.0375 These conditions produce the extreme normal operations relative displacement (6, the

,,,) between end of the clamp and the MNSA top plate which govern the MNSA cold gap settmg .

5.2.2 Cold Gap Setting vs. Calculated Displacements in accordance with Reference 6.8.1 (note 1), the gap of 0.043 0.005 in. has to be set between the clamp and top plate at cold condition to accommodate the extreme NOP conditions, as described in i

Section 5.2.1. It is recognized that the minimum cold gap of 0.038 in. is less than the design I displacement for the Sampling MNSA (6,. = 0.0721 in), which would mean that, at the conditions for the maximum relative displacement, the nozzle would be in direct contact with the top pl However, as noted before, the conditions used to obtain the design relative displacement are not anticipated during operation.

The maximum cold gap setting of 0.048 in. indicates that a gap of 0.048 in. - 0.0056 in. = 0.0424 in.

can exist during normal operating conditions for the Hot Leg Sampling MNSA.

A final gap value of 0.043 in for normal operating conditions will be used in the subsequent determination of the impact force.

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A IkIt M IFIF C-PENG-CALC-018, Kev. 00 Page 27 of 54 s

5.2.3 Determination ofImpact Force impact force is calculated in accordance with the methodology described in Section 2.

5.2.3.1 Net Ejection Force. F.

The net ejection force acting on the nozzle, F., is the difference between the " pressure force", F,, and the seal friction force, Fr:

F. = F, - Fr 5.2.3.1.1 Force Due to Ivemal Design Pressure Motion.of the nozzle at the moment at which there is an instantaneous break is due to the forc created by internal pressure pushing against the entire cross section of the nozzle. (From Section 2

3.1.6, the pressure area is 0.785 in ). This force, F,, is determined as follovts:

F, = (2500 psi)(0.785 in') = 1,963 lb 5.2.3.1.2 Frie: ion Force Fr = P A where:

P= radial seal load (pressure) = 3100 psi (Reference 6.10) p= coefficient of friction = 0.3 A= surface area of the seal in contact with the nozzle surface

=

xDh h= 0.25 in (Reference 6.8)

D= 1.00 in Therefore:

Fr = (3100 psi) (0.3) x (1.00 in) (0.25 in) = 730 lb Based upon the forces crJeulated above,;he net ejection force is:

F. = 1963 - 730 = 1233 lb.

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A It11 M IDIF C-PENG-CALC-018, Rev. 00 Page 28 of 54 Use F. = 1250 lb 5.2.3.2 Deflection ofComponents Due to No::le Ejection The re-written base equation for evaluating the total deflection is 5K,, Ar* - F, At - F, Gap = 0 2

In order to determine Ax, the stiffnesses of the tie rods and of the top plate are first calcu!ated and then the equivalent stiffness of the tie rods-top plate system is calculated.

5.2.3.2.1 Stiffness of 4 Tie Rods The total stiffness of the four (4) tie rods, Ka, is based upon an equation from Reference 6.25 (p.

31):

AE Ka=4 where:

A=

0.0878 in* cross-sectiod area of the tie rod, based on the basic pitch diameter E= 27.0 x 10' psi (conservatively chosen at 300*F)

L= 13.3 in, length between top plate and upper flange So:

2 K =4 (0.0878/n )(27.0X10' k* )

= 7.13ES f3-13.3m. m 5.2.3.2.2 Stiffness of the Top Plate The equations for calculating the deflection of the top plate are found in Reference 6.11, Table 2 Case 9:

w a' L' y = 2D (1 + y - 2L )3 where:

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k M IDIF C-PENG-CALC-018, Rev. 00 Page 29 of 54 D= Et 3

=

27.3X10' ' q"lb (0.75)'in' 12(1 -y ) 2 = 1054690in -Ib 12(1- 0.22) and L3 and L, are constants, and are calculated using the equations of Reference 6.11, pp. 398-399 where:

a = 2.375 in

r. = 2.25 in (conservatively) t = 0.75 in y = 0.3 E = 27.3 X 10' psi (at 250*F. This temperature is chosen due to the proximity of the Top Plate to the Clamp, and the Clamp NOP temperature is assumed in Section 5.2.1 to be 300*F)

L3 = 0.00002365 L, = 0.05029 Solving for the stiffness of the top plate:

K,, = 2s*

3 = 57.62Eh '"

1 2D(1+y - 2L ) 3 5.2.3.2.3 Equivalent Stiffness: Tie Rods-Top Plate System The equivalent stiffness of tie rod-top plate system is determined as follows:

K, =

Ka X,p Therefore:

I K" 704,000lb 1

+

1 in 7.13E5 57.62E6 l

l ABB Combustion Engineering Nuclear Power

M IDIF C-PENG-CALC-018. Rev. 00 Page 30 of 54 5.2.3.3 TotalDeflection, Ar The base equation developed previously is used to determine Ax:

1

-K,, Ar' - F, Ar - F, Gap = 0

=> Ax = [F. (F.2 + 2KyF. Gap)"] / K y Given:

F. = 1250 lbs Ky= 704,000 lb / in Gap = 0.043 in.

=> Ax = 0.0143 in.

5.2.4 Impact Force The impact force, F.pci, on the top plate and tie rods is then calculated:

F,,,,,, = K,, Ar Therefore:

F,,,,,, = 704,000 S(0.0143in) = 10.07 kips m

The value of Foyoi = 10,200 lb will be used in subsequent analysis of the MNSA components.

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M IDIN C-PENG-CALC-018 Rev. 00 Page 31 of 54 I 5.3 StructuralStress Analysis ofthe MNSA Components The Design Loads for the various MNSA components will be a function of either bolt pre impact load, and/or thermal expansion loads, depending upon the effects of the so particular component.

5.3.1 Tie Rod 5.3.1.1 Design Bolt LoadStresses The design bolt load for the tie rod is considered to be the hydrostatic load which results from Design Pressure only, since the tie rod is not used for gasket-joint purposes.

Section 5.4.2.1 determines the service stress in the tie rod for a pressure which bou Pressure; this stress value is compared to the design bolt load stress allowable:

6.45 ksi < 26.9 ksi (S. at 650*F)

The hydrostatic load stress is below the stress allowable, which indicates that the actual bo (A S) exceeds the minimum required bolt area (A.).

5.3.1.2 StressDue to ThermalExpansion The differential thennal expansion between the tie rod and the upper flange (or top plate an additional tensile load on the tie rod. For the analysis of the tie rod, this additional load is assumed to be completely taken up by deformation of the tie rod.

The stress effect of this differential thermal expansion is ' termined below.

From Reference 6.25:

ot.. = E a AT The a to be used is the differential in coefficients for the materials of the tie rod an Aa. Therefore:

ot.. = E Aa AT Given:

E = 25.25 E6 psi at 61l'F ABB Cornbustion Engineering Nuclear Power  !

~

A ItIt M IDIF C-PENG-CALC-018. Rev. 00 Page 32 of 54 Aa = an as,w = (9.55 E 8.95 E-6) = 0.6 E-6 (a at 61l'F)

AT = 611 - 70 = 541 *F I

ct = 8.2 ksi The effective tensile force due to this thermal expansion is determined by: '

2 P=or A = (8.2 ksi)(0.0775 in ) = 0.636 kips 5.3.1.3 Preload The tie rod and nuts are being preloaded to 75 in-lbs. To determine the load in each tie rod, the following equation is used (Reference 6.15):

T = 0.2 F d

=> F =T / 0.2 d Given:

T = the applied torque = 75 in-lbs d = is the nominal major tie rod diameter = 0.375 in.

l F = (75 in-lbs) / (0.20) (0.375 in) = 1.00 kips.

5.3.1.4 Impact Load Since the impact load is distributed evenly to each of 4 tie rods, the impact load on the tie r 10.2/4 = 2.55 kips l 5.3.1.5 Maximum Tie RodLoad l

The load on the tie rod will be the greater of the load due to preload plus thermal expansion and t load due to the impact:

Preload + thermal expasion = 1.00 + 0.636 = 1.636 kips impact = 2.55 kips (> 1.636 kips) c ABB Combustion Engineering Nuclear Power

MIDIF- C-PENG-CALC-018, Rev. 00 Page 33 of 54 Therefore, the maximum tie rod load is impact load P = 2.55 kips.

5.3.1.6 A verage Stress, cr, (NB-3232.1)

The average (axial) stress (oi) in the tie rod is due to the maximum tie rod load:

ci= P/A.

A = 0.0775 in2 i

P = 2.55 kips 2

ei = (2.55 kips /0.0775 in ) = 32.9 ksi < 2 Sm = 53.8 ksi (S. at 650*F) 5.3.1.7 Marimum Stress (NB-3232.2)

The maximum stress in the tie rod is essentially a stress intensity due to a combination of the average stress, bending stress from the OBE (Operating Basis Earthquake) event, and the torsional shear stress due to residual torque (from preload).

5.3.1.7.1 Seismic Bending Stress Prior to any (complete) weld failure, a seismic event will cause accelerations of the MNSA. Most.

components will experience little er very little effects from these seismic accelerations. However, i' because of motions associated with the top plate, the inboard end of each tie rod will be subjected to bendir.g stress. This bending stress (a s) will be conservatively added to the average and torsional shear stress for determming the maximum stress in the tie rod.

The bending stress at each of the tie rods is deternune,i by applying the maximum acceleration occurring at OBE event to the top plate, following with the even distribution of the resulting force to each of 4 tie rods, and then calculating the stress at the tie rod inboard end.

c = 1/4 Mc/I = 1/4 (65.4 x 0.149) / 3.82 x 10" = 6.38 ksi, where l

M=FL/2 = bending moment (Reference 6.11, Table 3, Case Ib)

M=FIJ2 = (9.56 x 13.68)/2 = 65.4 in-lbs - bending moment F = W(1+a) = 5.17 (1+0.85) = 9.56 lbs - acting force; W = pV = 0.29 x 17.82 = 5.17 lbs - weight of the top plate l ABB Combustion Engineering Nuclear Power l

l

A ItIt M IDIF C-PENG-CALC-018. Rev. 00 Page 34 of 54 2

a = }G,2 +G,2 = y0.6 + 0.6 = 0.85g - maximum acceleration at OBE, according to Reference 6.14, accelerations in any horizontal and vertical direction shall be applied simultaneously.

G, = G y= 0.6g - conservatively assumed in both horizontal directions and in the vertical direction OBE acceleration values (Reference 6.1 p = 0.29 lb/m~ ' - density of the stainless steel 2

V = x/4 x 5.5 x 0.75 = 17.82 in' - volume of the top plate (conservative) l L = (15.3 - 2 + 0.75/2) = 13.68 in - length of the tie rod from upper flange to the center ofgravity of top plate I = xd'/64 = 3.82 x 10" in' - moment ofinertia of the tie rod d = 0.297 in - basic minor diameter of the tie rod c = d/2 = 0.297/2 = 0.149 in 5 3.1.7.2 Residual Torque, Ta The residual torque due to preload may be calculated using the following equation for standard threads (Reference 6.15, Equation 6):

Ta = 0.5625 T = 0.5625 (75) = 0.042 in-kips where:

0.5625 = multiplier based upon a coefficient of friction of 0.15 and standard bolt dimensions T = 75 in-lb = applied torque 5.3.1.7.3 Torsional Shear Stress, tr The torsional shear stress is determined using the following equation derived from Equation 7 of Reference 6.15:

Tr = 16Ta / xd' =16 (0.042) / x (0.32)$ = 6.53 ksi where:

d = average of basic pitch diameter (0.3344 in.) and minor diameter (0.297 in.) =

0.32 in.

i 5.3.1.7.4 Maximum Stress, c.

i The maximum stress intensity (o.) is determined using the following equation (Reference 6.15, Equation 8):

l l

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A It R M IDIN C-PENG-CALC-018, Rev. 00 1

Page 35 of 54 l om =2(#' #' )2 +(rr ) = 41.39 ksi < 2.7 Sm = 72.6 ksi (S. at 650*F)

< 3.0 Sm = 80.7 ksi (S. at 650*F)  !

5.3.1.8 Shear Stress (r) - Threads i

i At Too Plate (hex nuts)

The tie rods pass through the top plate and are held in place with hex nuts at the top and at the bottom. The impact load, in directly loading the top plate and top nut, will create stresses in this section of the tie rod which are in addition to the tie rod / nut preload stresses. The nuts are of the same material as the rods. Therefore, the parameters associated with the external threads of the rod l

are used (i.e., because of the smaller shear area).

l From Reference 6.18:

AS,= n n Le Kn max [(1/2n) + 0.57735 (Es min - Kn max)] = 0.288 in2 where:

n = number of threads per inch = 16 Le = the length of engagement (nut thickness) = 0.5 in (Ref. 6.8)

Kn max = maximum minor diameter ofinternal thicad = 0.321 in Es min = minimum pitch diameter of external thread = 0.3287 in i

P = 2.00 idps 2

t = 2.55 kips / 0.288 in = 8.85 ksi < 0.6 Sm = 16.14 ksi

At Uoner Flance On the other side, the tie rods thread into the Upper Flange. The lower strength Upper Flange threads are evaluated below. (The external tie rod threads in the Upper Fi.uige have essentially the same stress as the external tie rod threads in the top plate region, which were evaluated previously.)

From Reference 6.18:

AS, = x n Le Ds min [(1/2n) + L57735(Ds min - En max)] = 0.414 in 2 where:

n = number of threads per inch = 16 ABB Combustion Engineering Nuclear Power l

J

A R It M IfIF C-PENG-CALC-018. Rev. 00 Page 36 of 54 Le = the length of engagement. Assume equal to 0.5 in En max = maximum pitch diameter ofinternal thread = 0.3401 in Ds min = minimum major diameter of exterdathread = 0.3643 in P = 2.55 kips 2

t = 2.55 kips / 0.414 in = 6.16 ksi < 0.6 Sm = 9.72 ksi The minimum allowable length of engagement of the tie rod into the Upper Flange may be calculated as a simple proportion:

Le., = (Shear Stress / Allowable Stress) x Assumed Length ofEngagement =

= (6.16/9.72) x 0.5 = 0.317 in.

5.3.2 Hex Head Bolt (into Hot Leg) 5.3.2.1 DesignBoltLoadStresses Design Bolt Load for the Design Pressure, W i W.i = H + H, H = 1.963 kips (from Section 5.2.3.1.1)- hydrostatic end force H, = 2b x 3.14GmP - compression load to ensure a tight joint (

b = 0.25 in (width of seal)

G = 1.24 in (average diameter of seal, Reference 6.8) m = 1.3 (from Reference 6.21, p. 47)

P = 2.500 ksi

=> H, = 6.33 kips

=> W.i =1.963 + 6.33 = 8.29 kips Stress due to W i = W.i / A.

= 8.29 kips / 4(0.1599 in')

= 12.96 ksi < 26.9 ksi (S. at 650*F)

Design Minimum Initial Bolt Load, W.2 ABB Combustion Engineering Nuclear Power m

A ItIt  !

1. "LIFIF C-PENG-CALC-018, Rev. 00 Page 37 of 54 i

Wm2 s taken as the total preload. Bolt stress due to preload only (cipi)is calculated in Section 5.4.3.1:

ci.pi = 22.51 ksi < 28.3 ksi (Sm at 100*F)

The stress due to 'W.i and W.2 are below their respective allowables, which indicates that the actua bolt area (A3) exceeds the minimum required area (Am). '

5.3.2.2 Stress Due to ThermalExpansion The differential thermal expansion between the hex head bolt and the upper flange and comp collar will create an additional tensile load on the bolt. For the analysis of the bolt, this additional load is assumed to be completely taken up by deformation of the bolt.

The stress effect of this differential thermal expansion is considered to be equivalent to that of the rod since the respective tic rod - top plate and bolt-flange materials are the same:

ot = 8.2 ksi The effective temile force due to this thermal expansion is determined by:

2 I = ct.. A = (8.2 ksi)(0.1599 in ) = 1.31 kips 5.3.2.3 Preload The bolts are being preloaded to . 9 ft Ib. To determine the load in each bolt, the following is used (Reference 6.15):

T = 0.2 F d

> F =T / 0.2 d Given:

T = the applied torque = 360 in-lbs d = is the nominal major bolt diameter = 0.50 in.

F = (360 in-lbs) / (0.20) (0.50 in) = 3.600 kips.

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A It11 M IDIF C-PENG-CALC-018, Rev. 00 Page 38 of 54 ,

5.3.2.4 Maximum Bolt Load As it was described in Section 2.3.2, the impact from ejection of the nozzle will increase the load on the bolts.

The total load on the bolt can be expressed as follows:

I e 8 F = Preload +

y" F,

<Kw+k,w In the above expression, Kno, is considered to be the equivalent stiffness of the components which are put in compression due to the torquing / tightening of the hex head bolts.

The stiffness of the components in the flanged connection between the MNSA and the Hot Leg is edculated below.

l l

Stiffness of Hex Head Bolts: 1 l

The stiffness of the bolts is calculated using the same methods described for the tie rods in Section 5.2.3.2.1. Dimensions are taken from Reference 6.8.

2 AE (0.172 in )(25.0X 10' g j K%=4 I =4 .

",) = 6,719,000 2.56 in m 2

l where: A = 0.172 in , cross-sectional area of the bolt, based on the basic pitch diameter 0.4675 in.

E = 25.0 x 10' psi (at 650*F) 1 = effective length of bolt, assuming 0.5 in of thread engagement

= thread engagement + lower flange + upper flange + washer

= 0.5 + 0.5 + 1.5 + 0.06 = 2.56 inch ABB Combustion Engineering Nuclear Power

A ItIk

1. "LIf W C-PENG-CALC-018, Rev. 00 Page 39 of 54 Stiffness of Overall Flange:

The Sampling MNSA has three components which represent the flanged connection to the Hot Leg, the upper flange (top), the upper flange (bottom), and the compression collar. The stiffness of each of these components is calculated with the use of Reference 6.11.

Upper flange (top and bottom):

The following equations are found in Reference 6.11, Table 24, Case la. All dimensions are taken from Reference 6.8.

w a' C, L, y= g ( (, - L,)

where:

3 Et 25.0X 10' #"2 (0.75)9n' D= = = 965,831/n -lb 12(1-y2 ) 2 12(1- 0.3 )

C i, C2, L9, and L 3are constants, and are calculated using the equations of Reference 6.11, using the following dimensicns. Since the flange does not have a rectangular cross section, the dimensions an selected to produces the lowest flange stiffness.

a = outer radius,1.906 in b = inner radius,0.781 in ro= radius of applied load, 0.781 in t = thickness,0.75 in y = Poisson's ratio, 0.3 E - elastic modulus, 25.0 X 10' psi (at 650*F)

C = 0.593 C, = 0.924 l L3 = 0.0215 L, = 0.2973 i

Solving for the stiffness of the upper flange, top:

\

K ,_ =W= y a, 2xr" Ib

=4,043,000 .f in D (C,L, C,

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M IF15 C-PENG-CALC-018, Rev. 00 l Page 40 of 54 Compression Collar:

i2 1 Soo" r' [

t I  !

1 169 ,

I -

l*

i V

7 033N kc d o250' f = 1 oo3- =ff 1

, m-

l

': 2 00o- =

a = C.75 in b 0.502 in l r, = 0.623 in l t = 0.839 in (conservatively) 7 = 0.3 l E = 25.0 X 10' psi I Ci = 0.3190 C7 = 0.3752 L 3= 0.0007 '

L, = 0.1452 Et 3 25.0X10' i" (0.839)'in' D- 2

= = 1,352,083i.7 - Ib 2

12(1 - 7 ) 12(1 - 0.3 )

K. = E = , 2n* = 102,222,000 bf 7 .

a' (C,L, in D C, Determination of equivalent flange stiffness:

The two upper flange pieces (top and bottom) are considered to act in parallel with each

~ $er (see Section 2.3.2). The overall flange and the compression collar act in series with the bolt. The effective stiffness of the components is calculated below.

l l

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A IkIt M IDIF C-PENG-CALC-018, Rev. 00 Page 41 of 54 I

K g,= = 7,493,000 (K,,,,,,,+K,,,,,,,) K a, l Therefore, the maximum bolt load is l 1

i

, 102 F"" = 3.6 + = 4 806 kips s 6,719,000 + 7,493.000 s 4 1 Use F,., = 4.82 kips. i 1

1 5.3.2.5 Average Stress, m (NB-3232.1) 8 The average (axial) stress (oi) in the bolt is due to a combination of stresses from the maximum bolt 1 load and from differential thermal expansion:

c = P/A + ot..

Ai = 0.1599 in 2 P = 4.82 kips ot.. = 8.2 ksi 2

ci = (4.82 kips /0.1599 in ) + 8.2 = 38.34 ksi < 2 Sm = 53.8 ksi (at 650*F) 5.3.2.6 Marimum Stress (NB-3232.2)

The maximum stress in the bolt is essentially a stress intensity due to a combination of the av stress and the torsional shear stress due to residual torque (from preload).

5.3.2.6.1 Residual Torque, Ta The residual torque due to preload may be calculated using the following equation for standard threads (Reference 6.15, Equation 6):

Ta = 0.5625 T = 0.5625 (360) = 0.203 in-kips where:

l

' 0.5625 = multiplier based upon a coefficient of friction of 0.15 and standard bolt dimea:; ions ABB Combustion Engineerin,) Nuclear Power

1. "tIfIF C-PENG-CALC-018 Rev. 00 l Page 42 of 54 T = 360 in-lb = applied torque l 5.3.2.6.2 Torsional Shear Stress, tr '

The torsional shear stress is determined using the following equation derived from Equation 7 of Reference 6.15:

t t = 16Ta / xd' =16 (0.203) / x (0.7 ~ = 11.16 ksi l

where:

d = average of basic pitch

-(0.4675 in.) and minor diameter (0.4374 in.) =

0.4525 in.

5.3.2.6.3 Maximum Stress, c The maximum stress intensity (c ) is determined using the following equation (Reference 6.15, Equation 8):

e =2( )2 + (tr)' = 44.36 ksi < 2.7 Sm = 72.6 ksi (S. at 650*F)

< 3.0 Sm = 80.7 ksi (S. at 650'F) 5.3.2.7 1hreadShearStress, r '

From Reference 6.18:

AS, = x n Le Kn max [(i/2n) + 0.57735(Es min - Kn max)] = 0.400 in 2 where:

n= number of threads perinch = 20 Le = the length of engagement. Assume equal to 0.5 in Kn max = maximum minor diameter ofinternal thread = 0.457 in Es min = minimum pitch diameter ofexternal thread = 0.4619 in P = 4.82 kips t = 4.82 kips / 0.400 in 2= 12.05 ksi < 0.6 Sm = 16.14 ksi The minimum allowable length of engagement of the hex head bolt into the Hot Leg pipe may be calculated as a simple proportion, based on the bolt threads.

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A ItIt HEBlF C-PENG rALC-018. Rev. 00 Page 43 of 54 Le

(Shear Stress / Allowable Stress) x Assumed Length of Engagement

= (12.05/16.14) x 0.5 = 0.373 in.

5.3.3 Top Plate 5.3.3.1 Shear Stress, r A. =(2a + 2b) t = (l5.00 in) (0.75 in) = 11.25 ir.2 where:

2a +2b = the perimeter of the clamp = 15.00 in (Reference 6.8) t = the thickness of the top plate = 0.75 in P = 10.2 kips 2

I= 16.2 kips /11.25 in = 0.91 ksi < 0.6 Sm = 9.72 ksi 5.3.3.2 Bendingstress, ob The impact load is distributed over the area of the top plate in contact with the clamp. Con the impact load is applied at the location of the radius equal to the smallest dimension of the rectangular contacting surface of the clamp (r = 1.5 in). From Reference 6.11, Table 24, Case 9a (dimensions are taken from Reference 6.8):

r = 1.5 in a= 2.375 in t = 0.75 in w = Lp., / (2 n) r. = 10.2 kips / (2 x) 1.5 in = 1.082 kips / in r./a= 0.6316 Conservatively K. = 0.26642 M = K. w a = 0.26642 (1.082 kips /in) (2.375 in) = 0.684 kips-in 2

o = 6M/t = 6 (0.684 kips-in) / (0.75)2 in2 = 7.296 ksi on = 7.30 ksi < l.5 Sm = 24.3 ksi ABB Combustion Engineering Nuclear Power

A It H i B lI kB C-PENG-CALC-018, Rev. 00 Page 44 of 54 5.3.4 Compression Collar 5.3.4.1 Shear Stress, r 2

A.= (x)(D)(t) = (x)(l.500 in) (0.335 in) = 1.579 in P = 3.6 kips /bc!t x 4 bolts = total preload from b31ts

= 14.4 kips t = 14.4 kips /1.579 in2 = 9.12 ksi < 0.6 Sm = 9.72 ksi 5.3.4.2 Bearingstress, s 2 2 A. = (x/4)(D -,wn,oo - d w.p unto) = (x/4)(l.49 2- 1.003 )2 in = 0.954 in 2 P = 14.4 kips ci, = 14.4 kips / 0.954 in2 = 15.09 ksi < Sy = 17.9 ksi 5.3.5 Upper Flange Shear s. tress, t z

A. = (x)(D) (t) = (n)(2.02 in) (0.325 in) = 2.062 in P = 14.4 kips t = 14.4 kips / 2.062 in z= 6.98 ksi < 0.6 Sm = 9.72 ksi Due to the proximity of the bolts and support surface, bending stresses are considered to be small and are neglected.

5.3.6 Clamp The clamp is designed to deliver the load acting on a nozzle symmetrically to a top plate. The cla is designed such that relative motion between the elbow and the clamp halves may occur during the operation (or impact), due to the gap existing between those parts at initial clamp assembly. This relative motion may cause a potential wedge situation, and therefore, the axial load may be carried by the clamp bolts under pressure (impact). Because of complexity of evaluating this load, it is conservatively assumed that full load acting on a nozzle is carried by the bolts.

ABB Combustion Engineering Nuclear Power

M IDIF C-PENG-CALC-018. Rev. 00 Page 45 of 54 5.3.6.1 Hex HeadBolt - Clamp 5.3.6.1.1 Design Bolt Load Stresses l

Section 5.4.1 determines the normal operating pressure force:

F, = 2.00 kips Using an assumption that tlus force is carried by the bolts, a service svesi m each bolt is calculated and compared to the design bolt load stress allowable:

on = (2.00 kips /3) / 0.1599 inz = 4.17 ksi 4.17 ksi < 26.9 ksi (S. at 650*F)

The load stress is below the stress allowable, which indicates that the actual6 bolt area (A ) excee the minimum required bolt area (A.).

5.3.6.1.2 Stress Due to Thermal Expansion l

The stress effect of the differential thermal expansion and the corresponding effective tensile force 1

due to this thermal expansion are assumed equivalent to ones determined in Section 5.3.2.2.

ot.. = 8.2 ksi P = ct.. A = 1.31 kips 5.3.6.1.3 Preload The bolts are being preloaded to 30 ft-lb. To determine the load in each bolt, the following eq is used (Reference 6.15):

j T = 0.2 F d i => F =T / 0.2 d Given:

T = the applied torque = 360 in-lbs d = is the nominal major bolt diameter = 0.50 in.

i ABB Combustion Engineering Nuclear Power

A ItIk MIDIN C-PENG-CALC-018, Rev. 00 Page 46 of 54 F = (360 in-lbs) / (0.20) (0.50 in) = 3.600 kips.

5.3.6.1.4 Impact Load l

Assuming that the impact load is applied evenly to each of the 3 bolts, the impact load on the bolt equals:

10.2/3 = 3.4 kips 5.3.6.1.5 Maxin um Load l l

The maximum load on each bolt will be greater of the load due to preload and thermal expansion and the load due to the impact:

Preload + thermal expansion = 3.6 + 1.31 = 4.91 kips:

l Impact = 3.4 kips (< 4.91 kips)

Therefore, the maximum bolt load is 4.91 kips.

5.3.6.1.6 Average Stress,c The average (axial) stress (oi) in the bolt is due to the maximum bolt load:

l or = P/A.

2 A. = 0.1599 in P = 4.91 kips 2

c = (4.91 kips /0.1599 in ) = 30.71 ksi < 2 Sm = 53.8 ksi (S. at 650 F) 5.3.6.1.7 Maximum Stress

The maximum stress in the boli is due to a combination of the average stress and the torsional shear i stress due to residual torque (from preload).

Torsional Shear Stress. n Using results from Sections 5.3.2.6.1 and 5.3.2.6.2, the residual torque due to preload and corresponding torsional shear stress are:

4 ABB Combustion Engineering Nuclear Power

A It11 M IDIF C-PENG-CALC-018, Rev. 00 Page 47 of 54 Ta = 0.203 in-kips tr = 11.16 ksi Maximum Stress. o m The maximum stress intensity (o ) is determined using the following equation (Reference 6.15, Equation 8):

c., = 2 (

)2 + (rr )2 = 37.96 ksi < 3.0 Sm = 80.7 ksi (S. at 650*F) 5.3.6.1.8 Thread Shear Stress, t From Reference 6.18:

AS, = n n Le Kn max [(1/2n) + 0.57735(Es min - Kn max)] = 0.800 in2 where:

n= number of threads per inch = 20 Le = the length of engagement. Assume equal to 1.0 in Kn max = maximum minor diameter ofinternal thread = 0.457 in Es min = minimum pitch diameter of external thread = 0.4619 in P = 4.91 kips z

t = 4.91 kips / 0.800 in = 6.14 ksi < 0.6 Sm = 16.14 ksi The minimum allowable length of engagement of the bolt into the clamp half may be calculated as simple proportion, based on the bolt threads.

Le e

(Shear Stress / Allowable Stress) x Assumed Length of Engagement

= (6.14/16.14) x 1.0 = 0.38 in.

5.3.6.2 ClampHalf 5.3.6.2.1 Thread Shear Stress, t On the other side, the bolts thread into the clamp half. The lower strength clamp threads are evaluated below.

ABB Combustion Engineering Nuclear Power

\

l PKEFEF C-PENG-CALC-018, Rev. 00 Page 48 of 54 l From Reference 6.18:

AS, = x n Le Ds min [(1/2n) + 0.57735(Ds min - En max)] = 1.082 in 2 where:

n = number of threads per inch = 20 Le = the length of engagement. Assume, equals to 1.0 in En max = maximum pitch diameter ofinternal thread = 0.4731 in Ds min = minimum major diameter of external thread = 0.4906 in P = 4.01 kips i

z t = 4.91 kips /1.082 in = 4.54 ksi < 0.6 Sm = 9.72 ksi The minimum allowable length of engagement of the bolt into the clamp half may be calculated as a simple proportion:

i Le

(Shear Stress / Allowable Stress) x Assumed Length of Engagement

= (4.54/9.72) x 1.0 = 0.47 in.

M FatigueAnalysis The fatigue analysis of the components will conservatively consider loads which may exist on the components after weld or nozzle failure has occurred. Prior to failure, components will be subject to loads due mainly to preload and thermal expansion. After failure, and assuming that the nozzle / valve is free to move, certain components will be additionally stressed because of the interna pressure forcing the nozzle / valve up against the top plate. The load on these components would be cyclical, given the change in pressure and temperature that occurs as the plant heats up and then cools down.

The critical components for fatigue analysis purposes are the tie rod and hex head bolt, on the basis of:

e preload tensile stresses l e thermal expension tensile stresses

! e stress concentrations in the threaded sections, and I e i for the levels of stresses involved, a more restnctive number of allowable cycles (versus the stainless steel MNSA components; see Table I-9.1 of Reference 6.9)

ABB Combustion Engineering Nuclear Power l

l kkk l M IDIF C-PENG-CALC-018, Rev 00 i Page 49 of 54 It is noted that, in the fatigue analyses below, the stresses produced by the one-time application impact load are not considered since the contribution to fatigue from this one occurrence is not significant.

5.4.1 Normal Operating Pressure Force i

The effect of the force acting on the MNSA components due to internal pressure is similar to that the impact load, only of a smaller magnitude, and it is a function of the internal pressure and the of the nozzle. The pressure used to detennine the force is based on the maximum interna for all Normal and Upset conditions from Reference 6.3, which is 2350 psia (for the 10%

Increase transient). The normal operating force is determined as a function of the Design Pres force calculated in Section 5.2.3.1.1:

F, = 1.963 kips (2350 psia / 2500 psia) = 1.85 kips 1

Use F, = 2.00 kips This force will be used to calculate normal operating tensile stresses in the tie rod and the Peak Stress Intensity calculation.

5.4.2 Tie Rods 5.4.2.1 Peak Stress The maximum Peak Stress in the tie rod is calculated, as follows:

Peak Stress = fsrf*(c .)

where:

fsrf = fatigue strength reduction factor = 4.0 (from NB-3232.3) c . = maximum stress using the load from normal operating pressure instead of the impa load:

com. = 2 (cr' + <r* )2 + (tr ),'

2 where:

ci = ci., + ci.pi+ c .a(conservatively), and ci., = tensile stress due to pressure ABB Combustion Engineering Nuclear Power

l l

M IDIN C-PENG-CALC-018, Rev. 00 Page 50 of 54

= (2.00 kips /4)] / 0.0775 in 2

= 6.45 ksi ci.pi = tensile stress due to preload (preload from Section 5.3.1.3)

= 1.0 kips / 0. 0775 in 2

= 12.9 ksi l c a. = tensile stress due to thermal expansion (stress from Section 5.3.1.2)

= (8.2 ksi)*(30E6/25.25E6) (E/E per NB-3232.3(d))

= 9.74 ksi on

= 6.38 ksi (stress from Section 5.3.1.7.1) tr = 6.53 ksi(stress from Section 5.3.1.7.3)

=> o x. = 37.80 ksi I

=> Peak Stress = 4 (37.80) = 151.2 ksi Based upon a minimum stress value of 0.0 ksi (this is a conservative approach, since preload goes away), the maximum Peak Stress Intensity Range (Sp ) in the tie rod is:

Sp = 151.2 - 0 = 151.2 ksi '

1 l

54.2.2 Usage Factor l The calculation of the usage factor for the tie rod is based upon the maximum Peak Str Range of 151.2 ksi:

S. = Sp/2 = 75.6 ksi This range is considered to occur a total number of 700 cycles, which is the sum of the numbers o cycles for Heatup/Cooldown (500) and Plant Leak Test (200), per Reference 6.20 (seismic loads conservatively included in a total number of cycles).

For a S. = pS /2 = 75.6 ksi, the allowable number of cycles per Table I-9.1 (Reference 6.9) is approximately 1750. The usage factor (U)is:

l U = 700/1750 = 0.4 There are other Normal and Upset transients which are defined for the Piping (per Referenc but their contribution to fatigue in the tie rod is not significant.

ABB Combustion Engineering Nuclear Power l

l

b 4

M IDIF C-PENG-CALC-018, Rev. 00 l Page 51 of 54 5.4.3 Hex Head Bolt 5.4.3.1 Peak Stress The maximum Peak Stress in the bolt is calculated, as follows:

l Peak Stress = fstf*(c.x.)

where:

fsrf = fatigue strength reduction factor = 4.0 (from NB-3232.3) c x. = maximum stress using the load from normal operating pressure instead of the impact load:

c.x. = 2 ( ):+(r )'

r where:

ci = ci, + c,i+ ct (conservatively), and ci, = tensile stress due to pressure

= (2.00 kips /4) / 0.1599 in 2

= 3.13 ksi ci.pi = tensile stress due to preload (preload from Section 5.3.2.3)

= 3.6 kips / 0.1599 in2

= 22.51 ksi

ot.. = tensile stress due to thermal expansion (stress from Section 5.3.2.2)

= (8.2 ksi)*(30E6/25.25E6) (E/E per NB-3232.3(d))

= 9.74 ksi tr = 11.16 ksi(stress from Section 5.3.2.6.2)

=> c.x. = 41.83 ksi *

=> Peak Stress = 4 (41.83) = 167.32 ksi Based upon a minimum stress value of 0.0 ksi (this is a conservative approach, since preload never goes away), the maximum Peak Stress Intensity Range (Sp ) in the bolt is:

S, = 167.32 - 0.0 = 167.32 ksi l ABB Combustion Engineering Nuclear Power i

f i

A ItIk

1. "t1915 C-PENG-CALC-018, Rev. 00 Page 52 of 54

5. 4.3. 2 Usage Factor ,

l The calculation of the usage factor for the hex head bolt is based upon the maximum Peak Stress Intensity Range of 167.32 ksi: '

S. = S,/2 = 83.7 ksi This range is considered to occur a total nun.ber of 700 cycles, which is the sum of the numbers of cycles for Heatup/Cooldown (500) and Plant Leak Test (200), per Reference 6.20.

For a S. = S,/2 = 83.7 ksi, the allowable number of cycles per Table I-9.1 (Reference 6.9) is ,

approximately 1400. The usage factor (U)is: '

U = 700/1400 = 0.5 There are other Normal and Upset transients which are defined for the Piping (per Reference 6.20),

but their contribution to fatigue in the bolt is not significant.

5.5 Consideration ofHydrostatic Test Pressure Conditions Per Paragraph 1.3.1 ofReference 6.7, the deliverable MNSA hardware is not required to be hydrostatically tested. However, it is noted that the pressure load created by the hydrostatic pressure of 3125 psi (Reference 6.20) is less than the impact load. Therefore, stresses resulting from hydrostatic testing would be acceptable.

5.6 Consideration ofFaulted Conditions Reference 6.20 li:.ts Faulted Conditions which are identical to the Design Condition except that they also include a Design Basis Earthquake and Pipe Rupture events. Pipe Rupture event has no effect on the MNSA components. An asse,ssment is made of the effect ofFaulted Conditions by reviewing the maximum stress results for the tie rod (Sections 5.3.1.7.1), which is the component most significantly affected by an earthquake event (either OBE or Maximum /DBE). Conservatively doubling the OBE bending stresses to simulate the effects of the Maximum Earthquake event results in stresses which meet the 3S. allowable for Maximum stress. Therefore, the stresses resulting from Faulted Conditions are acceptable.

ABB Combustion Engineering Nuclear Power

l N k

1. "LIFIF C-PENG-CALC-018, Rev. 00 Page 53 of 54 6 REFERENCES 6.1. ABB Project Plan No. C-NOME-IPQP-0263, Revision 0, "Waterford Mechanical Nozzle Seal Assemblies", March 1999. l i

6.2.

ABB Combustion Engineering Nuclear Power Quality Procedures Manual QPM-101, Revision 03.

6.3.

" Analytical Report for Waterford Unit No. 3 Piping," Report No. CENC-1444, May 1981. '

6.4. Dravo Cogoration Drawing E-3029-LW3-RC-10, Field Rev. 5.

6.5. ABB CE Drawing 74470-771-003, Revision 02, " Primary Pipe Assembly".

6.6.

ABB CE Drawing 74470-772-001, Revision 04, " Instrument Nozzles Waterford III Piping" 6.7.

" Design Specification for Mechanical Nozzle Seal Assembly (MNSA) Waterford Unit 3",

Specification No. C-NOME-SP-0067, Revision 01.

6.8. ABB Drawing No.

6.8.1. E-MNSAWFD-228-003, Revision 03, " Hot Leg Sampling MNSA" 6.8.2. E-MNSA-228-013, Revision 07, " Mechanical Nozzle Seal Assembly Details" 6.8.3. E-MNSA-228-004, Revision 05, " Mechanical Nozzle Seal Assembly Details" 6.9. American Society of Mechanical Engineers Boiler and Pressure Vessel Code,Section III,1989 Edition (No Addenda).

I 6.10. " Test Report for MNSA Hydrostatic and Thermal Cycle Tests," Test Report No. TR-PENG-042, Rev.00.

6.11. "Roark's Formulas for Stress and Strain," Warren C. Young, Sixth Edition,1989, McGraw-Hill.

6.12. " Heat Transfer A Basic Approach", M. Necati Ozisik,1985, McGraw-Hill.

6.13. Inter-Office Correspondence from J. T. McGarry to K. H. Haslinger, "Waterford MNSA Stress Analysis", Letter No. NOME-99-C-0122, Revision 01, dated March 18,1999.

l 6.14.

" General Specification for Reactor Coolant Pipe and Fittings", Specification No. 00000-PE-140, Rev. 02, July 1973.

.ABB Combustion Engineering Nuclear Power l

~

b niFEB C-PENG-CALC-018, Rev. 00 Page 54 of 54 6.15.

"How to Calculate and Design for Stress in Preloaded Bolts", A.G. Hopper and G.V. Thompson, P i Engineering,1964.

l 6.16.

Engineering Report No. C-NOME-ER-0120, Revision 00, " Design Evaluation of MNS A for Various Applications at Waterford Unit 3", March 1999. I 6.17.

Calculation No. C-PENG-CALC-021, Revision 00, "Detemunation of Waterford 3 Hot Leg Seism Response Spectra & Accelerations for Use in the Analyses ofMNSAs", March 1999.

6.18. ANSI Standards for Threads, Appendix B, Bl.1,1982.

6.19.

" Addendum to CENC-1365 and CENC-1507 Analytical Reports for Southem California Edison San Oncie Units 2 and 3 Piping", Design Report No. S-PENG-DR-005, Rev. 01 6.20.

" Project Specification for Reactor Coolant Pipe and Fittings for Entergy Operations, Inc. Wa Unit 3", Specification No. 09270-PE-140, Rev. 07, December 1993.

6.21.

Union Carbide Grafoil, " Engineering Design Manual, " Volume One, Sheet and Lanunated R.A. Howard,1987.

6.22.

• Inter-Office Correspondence from Dale Gallodoro to Gary Bundick, " Sample Line Operati Conditions", dated March 14,1999.

6.23.

"Engineenng Mechanics: Statics and Dynamics", F.L. Singer, Third Edition, Harper & Row, N 1975.

4 6.24.

" Mechanical Engineers' Handbook", M. Kutz, ed., John Wiley & Sons, Inc.,1986.

6.25.

" Strength of Materials", F. L. Singer, Second Edi+ ion, Harper & Row, New York,1962

• This Reference is used per heference 6.13.

l ABB Combustion Engineering Nuclear Power l

. _ ~ -

Nk!k M 1987 '**"*'; l,*J'ld APPENDIX A ASSEMBLY DRAWING ABB Combustion Engineering Nuclear Power

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ggg C-PENG-CALC-018, Rev. 00 Page B1 ofB2 l

1 I

APPENDIX B 1

CALCULATION OF THE TIE ROD AVERAGE TEMPERATURE

\

i ABB Combustion Engineering Nuclear Power i

gg g C-PENG-CALC-018, Rev. 00 Page B2 ofB2 Calculation of average temperature due to axial conduction infinde Fin (" Heat Transfer - a Basic Approach", M.N. Ozisik, McGraw-Hill Iric.,1985)

Reference 6.12 T - T. . , hP To - T. " ' ,,, "5 Ambient Temperature T,inifinite 120 *F Base Temperature T,o 500 *F Nominal dimension tie rod OD 0.375in Norninal dimension tie rod OD 0.03125 ft Tie Rod area A 0.000767 ft2 Tie Rod perimeter P 0.098175 ft Tie Rod therm cond (Ref 6.9,300*F) k 8.8 btu /hr-ft-F generic outside film coefficient h 1.7 btu /hr-ft2-F calulcated coefficient m 4.972652 Length x T 0 0 500 0.6 0.0500 416 1.2 0.1000 351 1.8 0.1500 *00 2.4 0.2000 261 3.0 0.2500 230 3.6 0.3000 205 4.2 0.3500 187 4.8 0.4000 172 5.4 0.4500 161 6.0 0.5000 152 6.6 0'5500

. 145 7.2 0.6000 139 7.8 0.6500 I35 8.4 0.7000 132l 2321 9.0 0.7500 129 Average temperature at 8.4 inches is 232*F ABB Combustion Engineering Nuclear Power 1

j

AL IkIk 4

ggg C-PENG-CALC-018, Rev. 00 Page C1 ofC6 l

APPENDIX C QUALITY ASSURANCE FORMS ABB Combustion Engineering Nuclear Pour

ggg C-PENG-CALC-018, Rev. 00 Page C2 ofC6 i

Design Analysis Verification Checidist i

l Instivetions: If a major topic area (generally unnumbered, bold face type such as Use of Computer Software) is not applicable, then N/A (not applicable) next to the topic may be checked and the check boxes for all items under it may  :

be left blank. Where there is no check box under N/A for a numbered item, such a response is generally I inappropriale. If N/A is checked in such a situation, document the basis at the end of this checklist in the Comments section. )

Author IR Overall Assessment Concur.

Yes l N/A i

1. Are the results/ conclusions correct and appropriate for their intended use? 1 {

2. Are alllimitations and contingencies on the results/ conclusions documented? E Assignment of C9gnizant Enf,ineers, Independent Reviewers and Mentors

1. If there are multiple Cognizant Engmeers, has their scope been d~~-red?

@ @ {

1 [f there are multiple th' Reviewers, has their scope been documented?

] l

3. If there will be multiple Menagement Approvers, has their scope been d~~a-arad?

] l

4. If an law Reviewer is the supervuor or Project Manager, has authonzahon as an Lad-aaadaa' Reviewer been documented?

5. If there is a Mentor, has their scope and resporsibilrties been adequately h-*=d?

l Use of Computer Software For software wtuch has been vahdated under QP 3.13:

1. Is the software hsted on an Approved QC 1 Software l#

2. la the software apphcable for this analysa?

For Codelike Construas vahdnand unda QP 3.14.

1. Is the Code-Lake Construct hsted on an Approved QC 1 Software List?

2. Is the Code Lake Construct apphcable for tius analysts?

3. Was the Code l.ake Construa used drectly in the controlled locanon? No

- If No above, is the ccpy idersacal to the wrnoon in the corarolled locauan? (leave blank ifnot appbcable.)

if changes were made to the Code Lake Construct to meet specific analysis needs, were such changes head as non-validated software following pa.W 3.3.3? (tenve blank if not applicable. Complete the next section of this Checklist if"Yes".)

f or sobare wtuch has not been validated under QP 3.13 or QP 3.14:

1 Is the computer type, program name and revision identificahon documented?

2. Is a copy of the software mciuded m the Design Analysis?

Y Tiave tests been documented wtuch are adequate to demonstrate correct operation for the software's intended use?

4. Is the output inom the tests included m the Design Analysis?

5.

O flas the Cogruzant Engineer documented the results of the tests and the basis for concludmg the software is operatmg correctt ' for its mtended use?

6. Did the software, as used m tius analysts, give correct results?

ABB Combustion Engineering Nuclear Pover

i I

C-PENG-CALC-018, Rev. 00 Page C3 of C6 l Design Analysis Verification Checidist i

Author IR Use of Computer Software (continued) Concur.

Yes l N/A

1. Were .p-A-e. used m thz. Design Analysts m any way - data display, plonung, computanons, etc.? No

. If data display gqy (no computauons or ploumg), check "Yes" and slup remauung quesuons.

. If used for computauons:

• Are the computauons adequately documented?

. Are the resuhs cornct?

1

. If used for plotmg:

e is the anta to be plotted correct?

e Are the plots correct in other respects? (titles, scales, labels, etc.)

O i

8. Is a copy of the spreadsheet included m the Demgn Analysts? (A copy of tne fde may be mcluded or sufficient dctad included in the analvsts hwa6on to pemut recreatmg the spreadsheet.) E w

Objective of the Design Analysis

1. Has informauon necessary to derme the taak been included or referenced?

2. Have the objectives been enumerated?

3. Has the applicabihty and intended use of the results been documented?

Assessment of Significant Design Changes

1. Have significant design related changes that might impact this analysis becn considered?

2. If any such changes have been identified, have they been adequately addressed?

O @

4 Analytical Techniques (Methods)

l. Are the analytical techruques (methods) desenbed m sufficient detett tojudge their appropnateness? l

2. Are the analytical techtuques used or their apphcauon governed by an NRC issued SER7 No If yes, have the applicable SERs been documaned?

If yes, has the basis for concludmg the analysis is in conformance been h-~4?

]

3. Have analyucal technaques incorporated by reference to genene analyses, lead plant analyses or previous cycle analyses been previously venfied? Tl lJ

4. Are any modificauons or departures frorn previously approved analyucal techruques or Convenuonal or Automated Procedures h-ad and justified?

3. If superseded approved analytscal technsques or engmeenng procedures are used, ts their use jusufied and approved?

6. Does the saue date of referenced approved Convenuona) or Automated Procedures predate their use m this analysts?

Selection of Design Inputs Are the design inputs hW 1.

2 Are the design inputs correctly selected and tramble to their source?

@ Q

3. Are the bases for selection of all design inputs documented?

4 Is previously unvenfied design input uised in this inalysis?

If Yes. is it treated in accordance with QP 3.2, paragraph 3.4 for use of unvenfied desiris information?

ABB Combustion Engineering Nuclear Pover

p gg C-PENG-CALC-018, Rev. 00 Page C4 ofC6 Design Analysis Verification Checklist l Author IR Selection of Design inputs (continued)

Yes l N/A Concur. l S.

Is the venficauon stanas of design inputs transnutted from customers or CENP Nuclear Svstems appropnate and documented? @ ]

6. I Is the use of customer. controlled sources such as Tech Specs LTSARs, etc. authonzed, and does the authonzatson spectfy ** level, revuion number, etc.? @ @

Assumptions

1. If *here are no assumpuons, is this documented?

2. Are local assumpuons d~'-'=d fullyjustified and verified?

3.

Are Intemal ard External M'~*% which must be cleard by CENP or the customer hsted on a Conungencies and Assumptions form? @ @

4. Is the Project Manager responsible for cleanng the Assumptions idecufied on the form?

Results/ Conclusions

1. Are all results contamed in or referenced in the Results/ Conclusion sectaon? (Where feasible, in the enumerated order of the objecuves.) @ Q

2. Are all linutations on the resuhs/ conclusions and their applicabihty Mead in this sectioti9 3.

Are all conungencies on the resuhu that must be cleared hsted in the Results/ Conclusion section or the Coniingencies and Assumpuons form referenced? @ @

4. Is the Project Manager responsible for clearing the Assumptions or Contagencies identdied on the form?

] @

Other Elements 1.

Has a companson of the results wrth those of a previous cycle or sicular analysts been d~~~ad and sigraficant ddferences explamed?

2. Have apphcable Codes (e.g., ASME Code) and standards teen appropnately referenced and apphed?

3. Is the information from relevant hterature searchas/ background data adequately documented and referenced?

4. Are hand calculabons correct and appropnately M= =**d?

5. is all apphcable computer output and input included?

6. Is all computer software used identdied by name and revision ida*h-?

References

1. Arc all references used to perfwm the analyes listed?

2. Are the references as direct as possible and appropnate to the source?

3. Is the refererix natanon space 6c to the informanen stilaed, erludmg revmon kwl or date of naae, and where appropiale, idersificatson of the locanon of the stormauen at the refereux:c such as page, table or paragraph number?

Independent Reviewer's Statement of Verification Activities:

Independent Reviewer to desenbe details of verification activities beyond the obvious on this checklist including, but not linoted to the review of new methods, use ofsoftware under paragraph 3.3.3, spreadsheet use, assessment of design and methodology changes, engue 3 judgments, the use of previousiv unvenfied inputs, etc.

None ABB Combustion Engineering Nuclear Power

l 1

kkkg glg C-PENG-CALC-018, Rev. 00 Page C5 ofC6 Design Analysis Verification Checklist The Form and Format section of the Checklist below may be completed by a Checker under the direction of the l Independent Reviewer.

l Author IR Form /Fonnat Yes N/A Concur.

1. Is the document legible, reproducible and in a form suitable for filing and retneving as a Quahty Record?

2. Except as permitted by 3.1.3.a. are all pages identified with the document number, meluding revision number?

3. Except as pemuned by 3.1.3.a. do all pages have u uruque page number?

Q

4. 51 ave all changes been =h~"A by the imtials and date of the Quahty Records Controller?

] X

5. Are all files on CD-ROM identified by the path name?

Q Q

6. Are all computer disks identified with the analysis number?

Q Q l

7. Are any unvenfied sections of an otherwue venfied analysis clearty indicated?

g Q For a " Memorandum Revision"to a completed Design Analysis:

Q

l. Have the trtle and document number been preserved without change?

2. Does this revision meet the entana for a " simple revision"?

] ]

3. Are the Author, the Reviewer and Management Approver and their roles identified?

] ]

For a revision to a completed analysis in the " Complete Revision" and "Page Change Package" formats:

1. Where practical, have changes and additions been identified by mechanisms such as veitical imes, etc.?

]

2. Where pracucal, have deletions been identified by i,mlwue such as strike outs etc.?

] ]

3. Have indications of change in previous revisions been removed?

]

4. Does the distribution of the revision include those on the distribution of the previous revision?

]

For a " Complete Revision":

Q

l. Have the title and document number been preserved without change?

]

2. Has the revnion number been Le.. .;ed by one?

]

For a "Page Change Package":

1. Are pages numbered in accordance with the original analysis?

]

2. Are instructions provided for the iroertion and deletion of revisico pages?

] ]

3. Has a new Title Page been prepared?

4. Does the Package Cornerts Page reflect the change package contents?

] ]

Form / Format section completed by the Independent Reviewer.

O Form / Format section completed by the Checker identified below:

Checker Name: Signature:

ABB Combustion Engineering Nuclear Pover

AIIk gg g C-PENG-CALC-018. Rev 00 l Page C6 ofCe l

l Reviewer's Comment Form

Title:

Analysis ofWaterford 3 Hot Leg Sampling MNSA Document Number: C-PENG-CALC-018 Revision Number: 00 Comment Renewer's Comment Response Author's Response Response l Number Required? Accepted?

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ABB Combustion Engineering Nuclear Power

l'

' C-PENG-DR-006 Rev 01 1 Page Cl of C82 j l

ATTACHMENT C 1

l C-PENG-CALC-019, Revisions 00 and 01,

" Analysis of Waterford 3

Hot Leg PDT MNSA"  !

(82 pages including cover) l l

l l

I

Inter Office Correspondence C-PENG-CALC-019 Revision 01 March 22,1999 To: QA File (2)

From: C.L. Mendrala cc: K.H. Haslinger TITLE: ANALYSIS OF WATERFORD 3 HOT LEG PDT MNSA The following typographical error is noted in Revision 00 of the subject calculation.

LOC _ATION CORRECTION NOTE p.34 Reference 6.10 should be Reference 6.11 NONE REVIEW AND APPROVAL This document is verified in accordance with ABB-CE QPM-101 Revision 03. Management authorizes the above editorial changes to the text.

Printed Name Signature Date Cognizant Engineer C.L. Mendrala 8/lmMM J/Jg95 Independent Reviewer S.T. Slowik .A . 3 TJ Management Approval R.O. Doney utt 3 F1 0

ABB Combustion Engineering Nuclear Power

Design Analys'a Title Pcge

Title:

Analysis of Watesfoni 3 Hot 14g PDT MNSA Document Number C-PENG-CALC-019 Revision Number: 00 i 1

1. VertRcanon Status: '

@Cee O NotRequued O Completew/continWAssumptions

2. Apprwalof Conspleted Analysis ThisDesign Analysisiscompleteandvenfied. M==------ tauthorizestheuseofitsresultsandatteststothequahfication of the Canni=nt Engmeer(s), Mentor and Independent Renewer (s).

Painted Name Signaeste Date Cognizant Engmeer(s) C.L. Mendarla ff)(,,]jp g/g '

i, Mentor @ None 2

":,--- :i--t Renewer

(s) S.T. Slowik [M 3[jd3 Ma=====ent Apprwal M Hashnga hl[g hg[g ,

$//j/j$

3. Package Contests (this section may be (-- ; ' --:-j aAer Management apprm31):

i Total page count, i-iding body, appendicies, attachments, etc. 80 List associated CD-ROM disk Volume Numbers and path names: @ Nome CD ROM Volume Numbers Path Names (to kmst directory wtuch uruquely appbes to ttus document)

Other metachments (specify): @ None l l

4. Di.,4h tio.

QA(2) BevBoys l

ABB Combustion Engineering Nuclear Power i l

l

ABB C-PENG-CALC-019 Rev. 00 l Page 2 of 53 RECORD OF REVISIONS Rev Date Pages Changed Prepared By Reviewed By Approved By l 00 03/19/99 Original C.L. Mendrala S.T. Slowik K.H. Haslinger I

l l

l l

)

ABB Cornbustion Engineering Nuclear Power

i ABB C-PENG-CALC-019 Rev. 00 Page 3 Of 53 TABLE OF CONTENTS Smion Pane NO.

1 INTRODUCTION . . . . .. . . . .5 1.1 OBJECTIVE . .

. .5 1.2 ASSESSMENT OF SIGNIFICANT DESIGN CHANGES . .5 2 METHODOLOGY-- .. ..6 2.1 GAP AT NORMAI.OPERAUNG CONDTUONS.. . . . . .6 2.2 DETERMINATION OF IMPACT FORCE.. .8 2.2.1 Net Ejection Force. F,.. . . .

.8 2.2.2 Deflection ofComponents Due to No=le Ejection.. .

.9 2.2.3 Impact Force.. . .

.10 2.3 STRUCTURAL ANA1.YSIS OFMNSA COMPONENT 3.. . . . . .1I 2.3. ) Tie rods.. .. . .

.11 2.3.2 Hex Head Bolts.. . .

. .. )1 2.3.3 Top Plate.. . ..

. .I2 2.3. 4 Compression Collar.. . .12 2.3.5 Upper Flange.. . .I2 3 DESIGN INPUTS. - - - ..... .. 13 3.I SELECTION OF DESIGN INPtJTS.. . . . . . .13 3.1.1 Design and Operating Pressures and Temperatures.. . .. . .13 3.1.2 MNS4 Materials.. . . .

.13 3.1.3 Nonle and Hot LegMaterials.. . . . . .13 3.1. 4 Material Properties.. .. . . . .14 3.1.5 MNSA ComponentDimensions.. .. . .

.16 3.1.6 Nonle andNonle Component Dimensions.. .17 3.2 ASSUMPDONS.. .. ..

.17 3.2.1 Loading Conditions.. . .

. .17 3.2.2 Consideration ofSeismic Loads.. .

.17 3.2.3 Friction Force. . .18 3.2.4 Sealingpressure. .. . . . . .

.19 3.2.3 Preload.. . . .19 3.2.6 Dimensions.. . . .

.I9 4 SIGNIFICANT RESULTS .

. . . . ... .. . .. 20 5 ANALYSIS -

- - . -. - .~--.-..-.....-.21 5.1 MNSA DESCRFDON.. . . .

. . 21 5.2 CONSIDERATION OF IMPACT LOAD., . 23 5.2.1 . Relative Displacements Due to ThermalExpansion. . . .23 5.2.2 Cold Gap Setting vs. Calculated Displacements.. . . 25 3.2.3 Determination offmpact Force.. . . 26 5.2.4 Impact Force.. .

. 29 5.3 STRESSES !N THE PDT MNSA COMPONENTS. . . . 29 3.3.1 Tie Rod Stresses... .

. . 30 i 5.3. 2 Hex Head Bolt Stresses.. .. . . . . . 36 5.3.3 Top Plate Stresses.. . . . . . . . .

. 43 ABB Combustion Engineering Nuclear Power

AB C-PENG-CALC-019 Rev. 00 Page 4 Of 53 5.34 Compression Collar Stresses.. . 44 5.3.5 Upper Flange Stresses.. . . . .45 5.4 FATIGUE ANALYSIS.. .

.46

5. 4.1 Normal Operating Pressure Force.. .

. 46

3. 4. 2 Tie Rod Fatigue... . .. .

. 47

5. 4.3 Hex HeadBolt Fatigue.. .. . .

. 49 5.5 CONSIDERATION OF HYDROSTATIC TEST PRESSURE CONDmONS . . 51 5.6 CONSIDERATION OF FAULTED CONDITIONS . . . . 51 6- RE FERENCES . .. .. . . ..... . .. .. ... .. .... ..... 52 LIST OF FIGURES FIGURE DESCRIPTION PAGE 1 Hot Leg PDT MNSA. . . . . . .. . .. . 22 LIST OF APPENDICES NO. Pages APPENDIX A: CALCULATION OF TIE ROD TEMPERATURE . . . . . .2 l APPENDIX B: ANSYS EVALUATION OF TOP PLATE . . . . . 17 APPENDIX C: HOT LEG PDT MNSA ASSEMBLY DRAWING . .. .2  ;

APPENDIX D: QUALITY ASSURANCE FORMS .. . .6 i l

ABB Combustion Engineering Nuclear Power l

A BB C.PENG-CALC-019 Rev. 00 Page 5 of 53 1 INTRODUCTION 1.1 Objective The objective of this calculation is to analyze the Mechanical Nozzle Seal Assembly (MNS A) to be installed on the Hot Leg (PDT) nozzle at the Waterford Unit 3.

I The MNSA is a mechanical device that acts as a complete replacement of the "J" weld between an Inconel 600 instrument nozzle and the Hot Leg pipe. Its function is to prevent leakage and restrain the nozzle from ejecting in the event of a through-wall crack or weld failure of a nozzle. The j

potential for these events exists due to Primary Water Stress Corrosion Cracking.

This calculation will be incorporated into a Design Stress Report for the RCS Piping.

1.2 Assessment ofSignificant DesigM Changes This report presents the detailed structural and thermal analyses required to substantiate the adequacy of the design of th: Waterford Unit 3 Mechanical Nozzle Seal Assembly as a replacement of the nozzle "J" weld. This analytical work encompasses the requirements set forth in Reference 6.1 and is perfonned in accordance with the requirements of the ABB CENO Quality Procedures Manual QPM-101 (Reference 6.2).

Addenda to the original Piping Design Repou (Reference 6.3) were reviewed and it was determined that their results have no impact on the current analysis and also that the current analysis does not impact their results. These Addenda Reports inclucc CENC-1460 (2/81)

C-MECH-DR-001 (12/93)

C-MECH-DR-004 (12/93)

ABB Combustion Engineering Nuclear Power

C-PENG-CALC-019 Rev. 00 Page 6 of 53 2 METHODOLOGY The objective of this calculation is to analyzed the Mechanical Nozzle Seal Assembly (MNS A) to be installed on the Hot Leg PDT nozzle at the Waterford Unit 3. The methodology used in this calculation is based on the method developed in Reference 6.8, for a similar design.

2.1 Gap at Normal Operating Conditions A cold gap between the valve and the top plate of has to be established to account for the relative thermal expansions of the components (see representative drawing below). l I

I l

, \

N 4 I i

\/

/lm veo '

The magnitude of the impact of the valve against the top plate is dependent upon this gap - the larger the gap, the greater the work done by the internal pressure, the greater the deflection of the components, and the greater the load on the components. l In order to determine the load impacting the MNSA components, if the nozzle ejects from the hot leg, the gap between the valve and top plate with the components at normal operating temperatures shall be determined. The thermal expansion displacements of each of the relevant components shall first be determined and then added to or subtracted from the cold gap setting to determine the final operating conditions gap.

ABB Combustion Engineering Nuclear Power

1 ABB C-PENG-CALC-019 Rev. 00 i Page 7 of 53 l

The linear thermal expansion displacements of each of the relevant components is calculated using i the following equation from Reference 6.14 (p. 53):

l l 6 = a L AT where:

6= the displacement (deformation) of a component caused by linear thermal expansion et = the coefficient oflinear thermal expansion L= the length of the component AT = the temperature change from a reference temperature of 70*F to the applicable operating temperature The relative displacement,6, is determined by adding or subtracting the displacement of each of the individual components, as follows:

6, = Soos. + 6 r. .a +6%. + Sa. - Sc. ,a - Sm ,*

I

(* Smo, in this analysis is taken to mean the combined thermal displacement of the lower and upper flanges) l I

For determining the maximum relative displacement, it is assumed that the temperature of the tie rod and flanges increases from a reference temperature of 70 F to the ambient temperature of 120 F, and that the nozzle reaches the normal operating temperature of 611 F. Because of respective component l lengths and coefficients of thermal expansion, these conditions produce the maximum relative

.!isplacement (6, ) between the nozzle and the MNSA top plate, such that the overall nozzle displacement exceeds the displacement of the top plate by a maximum amount. If there was no cold gap, the valve would (theoretically) extend beyond the inboard surface of the top plate by a distance of 6,..

The value of 6, sets an upper limit on the cold gap setting, though the extreme temperature j differences evaluated above would not be seen during plant operations. ,

1 ABB Combustion Engineering Nuclear Power

ABB C-PENG-CALC-019 Rev. 00 Page 8 of 53 For determining the normal operation relative displacement, a more reasonable temperature distribution is selected, based on more realistic and appropriately conservative conditions. These conditions produce the normal operations relative displacement (6,,) between the end of the valve and the MNSA top plate after heatup cf the plant. If there were no gap between the end of the valve ar d the top plate at cold conditions, the valve would (theoretically) extend beyond the inboard surface of the top plate by a distance of 5,,.

A final gap for the normal operating conditions will be determined by subtracting the value of normal operations relative displacement (6%) from the maximum cold gap setting.

2.2 Determination ofImpact Force At the moment at which an instantaneous break occurs, the internal fluid pressure in the hot leg will eject the nozzle outward, with the nozzle impacting the top plate. In order to determine the stress effects of this impact on the top plate, it will be necessary to first determine the net ejection force  ;

acting on the nozzle. Once this net force is known, a relation can be defined between the work performed by the ejection force and potential energy stored in the deflection of the affected components (at the point of maximum deflection). Once the deflection of the components is known, the impact force can be calculated and then used to determine stress effects. '

2.2.1 Net Ejection Force, F.

The nozzle will be forced out of the hot leg by the internal fluid pressure; this outward motion will be opposed by the friction force which the Grafoil Seal exerts on the external surface of the nozzle.

The net ejection force acting on the nozzle, F., is the difference between the " pressure force", F,, and the seal friction force, Fr:

F. = F, - Fr 2.2.1.1 Force Due toInternalPressure Motion of the nozzle at the moment at which there is an instantaneous break is due to the force created by internal pressure pushing against the entire cross section of the nozzle. This force, F,, is determined as follows:

i F, = p A l

ABB Combustion Engineering Nuclear Power

AB C-PENG-CALC-019 Rev. 00 Page 9 of 53 where: 1 p= design pressure l

A= pressure area 2.2.1.2 FrictionForce The determination of the friction force (Fr) provided by the Grafoil seal is made based upon the coefficient of friction for the seal against the nozzle and the radial load provided by the seal against the nozzle (produced by the compression of the seal).

1 Fr = P A i

where:

P= radial sealload (pressure) p= coeSicient of friction A= surface area of the seal in contact with the nozzle surface 2.2.2 Deflection of Components Due to Nozzle Ejection The total deflection of the impacted components due to the ejection of the nozzle will be determined based upon the conservative understanding that all of the work put into the system by the net ejection force is converted completely into the potential energy of the deflected components (i.e.,

there are no losses). The base equation for evaluating the total deflection is derived from Equation (a) on page 471 of Reference 6.14 and is presented, as follows:

F, s = 1 K ,Ar2 where:

F. = the net ejection force s= total distance traveled by nozzle Ax = total deflection of MNSA tie rods and top plate, and K, = the equivalent stiffness of MNSA tie rods and top plate The total distance traveled by the nozzle, s, is equal to the distance of the gap at-temperature

(" Gap") plus the total deflection of the tie rods and top plate (Ax), or s= Gap +Ax.

The base equation may be re-written as follows:

ABB Combustion Engineering Nuclear Power

ABB C-PENG-CALC-019 Rev. 00 Page 10 of 53 F, (Gap + &) = 1 X,, &*

2 1

=> - K,, & * - F, & - F, Gap = 0 In order to determine Ax, the stiffnesses of the tie rods and of the top plate are first calculated.

Because the top plate has a section cut out to acconunodate the valve, the standard plate solutions are not appropriate. Therefore, the stiffness of this component is calculated using ANSYS. Then, since the tie rods and top plate stiffness act in series against the impact load, the equivalent stiffness of tie rod-top plate system is calculated based upon a series stiffness equation from Reference 6.15 (p. 702):

1 K,= 3

+-

Ka K,, p The base equation developed previously is used to determine Ax:

Ax = [F, (F,2 + 2K, F. Gap)"] / Ky 2.2.3 Impact Force The impact force, F%, on the top plate and tie rods is then calculated:

F ,,,,,,, = K ,, &

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ABB C-PENG-CALC-019 Rev. 00 Page11of53 l

2.3 Structural Analysis ofMNSA Components l l

The stresses in the MNS A components are examined considering, pressure, thermal loads and the impact force defined above.

2.3.1 Tie rods The Design tie rod stresses are considered in accordance with NB-3231 and Appendix E.

The maximum tie rod load is determined by comparing the load created by preload and thermal '

expansion with the impact load. Whichever is greater is used throughout the remaining calculation as the tie rod loading. l

• The average and maximum stress in the tie rod and shear stresses in the threads are then calculated and compared to the corresponding ASME Code allowables.

A fatigue evaluation of the component is then performed 2.3.2 Hex Bead Bolts '

The Design bolt load stresses are considered in accordance with NB-3231 and Appendix E.

• Maximum Bolt Load Due to the flexibility in the design of the flanged connection between the MNSA and the Hot Leg, the impact from ejection of the nozzle will increase the load on the bolts. The stiffness of the flange relative to the stiffness of the bolts will determine what percentage of the impact load will be effectively transmitted to the bolts.

The total load on the bolt can be expressed by the following equation derived from Reference 6.13 (p. 579):

y" F'"" = Preload + F""""

<Km+Kw In the above expression, K% s considered i to be the equivalent stiffness of the components which are put in compression due to the torquing / tightening of the hex head bolts; these components include the upper flange (top and bottom pieces which are considered to act in parallel with each other), and the compression collar, which acts in series with the upper flange.

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l C-PENG-CALC-019 Rev. 00 l Page 12 of 53 l

l (The consideration of relationships between the top and bottom pieces of the upper flange l l represents a condition between two extremes that could be assumed for the components. The first, is that they act together as one solid piece. This would result in an unrealistically high stiffness. The second, is that they are both simply supported rings which act in series. Since l

the top ring is supported across the entire bottom surface, this would result in an i unrealistically low stiffness. The assumption that the two rings act in parallel provides a stiffness that is between these two extremes and is concluded to be reasonable).

It may be concluded (from the above equation) that the greater the stiffness of the bolts as compared to the stiffness of the flange components, the greater the increase in load on the bolts from the impact (i.e., as Kn ,/lb -+ 0, the multiplier for F4a -+ 1).

The stiffness of each component is considered in the detailed analysis to calculate the

! maximum hex head bolt load.

• The average and maximum stress in the bolt and shear stresses in the threads are then calculated l and compared to the corresponding ASME Code allowables. 1

! A fatigue analysis of the component is then performed.

i 2.3.3 Top Plate The shear and bending stresses in the top plate are calculated due to the impact load and compared to the corresponding ASME Code allowables. Because the shape of the top plate is not uniform, the i ANSYS model generated to calculate the stiffness of the component is also used to obtain the stress.

2.3.4 Compression Collar The shear stress and bearing stress due to the preload of the hex head bolts are calculated and compared to the corresponding ASME Code allowables.

2.3.5 Upper Flange l The shear stress due to the preload of the hex head bolts is calculated and compared to the corresponding ASME Code allowables.

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1 C-PENG-CALC-019 Rev. 00 l Page 13 of 53 l

3 DESIGN INPUTS l

3.1 Selection ofDesign Inputs 3.1.1 Design and Operating Pressures and Temperatures l

\

l The Mechanical Nozzle Seal Assembly is considered a pressure-retaining component. The Design Pressure is 2500 psia and Design Temperature is 650 F. The plant operating pressure and temperature are 2250 psia and 611*F, respectively (Reference 6.17). Ambient design temperature is 120 F (Reference 6.4).

3.1.2 MNSA Materials I MNSA materials are taken from Reference 6.5.

Itsm Material Compression Collar SA-479, Type 304 Lower Flange SA-479, Type 304 Upper Flange SA-479, Type 304 Top Plate SA-479, Type 304 Hex Bolt SA-453, Grade 660 '

Hex Nut S A-453, Grade 660 Tie Rod SA-453, Grade 660 3.1.3 Nozzle and Hot Leg Materials Hot Leg PDT nozzle and fitting materials are taken from the references indicated below.

lism Material Reference Nozzle Neck SB-166 6.3.1 Safe End SA-182, Type 316 6.3.1 Nipple SA-376, Type 304 6.4.1 Valve SA-182, Type 316 6.18.1 l

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C-PENG-CALC-019 Rev. 00 Page 14 of 53 3.1.4 Material Properties Material properties used in this analysis include coefficients of thermal expansion (a), moduli cf elasticity (E), design stress intensity values (S,,,) and Yield Strength Values (S,). These properties are presented below and are found in the Appendices of Reference 6.6. All tables noted below are i relative to Reference 6.6.

l 3.1.4.1 Coefficient ofLinear ThermalExpansion, a The following table presents the temperature-dependent coefficients oflinear thermal expansion for various materials:

temperature SB-166 SA-479 Type 304 316 SS SA-453, Grada CT

(*F) (Alloy 600) (304 SS) (/Zufo60) 100 6.90 8.55 8.54 8.24 200 7.20 8.79 8.76 8.39 300 7.40 9.00 8.97 8.54 400 7.57 9.19 9.21 8.69 500 7.70 9.37 9.42 8.82 600 7.82 9.53 9.60 8.94 611 7.83* 9.55* 9.62* 8.95*

650 7.88 9.61 9.69 9.00

• by interpolation All coefficients are Coefficient B values from Table I-5.0, where Coefficient B is the mean coefficient 4

of thermal expansion X 10 inlin./*F in going from 70*F to the indicated temperature.

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9 C-PENG-CALC-019 Rev. 00 Page 15 of 53 3.1.4.2 Modulus ofElasticity, E The following table presents the temperature-dependent moduli of elasticity for S A-479 Type 304 ,

and SA-453, Grade 660: l temperature

@)

70 28.3 200 27.6 300 27.0 I 400 26.5 500 25.8 600 25.3 650 25.0*  ;

700 24.8

• by interpolation l All moduli of elasticity values are from Table I-6.0, where E = value given X 10' psi.

3.1.4.3 Design StressIntensity Value, S.

The following table presents the temperature-dependent design stress intensity values for various materials:

temperature SA-479 304 S A-453, Grade 660

@)

100 20.0 28.3 200 20.0 27.6 300 20.0 27.3 400 18.7 27.2 500 17.5 27.1 600 16.4 27.0 650 16.2 26.9*

700 16.0 26.8 l

• by interpolation a

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l 6 C-PENG-CALC-019 Rev. 00 Page 16 of 53 The design stress intensity values for SA-479 Type 304 are from Table I-1.2; and the design stress inte ,*/ values for SA-453, Grade 660 are from Table I-1.3. All S., values are given in ksi.

3.1.4. 4 YieldStrength Value, Sy The following table presents the temperature-dependent yield strength values for SA-479 Type 304:

temperature SA-479 304

( F) 100 30.0 200 25.0 300 22.5 400 20.7 4

500 19.4 600 18.2 650 17.9 700 17.7 The yield strength values for SA-479 Type 304 are from Table I-2.2. All Sy values are given in ksi.

3.1.5 MNSA Component Dimensions The bolts and tie rods have the following dimensions (References 6.5, and 6.11):

. Bolts Tie Rods

[0.500-20 UNF-2A] [0.375-16 UNC-2A]

Basic major diameter 0.5000 in 0.3750 in Basic minor diameter 0 4374 in 0.297 in Basic pitch diameter 0.4675 in 0.3344 in 2

Tennile stress area J.1599 in 0.0775 in Kn max (max minor 0.457 in 0.321 in diam. ofinternal thread)

Es min (min pitch diam. 0.4619 in 0.3287 in of external thread)

En max (max pitch diam. 0.4731 in 0.3401 in ofinternal thread) -

Ds min (min major diam. 0.4906 in 0.3643 i'n of external thread)

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ABB C-PENG-CALC-019 Rev. 00 Page 17 of 53 ;

I 3.1.6 Nozzle and Nozzle Component Dimensions Various ccmponents dimensions are taken from the references indicated below:

PDT Nozzle Ref.

Pressure Diameter 1.00 in 6.18.2 Length of Safe End 3.00 in 6.I8.2 Length ofNozzle 1.75 in 6.18.2 j Length of Nipple 3.50 in 6.18.2 Length of Valve 5.188 in 6.18.2 2 2 Pressure Area = (x r ) 0.785 in _

l 3.2 Assumptions 3.2.1 Loading Conditions If no crack is present, it is assumed that, except for preload and thermal expansion, the MNSA components are not loaded during normal operating conditions. An impact load may be experienced if there is a complete and instantaneous failure in the J-weld or a 360* circumferential crack in the nozzle, such that the nozzle would be forced outward against the top plate, closing any gap between the two components. After this event occurs, a normal operating load, without impact, would exist, with the internal pressure holding the nozzle up against the top plate; this load would be cyclical -

from essentially zero at Cold Shutdown to a maximum at normal operating conditions.

For the purposes of this analysis, it is assumed that there is a complete and instantaneous failure of the J-weld (or a 360* circumferential crack in the nozzle) such that the nozzle is ejected outward and impacts against the top plate, which will also then load the tie rods and other components. The impact of the nozzle against the top plate conservatively represents the maximum load that the restraining components would experience.

3.2.2 Consideration of Seismic Loads Because of the nature of the accelerations from seismic events, only the tie rods will be evaluated for the stress effects of the seismic event. The remaining MNS A components will not be significantly affected. Separate seismic tests on similar MNSA configuration were performed to demonstrate an adequate seal performance (see Reference 6.19).

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ABB C-PENG-CALC-019 Rev. 00 Page 18 of 53 3.2.3 Friction Force .

The effects of any impact of the nozzle against the top plate are dependent upon certain assumptions regarding the determination of the ejection force acting on the nozzle.

In an " ideal" (and worst case) break scenario, the crack would be complete, instantaneous and oriented such that no base or weld metal could interfere with the motion of the nozzle. In this case, i

the only resistance offered to the nozzle motion would be provided by the attached piping and by the Grafoil seal.

In reality, the crack characteristics necessary for the " ideal" scenario would not exist, and, instead, e

there would be potentially significant resistance offered to the motion of the nozzle by the crack

surfaces and by integral material, if motion would be allowed at all.

In this analysis, a scenario which is somewhere "between" the " ideal" scenario and the "real" scenario will be evaluated: it will be conservatively assumed that motion will be allowed but in the presence of an opposing force provided by the crack metal and by the Grafoil seal. This opposing force will be accounted for by applying a coefficient of friction for the Grafoil-to-nozzle contact, as described below:

A coefficient of friction ( ) of 0.30 for Grafoil-to-nozzle contact will be used to determine the force which opposes motion of the nozzle. This value of 0.30 used for the (kinetic) motion of the nozzle ejection is higher than the values provided by the Grafoil seal manufacturer in Reference 6.12, which lists (static) coefficients in the range of 0.05 to 0.20 (see Table III of Reference 6.12). However, the application of a friction force based upon the coefficient value of 0.3 will be maintained on the basis that the actual force which would tend to limit or prevent motion in the "real" scenario would be higher.

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3.2.4 Sealing pressure Compression of the Grafoil creates a radial pressure against 'the nozzle surface of at least 3,100 psi for a preload of 30 fl-lb, based upon Reference 6.21. (This value will be used to determine a friction force on the nozzle from the Grafoil seal.)

3.2.5 Preload Nominal values of tie rod / bolt preload are used in this analysis since maximum values of preload will not significantly increase corresponding preload stresses. (A check of the results indicate that use of the maximum preload values will result in stresses which will remain below, or will be on the order of, their respective allowables. Therefore, use of the nominal values is acceptable).

3.2.6 Dimensions Referenced overall length of the assembly (13.438 inches) from Reference 6.5.1 was used in this report. Field data measurements of the parts, Reference 6.18, were used for the calculations of the relative displacement and for the stress analysis, except when noted. The use of the dimensions from design drawings may produce slightly different values but do not effect the results of the current analysis.

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A BB C-PENG-CALC-019 Rev. 00 Page 20 of 53 4 SIGNIFICANT RESULTS The results presented below were detennined using the assumptions defined and justified in Section 3.0. These results were compared to those of a similar design in Reference 6.8 and found to be reasonable. There are no additional contingencies or assumptions that are applicable to these results.

Results of this analysis are summarized below in the table below:

Component Stress Calculated Stress / Allowable Stress /

Category Usage factor Usage Factor (stress in ksi) (stress in ksi)

Tie Rod Design 8.58 26.9 Average 30.0 53.8 Maximum 37.4 80.7 Thread Shear 8.1 16.14 Usage Factor 0.424 1.00 Hex Head Bolt Design 22.51 28.3 Average 35) 53.8 Maximum 42.1 80.7 Thread Shear 11.0 16.14 Usage Factor 0.50 1.00 Top Plate Shear 2.64 9.72 Bending 22.4 24.3 Compression Collar Shear 9.26 9.72 Bearing 15.1 17.9 Upper Flange Thread Shear 5.60 9.72 Shear 6.98 9.72 Lengths of engagement used in analysis:

. Tie Rod - Upper Flange: 0.5 in (0.29 in. minimum) e Hex Head Bolt - Hot Leg Pipe: 0.5 in (based upon bolt thread shear; 0.34 in. minimum)

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ABB C-PENG-CALC-019 Rev. 00 Page 21 of 53 5 ANALYSIS 5.1 MNSA Description l

The MNS A is a mechanical device that acts as a complete replacement of the "J" weld between an Inconel 600 instrument nozzle and the hot leg pipe, Figure 1. It replaces the sealing function of the weld using a Grafoil seal compressed at the nozzle outside diameter to the outer hot leg surface. The MNSA also replaces the weld structurally by means of threaded fasteners engaged in tapped holes in the outer hot leg surface, and a restraining plate held in place by threaded tie rods. This feature  ;

prevents the nozzle from ejecting from the hot leg, should the "J" weld fail or the nozzle develop a circumferential crack.

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C.PENG. CALC- ge 0 753 FIGURE 1 Hot Leg PDT MNSA

/ f r,

'l b^

i r/ '

\

/ t, fQ in

,, i ! N l .d ll 3 12 io

@ ~~

-- r -

K /

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C-PENG-CALC-019 Rev. 00 Page 23 of 53 L2 Consideration ofImpact Load 5.2.1 Relative Displacements Due to Thermal Expansion i

According to Section 2, for detennining the maximum relative displacement, it is assumed that the temperature of the tie rod and flanges increases from a reference temperature of 70 F to the ambient temperature of 120 F, and that the nozzle and valve reach the normal operating temperature of 611 F. Because of respective component lengths and coefficients of thermal expansion, these ,

conditions produce the maximum relative displacement (6, ) between the valve and the MNSA top  !

plate, such that the overall nozzle / valve displacement exceeds the displacement of the top plate by a l maximum amount. If there were no cc'1 gap, the valve would (theoretically) extend beyond the inboard surface of the top plate by a distance of 6, .

Hot Len PDT MNSA Maximum Relative Disolacement Component Tenperature a L AT S 4

(*F) (10 inlinl*F) 'a.) ("F) (in.)

nozzle 611 7.83 1.75 541 0.0074 safe end 611 9.62 3 541 0.0156 nipple 611 9.55 3.5 541 0.0181 valve 611 9.62 5.188 541 0.0270 tie rod (lower)* 120 8.27 6.25 50 -0.0026 tie rod (upper)* 120 8.27 5.188 50 -0.0021 flange 120 8.60 2 50 -0.0009 6, = 0.0625

• the tie rod length is divided into two lengths - one up to the inboard valve face, one equal to the valve length. The overall tie rod length is based upon the overall assembly length less.the thickness of the lower flange and the thicknesses of the two upper flange pieces.

The value of 6, sets an upper limit on the cold gap setting, though the extreme Design temperature differences evaluated above would not be seen during plant operatioris.

For determining the normal operations relative displacement, a more reasonable temperature distribution is selected, based on more realistic and appropriately conservative conditions. In this case, the following temperatures are used:

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C-PENG-CALC-019 Rev. 00 Page 24 of 53 Component Operating Temperature 1

(*F) l nozzle neck / safe end 611 )

nipple 550 I vdve 350 tie rod (inside) 500 tie rod (outside) 350 flange 550 l

The temperatures used in the determination of the normal operations relative displacement are average component values and are based upon the following: I

• Nozzle Neck / Safe End - 61l'F: due to the nozzle location in the hot leg pipe, the average temperature of the nozzle (external to the pipe) would be at or very near the hot leg temperatuc (611*F).

• Nipple - 550*F: this component would receive some heat conducted through the nozzle. Some heat would be lost during conduction but, overall, the nipple average temperature would be elevated to a relatively high temperature, given its direct attachment to these hotter components.

l

• Upper Flange - 550*F: this component would receive some heat conducted through the pipe, seal retainer and lower flange, as well as heat by convection / radiation through gaps. Some heat ,

would be lost to the insulation ambient though, overall, the flange average temperature would be I I

elevated to a relatively high temperature, given its proximity to these hotter components. A high temperature for this component is conservative as it will increase the impact gap.

• Tie rod -inside - 500 F this component wo . receive some heat conducted through the center portion of the upper flange, with some heat lost to insulation ambient. A high temperature for ,

this component is conservative as it will increase the impact gap. l l

l

• Tie rod - outside - 350*F: this component would receive some heat conducted from the 6.25 inch portion of the tie rod inside of the insulation. Appendix A uses a fin equation to estimate the i average temperature of the rod assuming a 500 F heat source and heat lost to ambient at 120 F. l The estimated average temperature is 275 F. However, since a high temperature for this component is conservative as it will increase the impact gap, 350 F is used.

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l

1 C-l'ENG-CALC-019 Rev. 00 Page 25 of 53 i e

Valve - 350*F: this component would receive heat conductes ei,%ej from the nipple and nozzle fluid, with some heat lost to containment ambient at 120*I . It is reasonable that this component would be slightly hotter than the estimated 300 F temperature of the outer (upper) tie rods.

Hot Lea PDT MNSA NOP Relative Displaceme.nl Component temperature a L AT S 4

(*F) (10 in./in./*F) (in.) (*F) (in.)

nozzle neck 611 7.83 1.75 541 0.0074 safe end 611 9.62 3 541 0.0156 i Nipple 550 9.45 3.5 480 0.0159 Valve 350 9.09 5.188 280 0.0132 tie rod (lower) 500 8.82 6.25 430 -0.0237 tie rod (upper) 350 8.62 5.188 280 -0.0125 Flangn 550 9.45 2 480 -0.0091  !

6, ,, = 0.0068 These conditions produce the normal operations relative displacement (6, ,,,) between the end of the valve for the PDT Nozzle and the MNSA top plate after heatup of the plant. If there were no gap l

between the end of the nozzle and the top plate at cold conditions, the valve would (theoretically)  ;

extend beyond the inboard surface of the top plate by a distance of S w. I 5.2.2 Cold Gap Setting vs. Calculated Displacements It is recognized that the minimum cold gap of 0.015 in. is less than the maximum relative displacement for the Hot Leg PDT MNS A, 0.0625 in. If this occurred, the nozzle would be in direct 3 contact with the top plate. However, as noted before, the conditions used to obtain the maximum l

- _ relative displacement are not anticipated during operation.

The maximum cold gap setting of 0.025 in. indicates that a gap of 0.025 in. - 0.0068 in. = 0.0182 in.

can exist during normal operating conditions for the long Hot Leg PDT MNSA. A gap value of j 0.019 inches for normal operating conditions will be used in the subsequent determination of the (

impact force.

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D'1l0EF C-PENG-CALC-019 Rev. 00 i l

Page 26 of 53 5.2.3 Determination ofImpact Force Impact force is calculated in accordance with the methodology described in Section 2.

5.2.3.1 Net Ejection Force, F.

i The nozzle will be forced out of the hot leg by the internal fluid pressure; this outward motion will be opposed by the friction force which the Grafoil Seal exerts on the external surface of the nozzle.

The net ejection force acting on the nozzle, F., is the difference between the " pressure force", F,, and the seal friction force, Fr:

F. = F, - Fr 5.2.3.1.1 Force Due to Internal Design Pressure Motion of the nozzle at the moment at which there is an instantaneous break is due to the force created by internal pressure pushing against the entire cross section of the nozzle. From Section 2

3.1.6, the pressure area of the Hot Leg PDT nozzle is 0.785 in . This force, F,, is determined as follows:

2 F, = (2500 psi) (0.785 in ) = 1,963 lb 5.2.3.1.2 Friction Force The determination of the friction force (Fr) provided by the Grafoil seal is made based upon the coefficient of friction for the seal against the nozzle and the radial load provided by the seal against the nozzle (produced by the compression of the seal).

Fr = P A where:

P =' radial seal load (pressure) = 3100 psi (Reference 6.21)

= coefficient of friction = 0.3 A= surface area of the seal in contact with the nozzle surface

= xDh h= 0.25 inch D= 1.0 inch (Grafoil seal ID is defined as 0.993 inch, but OD of nozzle, sealing surface, is 1.0 inch)

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C-PENG-CALC-019 Rev. 00 Page 27 of 53 Therefore:

Fr = (3100 psi) (0.3) n (1.0 in) (0.25 in) = 730 lb Based upon the forces calculated above, the net ejection force is: i l

F. = 1963 - 730 = 1233 lb 5.2.3.2 Defection of Components Due to No::le Ejection The re-written base equation for evaluating the total deflection is 1

-K y Ar - F, Ar - F, Gap = 0 2

In order to determine Ax, the stiffnesses of the tie rods and of the top plate are first calculated and then the equivalent stiffness of the tie rods-top plate system is calculated.

5.2.3.2.1 Stiffness of 4 Tie Rods The total stiffness of the four (4) tie rods, Ka, is based upon an equation from Reference 6.14 (p.

31):

K a-4^-

y where:

2 A= 0.0878 in cross-sectional area of the tie rod, based on the basic pitch dbneter E= 26.8 x 10' psi (at 350*F) 1= length between top plate and upper flange = 13.438 - 2.00 = 11.438 in So:

(0.0878 in')(26.8X10' lb )

, K =4 '"

= 8.23E5 Ib-i l1.438 in in ABB Combustion Engineering Nuclear Power l

ABB C-PENG-CALC-019 Rev. 00 Page 28 of 53 5.2.3.2.2 Stiffness of the Top Plate The ANSYS finite element analysis code (Reference 6.9) was used to determine the stiffness of the top plate, which has an irregular shape. A half-symmetry model of the plate is generated using the SHELL93 type element. The model was restrained at the locations of the tie rods in all directions, with symmetry boundary conditions applied to those nodes on the plane of symmetry, Figure B1 of Appendix B.

To assess the effects of the impact load, a distributed load of 500 lb was applied at an outer diameter edge where the valve contacts the plate - at a radius of 1.25 inch. This radius is slightly lower than the 1.259 inch diameter from the as-measured data shown in Reference 6.18, but will not affect the results. (Due to the model symmetry, the applied 500 lb load is equivalent to applying 1000 lb. for a full model.) The maximum deflection of the top plate was 0.000398 in.

The d*iffness is determined as follows:

K %% = F / d = 1000 lbs / 0.000398 in = 2.51 E6 lb/in The output file from the ANSYS evaluation is presented in Appendix B.

i 1

5.2.3.2.3 Equivalent Stiffness: Tie Rods-Top Plate System 1

The tie rods and top plate stiffness act in series against the impact load. The equivalent stiffness of tie rod-top plate system is based upon a series stiffness equation from Reference 6.15 (p. 702):

K, =

+ 1 Ka K, p 1

619,900 lb K" l 1 in l

l

+

8.23E5 2.51E6 5.2.3.3 TotalDeflection, At 1

The base equation developed previously is used to determine Ax:

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ABB C-PENG-CALC-019 Rev. 00 Page 29 of 53 ;

}

1

-K,, & * - F, & - F, Gap = 0 l 1

=> Ax = [F. (F.' + 2Ky F. Gap)"'] / K.,

i Given: I F. = 1233 lbs l K., = 619,900 lb / in Gap = .019 in. ,

l

=> Ax = 0.0109 in.

5.2.4 Impact Force l

The impact force, F,a, on the top plate and tie rods is then calculated:

l F,, , = K ,, & '

l Therefore:

F,,, = 619,900 h(0.0109 in) = 6.8 kips in The following value of the impact force will be used in subsequent analysis of the MNSA:

F,a = 7,000 lb 5.3 Stresses in the PDTMNSA Components The Design Loads for the various MNSA components will be a function of either bolt preload, the impact load, and/or thermal expansion loads, depending upon the effects of the source load upon a panicular component.

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A BB C-PENG-CALC-019 Rev. 00 Page 30 of 53 5.3.1 Tie Rod Stresses 5.3.1.1 Design Bolt Load Stresses (NB-3231 and Appendix E)

The design bolt load for the tie rod is considered to be the hydrostatic load which results from Design Pressure only, since the tie rod is not used for gasket-joint purposes.

Section 5.4.2.1 determines the servic. stress in the tie rod for a pressure which bounds the Design Pressure; this stress value is compared to the design bolt load stress allowable:

8.58 ksi < 26.9 ksi (S. at 650*F)

The hydrostatic load stress is below the stress allowable, which indicates that the actual bolt area (AS) exceeds the minimum required bolt area (A.).  ;

5.3.1.2 StressDue to ThermalExpansion The differential thermal expansion between the tie rod and the upper flange (or top plate) will create l l an additional tensile load en the tie rod. For the analysis of the tie rod, this additional load is assumed to be completely taken up by deformation of the tie rod. The stress effect of this differential thermal expansion is determined below. -

From Reference 6.14:

ot.. = E ot AT 1

The a to be used is the differential in coefficients for the materials of the tie rod and flanges, Aa. Therefore:

at.. = E ha AT t

I Given:

E = 25.25 E6 psi at 611*F Aa = otn - %,a = (9.55 E 8.95 E-6) = 0.6 E-6 (et at 61l'F)

AT = 611 - 70 = 541 "F ct.. = 8.2 ksi ABB Combustion Engineering Nuclear Power

AIB C-PENG-CALC-019 Rev. 00 Page 31 of 53 l

The effective tensile force due to this thermal expansion is determined by:  !

2 P=ca A, = (8.2 ksi)(0.0775 in ) = 0.636 kips I 1

5.3.1.3 Preload 1

l The tie rod and nuts are being preloaded to 75 in-lbs. To determine the load in each tie rod, the following equation is used (Reference 6.10):

i T = 0.2 F d f

o F=T/0.2d l

Given: '

l T = the applied torque = 75 in-lbs d = is the nominal major tie rod diameter = 0.375 in.

F = (75 in-lbs) / (0.20) (0.375 in) = 1.00 kips.

i 5.3.1.4 Impact Load The results of the ANSYS analysis of the top plate indicate that the reaction loads at the tie rod locations are not equal and that the load on one pair of rods may be approximately twice as high as the load on the other pair, see output in Appendix B.

l For the purposes of analyzing stress in the tie rod (and upper flange threads), the impact load on the I tie rod will be taken to be 1.33 times the average tie rod load, or: l 1.33 (7.0/4) = 2.33 kips 5.3.1.5 Maximum Tie RodLoad l The load on the tie rod will be the greater of the load due to preload and thermal expansion and the load due to the impact:

Preload + thermal expansion = 1.00 + 0.636 = 1.636 kips Impact = 2.33 kips (> 1.636 kips)

Therefore, the maximum tie rod load is 2.33 kips.

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ABB C-PENG-CALC-019 Rev. 00 Page 32 of 53 l

5.3.1.6 Average Stress,0;(NB-3232.1)

The average (axial) stress (ci) in the tie rod is due to the maximum tie rod load.

ci = P/A, A = 0.0775 in2 i

P = 2.33 kips

)

i 2

ci = (2.33 kips /0.0775 in ) = 30.0 ksi < 2 Sm = 53.8 ksi(S. at 650 F) l 1

5.3.1.7 Maximum Stress (NB-3232.2) l l

l The maximum stress in the tie rod is essentially a stress intensity due to a combination of the average l stress, bending stress from the OBE (Design Earthquake) event, and the torsional shear stress due to  ;

residual torque (from preload).

l 1

5.3.1.7.1 Seismic Bending Stress I l

Prior to any (complete) weld failure, a seismic event will cause accelerations of the MNS A. Most components will experience little or very little effects from these seismic accelerations. However, l because of motions associated with the top plate, the inboard end of each tie rod will be subjected to l bending stress. This bending stress (c ) will be conservatively added to the average and torsional i shear stress for determining the maximum stress in the tie rod.

The bending stress at each of the tie rods is determined by applying the maximum acceleration l occurring at OBE event to the top plate, following with the even distribution of the resulting force to each of 4 tie rods, and then calculating the stress at the tie rod inboard end. 1 on = 1/4 Mc/I = 1/4 (51.3 x 0.149) / 3.82 x 10" = 5.0 ksi, 1

where l M=FIJ2, bending moment (Reference 6.7, Table 3, Case Ib)

M=FIJ2 = (8.7 x 11.813)/2 = 51.3 in-lb F = W(1+a) = 4.7 (1+0.85) = 8.7 lbs - acting force; W = pV = 0.29 x 16.2 = 4.7 lbs - weight of the top plate 2 2 2 a=}G 4g,2 = }0.6 + 0.6 = 0.85g - maximum acceleration at OBE, according to Reference 6.20, accelerations in any horizontal and l

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ABB C-PENG-CALC-019 Rev. 00 Page 33 of 53 vertical direction shall be applied simultaneously. Gx = Gy = 0.6 -

conservative horizontal and venical OBE acceleration values.

p = 0.29 lb/in) - density of the stainless steel i 2 2 V = n/4 (5.5 - 1.66 ) 0.75 = 16.2 in' - volume of the top plate l L = (13.438 - 2 + 0.75/2) = 11.813 in; length of the tie rod from upper flange to the center ofgravity of top plate I = xd'/64 = 3.82 x 10" in' - moment ofinertia of the tie rod d = 0.297 in - basic minor diameter of the tie rod c = d/2 = 0.297/2 = 0.149 in 1

5.3.1.7.2 Residual Torque, Ta The residual torque due to preload may be calculated using the following equation for standard threads (Reference 6.10, Equation 6):

Ta = 0.562< T = 0.5625 (75) = 0.042 in-kips where:

0.5625 = multiplier based upon a coefficient of friction of 0.15 and standard bolt dimensions T = 75 in-lb = applied torque 1

5.3.1.7.3 Torsional Shear Stress, tr i I

The torsional shear stress is determined using the following equation derived from Equation 7 of Reference 6.10:

tr = 16Ta / xd' =16 (0.042) / x (0.32)' = 6.53 ksi where:

I d = average of basic pitch diameter (0.3344 in.) and minor diameter (0.297 in.) =

0.32 in.

l l 5.3.1.7.4 Maximum Stress, om .,

l The maximum stress intensity (c.,x) is determined using the following equation (Reference 6.10, i

Equation 8):

om,x = 2 ( #' #')'+(r )2 7 = 37.4 ksi < 2.7 Sm = 72.6 ksi (S. at 650*F)

< 3.0 Sm = 80.7 ksi (S. at 650*F)

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C-PENG CALC-019 Rev. 00 l Page 34 of 53 l 5.3.1.8 Shear Stress (r) - Threads 1

l At Too Plate (hex nuts)

The tie rods pass through the top plate and are held in place with hex nuts at the top and at the bottom. The impact load, in directly loading the top plate and top nut, will create stresses in this l

l section of the tie rod which are in addition to the tie rod / nut preload stresses. The nuts are of the  !

same material as the rods. Therefore, the parameters associated with the external threads of the rod are used (i.e., because of the smaller shear area).  !

From Reference 6.10:

AS,= x n Le Kn max [(1/2n) + 0.57735 (Es min - Kn max)] = 0.288 in2 where:

n = number of threads per inch = 16 Le = the length of engagement (nut thickness) = 0.5 in (Ref. 6.5) l Kn max' = maximum minor diameter ofinternal thread = 0.321 in Es min = minimum pitch diameter of extemal thread = 0.3287 in P = 2.33 kips T = 2.33 kips / 0.283 in2= 8.1 ksi < 0.6 Sm = 16.14 ksi At Uoner Flange On the other side, the tie rods thread into the Upper Flange. The lower strength Upper Flange threads are evaluated below. (The extemal tie rod threads in the Upper Flange have essentially the same stress as the external tie rod threads in the top plate region, which were evaluated previously.)

From Reference 6.11:

2 AS = x n Le Ds min [(1/2n) + 0.57735(Ds min - En max)] = 0.414 in l where:

l n = number of threads per inch = 16 l Le = the length of engagement. Assume equal to 0.5 in En max = maximum pitch diameter ofinternal thread = 0.3401 in l

Ds min = minimum major diameter of external thread = 0.3643 in P = 2.33 kips ABB Combustion Engineering Nuclear Power l

l

ABB C-PENG-CALC-019 Rev. 00 Page 35 of 53 T = 2.33 kips / 0.414 in2= 5.6 ksi < 0.6 Sm = 9.72 ksi The minimum allowable length of engagement of the tie rod into the Upper Flange may be calculated as a simple proportion:

Le mm = (Shear Stress / Allowable Stress) x Assumed Length of Engagemer.t =

= (5.6/9.72) x 0.5 = 0.289 in.

l l

l l

i k

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ABB C-PENG-CALC-019 Rev. 00 Page 36 of 53 5.3.2 Hex Head Bolt Stresses 5.3.2.1 Design Bolt Load Stresses (NB-3231 and Appendix E)

Design Bolt Load for the Design Pressure, W.i W.3 = H + H, H = 1.963 kips (from Section 5.2.3.1.!)- hydrostatic end force H, = 2b x 3.14GmP - compression load to ensure a tight joint b = 0.25 in (width of seal)

G = 1.24 in (average diameter of seal, Reference 6.5) m = 1.3 (from Reference 6.12, p. 47)

P = 2.500 ksi

=> H, = 6.33 kips

=> W i =1.963 + 6.33 = 8.29 kips Stress due to W.i = W.i/ A

= 8.29 kips / [4(0.1599 in2 ))

= 12.96 ksi < 26.9 ksi (S. at 650'F)

Design Minimum Initial Bolt Load, W.2 W 2 si taken as the total preload. Bolt stress due to preload only (ani)is calculated in Section 5.4.3.1:

ci.pi = 22.51 ksi < 28.3 ksi (S. at 100'F)

The stress due to W.i and W.2 are below their respective allowables, which indicates that the actual bolt area (A6) exceeds the minimum required area (A.).

5.3.2.2 Stress Due to ThermalExpansion The differential thermal expansion between the hex head bolt and the upper flange and compression collar will create an additional tensile load on the bolt. For the analysis of the bolt, this additional load is assumed to be completely taken up by deformation of the bolt.

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1 l

A5B C-PENG-CALC-019 Rev. 00 Page 37 of 53 The stress effect of this differential thermal expansion is considered to be equivalent to that of the tie rod since the respective tie rod - top plate and bolt-flange materials are the same:

ot.. = 8.2 ksi The effective tensile force due to this thermal expansion is determined by:

2 P = at . A, = (8.2 ksi)(0.1599 in ) = 1.31 kips 5.3.2.3 Preload The bolts are being preloaded to 30 ft-lb. To determine the load in each bolt, the following equation is used (Reference 6.10):

T = 0.2 P d

=> F =T / 0.2 d Given:

T = the applied torque = 360 in-lbs d = is the nominal major bolt diameter = 0.50 in.

F = (360 in-lbs) / (0.20) (0.50 in) = 3.600 kips.

5.3.2.4 Maximum Bolt Load As discussed in Section 2.3.2, the stiffness of the components in the flanged connection contribute to the maximum hex head bolt load. The stiffness of the components is calculated below Stiffness of Hex Head Bolts:

The stiffness of the bolts is calculated using the same methods described for the tie rods in Section 5.2.3.2.1. Dimensions are taken from Reference 6.5.

2 AE (0.172 in )(25.0X 10' f,) g

• = 6,719,000 .j Ka=4 /

=4 2.56 m. m ABB Combustion Engineering Nuclear Power

AIB C-PENG-CALC-019 Rev. 00 Page 38 of 53 i where: A = 0.172 in 2, cross-sectional area of the bolt, based on the basic pitch diameter 0.4675 in.

E = 25.0 x 10' psi (at 650*F) 1 = effective length of bolt, assuming 0.5 in of thread engagement

= thread engagement + lower flange + upper flange + washer

= 0.5 + 0.5 + 1.5 + 0.06 = 2.56 inch Stiffness of Overall Flanne:

The PDT MNSA has three components which represent the flanged connection to the Hot Leg, the upper flange (top), the upper flange (bottom), and the compression collar. The stiffness of each of these components is calculated with the use of Reference 6.7. Both the top and bottom Upper Flange have the same dimensions.

Upper flange (top / bottom): )

The following equations are found in Reference 6.7, Table 24, Case la. All dimensions are taken from Reference 6.5. ,

l w a' C l

.Y = g ( ,C,L, ' ' l 1

where:

Et 3 25.0X 10' I" ,bf (0.75)'in' D= = = 965,831 in -Ib I 12(1-y ) 2 2 12(1 - 0.3 )

Ci, C7, L,, and L3 are constants, and are calculated using the equations of Reference 6.7, using the following dimensions. Since the flange daes not have a rectangular cross section, the dimensions are selected to produces the lowest nage stiffness.

a = outer radius,1.906 in b = inner radius,0.781 in

r. = radius of applied load,0.781 in t = thickness,0.75 in y = Poisson's ratio,0.3 E = elastic modulus,25.0 X 10' psi l Ci = 0.593

C7 = 0.924 L3 = 0.0215 L, = 0.2973 ABB Combustion Engineering Nuclear Power

ABB C-PENG-CALC-019 Rev. 00 Page 39 of 53 Solving for the stiffness of the upper flange, top and bottom:

K, = E = a, '"* =4,043,000 E y '"

D (C,L' C,

-L)

Compression Collar:

COMPRESSION COLLAR lC 1.500" =l l 0.335"

= 1.0er = +-* - 0.250" l ll

1.4 w

A i* 2.000" =

a = 1.5/2 = 0.75 in

, b = 1.003/2 = 0.502 in

r. = 0.621 in t = 1.169 - 0.330 = 0.839 in (conservatively) y = 0.3 E = 25.0 X 10' psi Ci = 0.3190 C, = 0.3752 L3 = 0.0008 L, = 0.1471 D=

Et 3 = 25.0X10' I k" (0.839)'in'

= 1,352,100 in -Ib 12(1 - y2 ) 12(1 - 0.3 )

2 4

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A BB C-PENG-CALC-019 Rev. 00 Page 40 of 53 K.=b= J' a' C,L, _

= 100,594,300 m

.D ( C.,

Determination of equivalent flange stiffness:

The two upper flange pieces (top and bottom) are considered to act in parallel with each other (see Section 2.3.2). The overall flange and the compression collar act in series with the bolt. The effective stiffness of the components is calculated below.

K%= = 7,484,330$

, m (K,,,,,,,,+K,,,,,.w) Ka Therefore, the maximum bolt load is 6,719,000 '7.0 F"" = 3.6 + = 4.43 la.ps (6,719,000 + 7,484,330j 4.

5.3.2.5 Average Stress, m (NB-3232.1)

The average (axial) stress (ci) in the bolt is due to a combination of stresses from the maximum bolt load and from differential thermal expansion:

ci = P/A + ot..

A = 0.1599 in2 i

P = 4.43 kips at == 8.2 ksi 2

ai = (4.43 kips /0.1599 in ) + 8.2 = 35.9 ksi < 2 Sm = 53.8 ksi (at 650*F) 5.3.2.6 Maximum Stress (NB-3232.2)

The maximum stress in the bolt is essentially a stress intensity due to a combination of the average stress and the torsional shear stress due to residual torque (from preload).

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AIB C-PENG-CALC-019 Rev. 00 Page 41 of 53 5.3.2.7 Residual Torque Ta l

The residual torque due to preload may be calculated using the following equation for standard l threads (Reference 6.10, Equation 6):

Ta = 0.5625 T = 0.5625 (360) = 0.203 in-kips i l

where:

0.5625 = multiplier based upon a coefficient of fiiction of 0.15 and standard bolt dimensions T = 360 in lb = applied torque 5.3.2.7.1 Torsional Shear Stress, TT i

The torsional shear stress is determined using the following equation derived from Equation 7 of Reference 6.10:

3 tr = 16Ta / xd =16 (0.205) / x (0.4525)' = 11.16 ksi where:

d = average of basic pitch diam. (0.4675 in.) and minor diam. (0.4374 in.) = 0. 4525 in.

5.3.2.7.2 Maximum Stress, c.

The maximum stress intensity (o.) is detennined using the following equation (Reference 6.10, Equation 8):

a.=2( )2+(tr)2 = 42.1 ksi < 2.7 Sm = 72.6 ksi (S. at 650*F)

< 3.0 Sm = 80.7 ksi (S. at 650*F) 5.3.2.8 ThreadShear Stress, r From Reference 6.11:

AS, = x n Le Kn max [(1/2n) + 0.57735(Es min - Kn max)] = 0.400 in2 where:

n= number of threads per inch = 20 Le = the length of engagement. Assume equal to 0.5 in ABB Combustion Engineering Nuclear Power

AIB C-PENG-CALC-019 Rev. 00 Page 42 of 53 Kn max = maximum minor diameter ofinternal thread = 0.457 in Es min = minimum pitch diameter of external thread = 0.4619 in P = 4.43 kips t = 4.43 kips / 0.400 in 2= 11.I ksi < 0.6 Sm = 16.14 ksi The minimum allowable length of engagement of the hex head bolt into the Hot Leg pipe may be calculated as a simple proportion, based on the bolt threads.

Le.i. = (Shear Stress / Allowable Stress) x Assumed Length ofEngagement =

= (11.1/16.14) x 0.5 = 0.34 in.

l l

l l

l l ABB Combustion Engineering Nuclear Power

ABB C-PENG-CALC-019 Rev. 00 Page 43 of 53 l

5.3.3 Top Plate Stresses I l

5.3.3.1 Shear Stress, r 2

A. = (f) x (D) t =0.5 x (2.25 in) (0.75 in) = 2.65 in where:

f = factor to account for slot in top plate (reduction in area).

Conservatively, use f = 0.5 )

l D = the diameter of the valve = 2.25 in (Reference 6.18) t = the thickness of the top plate = 0.75 in P = 7.0 kips t = 7.0 kips / 2.65 in 2= 2.64 ksi < 0.6 Sm = 9.72 ksi 5.3.3.2 Bendingstress, oy The top plate finite element model discussed in Section 5.2.3.2.2 was used to determine the stresses in the top plate. The effective applied load of 1000 lbs generates a stress distribution in the plate, see Figures B1 and B2 in Appendix B. The maximum stress intensity in the model is 2915 psi and occurs at the inner radius of the plate on the symmetry plane. Scaling this value produces:

2915 psi X (7000 lbs /1000 lbs) = 20.4 ksi os = 20.4 ksi < l.5 Sm = 24.3 ksi ABB Combustion Engineering Nuclear Power

AIB C-PENG-CALC-019 Rev. 00 Page 44 of 53 l

l 5.3.4 Compression Collar Stresses 5.3.4.1 Shear Stress, r 2

A.= (x)(D)(t) = (x)(l.500 in) (0.330 in) = 1.555 in l P = 3.6 kips / bolt x 4 bolts = total preload from bolts

= 14.4 kips

= 14.4 kips /1.555 in2 = 9.26 kd < 0.6 Sm = 9.72 ksi 5.3.4.2 Bearings:ress, crs 2 2 Ae = (x/4)(D .o.p.,n,.co - d ,,,p,,,,,g) , (gf4)(3,492- 1.003 )2 in = 0.954 in2 P = 14.4 kips l 2

c6 = 14.4 kips / 0.954 in = 15.1 ksi < Sy = 17.9 ksi l

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AB C-PENG-CALC-019 Rev. 00 Page 45 of 53 1

5.3.5 Upper Flange Stresses 1 1

Shear stress, t 2

A = (n)(D) (t) = (n)(2.02 in) (0.325 in) = 2.062 in P = 14.4 kips l

t = 14.4 kips / 2.062 in 2= 6.98 ksi < 0.6 Sm = 9.72 ksi l Due to the proximity of the bolts and support surface, bending stresses are considered to be small and are neglected.

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ABB C-PENG-CALC-019 Rev. 00 Page 46 of 53 S.4 Fatigue Analysis The fatigue analysis of ti.s components will conservatively consider loads which may exist on the components after weld or nozzle failure has occurred. Prior to failure, components will be subjected to loads due mainly to preload and thermal expansion. After failure, and assuming that the nozzle / valve is free to move, certain components will be additionally stressed because of the internal pressure forcing the nozzle / valve up against the top plate. The load on these components would be cyclical, given the change in pressure and temperature that occurs as the plant heats up and then cools down.

The critical components for fatigue analysis purposes are the tie rod and hex head bolt, on the basis of:

e preload tensile stresses e thermal expansion tensile stresses e stress concentrations in the threaded sections, and

= for the levels of stresses involved, a more restrictive number of allowable cycles (versus the stainless steel MNSA components; see Table I-9.1 of Reference 6.6 It is noted that, in the fatigue analyses below, the stresses produced by the one-time application of an impact load are not considered since the contribution to fatigue from this one occurrence is not significant.

5.4.1 Normal Operating Pressure Force The effect of the force acting on the MNSA components due to internal pressure is similar to that of the impact load, only of a smaller magnitude, and it is a function of the internal pressure and the area of the nozzle. The pressure used to determine the force is based on the maximum internal pressure for all Normal and Upset conditions from References 6.3, which is 2350 psia (for the Reactor Trip, Loss of Load / Loss of Flow transient). However, the Design Pressure of 2500 psia is greater, and is used here. The force due to 2500 psiis calculated in Section 5.2.3.1.1 to be 1.963 kips and is rounded to 2.0 kips for use in this section.

This force will be used to calculate normal operating tensile struses in the tie rod and bolt as part of the Peak Stress Intensity calculation.

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ABB C-PENG-CALC-019 Rev. 00 Page 47 of 53 5.4.2 Tie Rod Fatigue S. ./.2.1 Tie RodPeak Stress The maximum Peak Stress in the tie rod is calculated, as follows:

Peak Stress = fsrf*(om.)

where:

fsrf = fatigue strength reduction factor = 4.0 (from NB-3232.3) om. = maximum stress using the load from normal operating pressure instead of the impact load:

om. = 2 ( #' #' )2 + (rr )'

where:

ci = ci, + ca pi+ c .o (conservatively), and ci, = tensile stress due to pressure

= [1.33(2.00 kips /4)] / 0.0775 in 2

= 8.58 ksi ci.,i = tensile stress due to preload (creload from Section 5.3.1.3)

= 1.0 kips / 0. 0775 in 2

= 12.9 ksi ot . = tensile stress due to thermal expansion (stress from Section 5.3.1.2)

= (8.2 ksi)*(30E6/25.25E6) (E/E per NB-3232.3)

= 9.74 ksi on = 5.0 ksi (stress from Section 5.3.1.7.1) tr = 6.53 ksi(stress from Section 5.3.1.7.2)

=> om.= 38.5 ksi o Peak Stress = 4 (38.5) = 154.0 ksi Based upon a minimum stress value of 0.0 ksi (this is a conservative approach, since preload never goes away), the maximum Peak Stress Intensity Range (S,) in the tie rod is:

S, = 154.0 - 0 = 154.0 ksi ABB Combustion Engineering Nuclear Power

ABB C-PENG-CALC-019 Rev. 00 Page 48 of 53

5. 4.2.2 Tie Rod Usage Factor The calculation of the usage factor for the tie rod is based upon the maximum Peak Stress Intensity l Range of 154.0 ksi: l S. = Sp/2 = 77.0 ksi This range is considered to occur a total number of 700 cycles, which is the sum of the numbers of cycles for Heatup/Cooldown (500) and Plant Leak Test (200), per Reference 6.17 (seismic loads are conservatively included in a total number of cycles). '

For a S. = S,/2 = 77.0 ksi, the allowable number of cycles per Table I-9.1 (Reference 6.6) is approximately 1650. The usage factor (U)is:

U = 700/1650 = 0. 424 l

There are other Normal and Upset transients which are defined for the Piping (per Reference 6.17),

but their contribution to fatigue in the tie rod is not significant.

ABB Combustion Engineering Nuclear Power

r l

i AIB C-PENG-CALC-019 Rev. 00 Page 49 of 53 5.4.3 Hex Head Bolt Fatigue 5.4.3.1 Hex HeadBolt Peak Stress The maximum Peak Stress in the bolt is calculated, as follows:

Peak Stress = fsrf*(cmx.)

where:

fsrf = fatigue strength reduction factor = 4.0 (from NB-3232.3) ou = maximum stress using the load from normal operating pressure instead of the impact load:

omx. = 2 ( )2 +(rr )*

where:

ci = ca., + c.,i+

a ct.. (conservatively), and ci., = tensile stress due to pressure

= [(2.00 kips /4)] / 0.1599 in2

= 3.13 ksi ci.,i = tensile stress due to preload (preload from Section 5.3.2.3)

= 3.6 kips / 0.1599 in 2

= 22.51 ksi ot = tensile stress due to thermal expansion (stress from Section 5.3.2.2)

= (8.2 ksi)*(30E6/25.25E6) (E/E per NB-3232.3)

= 9.74 ksi tr = 11.16 ksi(stress from Section 5.3.2.7.1)

=> ow = 41.83 ksi

=> Peak Stress = 4 (41.83) = 167.3 ksi Based upon a minimum stress value of 0.0 Lsi (this is a conservative approach, since preload never goes awa.r), the maximum Peak Stress Intensity Range (Sp ) in the bolt is:

S, = 167.3 - 0.0 = 167.3 ksi l

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I ABB C-PENG-CALC-019 Rev. 00 Page 50 of 53 U.3.2 Hex HeadBolt Usage Factor The calculation of the usage factor for the hex head bolt is based upon the maximum Peah 5 tress I Intensity Range of 167.3 ksi:

1 S = S,/2 = 83.7 ksi This range is considered to occur a total number of 700 cycles, which is the sum of the numbers of cycles for Heatup/Cooldown (500) and Plant Leak Test (200), per Reference 6.17.

For a S. = S,/2 = 83.7 ksi, the allowable number of cycles per Table I-9.1 (Reference 6.6) is approximately 1400. The usage factor (U)is:

l i

U = 700/1400 = 0.500 l

There are other Normal and Upset transients which are defmed for the Fiping (per Reference 6.17), ,

but their contribution to fatigue in the bolt is not significant.

1 ABB Combustion Engineering Nuclear Power

-- . ~

ABB C-PENG-CALC-019 Rev. 00 Page 51 of 53 S.5 Consideration ofHydrostatic Test Pressure Conditions Per Paragraph 1.3.1 of Reference 6.4, the deliverable MNSA hardware is not required to be hydrostatically tested. However, it is noted that the pressure load created by the hydrostatic pressure of 3125 psi (Reference 6.17) is less than the impact load. Therefore, stresses resulting from hydrostatic testing would be acceptable.

S.6 Consideration ofFaulted Conditions Reference 6.17 lists a Faulted Condition which, by dermition, is identical to the Design Condition except that it also includes a Maximum Earthquake event. An assessment is made of the effect of Faulted Conditions by reviewing the maximum stress results for the tie rod (Sections 5.3.1.7), which is the component most significantly affected by an eanhquake event (either OBE or Maximum /DBE).

Conservatively doubling the OBE bending stresses to simulate the effects of the Maximum Earthquake event results in stresses which meet the 3S. allowable for Maximum stress. Therefore, the stresses resulting from Faulted Conditions are acceptable.

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ABB C-PENG-CALC-019 Rev. 00 Page 52 of 53 6 REFERENCES 6.1. ABB Project Plan No. C-NONE-IPQP-0263, "Waterford Mechanical Nozzle Seal Assemblies (MNSA)", Revision 00, March 1999.

6.2. ABB Combustion Engineering Nuclear Power Quality Procedures Manual QPM-101, Revision 03.

! 6.3. Report No. CENC-1444, " Analytical Report for Waterford Unit No. 3 Piping", May 1981.

6.3.1 Drawing 74470-772-001, Revision 04, " Instrument Nozzles Waterford III Piping" 6.4. Specification No. C-NOME-SP-0067, Revision 01, " Design Specification for Mechanical Nozzle Seal Assembly (MNSA) Waterford Unit 3", March 1999.

6.4.1 Dravo Drawing E-3023-LW3-RC-34, Revision 7.

6.5. MNSA Drawings 6.5.1. E-MNSAWFD-228-002, Revision 02, " Hot Leg PDT MNSA" 6.5.2. E-MNSAWFD-228-003, Revision 03, " Hot Leg Sampling MNSA" 6.5.3. E-MNSA-228-013, Revision 06, " Mechanical Nozzle Seal Assembly Details" 6.5.4. E-MNSA-228-004, Revision 05, " Mechanical Nozzle Seal Assembly Details" 6.6. American Society of Mechanical Engineers Boiler and Pressure Vessel Code,Section III,1989 Edition (No Addenda).

6.7. "Roark's Formulas for Stress and Strain", Warren C. Young, Sixth Edition,1989, McGraw-Hill.

6.8. S-PENG-DR-005, Revision 01, " Addendum to CENC-1365 and CENC-1507 Analytical Reports for Southern California Edison San Onofre Units 2 and 3 Piping", February 1998.

6.9. ANSYS Engineering Analysis System computer code, Revision 5.3.

6.10. "How to Calculate and Design for Stress in Preloaded Bolts", A.G. Hopper and G.V. Thompson, Product Engineering,1964.

6.11. ANSI Standards for Threads, Appendix B, Bl.1,1982.

6.12. Union Carbide Grafoil, " Engineering Design Manual", Volume One, Sheet and Laminated Products, by R.A. Howard,1987.

l ABB Combustion Engineering Nuclear Power 1

ABB C-PENG-CALC-019 Rev. 00 Page 53 of 53 6.13. " Mechanical Engineers' Handbook", M. Kutz, ed., John Wiley & Sons, Inc.,1986.

6.14. " Strength of Materials", F. L. Singer, Second Edition, Harper & Row, New York,1962 6.15. " Engineering Mechanics: Statics and Dynamics", F.L. Singer, Third Edition, Harper & Row, New York, 1975.

6.16. Specification No. 00000-PE-140, Revision 01, " General Specification for Reactor Coolant Pipe and Fittings", November 1972.

6.17. Specification No. 09270-PE-140, Revision 07, " Project General Specification for Reactor Coolant Pipe and Fittings for Entergy Operations, Inc., Waterford Unit 3", December 1993.

6.18. ABB Inter-Office Correspondence NOME-99-C-0122, Revision 1, Waterford MNSA Stress Analsyis",

March 18,1999.

6.18.1. Velan Drawing PI-4562-N-1, Sheet 1 of 4, Revision Y, " %" - 2" Bonnetless Inclined Valve" 6.18.2. As-Built Measurements 6.19. Engineering Report C-NOME-ER-0120, Revision 00, " Design Evaluation ofMNSA for Various Applications at Waterford Unit 3", March 1999.

6.20. Calculation C-PENG-CALC-021, Revision 00, " Determination ofWaterford 3 Hot Leg Seismic Response Spectra & Accelerations for Use in the Analyses ofMNSAs", dated March 1999.

6.21. Test Report TR-PENG-042, Revision 00, " Test Report for MNSA Hydrostatic and Thermal Cycle Tests" ABB Combustion Engineering Nuclear Power

. . - - . . ~ . - - _ ._ . -...- - .- . . .. _. - - . . . _ _ _. -. .--.. -

C-PENG-CALC-019 Rev. 00 Page Al of A2 APPENDIX A 4

CALCULATION OF TIE ROD TEMPERATURE

+

ABB Combustion Engineering Nuclear Power

L C-PENG-CALC-019 Rev. 00 Page A2 of A2 Calculation of average temperature due to axial conduction Solution for Temperature over Length of an Infmite Fin

(" Heat Transfer - a Basic Approach", M.N. Ozisik, McGraw-Hill Inc.,1985)

T-T._., 2 _ hP To - T, kA Ambient Temperature *F T,infmite = 120 i Base Temperature, 'F T,o = 500 Nominal dimension, in tie rod OD = 0.375 Nominal dimension, ft tie rod OD = 0.03125 Tie Rod area, ft2 A = 0.00077 Tie Rod perimeter, ft F = 0.09817 Tie Rod therm cond (Ref 6.6,300*F), Btu /hr-fi-F k = 8.8 Generic outside film coefficient, Btu /hr-ft2-F h = 1.7 Calulcated coefficient m = 4.97265 Length x T (inch) (ft) ('F) 0 0 500 l 0.5 0.0417 429 1.0 0.0833 371 1.5 0.1250 324 2.0 0.1667 286 2.5 0.2083 255 3.0 0.2500 230 3.5 0.2917 209 4.0 0.3333 192 4.5 0.3750 179 5.0 0.4167 168 5.3 0.4375 163 275 5.5 0.4583 159 Average temperature at 5.3 inches is 275 F l ABB Combustion Engineering Nuclear Power 1

l

C-PENG-CALC-019 Rev. 00 l Page B1 of B17 APPENDIX B ANSYS EVALUATION OF TOP PLATE This appendix contains figures of the ANSYS model used to obtain the stiffness and stress of the top ,

plate, as well as the resulting output file from the computer run. The output file print.:d here contains an I echo print of the input file used.

1 l

l l

ABB Combustion Engineering Nuclear Power

C-PENG-CALC-019 Rev 00 Page B2 of B17 Figure Bl. Element Plot of ANSYS Top Plate Model  :

i

@ Re.sta:uned loca+ tons - Ioeally high stresses aroud 4hese ponnis are tynoud

+ Dist<ibwled aeplied load l l

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1 '

T' i

1. 206 I l 1

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,;zr ,//// m /

I L 1 +, # .

Locahm of stx dveu s dicp/acemed 1  :., ,, u,2 J/ress brieoret Afy l'ne of symmedry Z _.x ABB Combustion Engineering Nuclear Power

C-PENG-CALC.019 Rev. 00 Page B3 of B17 ANSYS OUTPUT FILE l

AA AA BBB BBB BBB BBB i AAA AAA BBB BBBB BBB BBBB i AAAA AAAA BBB BBB BBB BBB AAAAA AAAAA BBB BBBB BBB BBBB AAAAAA AAAAAA BBB BBBBB BBB BP,dBB AAAAAAA AAAAAAA BBB BBBBBB BBB dBBBBB AAAAAA AAAAAA BBB BBBBB BBL BBBBB AAAAAA AAAAAA BBB BBB SeB BBB ASEA BROWN BOVERI l ....

l COMBUSTION ****

ENGINEERING l

I NUCLEAR POWER WORKSTATION ENVIRONMENT Date/ Time Fri Mar 12 17:02:13 EST 1999 User . mendrala Group + v9421me Project, Task : default _ charge Job Id  : Ou9gpghs Job Name Ou9gpghs Node Name  : twister Cpu Type HP-UX Process ID  : 2061 Priority  : 8 SuDmit Node  : tornado l

Directory a twisters /PS/mendrela/mnsa '

l 1 l

Job Stream Commands From the File: twisters /PS/mendrala/mnsa/sub 1 ansys -a. 48 < cim1.inp 1 Executing /ansys53/cin/hp100/ansys.e53 j

I I I WELCOME TO .T H E ANSYS PROGRAM i I

1-

............. .......................................1

?

ANSYS 5.3 NOTICES *

• THIS AN3YS, INC. SOITWARE PRODUCT ITHE PROGRAM) AND PROGRAM

• I

• DOCUMENTATION IDOCUMENTATION) ARE filRNISHED BY ANSYS, INC.

• UNDER AN ANSYS SOFTWARE LICENSE AGREEMENT THAT CONTAINS

• PROVISIONS CONCERNING NON-DISCLOSURE, COPYING, LENGTH AND

• ABB Combustion Engineering Nuclear Power

)

4 s

l C-PENG-CALC-019 Rev. 00 u

Page B4 of B17

• NATURE OF USE, WARRANTIES, DISCLAIMERS AND REMEDIES, AND *

• OTHER PROVISIONS. THE PROGRAM AND DOCUMENTATION MAY BE

• l

• USED OR COPIED ONLY IN ACCORDANCE WITH THE TERMS OF THAT *

• LICENSE AGREEMENT. *

• Copyright 1996 SAS IP, Inc. Proprietary Data.

• l

• Unauthorized use, distribution, or duplication is *

• prohibited. All Rights Reserved.

• The Program also contains the following licensed software:

• PCGLSS: Linear Equations Solver *

(C) Copyright 1992-1995 Computational Applications and

• System Integration Inc.

• All rights Reserved.

CA&SI, 2004 S. Wright' Street,-Urbana, IL 61821

• Ph (217)244-7875

• Fax (217)244-7874 *

• CA&SI DOES NOT GUARANTEE THE CORRECTNESS OR USEFULNESS OF *

• • THE RESULTS OBTAINED USING PCGLSS. CAESI IS NOT LIABLE FOR

• ANY CONCLUSIONS OR ACTIONS BASED ON THE RESULTS. IT IS THE

• RESPONSIBILITY OF THE USER TO CONFIRM THE ACCURACY AND +  !

• USEFULNESS OF THE RESULTS.

• ANSYS/ Mechanical AFTER YOU HAVE READ AND UNDERSTOOD THE PREVIOUS NOTICES, PRESS <CR> OR < ENTER > TO CONTINUE ANSYS COMMAND LINE ARGUMENTS "***

MEMORY REQUESTED (MB) = 48.0 GRAPHICS DEVICE REQUESTED = X11

• • • NOTE *** CP= 0.590 TIME = 17:02:17 There are no parameters and no abbreviations defined.

10158-5.3 VERSION =HP 9000/700 RELEASE = 5.3 UP071096 FOR SUPPORT CALL L.L. Beaudreau PHONE 860-285-3991 FAX 860-285-2901 CURRENT JOBNAME-file 17:02:17 MAR 12, 1999 CP= 0.590 l

/SHOW SET WITH DRIVER NAME= X11 , RASTER MODE, GRAPHIC PLANES = 8  !

RUN SETUP PROCEDURE FROM FILE = /ansys53/docu/ start.ans

/ INPUT FILE = /ansys53/docu/ start.ans LINE= 0 ABBREVIATION = VED, EDIT /SYS,/usr/ bin /X11/ved &

ABBREVIATION = ANSYSWEB DELETED.

BEGIN:

1 / BATCH, list 2 ! cimi.inp 3 t Non-standard plate for MNSA 4 /filnam,c1 5' / PREP 7 6 / TITLE,WSES-3 MNSA Hot Leg PDT Top plate 7

8 tVariables 9 R1=0.83 f IR of center hole I 10 1R2=2.06/2 1 radius of SONGS 11 R2=2.50/2 1 radius for applied force 12 R3=4.75/2  ! bolt center radius 13 R4=5.50/2 i OR of plate 14 th=0.75 1 plate thickness 15.

16 tMaterials 11 M PTEM P,1,100, 2 00, 300, 4 00, 500, 600 18 t S.S. TYPE 304, 1989 ASME 19 MPDATA,ALPX,1,,8.55E-6,8.79E-6,9.00E-6,9.19E-6,9.37E-6,9.53E-6 j 20 M PDATA, EX,1, , 2 8.1 E 6,27. 6E 6, 27. 0E6,2 6. 5 E 6,2 5. 8 E 6,2 5. 3 E6 J 21 TRE F,70 22

, 23 ! Graphical options i j 24 /pnum, area,1 25 /pnum,11ne,1 li j 26 / triad,1 bot

27 28 tElements

! ABB Combustion Engineering Nuclear Pover .

i

C-PENG-CALC-019 Rev. 00 Page B5 of B17 29 ET,1,SHELL93 30 R,1, *.h 31 eshape,3 32 esize,,15 33~

34 35 ! Geometry 36 K,1,0,0,1 37 K,2,0,0,0 1

38 K,3,R1 1 39 K,4,R2  !

40 K,5,R3 41 K,6,R4 )

4 42 L,2,3 43 L,3,4 l

44 L,4,5 1 45 L , 5, 6 46 AROTAT,l',2,3,4,,,1,2,-(90-21) I 47 AROTAT,5,6,7,8,,,1,2,-21 48 AROTAT,13,14,15, :. 6, , ,1, 2, -21 l

1

-49 AROTAT,21,22,23,34,,,1,2,-(90-21) 1 50 ADELE,1,5,4 l 51 52 !Make rectangle no subtract {

j 53 K,50,R4+1 i 54 'L,2,50 55 *get,know 1,11ne ,num, max m

56 ADRAG,13,7,,,,know 1 ,

57 58 'I remove slot 59 AGEN,2,I t MNGS A5 60 ASBA,2,5  ! MAKES A17 61 AGEN,2,1 1 MAKES A2 62 ASBA,3,2 i MAKES A5 63 AGEN,2,1  ! MAKES A2 64 ASBA,4,2  ! MAKES A3 65 ASBA,6,1  ! MAKES A5 66 ! remove 60-deg " corners" from end of slot 67 L1=1.75 68 k,100,L1,R1 69 k,101,R4,R1+{R4-l,1)*1.73205  ! 1,73205= tan 60-deg 70 k,102,R4,R1 71 a,100,101,102 12 AGEN,2,1-73 ASM,5,1 74 ASBA,3,4 75 76 1 delete semi-circle 17 adele,9,13,4 78 ! merge all "labols" 79 nummrg,all 80

.81  ! ready fc; mesh:.ng 82 83 asel,all 84 - lesize,30,,,3 85 lesize,33,,,18

'86 Iccat,38,43 87' amesh,14 88 amesh,1,13

~89 amesh,15,17 90 tamesh,all 91 92 !asel,,,,6 93 teshape,0 94 tamesh,all 95 -allsel 96 97 fApply B.C.

98 isymmetry 99 1sel,,,,30,32 100 .ns11,s,1 101 DSYM,aymm,y,0 102

'103 ffix bolt locations 104 ks e l , s , , , 9,17, 8 105 nsik,s 106 d,all,all 107 108 allsel,all 109 110 tsolution - apply load 111 / solu 112 antype, static 113 outres,all,all 114 115 1sel,s,,,26,34,8 ABB Combustion Engineering Nuclear Pour

r- 1

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• l C-PENG-CALC-019 Rev. 00 Page B6 of B17 116 1sel,a,,,18,43,25 117 ns11,s,1 1 118 *get,nnum, node,,ccunt 1 119 f,all,ft,-500/nnum t approx.-equal force dis *.rloution 120 allsel 121 solve 122 Spost-process results ,

. -123 / pos ti )

! 124 /page,300,,300 125 set 126 lthese are the nojes with displacements more than 0.00035 ,

127 - nsel, s u, z,-0. 000% -0. 0003 5 . I 128 prnsol,u,z I 129 allsel l 130 f these are the stress for nodes on the symmetry line J 131 1sel,s,,,30,32

.132 ns11,s,1 l 133 prnsoles,prin 134 allsel 135 !these are the reaction forces at the tie rod locations 136 presol,fz 137 finish  ;

I

- CURRENT JOBNAME REDEFINED AS cl l 1 1

• • • • • ANSYS - ENGINEERING ANALYSIS SYSTEM RELEASE 5.3 *****

i ANSYS/ Mechanical  !

10158-5.3 VERSION *HP 9000/700 17:02:17 MAR 12, 1999 CP= 0.670 -l FOR SUPPORT CALL L.L. Beaudreau PHONE 860-285-3991 FAX 860-285-2901 l

• • • • • ANSYS ANALYSIS DEFINITION (PREP 7) *****

TITLE =

WSES-3 MNSA Hot Leg PDT Top plate l PARAMETER R1 = 0.8300000  ;

PARAMETER R2 = 1.250000 i PARAMETER R3 = 2.375000 PARAMETER R4 = 2.750000 PARAMETER TH = 0.7500000

]

• • • PROPERTY TEMPERATURE TABLE NUM. TEMPS = 6 ***  !

SLOC= 1 100.0000 200.0000 300.0000 400.0000 500.0000 600.0000 PROPERTY TABLE ALPX HAT = 1 NUM. POINTS = 6 1 SLOC= 1 0.8550000E-05 0.8790000E-05 0.9000000E-05 0.9190000E-05 '

O.9370000E-05 0.9530000E-05

]

PROPERTY TABLE EX MAT = 1 NUM. POINTS = 4 I SLOC= 1 0.2810000E+08 0.2760000E+08 0.2700000E+08 0.2650000E+08 0.2580000E+08 0.2530000E+08 REFERENCE TEMPERATURE = 70.000 (TUNIF= 70.000)

AREA NUMEERING KEY = 1 i I

LINE NUMBERING KEY = 1 l XYZ TRIAC DISPLAY SET TO LEFT BOTTOM ELEMENT TYPE '1 IS SHELL93 8-NODE STRUCTURAL SHELL KEYOPT (1-12 ) = 0 0 0 0 0 0 0 0 0 0 0 0 CURRENT NODAL DOF SET IS UX UY UZ ROTX ROTY ROTZ THREE-DIMENSIONAL MODEL REAL CONSTANT SET l' ITEMS 1 TO 6 0.75000 0.00000E+00 0.00000E+00 0.00000E+00 0.00000E+C0 0.00000E+00 FOR ELEMENT TYPE (S) ALLOWING MULTIPLE SHAPES: l PRODUCE ALL QUADRILATERAL OR BRICK ELEMENTS IF POSSIBLE. I OTHERWISE PRODUCE MIXED ELEMENT SHAPES.

DEFAULT ELEMENT DIVISIONS PER LINE BASED ON ELEMENT SIZE = 0.150 l

l. KEY POINT 1 X,Y,2= 0.000000E+00 0.000000E+00 1.00000 IN CSYS= 0 l-  !

KEYPOINT 2 X,Y,2= 0.000000E+00 0.000000E+00 0.000000E+00 IN CSYS= 0  ;

a L ABB Combustion Engineering Nuclear Pover  !

1

- r _ -

C-PENG-CALC-019 Rev. 00 Page B7 of B17 KEYPOINT 3 X,Y,Z= 0.830000 0.000000E+00 0.00C000E-00 IN CSYS. O KEYPOINT 4 X,Y,Z= 1.25000 0.000000E+00 .0.000C00E+00 IN CSYS= 0 KEYPOINT 5 X,Y,Z= 2.37500 0.000000E+00 0.000000E+00 IN CSYS= 0 KEYPOINT 6 X,Y,Z= 2.75000 0.000000E+00 0.000000E+00 IN CSYS= 0 LINE CONNECTS KEYPOINTS 2 3 LINE NO.= 1 K Pl= 2 TAN 1= -1.0000 0.0000 0.0000 KP2= 3 TAN 2= 1.0000 0.0000 0.0000 LINE CONNECTS KEYPOINTS 3 4 LINE NO.= 2 KPl= 3 TANie L+1.0000 0.0000 0.0000 KP2= 4 TAN 2= 1.0000 0.0000 0.0000

.LINE CONNECTS KEYPOINTS 4 5 LINE NO.= 3 KPl= 4 TAN 1= -1.0000 0.0000 0.0000 KP2= 5 TAN 2= 1.0000 0.0000 0.0000 i 1

LINE CONNECTS KEYPOINTS 5 6 i LINE NO.= 4 KPl= 5 TAN 1= -1.0000 0.0000 0.0000 l KP2= 6 TAN 2= 1.0000 0.0000 0.0000 i ROTATE LINES 1, 2, 3, 4, ABOUT THE AXIS DEFINED BY KEYPOINTS 1 2 DEGREES OF ARC = -69.00 NUMBER OF SEGMENTS = 1 ROTATE LINES 5, 6, 7, 8, i ABOUT THE AXIS DEFINED BY KEYPOINTS 1 2 1 DEGREES OF ARC = -21.00 NUMBER OF SEGMENTS = 1 ROTATE LINES 13, 14, 15, 16, ABOUT THE AXIS DEFINED BY KEYPOINTS 1 2 DEGREES OF ARC = -21.00 NUMBER OF SEGMENTS = 1 ROTATE LINES 21, 22, 23, 24, ABOUT THE AXIS DEFINED BY KEYPOINTS 1 2 DEGREES OF ARC = -69.00 NUMBER OF SEGMENTS = 1 DELETE SELECTED AREAS FROM 1 TO 5 BY 4 i DELETED 2 AREAS KEYPOINT 50 X,Y,2= 3.75000 0. 7000E+00 0.000000E+00 IN CSYS= 0

.LINE CONNECTS KEYPOINTS' 2 50 LINE NO.= 37 KPl= 2 TAN 1= -1.0000 0.0000 0.0000 KP2= 50 TAN 2= 1.0000 0.0000 0.0000

• GET know 1 ,

FROM LINE ITEM =NUM MAX VALUE- 37.0000000 DRAG LINES:

13, ALONG LINES 37, GENERATE 2 TOTAL SETS OF AREAS -

SET IS FROM 1 TO 1 IN STEPS OF 1 DX, DY,0Z= 0.000E+00 0.000E+00 0.000E+00 CSYS= 0 SuaTRACT AREAS AREA NUMBERS TO BE OPERATED ON = 2 AREAS OPERATED ON WILL BE DELETED AREA NUMBERS TO BE SUBTRACTED = 5 AREAS SUBTRACTED WILL BE DELETED OUTPUT AREAS = 17 GENERATE 2 TOTAL SETS OF AREAS SET IS FROM 1 TO 1 IN STEPS OF 1 DX,0Y,DZ= 0.000E+00 0.000E+00 0.000E+00 CSYS- 0 SUBTRACT AREAS AREA NUMBERS TO BE OPERATED ON = 3 AREAS OPERATED ON WILL BE DELETED AREA NUMBERS TO BE SUBTRACTED = 2 AREAS SUBTRACTED WILL BE DELETED OUTPUT AREAS = 5 GENERATE 2 TOTAL S'TS 4 OF AREAS SET IS FROM 1 TO 1 IN STEPS OF 1 DX,DY,DZ= 0.000E+00 0.000E+00 0.000E+00 CSYS= 0 SUBTRACT AREAS ABB Combustion Engineering Nuclear Pover i

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C-PENG-CALC-019 Rev. 00 Page B8 of B17 AREA NUMPERS TO BE OPERATED ON = .4 AREAS OPEllATED ON WILL BE DELETED AREA NUMBERS TO BE SUBTRACTED = 2 AREAS SUBTRACTED WILL BE DELETED OUTPUT AREAS = 3 *

' SUBTRACT AREAS AREA NUMBERS TO BE OPERATED ON = 6

' AREAS OPERATED ON WILL BE DELETED AREA NUMBERS TO BE SUBTRACTED

• 1 AREAS SUBTRACTED WILL BE DELETED-OUTPUT AREAS = 2

~ PARAMETER L1' = 1.750000- -

KEYPOINT 100 X,Y,Z= 1.75000 0.830000 0.000000E+00 IN CSYS= 0 KEYPOINT 101 X,Y,Z= 2.75000 2.56205. 0.000000E+00 IN CSYS= 0 KEYPOINT 102 X,Y,Z= 2.75000 0.830000 0.000000E+00 IN CSYS= 0 DEFINE AREA BY LIST OF KEYPOINTS KEYPOINT LIST = 100 101 102 AREA NUMBER = 1 GENERATE . 2 TOTAL SETS OF AREAS SET IS FROM 1 TO 1 IN STEPS OF 1

-DX,DY,DZ= 0.000E+00 0.000E+00 0.000E+00 CSYS= 0 SUBTRACT AREAS AREA NUMBERS TO BE OPERATED ON = 5

. AREAS OPERATED ON WILL BE DELETED.

AREA NUMBERS TO BE SUBTRACTED = 1 AREAS SUBTRACTED WILL BE DELETED OUTPUT AREAS = 6 SUBTRACT AREAS AREA NUMBERS TO BE OPERATED ON = 3 AREAS OPERATED ON WILL BE DELETED AREA NUMBERS TO BE SUBTRACTED = 4 AREAS SUBTRACTED WILL BE DELETED OUTPUT AREAS = 1 DELETE SELECTED AREAS FROM . 9 TO 13 BY 4 DELETED 2 AREAS ,

MERGE COINCIDENT NODES WITHIN TOLERANCE OF 0.10000E-03 MERGE IDENTICAL MATERIALS WITHIN TOLERANCE OF 0.10000E-06 MERGE IDENTICAL ELEMENT TYPES MERGE IDENTICAL REAL CONSTANT SETS WITHIN TOLERANCE OF 0.10000E-06 ,

MERGE IDENTICAL ELEMENTS MERGE IDENTICAL COUPLED DOF SETS MERGE IDENTICAL CONSTRAINT EQUATIONS WITHIN TOLERANCE OF 0.10000E-06 MERGE COINCIDENT KEYPOINTS WITHIN TOLERANCE OF 0.10000E-03 KEY POINT .

4 USED FOR KEYPOINT(S) 29 KEYPOINT 24 USED FOR KEYPOINTIS) 25 KEYPOINT ' 30 USED FOR KEYPOINT(S) 31 LINE 2 USED FOR LINE(S) 45 LINE 4 USED FOR LINE(S) 11 LINE. 43 USED FOR LINE(S) 46 ALL SELECT FOR ITEM = AREA COMPONENT =

IN RANGE 1 TO 17 STEP 1 12 AREAS (OF ' 12 DEFINED) SELECTED BY ASEL COMMAND.

SET DIVISIONS ON LINE 30 (IF SELECTED AND UNMESHED)

TO NDIV = 3, SPACING RATIO = 1.000 SET DIVISIONS ON LINE 33 (IF SELECTED AND UNMESHED)

-TO NDIV = 18, SPACING RATIO = 1.000 I

l CONCATENATE LINES-

-LINES = 38 43 ABB Combustion Engineering Nuclear Pover i

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l l

C-PENG-CALC-019 Rev. 00 Page B9 of B17 l i

. COMPOSITE.LINE NO.= 10 KPl= 100 KP2= 8 l GENERATE NODES AND ELEMENTS IN AREAS 14 TO 14 IN STEPS OF 1 NUMBER OF AREAS MESHED = 1

• iAXIMUM NODE NUMBER = 205 1 MAXIMUM ELEMENT NUMBER = 54 j GENERATE NODES AND ELEMENTS IN AREAS 1 TO 13 IN STEPS OF 1 NUMBER OF AREAS MESHED = 8 MAXIMUM NODE NUMBER = 1228 '

MAXIMUM ELEMENT NUMBER e 371

. GENERATE NODES AND ELEMENTS IN AREAS 15 TO 17 IN STEPS OF 1

" AREA 17 MESHED WITH 6 QUADRILATERALS, 1 TRIANGLES "

NUMBER OF AREAS MESHED = 3 MAXIMUM NODE NUMBER- = 1839 I i

MAXIMUM ELEMENT NUMBER = 576 SELECT ALL ENTITIES OF TYPE = ALL AND EELOW ALL SELECT ~ FOR ITEM =VOLU COMPONENT =

IN RANGE 0 TO O STEP 1 0 YOLUMES (or 0 DEFINED) SELECTED BY VSEL COMMAND.

ALL SELEC* FOR ITEM = AREA COMPONENT =

IN RANGE 1 TO 17 STEP 1 i

I 12 AREAS (OF 12 DEFINED) SELECTED BY ASEL COMMAND.

ALL SELECT FOR ITEM =LINE COMPONENT =

IN P1.NGE 1 TO 49 STEP 1.

30 LINES (OF 38 DEFINED) SLLECTED Da LSEL COMMAND.

ALL SELECT FOR ITEM =KP COMPONENT =

IN RANGE 1 TO 102 STEP 1 25 KEYPOINTS (OF 25 DEFINED) SELECTED BY KSEL COMMAND.

ALL SELECT FOR ITEM =ELEM COMPONENT =

IN RANGE 1 TO 576 STEP 1 576 ~ ELEMENTS (OF 576 DEFINED) SELECTED BY ESEL COMMAND.

. ALL. SELECT FOR ITEH= NODE COMPONENT =

IN RANGE 1 TO 1839 STEP 1

-1839 NODES (OF- 1839 DEFINED) SELECTED BY NSEL COMMAND.

SELECT FOR ITEM =LINE COMPONENT =

IN RANGE -30 TO 32 STEP 1

)

i

.3 LINES (or )

30 DEFINED) SELECTED BY LSEL CCM AND. '

i SELECT ALL NODES (INTERIOR TO LINE, AND AT KEYPOINTS)

RELATED TO SELECTED LINE SET. l i

29 NODES (OF 1839 DEFINED) SELECTED FROM 3 SELECTED LINES BY NSLL Com AND.

SYMMETRY CONSTRAINTS FOR COORDINATE SYSTEM ' O IN DIRECTION Y l ON. SURFACE DEFINED BY ALL SELECTED NODES '

• • • NOTE *** CP= 3.010 TIME 17:02:24 Nodes on symmetry surf aces are rotated into coordinate systein 0.

TOTAL SPECIFIED CONSTRAINTS = 87 SELECT FOR ITEM =KP COMPONENT ='

IN RANGE . 9 TO 17 STEP 8 4

1 2 KEYPOINTS (OF 25 DEFINED) SELECTED BY KSEL COMMAND.

SELECT NODES ASSOCIATED WITH SELECTED KEYPOINTS 2 NODES (OF 1839 . DEFINED) SELECTED FRCN 2 SELECTED KEYPOINTS BY NSLK COWAND. I SPECIFIED CONSTRAINT UX FOR SELECTED NODES 1 TO 1839 BY 1 ABB Combustion Engineering Nuclear Pover I

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_ . . _ . _ . _ __ _ _ _ -. _ .. . . _ . ______.m __ _ ._ _

C-PENG-CALC-019 Rev. 00 Page B10 of B17 REAL= 0.000000000E+00- IMAG= 0.000000000E+00 ADDITIONAL DOFS= UY- U* ROTX' ROTY- ROT 2 SELECT ALL ENTITIES OF TYPE = ALL AND BELOW ALL SELECT FOR ITEM =VOLU COMPONENT = i IN RANGE O TO O STEP 1 0 VOLUMES IOF 0 DEFINED) SELECTED BY VSEL COMMAND. .

ALL SELECT FOR ITEM = AREA COMPONENT ='

IN RANGE 1 TO 17 STEP 1 12 AREAS (Or - 12 DEFINED) SELECTED BY ASEL COMMAND.

ALL SELECT FOR ITEM =LINE COMPONENT =

IN RANGE 1 TO 49 STEP 1 38 LINES (OF 38 DEFINED) SELECTED BY-LSEL COMMAND.

ALL SELECT FOR ITEM =KP COMPONENT =

-IN RANGE 1 TO 102 STEP 1 25 KEYPOINTS (Or ~ 25 DEFINED) SELECTED BY KSEL COMMAND.

ALL SELECT FOR ITEM =ELEM COMPONENT =

IN RANGE 1 TO 576 STEP 1 576 ELEMENTS (OF 576 DEFINED) SELECTED BY ESEL COMMAND.

ALL SELECT FOR ITEM = NODE COMPONENT =

IN RANGE 1 TO 1839 STEP 1 1839 NODES (OF 1839 DEFINED) SELECTED BY NSEL COMMAND.

"*" ROUTINE COMPLETED ""* CP = 3.030 i

ANSYS SOLUTION ROUTINE PERFORM A STATIC ANALYSIS l THIS WILL BE A NEW ANALYSIS WRITE ALL ITEMS TO THE DATABASE WITH A FREQUENCY OF ALL FOR ALL APPLICABLE ENTITIES SELECT FOR ITEM =LINE COMPONENT =

IN RANGE 26 TO 1 34 STEP 8 i

2 LINES (OF 38 DEFINED) SELECTED BY LSEL COMMAND.

ALSO SELECT FOR ITEM-LINE COMPONENT =

IN RANGE 18 TO 43 STEP 25 4 LINES (OF 38 DEFINED) SELECTED BY LSEL COtt4AND.

SELECT ALL NODES (INTERIOR TO LINE, AND AT KEYPOINTS)

RELATED TO SELECTED LINE SET.

73 NODES (OF 1839 DEFINED) SELECTED FROM 4 SELECTED LINES BY NSLL COMMMD.

'GET nnum FROM NODE ITEM =COUN VALUE= 73.0000000 i

SPECIFIED NODAL LOAD FZ FOR SELECTED NODES 1 TO 1839 BY 1 '

REAL= -6.84931507 IMAG= 0.000000000E+00 SELECT ALL ENTITIES OF TYPE = ALL AND BELOW l ALL SELEC* FOR ITEM =VOLU COMPONENT =

(. IN PANGE o TO 0 STEP 1 0 VOLUMES (cF- 0 DEFINED) SELECTED BY VSEL COMMAND.

ALL SELECT, FOR ITEM =APEA COMPONENT =

IN RANGE 1 TO 17 STEP 1 12 AREAS (Or -12 DEFINED) SELECTED BY ASEL COMMAND.

ALL SELECT -FOR ITEM =LINE COMPONENT =

.:- IN RANGE 1 TO 49 STEP- 1 38 LINES (OF 38 DEFINED) SELECTED BY LSEL COMMAND.

ABB Combustion Engineering Nuclear Pover 1

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e

E: . 1 C-PENG-CALC-019 Rev. 00 Page Bil of B17 ALL SELECT FOR ITEM =KP OOMPONENT=

IN RANGE 1 TO 102 STEP 1 25 KEYPOINTS (or 25 ' DEFINED) SELECTED BY KSEL COMMAND.

ALL SELECT FOR ITEM =ELEM COMPONENT =

.IN RANGE. 1 TO . -576 STEP 1

! 576 ELEMENTS (OF 576 DEFINED) SELECTED BY ESEL COMMAND.

ALL SELECT FOR ITEM = NODE COMPONENT =

IN RANGE 1 TO 1839 STEP 1

,1839 ' NODES (OF 1839 DEFINED) SELECTED BY NSEL COMMAND.

ANSYS SOLVE COMMAND *""

1 l' ** **

• ANSYS - ENGINEERING ANALYSIS SYSTEM RELEASE 5.3 *****

ANSYS/ Mechanical 10158-5.3 VERSION =HP 9000/700 17:02:24 MAR 12, 1999 CP= 3.190 FOk SUPPORT CALL L.L. Beaudreau PHONE 860-285-3991 FAX 860-285-2901 WSES-3 MNSA Het Leg PDT Top plate S O L U T I.O N OPTIONS PROBLEM DIMENSIONALITY. . . . . . . . . . . . .3-D DEGREES OF FREEDOM. . . . . UX UY UZ ROTX ROTY ROTZ ANALYSIS TYPE . . . . . . . . . . . . . . . STATIC ( STEADY-STATE) f**' NOTE *** .

CP= 3.240 TIME = 17:02:25 Present time 0 is less than or equal to the previous time.

Time will default to 1.

L.O A D STEP OPTIONS I LOAD STEP NUMBER. . . ...... . ... . 1 TIME AT END OF THE LOAD STEP. . . . . . ... . 1.0000 NUMBER OF SUBSTEPS. . ............. 1 STEP CHANGE BOUNDARY CONDITIONS . . . . .... NO PRINT OUTPUT CONTROLS . . . . . . . . . . . . .NO PRINTOUT DATABASE OUTPUT CONTROLS ITEM FREQUENCY COMPONENT ALL ALL Range of element maximum matrix coefficients in global coordinates Maximum = 128820495 at element 274.

Minimum = 36160910.7 at element 486.

"* ELEMENT MATRIX FORMULATION TIMES TYPE NUMBER ENAME TOTAL CP AVE CP 1 576 SHELL93 3.110 0.005 Time at end of element matrix formulation CP= 6.82999981.

Estimated number of active DOF= 10935. t Maximum wavefront = 212.

Time at end of matrix triangularization CP= 14.68.

Equation solver maximum pivot = 84753222.3 at node 713 UX.

Equation solver minimum pivot = 1.19270276 at node 361 ROTZ.

"* ELEMENT RESULT CALCULATION TIMES TYPE NUMBER ENAME TOTAL CP AVE CP 1 576 SHELL93 1.690 0.003

• " NODAL LOAD CA1CULATION TIMES TYPE NUMBER ENAME TOTAL CP- AVE CP 1 576 SHELL93 0.-110 0.00,

• • • LOAD STEP 1- SUBSTEP 1 COMPl.af ED. CUM ITER = 1

• • • TIME = 1.00000, TIME INC = 1.00000 NEW TRIANG MATRIX

• • • PROBLEM STATISTICS .

ACTUAL NO. OF ACTIVE DEGREES OF FREEDOM = 10935 R.M.S. WAVEFRONT SIZE = 188.6 1'

l I ***

ANSYS BINARY FILE STATISTICS f BUFFER SIZE USED= 4096 r 11.016 MB WRITTEN ON ELEMENT MATRIX FILE: cl.emat L ABB Combustion Engineering Nuclear Power l

~ w

\

C-PENG-CALC-019 Rev. 00 Page B12 of B17 1.359 MB WRITTEN ON ELEMENT SAVED DATA FILE: cl.esav 16.141 MB WRITTEN ON TRIANGULARIZED MATRIX r!LE: cl.trl 1 4 53 Pt. WRITTEN ON RESULTS FILE: cl.rst FINISH SOLUTION /ROCESSING

""" ROUTINE CCNPLETED ""* CP = 17.170

-1

""* ANSYS - ENGINEERING ANALYSIS SYSTEM RELEASE 5.3 *""

ANSYS/ Mechanical 10158-5.3 VERSION =HP 9000/100 17:02:45 MAR 12, 1999 CP= 17.170 FOR SUPPORT CALL L.L. Beaudreau PHONE 860-285-3991 FAX 860-285-2901 WSES-3 MNSA Hot Leg PDT Top plate

• "" ANSYS RESULTS INTERPRETATION ( POST 11 * * * "

INTERACTIVE LINES PER PAGE = 300

. INTERACTIVE CHARACTERS PER LINE= 80 FILE OUTPUT LINES PER PACE' = 300 FILE OUTPUT CHARACTERS PER LINE= 132 USE LOAD STEP 1 SUBSTEP O FOR LOAD CASE O SET COMMAND GOT LOAD STEP = 1 SUBSTEP= 1 CUMULATIVE ITERATION = 1 TIME / FREQUENCY = 1.0000 TITLE = WSES-3 MNSA Hot Leg PCrr Top plate SELECT FOR ITEM =U COMPONENT =Z BETWEEN-0.50000E-03 AND -0.35000E-03 KABS= 0. TOLERANCE = 0.150000E-11

~235 NODES (Or 1839 DEFINED! SELECTED BY NSEL COMMAND.

PRINT U NODAL SOLUTION PER NODE 1

'"" ANSYS - ENGINEERING ANALYSIS SYSTEM RELEASE 5.3 *""

ANSYS/ Mechanical 10158-5.3 VERSION =HP 9000/700 17:02:46 MAR 12, 1999 CP= 17.520 FOR SUPPORT CALL L.L. Beaudreau PHONE 860-285-3991 FAX 860-285-2901 WSES-3 MNSA Hot Leg PDT Top plate

• "" POST 1 NODAL DEGREE Or FREEDOM LISTING ** *"

LOAD STEP = 1 SUBSTEP= 1 TIME =. 1.0000 LOAD CASE = 0 THE FOLLOWING DEGREE OF FREEDOM RESULTS ARE IN GLOBAL COORDINATES NODE . UZ 2 -0.39771E-03 6 -0.35025E-03 7 -0.35202E-03 8 -0.35385E-03 9 -0.35574E-03 10 -0.35768E-03 11 -0.35967E-03 12 -0.36168E-03 13 -0.36372E-03 14 -0.36577E-03 15 -0.36784E-03 16 -0.36989E-03 17 -0.37193E-03 18 -0.37395E-03 19 -0.37594E-03 20 -0.37789E-03 21 -0.37979E-03 22 -0.38163E-03 23 -0.38341E-03 24 -0.38511E-03 25 -0.38674E-03 26 -0.38827E-03 27 -0.38970E-03

. 28 -0.39104E-03 29 -0.39227E-03 30 -0.39338E-03 31 -0.39437E-03 4 32 -0.39524E-03 33 -0.39599E-03 ABB Combustion Engineering Nuclear Pover

. . . _ .. _ ._. _. ._. __ _._ ~ _ _ . - _ _ . _ _ _ . _ _ _ _ _

t' C-PENG-CALC-019 Rev. 00

^

Page B13 of B17 34 -0.39660E-03 35 -0.39709E-03 36 -0.39743E-03 37 -0.39764E-03 38 -0.38307E-03 39 -0.39471E-03 40 -0.39192E-03 41.=0.38931E-03

( 42 -0.38693E-03' l 43 -0.38477E-03 65 -0.35111E-03 66 -0.35491E-03' 67 -0.35840E-03 68 -0.36182E-03 69 -0.36490E-03

70 -0.36789E-03 l 71 -0.37051E-03

'72 -0.37302E-03 73 -0.37513E-03 74 -0.37712E-03

• I5 -0.37870E-03 76 -0.38014E-03 77 -0.38114E-03 78 -0.38200E-03 79 -0.38257E-03 115 -0.35184E-03 122 -0.35670E-03 127 -0.35061E-03 129 -0.36159E-03 130 -0.35345E-03 -

134 -0.35627E-03 136 -0.36643E-03 137 -0.35906E-03 138 -0.35180E-03 141 -0.36180E-03 143 -0.37111E-03 144 -0.36450E-03 145 -0.35bO1E-03 146 -0.35161E-03 148 -0.36712E-03 149 -0.35501E-03 150 -0.37557E-03 151 -0.36967E-03 152 -0.36391E-03 153 -0.35827E 154 -0.35274E-03 155 -0.37212E-03 156 -0.36144E-03 157 -0.37971E-03 158 -0.37448E-03 159 -0.36940E-03 160 -0.36445E-03 161 -0.35964E-03

.162 -0.37672E-03 163 -0.36735E-03 164 -0.38346E-03 165 -0.37884E-03 166 -0.37437E-03 167 -0.3700$E-03 168 -0.36589E-03 169 -0.38082E-03 170 -0.37262E-03 171 -0.38676E-03 172 -0.38267E-03 173 -0.37875E-03 174 -0.37498E-03 175.-0.37138E-03 176 -0.38436E-03 177 -0.37718E-03 178 -0.38955E-03 179 -0.38591E-03 180 -0.38244E-03

'181 -0.37914E-03 182 -0.37602E-03

<183 -0.38728E-03 184 -0.38093E-03 185 -0.39177E-03 i

186 -0.38849E-03 187.-0.38540E-03 188 -0.38247E 189 -0.37973E-03 190 -0.38952E-03

~191 -0.38382E-03 192 -0.39339E 191 -0.39038 -03 194 -u.Joen5E-03 l 195 -0.38491E-03

j. ABB Combustion Engineering Nuclear Pover

-__ _. _ _ . _ _ . _ _ . . .__ _ __ _ . ._._ _ _ . . _ . _ .m m_.._.__._

l e C-PENG-CALC-019 Rev. 00 1:

L . . .

Page B14 of B17 1

196 -0.3824 5E '.19' -0.39104E-03

' 198 -0.38579E-03 .I 199 -0. 394 38E-03 200 -0.39152E-03 201 -0.38887E-03 202 -0.38638E-03 203 -0.38418E-03 204 -0.39181E-03 205 -0.38691E-03' 1229 -0.35529E-03

'1230 -0.37896E-03 1231 -0.37540E-03 1232 -0.37215E-03 1233 -0.36920E-03' 1234 -0.36651E-03 1235 -0.36410E-03 1236'-0.36198E-03 1237 -0.36014E-03 1238 -0.35858E-03 '

1239 -0.35730E-03 1240 -0.35629E-03 1241 -0.35556E-03 1242 -0.35509E-03 1243 -0.35490E-03 1244 -0.35497E-03 1277 -0.35155E-03 1278 -0.35362E-03 1279 -0.35487E-03~

1529 -0.35611E-03 1530 -0.35059E-03 1544 -0.35438E-03

-1551 -0.36217E-03 1552 -0.35786E-03 1553 -0.35320E-03 1566 -0.36109E-03 i

' 1567 -0.35253E-03 1573 -0.36840E-03 1574 -0.36399E-0.3 1575 -0.35985E-03 1576 -0.35595E-03 1577 -0.35229E-03 1588 -0.36661E-03 1589 -0.35899E-03 1590 -0.35240E-03 1595 -0.37289E-03 1596 -0.36888E-03 1597 -0.36515E-03

=

-1598 -0.16166E-03 1599 -0.35843E-03 1600 -0.35546E-03 1601 -0.35276E-03 l 1602 -0.35034E-03 ]

1610 -0.37086E-03 1611 -0.36394E-03 1612 -0.35808E-03 l 1613 -0.35331E-03 1617 -0.37618E-03 1618 -0.37246E-03 1619 -0.36901E-03 1620 -0.36582E-03 1621 -0.36289E-03 1622 -0.36023E-03 1623 -0.35786E-03 1624 -0.35575E-03 1625 -0.35392E-03 1626 -0.35237E-03 1627 -0.35109E-03 i i

1628 -0.35008E-03 1632 -0.37374E-03 l i

1633 -0.36729E-03 1634'-0.36192E-03 1635 -0.35767E-03 1636 -0.35452E-03 1637 -0.35247E-03 1638 -0.35149E-03 1639 -0.31825E-03 1640 -0.37463E-03 4 1641 -0.37136E-03 .j

.1642 -0.36835E-03 1643 -0.36560E-03 I i

1644 -0.36313E-03 l 1645 -0.36095E-03 1646 -0.35904E-03 1647 -0.35741E-03 1640 -0.35606E-03 1649 -0.35498E-03 .

1 ABB Combustion Engineering Nuclear Power 1

1 C PENG-CALC-019 Rev. 00 Page BIS of B17 1650 -0.35418E-03 1651 -0.35365E-03 1652 -0.3533BE-03 1653 -0.35337E '1654 -0.37523E-03 1655 -0.36898E-03 1656'-0.36386E-03 1657 -0.35986E-03 1658 -0.35699E-03 1659 -0.35521E-03 1660 -0.35452E-03 1661 -0.36118E-03 1662 -0.35579E-03 1663 -0 35449E-03 1664 -0.35738E-03 1665 -0.35846E-03 1666 -0.35973E-03 1698'-0.35288E-03 1699 -0.35650E-03 1700 -0.35910E-03 1701 -0.36066E-03 1812 -0.35073E-03 1813 -0.35172E-03 l 1814 -0.35244E-03 '

1815 -0.35410E-03 1816 -0.35406E-03

• 1917 -0.35469E-03 ,

1818 -0.35551E-03 )

1819 -0.35652E 1820 -0.35772E-03. '

1821 -0.35604E-03 1822 -0.3579BE-03 MAXIMUM ABSOLUTE VALUES.

NODE 2 VALUE -0.39171E-03 )

SELECT ALL ENTITIES OF TYPE = ALL AND BELOW ALL SELECT FOR ITEM =VOLU COMPONENT =

IN RANGE O TO O STEP 1 0 VOLUMES (OF 0 DEFINED) SELECTED BY VSEL COtNAND.

ALL SELECT FOR ITEM = AREA COMPONENT =

IN RANGE 1 TO 17 STEP 1 12 AREAS (OF 12 DEFINED) SELECTED BY ASEL' ' NAND.

ALL SELECT FOR ITEM =LINE COMPONENT =

IN RANGE 1 TO 49 STEP 1 38 LINES (OF 38 DEFINED) SELECTED BY LSEL COtNAND ALL SELECT FOR ITEM *KP COMPONENT =

IN RANGE 1 TO 102 STEP 1 25 KEYPOINTS (OF 25 DEFINED) SELECTED BY KSEL COMAND.

ALL SELECT FOR ITEM =ELEM COMPONENT =

IN RANGE. 1 TO 576 STEP 1 576. ELEMENTS (OF 576 DEFINED) SELECTED BY ESEL COMMAND.

ALL SELECT FOR ITEM = NODE COMPONENT = l IN RANGE 1 TO 1839 STEP 1 1 1839 NODES (OF- 1839 DEFINED) SELECTED BY NSEL COMMAND.

SELECT FOR ITEM =LINE COMPONENT =

IN RANGE 30 TO 32 STEP 1

)

i 3 LINES (OF 30 DEFINEDI SELECTED BY LSEL COMMAND.

SELECT ALL NODES (INTERIOR TO LINE, AND AT KEYPOINTS)

RELATED TO SELECTED LINE SET.

29 NODES (OF 1839 DE FINED) SELECTED FROM l 3' SELECTED LINES BY NSLL COMMAND. '

PRINT S NODAL SOLUTION PER NODE

-1

'"" ANSYS - ENGINEERING ANALYSIS SYWEM RELEASE 5.3 *""

ANSYS/ Mechanical i 10158-5.3 VERSION =HP 9000/*JO- 17 : 02 : 50 ' MAR 12, 1999 CP= 18.170 FOR SUPPORT CALL L.L. Beaudreau PHONE 860-285-3991 FAX 860-285-2901 ABB Combustion Engineering Nuclear Pover e - --.

agJ, ,,g2 24.44 L._. -ae4 tJ.-as.+.

F C-PENG CALC-019 Rev. 00 Page B16 of B17 i WSES-3 MNSA Hot Leg POT Top plate

-***** POST 1 NODAL STRESS LISTING *****-

l LOAD STEP = 1 SUBSTEP= 1 r

TIME = 1.0000 LOAD CASE = 0 SHELL NODAL RESULTS ARE AT TOP NODE S1 S2 S3 SINT SEQV 2 2914.2 27.832 -1.2574 2915.5 2901.1 38 2078.9 529.30 -25.456 2104.4 1889.1 40 2395.0 316.12 -2.1311 2397.2 2254.9 42 2176.3 420.94 -0.88102 2177.2 2000.0 -

1229 1240.3 20.750 -82.820 1323.1 1274.5 1231 1910.6 405.99 -82.490 1993.1 1799.3 1233 1779.4 282.32 -82.829' 1862.3 1709.2 1235 1664.6 189.95 -82.577 1747.2 1628.1 1237 1562.9 124.80 -84.401 1647.3 1553.3 1239 '1471.7 80.605 -87.429 1559.1 1482.3 1241 1389.0 51.667 -89.934 1478.9 1413.5 1243 1312.8 33.019 -89.418 1402.2 1345.1 1661 1031.3 0.16548 -2.5790 1033.9 1032.5 1663 '1176.3 12.886 -69.046 1245.3 1206.4 1665 1108.6- 6.3994 -44.293 1152.9 1128.4 MINIMUM VALUES

-NODE 1661 1661 1241 1661 1661 VALUE- 1031.3 0.16548 -89.934 1033.9 1032.5 MAXIMUM VALUES NODE- 2 38 42 2 2 VALUE - 2914.2 529.30 -0.98102 2915.5 2901.1

• • • • • ESTIMATED BOUNDS CONSIDERING THE EFFECT OF DISCRETIZATION ERROR *****

' MINIMUM VALUES NODE -1661 1661 1231 1661 1661 VALUE 1030.9 -0.18606 -208.54 1033.5 1032.2 MAXIMUM VALUES NODE 2 38 38 2 2

' V. A.L.U.E. . . 2 92 4. 7....

. .. . . . . . 7 0 7 ... 0 2. . . . . . 15

...2 2 6. . . .... .. 2 9 2 6 0. . . . . 2 911 6. . . . . . . . . . . . .

SELECT ALL ENTITIES OF TYPE = J..L AND BELOW ALL SELECT' FOR ITEM =VOLU COMPONENT =

IN RANGE O TO . O STEP ~ 1 0 VOLUMES (OF 0 DEFINED) SELECTED BY VSEL COMMAND.

ALL SELECT FOR ITEM = AREA COMPONENT = -

.IN RANGE 1 TO 17 STEP 1 12 AREAS (Or 12 DEFINED) SELECTED PY ASEL COMMAND.

ALL SELECT- FOR ITEH=LINE COMPONENT =

IN RANGE 1 TO 49 STEP. 1 38 LINES (OF 38 DEFINED) SELECTED BY LSEL COMMAND.

ALL' SELECT FOR ITEM =KP CCEPONENT=

IN RANGE- 1 TO 102 STEP 1 25 KEYPOINTS (OF 25 DEFINED) SELECTED BY KSEL COMMAND.

ALL SELECT FOR ITEM =ELEM COMPONENT =

IN RANGE 1 TO 576 STEP 1 576 ELEMENTS (OF 576 DEFINED) SELECTED BY ESEL COMMAND.

ALL SELECT -FOR ITEM = NODE COMPONENT =

IN RANGE ~ 1 TO 1839 STEP 1 1839 NODES (or 1839 DEFINED) SELECTED BY NSEL COMMAND.

PRINT F2 REACTION SOLUTIONS PER NODE 1

• • * *

• ANSYS - ENGINEERING ANALYSIS SYSTEM RELEASE 5.3 *****

l' ANSYS/ Mechanical 10158-5.3 VERSION =HP 9000/700 17:02:50 MAR 12, 1999 CP= 18.380 FOR SUPPORT CALL L.L. Beaudreau PHONE 860-285-3991 FAX 860-285-2901 l '

! . WSES-3 MNSA Hot Leg PDT Top platt 1

,. ABB Combustion Engineering Nuclear Pover l

l

. . _ . . --.- - . . ,, - . - - .. - . . ~. . - . - . . - -_

C.PENG-CALC-019 Rev. 00 Page B17 oflB17

• ** ** POST 1 TOTAL REI.CTION SOLUTION LISTING * * ***

LOAD S*EP= 1 SUBSTEP= 1 TIME = . 0000 LOAD CASE =- 0 1

THE FOLLOWING'X,Y,Z SOLUTIONS ARE IN GLOBAL COORDINATES

{

! NCDE F2 206- 152.95 i' 998 -347.05 TOTAL VALUES VALUE 500.00 EXIT THE ANSYS POSTI Di:2 ABASE PROCESSOR ROUTINE COMPLETED * * **

• CP = 18.400 END OF INPUT ENCOUNTERED * ****

PURGE ALL SOLUTION AND POST DATA SAVE ALL MODEL DATA $'

ALL CURRENT ANSYS DATA WRITTEN TO FILE NAME= cl.db I FOR POSSIBLE RESUME FROM THIS POINT NUMBER OF WARNING MESSAGES ENCOUNTERED. 0 NUMBER OF ERROR- MESSAGES ENCOUNTERED = 0

• ..............................................................._ - ..+

t l i I ANSYS RUN COMPLETED I I )

g.............................. \

=......g l l j l

l RELEASE 5.3 UP071096 HP 9000/700 l l

i l

)

i CP TIME (sec) = 18.810 TIME = 17:02:51 l 1- ELAPSED TIME (sec) = 36.000 DATE = 03/12/99 l t

1

• ......................................................... ...........I Start of Job Lopt Fri Mar 12 17:02:13 EST 1999 1 .ansys -m 48 I j

End of Job Logs Fri Mar 12 17:02:51 EST 1999 1-Date/ Times Fri Mar 12 17:02:52 EST 1999 Job Statistics:

CPU Time (liser) = 14.5 Socs.

CPU Time (System) = 1.0 Sees.

CPU Time ITotal) = 21.5 Sees.

Total Elapsed Time = 0:0:38 hete.inssee CPU Utilization = 566 l

L ABB Combusson Erigineering Nudear Power

C-PENG CALC-019 Rev. 00 Page C1 of C2 APPENDIX C HOT LEG PDT MNSA ASSEMBLY DRAWING

.I ABB Combustion Engineering Nuclear Pover

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C-PENG. CALC.ol9 Rev. og l

Page 0;oggg APPENDIX D QUALITY ASSURANCE FORMS l l

l i ABB Combustion Engineering Nuclear Power

C-PENG-CALC-019 Rev. 00 Page D2 of D6 Design Analysis Verification Checklist lastructions: If a major topic area (generally unnumberev bold face type such as Use of Cosaputer Software) is not applicable, then N/A (not applicable) next to the topic may x: checked and the check boxes for all items under it may be left blank.

Where there is no check box under N A for a numbered item, such a response is generally inappropriate. If N/A is checked in such a situation, docut tent the basis at the end of this checklist in the Comments section.

Author IR Overall Assessment Yes l N/A Concur.

1. Are the results/ conclusions correct and appropriate for their intended use? @ @

2. Arc alllimitations and contingencies on the results/ conclusions documented?

Assignment of Cognizant Engineers Independent Reviewen and Mentors 2.

If there m mukiple Cosmara Engmaars, has dwar scope bom Ifibere are muhiple ' ' '

??

O O'M

3. If there will be muhipie " .

- Reviewers, has their scope been documented?

g Approvers, has their scope been F-

4. If an ' '

??

' Revwww is the supervisor er Project Manneer, has authoruation as an b'-

O E E i documented?

' . . Reviewer been Q  % f 5.

If there is a Menter, has their scope and responsibildier been adequately h - '?

] Q g Use of Cosaputer Schware For soAware which has been validated under QP 3.13:

1. Is the soAware listed on an Approved QC.! SoAware List?

2.

@ ] g Is the soAware applicable for this analysis?

@ Q For Codn-IAe Canseruca vahdssed under QP 3.14.

1. Is the Code-Like Construct hated on an Approved QC 1 SoAware List?

@ M

]

2. Is the Code Like Construct applicable for this analysis?

]

3. Was the Code-tae Consruct used dressty a the casaround l=es=7 No

- IfNo above, is the copy idsreaal to the versson in the oceeround location? (Imave blank ifnot applumbic.)

] ]

Ifchanges were made to the Code Like Construct to meet specinc analysis needs, were such changes documented as non validstad soAware followeg paragraph 3.3.37 (lasve blank if not applicable. Complete the next section ]

of this Checkhst if *Yes".)

For soAware which has not been vahdated under QP 3.13 or QP 3.14.

1. In the computertype, proyaan name and revision idanuncation documented?

3.

]

Is a copy of the sonware included in the Design Analysis?

3. Have tests been - ' which are adaran* to demonstrate correct operation for the soAware's intended use?

4. Is the output from the tests included in the Design Analysis?

]

9.

Has the Coy.,.nt Engineer documented the resuha of the tests and the basis for concluding the soAware is operatag correctly for its intended use?

]

6. Did the soAware, as used in this analysis, pve correct resuha?

. O ABB Combustion Engineering Nuclear Pover l

l

1 1

C-PENG. CALC-019 Rev. 00 '

Page D3 of D6 Design Analysis Verification Checklist Author IR Use of Computer Software (contissed) 1.

YeslN/A Concur.

Were ., ' ' _ _ used in this Desip Analysm in any way - data display, plouang, computations. etc.? No

. If data duplay gely (no computations or plouing), check "Yes" and skip remauung questions.

. If used for - ._, -:-s:

e Are the - , "_ ' _ _-_ adequately documented?

e Are the resuhs correct?

. If used for plating-

]

e is the data to be ploemd corre.t?

e Are the plots correct in other respects? (titles, scales, labels, etc.)

s. Is a copy of the e ' ^ included in the Desip Analysas? (A copy of the file may be included or sufficient detail

]

included in the analyus d===anaam to pamut recreasang the spreadsheet) O O Objective of the Design Analysis

i. Ha wonnehon -io derme ihe ia.k been inauded - rs=wt a g

2. Haw the objectives been enumerated?

@ g

3. lies the artid1*/ and h use of the resuhs been d==aanad? g Assessment of Sismineast nesism chasses j

1. Have siendicant ' _ diences that might impact this analysis been consulared?

@ [

2. If any such chaness have been identdled, beve they been adequately ad&sened?

@ g Analytical Techeiagues (Methods) 1.

2.

Are the analytical eschasques (==sM) duecnhed in adlicient detail tojudge their appropnatenses?

@ Q {

Are the analyucal techniques used or their apphcation soverned by an NRC issued SER?

g J If yes, have the appbcable SERs been documsened?

O If yes, has the basis for concluding the analynas is in conformaam been d===nad?

3.

Have analytical techasques _- _ _ _ r __ ' by reference to esauric analyses, lead plant analyses or previous cycle analyses been previously venfied?

@ g a

4.

)

Are any niodificehons or departures froen previously approved analyuant techniques or Conv-e-I or Aue== sad Procedures documented andjustdied? @ $

5. If - approved analyucal *T= or ensamenns , _

l ]

_ are used, is their usejustafud an.* approved?

] 3 j 6.

j Does the issue date ofrufsrunced approved Conwatsonal or Automated Procedures predate their use ir. this analysis?

[ g Q

{ Selecties of nesism laputs-l 1. Are the desip inputs documented?

Q Q

2. Are the domic inputs correctly selected and traceable to their neuros?

Q

3. - Are the bases for miscsaan of all desip inputs documemed?

4. - Is previously unverified desip input uised in this analysis?

i M

If Yes, is at treated in accordance with QP 3.2, paragraph 3.4 for use of unverified desmg informsuon?

4 O

4 ~

ABB Combustion Engineering Nuclear Pover

C-PENG-CALC-019 Rev. 00 Page D4 of D6 Design Analysis Verification Checklist Author IR Selection of Design lapets (costimmed)

Yes l N/A Concur.

3.

l h the venncation stesus ofdesign inputs transmstsed from customers or CENP Nuclear Symems appropnate and documansed' g g 6.

Is the use af customarantrolled sources such as Tech Specs, LTSARs, etc. authorned, and does the authornataan l

specify amen 6nent level, revision number, etc.? @ Q

) *y_ -

1.

l If absre are no - rs= is this documsmed?

! 2.

] @ Q ,

I Are local === rana == '- " Aillyju=4=d and ven6ed?

3.

Q ] Q l Are lasernal and Emernal ."

, 1 fann?

__which must be clamrod by CENP or the customer listed on a Contagenews and

] @ g

4. Is the Project Manager r=Ta= ha for cleanns the .^ -

l id-ad=1 on the fann?

O t 9 1 m sam iCesclusions

1.

Are all resuhs aa====d in or retsenced in the Resulte/Concirmian section? (Where feasibis, in the ====esed order

! of the objectives.) Q Q

2. Art all Innetstacas on the resuhs/aa=banaa== and their appleanbility daaa===8-1 in this section?

i j 3.

Are all cesa=y=a== en the resubs that snust be cleared listed in the Rasuks/r'aa h- sostion or thera =*==g==a=

and." ,

fann referenced? @ g

4. Is the Project Manager respannbis for channs the ." , . _ _ or

Other Pb's===**

. adsudied on the fann?

] @ [

1.

Has a compenson of the resuhs whh those of a previous cycle or similar analysis been darv=wa=1 and sipuficant enn.on expiamsat @ g

a. - Have applicabie Code (e.g., ASME Cods) and mandenh bun apprepnesty refwenced and appind?

3.

@ Q h the infonneuen Dan roleant lasraem seenhenteckpound data adequeuly daa==ameand referenood' 4.

@ Q Are hand calculations correct and apprepnately daa====f?

3.

@ g h an appucahie =mpen amput and input inciuded?

i g g

6. Is all o@ PoOwWe used idNddied by nm mId reYugori edWW8[icWHR7 Itere m

1. Are au refer.m und to peinnn du mulysis lisud?

3.

Q ] Q Are the referomons as dirut as posesbie and appropnme to the source?

3.

Is the referunne natslian span 6c to en informanan imin==l aciudag revaian lowl or dune of neue, and where appropnsee, h i of the inomaan ofdu idananen a the refanns, mas as page, table or perupaph nunder? @ Q Independent Rcviewer's Sen**=need of Verification Activities:

1 i=dar== fait Reviewer to dancnbe details of vanfuntaan activities beyond the atmous on this checkhst inclueng, but not innhed to the review of asw methods, une of software under parayaph 3.3.3, spreadsbest uns, amassument of design and ' ' ., changes, ensmesnnsjudynsnes, the use of L

-viousiv unvenr.ed w i

L i

boC 4

l ABB Combustion Engineering Nuclear Power 1

i l ,

C-PENG-CALC-019 Rev. 00 Page D5 of D6 Design Analysis Verification Checklist The Form s.ad Format section of the Checklist below may be completed by a Checker under the direction of the Independent Reviewer.

Author IR Fonn/ Format Yes N/A Concur.

1.

2.

Is the document legible, reproducible and in a form sunatie for filmg and retneving as a Quainy Record?

@ Q Except as permrued by 3.1.3.a. are all pages identified with the document number, including revision number?

3.

Q Except as pomuned by 3.1.3.a. do all pages have a usuque page number?

Q l

4. Have all changes been authenticated by the irutials and date of the Quainy Records Controller?

@ g

5. Are all files on CD ROM identified by the path name?

O @ g

6. Are all computer deks identifwd with the analysis number?

g

7. Are any unverified sections of an otherwise venfied analysis clearly indicated?

@ g For a " Memorandum Revision" to a completed Design Analysis:

Q Q I

1. Have the title and hwnumber been preserved without change?

l 2. Does this revision most the cnteria for a " simple revision"?

l

3. Are the Author," ^ Reviewer and Management Approver and their roles identified? \

I For a revision to a compisted analyons in the " Complete Revision" and "Page Change Package" formats:

5 @ l

1. Where practacal, have changes and additions been identtfied by ==chaamma such as vertical lines, etc.?

2. Where precucal, have deletions been identified by mechamsms such as strike outs, etc.?

3. Have indications of change in previous revisions been removed?

4. Does the detribution of the revision include those on the detribution of the previous revision?

[

For a " Complete Revision *:

g

1. Have tM title and document number been preserved without change?

]

2. Has the revision number been incremented by one?

For a "Page Change Package":

1. Are pages numbered in accordance wnh the original analysis?

2. Are instructions provuled for the insertion and deletion ofrevaied pages?

3. Has a new Title Page been prepared?

4. Does the Package Contents Page reflect the change package contents?

]

! k Form / Format section completed by the Independent Reviewer.

Form / Format section completed by the Checker identified below:

Checker Name: Signature: _

ABB Combustion Engineering Nuclear Pover L

C-PENG-CALC-019 Rev. 00 Page D6 of D6 Reviewer's Comment Form

Title:

Analysis ofWaterford 3 Hot Leg PDT MNSA Document Number: C-PENG-CALC-019 Revision Number: 00 Comment Reviewer's Commen: Response Author's Response Response Number Reqmred? Accepted?

] W & Eu NO 2.I Y Det3W/W Y to abalu 5 PD'r MA)SA "# '

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ABB Combustion Engineering h Jclear Pover

C-PENG-DR-006 Rev. 01 Page D1 of DS9 l I ,

l ATTACHMENT D C-PENG-CALC-020, Revisions 00 and 01,

" Analysis of Waterford Unit 3 Hot Leg RTD MNSA" (59 pages including cover) l l

l l

l

A15 inter Office Correspondence l C-PENG-CALC-020 Revision 01 March 22,1999  !

To: QA File (2)

From: B. Nadgor cc: K.H. Haslinger TITLE: ANALYSIS OF WATERFORD UNIT 3 HOT LEG RTD MNSA The following typographical errors are noted in Revision 00 of the subject calculation.

LOCATION CORRECTION NOTE p.22 The gap of 0.020 t 0.005 in. should be set None instead ofincorrectly typed 0.20 t 0.005 in.

p.43 The allowaole number of cycles in the text This change does not affect should be 1950. the results of the analysis.

REVIEW AND APPROVAL This document is verified in accordance with ABB-CE QPM-101 Revision 03. Management authorizes the above editorial changes to the text.

Printed Name Signature Date Cognizant Engineer B Nadgor bcW 0 2[89' Independent Reviewer M.S. Mcdonald Wh,/I172e /d $[?R/99 Management Approval R.O. Doney . O hW, $~M -99 6

ABB Combustion Engineering Nuclear Power l

Design Analysis Title P ge

Title:

Analysis of Waterford Unit 3 Hot Leg RTD MNSA Document Number: C-PENG-CALC-020 Revision Number: 00

1. Verification Status:

@ Complete O NotRequhed O CompletewithContingencies/ Assumptions

2. Approval of Completed Analysis This Design Analysis is complete and verified. Management authorizes the use ofits results and attests of the Cogruzant Engineer (s), Mentor and Independent Reviewer (s).

Pnated Name Signature Date Cognazant Engineer (s) B. Nadgor hL/ g O g/pj Mentor @ None

""4 -"_- t Renewer (s)

M.S. Mcdonald 9gjpgg g3[g[9g

===* mat Appmal M Hashnge hg3((f g]jp44p gj//g/p o

3.

Package Contents (thi . section may be completed aAer Management apprmal):

Total page count, inc. wiing body, =g- - h. attachments, etc. 57 List a=caciwM CD-ROM disk Volume Numbers and path names: @ None CD-ROM Volume Numbers Path Names (to lowest durctory wiuch uniquely applies to tius document)

Other attachments (spectfy): @ None

4. Distnbuten QA(2) BevBoya K.H. Hashnger CL Mendrala M.S. Mcdonald l

l l

L ABB Combustion Engineering Nuclear Power

Page 2 of 47 i

RECORD OF REVISIONS l

Rev Date Pages Changed Prepared By Reviewed By Approved By 00 03/18/99 Original B. Nadgor M. S. Mcdonald K.H. Haslinger i

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i ABB Combustion Engineering Nuclear Power 1

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A ItIk F1IfIF C-PENG-CALC-020, Rev. 00 l

Page 3 of 47  ;

1 1

TABLE OF CONTENTS 1 Secuon Pane NO.

1 INTRODUCTION . . . . . .

.. ... . .. ... . . .. 5 1.1 OBJECTIVE . ,

1.2 ASSESSMENT OF S10NmCANT DESIGN CHANGES . l

.5 l 2 METHODOLOGY.. ..

l

_................6 2.1 GAP AT NORMAL OPERAUNO CONDmONS.. ..

.6 2.2 DETERMINATION OF IMPACT FORCE..

2.2.1 Net Ejection Force, F,.. .8 2.2.2 .8 Deflection ofComponents Due to Nonle Ejection.

2.2.3 Impact Force.. .9

.10 2.3 STRESS EVALUADON OF THE MNSA COMPONENTS..

2.3.1 Tie rods.. .10 2.3.2 Hex Head Bolts..

.10 2.3.3 Top Plate..

.10 2.3.4 Compression Collar..

.11 \

2.3.5 Upper Flange ..

.11

.I1 3 BASIC DATA AND ASSUMPTIONS .

.. . 12 3.1 SELECTION OF DESIGN INPtJTS., ..

3.1.1

. .12 Design and Operating Pressures and Temperatures..

3.1. 2 AfNSA Afaterials.. .

. . .)2

.12 3.1.3 Nonle andHot Leg Afaterials.. .

3.1.4 Afaterial Properties..

.12 3.1.5 AihSt ComponentDimensions..

.12

.25 3.1. 6 Nonle andNonle Component Dimensions.. ..

3.2 ASSUMPDONS., .

. . . . . . .15 3.2.1 Loading Conditions.. .16 3.2.2

.16 Consideration ofSeismic Loads.. .

.16 3.2.3 Friction Force.. .

3.2.4 Seahng pressure..

. .16 3.2.5

.I7 Preload... .

3.2.6 Dimensions..

.17

.17 4 SIGNIFICANT RESULTS . .

... .... . ..... I8 5 ANALYSIS. _

.. _ . .. . .. 19 5.1 MNSA DESCRIPTION..

.19 5.2 CONSIDERADON OF IMPACTLOAD..

3. 2.1

. . 20 Relative Displacements Due to ThermalExpansion..

5. 2.2 . 20 Cold Gap Setting vs. CalculatedDisplacements.. ..

5.2.3

. 22 Determination offmpact Force..

5.2.4 impact Force., . 22

. 26 5.3 STRUCTURAL STRESS ANALYSIS OF THE MNSA COMPONENTS . .,

5.3.1 Tie Rod..

. 27 5.3.2 . 27 Hex Head Bolt.. .

5.3.3 Top Plate.. . 32 l

. . 39 I

ABB Combustion Engineering Nuclear Power

A Ik k 7"EIf17 C-PENG-CALC-020, Rev. 00 Page 4 of 47 5.3. 4 Compression Collar..

5.3.5

. 40 Upper Flange.

.40 5.4 FATIGUE ANALYSIS..

. 41

5. 4.1 Normal Operating Pressure Force.. .

. 41 3.4.2 Tie Rods ..

5. 4.3

.42 Hex Head Bolt..

. 43 5.5 CONSIDERATION OF HYDROSTATIC TEST PRESSURE CONDmONS . . 45 5.6 CONSIDERATION OF FAULTED CONDmONS .. . . 45 6 REFEREN CES - . _ .. . . ... - . . . - ..... . . ... .. 46 LIST OF FIGURES j

FIGURE DESCRIPTION PAGE  !

1 i MNSA Concept Sketch for hot Leg RTD Nozzle. . .19 2 l Upper Flange . . .

.. .. .35 3 Compression Collar.. . .

. .36 LIST OF APPENDICES l

APPENDIX A: ASSEMBLY DRAWING . . .. .. .A-1 .

APPENDIX B:

CALCULATION OF THE TIE ROD AVERAGE TEMPERATURE .B-1 APPENDIX C: QUALITY ASSURANCE FORMS . ..

. C- 1 l

l ABB Combustion Engineering Nuclear Power l

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4 A ItIt M INIP C-PENG-CALC-020, Rev. 00 Page 5 of47 1 INTRGDUCTION I

1.1 Objective The objectice of this design report is to present the results of the evaluation of the Mechanical Nozzle Seal Asembly (MNSA) to be installed on the Hot Leg RTD nozzle at the Waterford Unit 3.

The MNSA is a mechanical device that acts as a complete replacement of the "J" weld between an l

Inconel 600 instrument nozzle and the Hot Leg pipe. Its function is to prevent leakage and restrain l the nozzle from ejecting in the event of a through-wall crack or weld failure of a nozzle. The potential for these events exists due to Primary Water Stress Corrosion Cracking.

1.2 Assessment ofSignificant Design Changes This report presents the detailed structural and thermal analyses required to substantiate the adequacy of the design of the Waterford Unit 3 Mechanical Nozzle Seal Assembly as a replacement of the nozzle "J" weld. This analytical work encompasses the requirements set forth in Reference 6.1 and is performed in accordance with the requirements of the ABB CENP Quality Procedures Manual QPM-101 (Reference 6.2).

Addenda to the original Piping Design Report (Reference 6.3) were reviewed and it was determined that their results have no impact on the current analysis and also that the current analysis does not impact their results. These Addenda Reports include:

CENC-1460 (2/81)

C-MECH-DR-001, Rev. 00 (12/93)

C-MECH-DR-004, Rev. 00 (12/93)

ABB Combustion Engineering Nuclear Power

i M INIP C-PENG-CALC-020 Rev. 00 Page 6 of 47 l 2 METHODOLOGY  !

The objective of this calculation is to analyze the Mechanical Nozzle Seal Assembly (MNSA) to be installed on the Hot Leg RTD nozzle at the Waterford Unit 3. The methodology used in this calculation is based on the method developed in Reference 6.19, for a similar design.

2.1 Gap at Normal Operating Conditions A cold gap between the clamp and the top plate has to be established to account for the relative thermal expansions of the components (see representative sketch below).

i I I i  !

N/

i I

v6D

/m The magnitude of the impact of the nozzle against the top plate is dependent upon this gap - the larger the gap, the greater the work done by the internal pressure, the greater the deflection of the components, and the greater the load on the components.

In order to determine the load impacting the MNSA components, if the nozzle ejects from the hot leg, the gap between the nozzle /thermowell and top plate with the components at normal operating temperatures shall be determined. The thermal expansion displacements of each of the relevant components shall first be determined and then added to or subtracted from the cold gap setting to determine the final operating conditions gap.

The linear thermal expansion displacements of each of the relevant components is calculated using the following equation from Reference 6.22 (p. 53):

S = u L AT where:

ABB Combustion Engineering Nuclear Power

A ItIt  !

M IfIF C-PENG-CALC-020, Rev. 00

)

Page 7 of C j S= the displacement (deformation) of a component caused by linear thermal expansion a= the coefficient oflinear thermal expansion L= the length of the component  ;

AT =

the temperature change from a reference temperature of 70*F to the applicable operating temperature The relative displacement, S,is determined by adding or subtracting the displacement of each of the individual components, as follows:

S, = Somi. + Somio u - Su. ,w - Sn.n,.* I

(* Sn in this analysis is taken to mean the combined thermal displacement of the lower and upper flanges) l j

For determining the maximum relative displacement, it is assumed that the temperature of the tie rod j and flanges increases from a reference temperature of 70*F to the ambient temperature of 120 F,i that the nozzle and thermowell reach the normal operating temperature of 611 F. Because of respective component lengths and coefficients of thermal expansion, these conoitions produce the maximum relative displacement (6, ) between the nozzle and the MNSA top plate, such that the overall nozzle /thermowell displacement exceeds the displacement of the top plate by a m amount. If there was no cold gap, the nozzle /thermowell would (theoretically) extend beyond the inboard surface of the top plate by a distance of 6, {

l The value ofS sets an upper limit on the cold gap setting, though the extreme temperature i differences evaluated above would not be seen during plant operations.

For determining the normal operations relative displacement, a more reasonable temperature distribution is selected, based on more realistic and appropriately conservative conditions. These conditions produce the normal operations relative displacement (6, op.) between the end of the thermowell and the MNSA top plate after heatup of the plant. If there were no gap between the end of the nozzle /thermowell and the top plate at cold conditions, the nozzle /thermowell would (theoretically) extend beyond the inboard surface of the top plate by a distance of Sr op,.

A final gap for the normal operating conditions will be determined by subtracting the value of normal operations relative displacement (6, op,) from the maximum cold gap setting.

ABB Combustion Engineering Nuclear Power

A ItIt M IDIF C-PENG-CALC-020, Rev. 00 Page 8 of 47 2.2 Determination ofImpact Force I

i At the moment at which an instantaneous break occurs, the internal fluid pressure in the h eject the nozzle outward, with the nozzle impacting the top p!ste. In order to determine the stress effects of this impact on the top plate, it will be necessary to first determine the net ejection fo acting on the nozzle. Once this net force is known, a relation can be defined between the work )

performed by the ejection force and potential energy stored in the deflection of the affected components (at the point of maximum deflection). Once the deflection of the components is know the impact force can be calculated and then used to determine stress effects.

2.2.1 Net Ejection Force, F.

l The nozzle will be forced out of the hot leg by the internal fluid pressure; this outward motion w opposed by the fricticn force which the Grafoil Seal exerts on the external surface of the nozzle. T net ejection force acting on the nozzle, F., is the difference between the " pressure force seal friction force, Fr:

F. = F, - Fr 2.2.1.1 Force Due toInternalPressure Motion of the nozzle at the moment at which there is an instantaneous brea created by internal pressure pushing against the entire cross section of the nozzle. Th determined as follows:

F, = p A where:

p= design pressure A= pressure area 2.2.1.2 FrictionForce The determination of the friction force (Fr) provided by the Grafoil seal is made based upo coefficient of friction for the seal against the nozzle and the radial load provided by the se the nozzle (produced by the compression of the seal).

Fr = P A l

t ABB Combustion Engineering Nuclear Power

M IDIF C-PENG-CALC-020, Rev. 00

l Page 9 of 47 l

where:

P= radial seal load (pressure)

=

coefficient of friction A= surface area of the seal in contact with the nozzle surface 2.2.2 Deflection of Components Due to Nozzle Ejection The total deflection of the impacted components due to the ejection of the nozzle will be determined  !

based upon the conservative understanding that all of the work put into the system by the net ejection force is converted completely into the potential energy of the deflected components (i.e.,

there are no losses). The base equation for evaluating the total deflection is derived from Equation (a) on page 471 of Reference 6 22 and is presented, as follows:

F, s = 1 K,, k '

where:

F. = the net ejection force )

s= total distance traveled by nozzle Ax = total deflection ofMNSA tie rods and top plate, and F=y the equivalent stiffness ofMNSA tie rods and top plate The total distance traveled by the nozzle, s, is equal to the distance of the gap at-temperature

(" Gap") plus the total deflection of the tie rods and top plate (Ax), or s = Gap +Ax.

The base equation may be re-written as follows:

F, (Gap + h) = 1 K,, & 2 1

1

=> -K,, & - 2F, & - F, Gap = 0 2

l In order to determine Ax, the stiffnesses of the tie rods and of the top plate are first calculated. Then, '

since the tie rods and top plate stiffness act in series against the impact load, the equivalent stiffness of tie rod-top plate system is calculated based upon a series stiffness equation from Reference 6.23 (p. 702):

1 l

1 1

+

, Ka K w, ABB Cornbustion Engineering Nuclear Power l

l

A ItIk l C-PENG-CALC-020, Rev. 00 Page 10 of 47 l The base equation developed previously is used to determine Ax:

Ax = [F. i (F.2 t 2K., F Gap)*] / K.,

2.2.3 Impact Force l l

The impact force, F,a, on the top plate and tie rods is then calculated: l F ,=K,,At l 2.3 Stress Evaluation ofthe MNSA Components \

The stresses and fatigue in the MNSA components are examined considering, pressure, thermal loads and the impact force defined above. '

l 2.3.1 Tie rods e

The Design tie rod load stresses are considered in accordance with NB-3231 and Appendix E. l The maximum tie rod load is determined by comparing the load created by preload and thermal expansion with the impact load. Whichever is greater is used throughout the remaining calculation as the tie rod loading.

The average and maximum stress in the tie rod and shear stresses in the threads are then calculated and compared to the corresponding ASME Code allowables.

e The fatigue analysis of the tie rod is then performed, considering loads which may exist after weld or nozzle failure has occurred. The calculated usage factor is compared to the ASME Code allowable of 1.0.

2.3.2 Hex Head Bolts The Design bolt load stresses are considered in accordance with NB-3231 and Appendix E.

• Maximum Bolt Load Due to the flexibility in the design of the flanged connection between the MNSA and the Hot Leg, the impact from ejection of the nozzle will increase the load on the bolts. The stiffness of the flange relative to the stiffness of the bolts will determine what percentage of the impact load will be effectively transmitted to the bolts.

l

! The total load on the bolt can be expressed by the following equation derived from Reference 6.24 (p. 579):

ABB Combustion Engineoring Nuclear Power l

7"EIfIF C-PENG-CALC-020, Rev. 00 Page11of47 e , 3 F,, = Preload + 'd

< Ksa , + A,' p,,,F j,,,

in the above expression, Kn.w. is considered to be the equivalent stiffness of the components which are put in compression due to the torquing / tightening of the hex head bolts; these components include the upper flange and the compression collar, which act in series with each other.

it may be concluded (from the above equation) that the greater the stiffness of the bolts as compared to the stiffness of the flange components, the greater the increase in load on the bolts from the impact (i.e., as Kn.,ihn

• 0, the multiplier for F%a -+ 1).

The stiffness of each component is considered in the detailed analysis to calculate the maximum hex head bolt load.

The average and maximum stress in the hex head bolt and shear stresses in the threads are then calculated and compared to the corresponding ASME Code allowables.

The fatigue analysis of the hex head bolt is then performed, considering loads which may after weld or nozzle failure has occurred. The calculated usage factor is compared to the ASM Code allowable of 1.0.

2.3.3 Top Plate The shear and bending stresses in the top plate are calculated due to the impact load a to the corresponding ASME Code allowables.

2.3.4 Compression Collar The shear stress and bearing stress due to the preload of the hex head bolts are calculated and compared to the corresponding ASME Code allowables.

2.3.5 Upper Flange The shear stress due to the preload of the hex head bolts is calculated and compared to the corresponding ASME Code allowables.

ABB Combustion Engineering Nuclear Power

A ItIk M IDIF C-PENG-CALC-020, Rev. 00 l Page11 of47 i ' ~

F,,, = Preload + K*"

< K,a, + K m, ,

F ,,,

l In the above expression, Kn.. is considered to be the equivalent stiffness of the components l

which are put in compression due to the torquing / tightening of the hex head bolts; these components include the upper flange and the compression collar, which act in series with each other, it may be concluded (from the above equation) that the greater the stiffness of the bolts as compared to the stiffness of the flange components, the greater the increase in load on the bolts from the impact (i.e., as Kn /lb -+ 0, the multiplier for F, a -+ 1).

The stiffness of each component is considered in the detailed analysis to calculate the maximum hex head bolt load.

The average and maximum stress in the hex head bolt and shear stresses in the threads are then calculated and compared to the corresponding ASME Code allowables.

! e The fatigue analysis of the hex head bolt is then performed, considering loads which may exist after weld or nozzle failure has occurred. The calculated usage factor is compared to the ASME Code allowable of 1.0.

2.3.3 Top Plate The shear and bending stresses in the top plate are calculated due to the impact load and compared 1

to the corresponding ASME Code allowables.

2.3.4 Compression Collar The shear stress and bearing stress due to the preload of the hex head bolts are calculated and compared to the corresponding ASME Code allowables.

2.3.5 Upper Flange The shear stress due to the preload of the hex head bolts is calculated and compared to the corresponding ASME Code allowables.

l l

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A ItIt M IDIF C-PENG-CALC-020, Rev. 00 Page 12 of 47 3

BASIC DATA AND ASSUMPTIONS 1 3.1 Selection ofDesign Inputs i

3.1.1 Design and Operating Pressures and Temperatures The Mechanical Nozzle Seal Assembly is considered a pressure-retaining component. T Pressure is 2500 psia and Design Temperature is 650 F. Operating pressure and temperatu 2250 psia and 611 F, respectively (Reference 6.20). Ambient design temperature is 120 F (Reference 6.7).

3.1.2 MNSA Materials MNS A materials are taken from Reference 6.8.

i item Material 1 Compression Collar SA-479, Type 304 Flat Lower Flange l SA-479, Type 304 Upper Flange SA-479, Type 304 Top Plate SA-479, Type 304 Hex Bolt SA-453, Grade 660 Hex Nut SA-453, Grade 660 Tie Rod SA-453, Grade 660 3.1.3 Nozzle and Hot Leg Materials Hot Leg RTD nozzle and fitting materials are taken from References 6.4,6.5,6.6.

11sm Material Reference Nozzle SB-166 6.5, 6.6 Thermowell SB-166 6.4 3.1.4 Material Properties Material properties used in this analysis include coefficients of thermal expansion (a), moduli of elasticity (E), design stress intensity values (S.) and Yield Strength y Values (S ). These prop presented below and are found in the Appendices ofReference 6.9.

ABB Combustion Engineering Nuclear Power

ARR M IDIF C-PENG-CALC-020, Rev. 00 Page 13 of 47 3.1. l.1 Coefficient ofLinear ThermalExpansion, a The following table presents the temperature-dependent coefficients oflinear thermal expansion various materials:

temperature SB-166 SA-479 Type 304 SA-453, Grade 660

(*F) (Alloy 600) (304 SS) (Alloy 660) 100 6.90 8.55 8.24 200 7.20 8.79 8.39 300 7.40 9.00 8.54 400 7.57 9.19 8.69 500 7.70 9.37 8.82 600 7.82 9.53 8.94 611 7.83* 9.55* 8.95*

650 7.88 9.61 9.00

• by interpolation l

{

All coefficients are Coefficient 4

B values from Table I-5.0, where Coefficient B is the mean coefficien of thermal expansion X 10 in./inJ'F in going from 70*F to the indicated temperature.

l 3.1.4.2 Modulus ofElasticity, E l

The following table presents the temperature-dependent moduli of elasticity for SA-479 Type and SA-453, Grade 660:

temperature E

("F) r 70 28.3 200 27.6 300 27.0 400 26.5 500 25.8 i

600 25.3 650 25.0*

, 700 24.8

• by interpolation ABB Combustion Engineering Nuclear Power

1 M IDIF C-PENG-CALC-020 Rev. 00 Page 14 of 47 All moduli of elasticity values are from Table I-6.0, where E = value given X 10' psi.

3.1.4.3 Design StressIntensity Value, S,,,

l i

The following table presents the temperature-dependent design stress intensity values for various materials:

temperature SA-479 304 SA-453, Grade 660 l

(*F) S. S.

100 20.0 28.3 200 20.0 27.6 300 20.0 27.3 400 18.7 27.2 500 17.5 27.I 600 16.4 27.0 650 16.2 26.9*

700 16.0 26.8

• by interpolation The design stress intensity values for SA-479 Type 304 are from Table I-1.2; and the design s intensity values for SA-453, Grade 660 are from Table I-1.3. All S. values are given in ksi.

3.1. 4.4 YieldStrength Value, Sy The following table presents the temperature-dependent yield strength values for SA-479 T temperature SA-479 304

( F) Sy 100 30.0 200 25.0 300 22.5 400 20.7 500 19.4 600 18.2 650 17.9 700 17.7 ABB Combustion Engineering Nuclear Power

A IkIt M INIP C-PENG-CALC-020, Rev. 00 Page 15 of 47 The yield strength values for SA-479 Type 304 are from Table 1-2.2. AllyS values are given in ksi.

3.1.5 MNSA Component Dimensions The bolts and tie rods have the following dimensions (References 6.8, and 6.18):

Bolts [0.500-20 UNF-2A] Tie Rods [0.375-16 UNC-2A]

Basic major diameter 0.5000 in 0.3750 in Basic minor diameter 0.4374 in 0.297 in Basic pitch diameter 0.4675 in 0.3344 in Tensile stress area 0.1599 in' O.0775 in 2

Kn max (max minor 0.457 in 0.321 in diam. of intemal thread)

Es min (min pitch diam. 0.4619 in 0.3287 in ofextemal thread)

En max (max pitch diam. 0.4731 in 0.3401 in ofintemal thread)

Ds min (min major diam. 0.4906 in 0.3643' in of extemal thread) 3.1.6 Nozzle and Nozzle Component Dimensions ,

Various component dimensions are taken from Reference 6.4, 6.5,6.6, and 6.13 as indicated below.

RTD Nozzle Ref.

Pressure Diameter 0.997 in 6.13 Length of Nozzle (overall) 8.375 in 6.5, 6.6 Length of Thermowell 0.5 in 6.4 2

Pressure Area = (x r ) 0.781 in' -

i a

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M IfIF C-PENG-CALC-020 Rev. 00 Page 16 of 47 3.2 Assumptions 3.2.1 Loading Conditions If no crack is present, it is assumed that, except for preload and thermal expansion, the MNSA components are not loaded during normal operating conditions. An impact load may be experienced if there is a complete and instantaneous failure in the J-weld or a 360* circumferential crack in the nozzle, such that the nozzle would be forced outward against the top plate, closing any gap between the two components. After this event occurs, a normal operating load, without impact, would exist, with the internal pressure holding the nozzle up against the top plate; this load would be cyclical -

from essentially zero at Cold Shutdown to a maximum at normal operating conditions.

For the purposes of this analysis, it is assumed that there is a complete and instantaneous failure of the J-weld (or a 360* circumferential crack in the nozzle) such that the nozzle is ejected outward and impacts against the top plate, which will also then load the tie rods and other components. The impact of the nozzle against the top plate conservatively represents the maximum load that the restr-ining components would experience.

3.2.2 Consideration of Seismic Loads Because of the nature of the accelerations from seismic events, only the tie 5ods will be evaluated for the stress effects of the seismic event. The remaining MNSA components will not be significan affected. Separate seismic tests on similar MNSA configuration were performed to demonstrate an adequate seal performance (see Reference 6.16).

3.2.3 Friction Force The effects of any impact of the nozzle against the top plate are dependent upon certain assum regarding the determination of the ejection force acting on the nozzle.

In an " ideal" (and worst case) break scenario, the crack would be complete, instantaneous and oriented such that no base or weld metal could interfere with the motion of the nozzle. In this case, the only resistance offered to the nozzle motion would be provided by the attached piping and b Grafoil seal.

In reality, the crack characteristics necessary for the " ideal" scenario would not exist, and, instead, there would be potentially significant resistance offered to the motion of the nozzle by the crack surfaces and by integral material, if motion would be allowed at all.

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A ItIk M IDIF C-PENG-CALC-020 Rev. 00 I

Page 17 of 47 In this analysis, a scenario which is somewhere "between" the " ideal" scenario and the "rea will be evaluated: it will be conservatively assumed that motion will be allowed but in the an opposing force provided by the crack metal and by the Grafoil seal. This opposing force will accounted for by applying a coefficient of friction for the Grafoil-to-nozzle contact, as described below:

A coefficient of friction (p) of 0.30 for Grafoil-to-nozzle contact will be used to determine the forc which opposes motion of the nozzle. This value of 0.30 used for the (kinetic) motion of the no ejection is higher than the values piovided by the Grafoil seal manufacturer in Reference 6.2 lists (static) coefficients in the range of 0.05 to 0.20 (see Table III of Reference 6.21). Ho application of a friction force based upen the coefficient value of 0.3 will be maintained on the ba that the actual force which would tend to Umit or prevent motion in the "real" scenario would be lugher.

3.2.4 Scaling pressure Compression of the Grafoil creates a radial pressure against the nozzle surface of at for a preload of 30 ft-lb, based upon Reference 6.10. (This value will be used to determine a force on the nozzle from the Grafoil seal.)

3.2.5 Preload Nominal values of tie rod / bolt preload are used in this analysis since maximum values not significantly increase corresponding preload stresses. (A check of the results indicate that the maximum preload values will result in stresses which will remain below, or will be on the o of, their respective allowables. Therefore, use of the nominal values is acceptable).

3.2.6 Dimensions Referenced overall length of the assembly (4.75 inches) from Reference 6.8.1 was used Nominal design dimensions of the parts were used for the calculations of the relative displa and for the stress analysis, except when noted. The use of the as-measured dimensions from Reference 6.13 produces slightly different values but does not affect the results of the current analysis.

l l

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ARR M IDIF C-PENG-CALC-020, Rev. 00 Page 18 of 47 4

SIGNIFICANT RESULTS The results presented below were determined using the assumptions defined an 3.2. There are no additional contingenc es or assumptions that are applicable to these All stresses are satisfactory and meet the appropriate allowable limits set forth in ASME Boiler and Pressure Vessel Code (Reference 6.9).

Results ofthis analysis are summarized below:

Component Stress Calculated Stress / Allowable Stress /

Category Usage factor Usage Factor (stress in ksi) (stress in ksi)

Tie Rod Design  !

6.45 26.9 Average 35.48 53.8 Maximum 39.33 80.7 Thread Shear 9.55 16.14 Usage Factor 0.359 1.00 '

Hex Head Bolt Design 22.51 28.3 Average 43.6 53.8 Maximum 48.98 80.7 Thread Shear 14.15 16.14  !

Usage. Pactor 0.5 Top Plate 1.00 Shear 3.00 9.72 '

Bending 23.88 i

Compression Collar 24.3 Shear 7.02 9.72 Bearing 15.08 Upper Flange 17.9 Shear 4.16 1 9.72 i Thread Shear 6.64 9.72 Lengths of engagement used in analysis:

Tie Rod - Upper Flange: 0.5 in, (0.34 in. minimum)

Hex Head Bolt - Hot Leg Pipe: 0.5 in. (based upon bolt thread shear; 0.44 ABB Combustion Engineering Nuclear Power

A IkIn M IDIF C-PENG-CALC-020, Rev. 00 Page 19 of 47 5 ANALYSIS 5.1 MNSA Description The MNSA is a mechanical device that acts as a complete replacement of the "T' weld between a inconel 600 instrument nozzle and the hot leg pipe. It replaces the sealing function of t a Grafoil seal compressed at the nozzle outside diameter to the outer hot leg surface. The MNSA

' also replaces the weld stmeturally by means of threaded fasteners engaged in tapped holes in outer not leg surface, and a restraining plate held in place by threaded tie rods. This feature p the n'ozzle from ejecting from the hot leg, should the "T' weld fail or the nozzle develo circumferential crack.

A general sketch ofRTD MNSA is depicted in Figure ! below.

FIGUPE 1 MNSA Concept Sketch for Hot Leg PTD Nozzle

~ ~e rEh -

T h: T&"s h

llij hM i h !k c orn p res sion conor q g L4j (SP"*)

l i N

Upper Flan g e ' Hex Heod B ol t s 7

q  :

Lower FIo n g e rh,"1 >

lM9&f8% l, N QQ I . Ml N

NI N / ,/ / / /,

l/b/4.w'/ /-

k h'N'Ah//

s LU Seol Re(split) t oin er ] Gr of oil Se ol (split)

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~

D%lElF C-PENG-CALC-020. Rev. 00 Page 20 of 47 5.2 Consideration ofImpact Lead 5.2.1 Relative Displacements Due to Thermal Expansion l According to Section 2.1, the determination of the maximum relative displacement is based on the assumption that the temperature of the tie rod arv' "inges increases from a reference temperature of l 70 F to the ambient temperature of 120*F, ant : u the nozzle and thermowell reach the normal operating temperature of 611 F. Because of respective component lengths and coefficients of thermal expansion, these conditions produce the maximum relative displacement (6, ) between the nozzle and the MNSA top plate, such that the overall nozzle /thermowell displacement exceeds the displacement of the top plee by a maximum amount. If there was no cold gap, the nozzle would (theoretically) extend beyond the inboard surface of the top plate by a distance of S, .

Hot Lee RTD MNSA Maximum Relative Disolacement component temperature cx L AT S 4

(*F) (10 inlin]"F) (in.) ('F) (in.)

nozzle 611 7.83 4.25* 541 0.0180 thermowell 611 7.83 0.5 541 0.0021 tie rod 120 8.27 3.635** 50 (-)0.0015 flange 120 8.60 1.I15 50 (-)0.0005 6, = 0.018

• nozzle length = 4.75" (Ref. 6.8) - 0.5" (thermowell) = 4.25" "the tie rod length is based upon the overall assembly length less the thickness of the both lower and upper flanges.

The value of 6, sets an upper limit on the cold gap setting, though the extreme design temperature differences evaluated above would not be seen during plant operations.

For determining the normal operations relative displacement, a more reasonable temperature distribution is selected, based on more realistic and appropriately conservative conditions. In this case, the following temperatures are assumed:

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A IkIt HEFEB C-PENG-CALC-020, Rev. 00 Page 21 of 47 component temperature (T) nozzle 611 thermowell 500 tie rod 350 flange 550 The temperatures used in the determination of the normal operations relative displacement are average component values and are based upon the following:

Nozzle - 61l'F: due to the nozzle location in the hot leg pipe, direct exposure to the primary coolant and to conduction, the average temperature of the nozzle (external to the pipe) would be i at or very near the hot leg temperature (611*F).

Thermowell - 500*F: this component would receive heat conducted through the nozzle and the nozzle fluid. Some heat would be lost during conduction but, c,verall, the thermowell average temperature would be elevated to a relatively high temperature. l Flange - 550*F: this component would receive heat conducted through the pip . seal retainer and lower flange, as well as by convection / radiation through gaps. Some heat would be lost tp t insulation ambient though, overall, the flange average temperature would be elevated to a relatively high temperature, given its proximity to these hotter components. A high temperature for this component is conservative as it will increase the impact gap.

Tie rod - 350*F: this component would receive some heat conducted through the upper fla but it also exposed to ambient temperature. The average temperature of 350T is based upon i heat transfer evaluation which assumes a 500T heat source temperature (the temperature of th top of the upper flange) and heat lost to ambient at 120*F (see Appendix B). A high temperature for this component is conservative as it will increase the impact gap.

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k

1. "tIfIF C-PENG-CALC-020, Rev. 00 i Pige 22 of 47 Hot Lee RTD MNSA NOP Relative Disolacement <

l component temperature L

' cx AT 6 4

(*F) (10 in/m/F) (in.) (*F) (in.)

nozzle 611 7.83 4.25 541 0.0180 thermowell 500 7.70 0.5 430 0.0017 tie rod 350 8.62 3.635 280 (-)0.0088 flange 550 9.45 1.115 480 (-)0.0051 l 6, - 0.006 These conditions produce the normal operations relative displacement (6,,) between the nozzle /thermowell and the MNSA top plate after heatup of the plant.

5.2.2 Cold Gap Setting vs. Calculated Displacements I in accordance with Reference 6.8.1 (note 1), the gap of 0.20 0.005 in. has to be set between the thermowell and top plate at cold condition. It is recognized that the muumum cold gap of 0.015 less than the maximum relative displacement for the RTD MNSA (6,. = 0.018 in)., which would mean that, at the conditions for the maximum relative displacement, the thermowell would be in direct contact with the top plate. However, as noted before, the conditions used to obtain the maximum relative displacement are not anticipated during operation.

The maximum cold gap setting of 0.025 in. indicates that a gap of 0.025 in. - 0.006 in. = 0.019 in.

l can exist during normal operating conditions for the Hot Leg RTD MNSA.

I A final gap value of 0.019 in for normal operating conditions will be used in the subsequent determination of the impai.t force.

5.2.3 Determination ofImpact Force Impact force is calculated in accordance with the methodology described in Section 2.

5.2.3.1 Net Ejection Force, F.

The net ejection force acting on the nozzle, F., is the difference between the " pressure for the seal friction force, Fr:

l F. = F, - Fr ABB Combustion Engineering Nuclear Power

A ItIt M IDIF C-PENG-CALC-020, Rev. 00 Page 23 of 47 5.2.3.1.1 Force Due to Internal Design Pressure Motion of the nozzle at the moment at which there is an instantaneous break is due to th created by internal pressure pushing against the entire cross section of the nozzle. (From Section 2

3.1.6, the pressure area is 0.781 in ). This force, F,, is determined as follows:

F, = (2500 psi) (0.781 in2 ) = 1,953 lb 5.2.3.1.2 Friction Force Fr= P A where:

P= radial seal load (pressure) = 3100 psi (Reference 6.10) p= coefficient of friction = 0.3 A= surface area of the seal in contact with the nozzle surface

= xDh h= 0.25 in (Reference 6.8)

D= 0.997 in Therefore:

Fr = (3100 psi) (0.3) x (0.997 in) (0.25 in) = 728 lb Based upon the forces calculated above, the net ejection force is:

F. = 1953 - 728 = 1225 lb.

Use F. = 1250 lb 5.2.3.2 Defection ofComponents Due to No::le Ejection The re-written base equation for evaluating the total deflection is 2

f K,, Ar - F, Ar - F, Gap = 0 in order to determine Ax, the stiffnesses of the tie rods and of the top plate are first calculated and then the equivalent stiffness of the tie rods-top plate system is calculated.

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FLIFID C-PENG-CALC-020 Rev. 00 Page 24 of 47 5.2.3.2.1 Stiffness of 4 Tie Rods The total stiffness of the four (4) tie rods, Ka, is based upon an equation from Reference 6.22 (p.

31):

AE K=4 where:

A= z 0.0878 in , cross-sectional area of the tie rod, based on the basic pitch diameter E= 26.8 x 10' psi (at 350*F)

L= 4.515 in, length between the top of the top plate and upper flange So:

2 (0.0878in )(26.8XI0'-lb -

)

K =4 '"'

= 2.085E6 5 4.51 Sin in l

l 5.2.3.2.2 Stiffness of the Top Plate The equations for calculating the deflection of the top plate are found in Reference 6.11, Table 24, Case la: ,

w a' C y= g ( ,4 (, -4) where:

Et 3 26.8X 10, Ib( (0.88)'in' D= = I"

= 1672500in -Il 12(1-y ) 2 2 12(1 - 03 )

and Ci , C7, L,, and L3 are constants, and are calculated using the equations of Reference 6.11, where:

a = 1.906 in b = 0.56 in ro= 0.663 in - radius of the thermowell contacting with the top plate t = 0.88 in y = 0.3 E = 26.8 X 10' psi (at 350*F)

Ci = 0.7781 ABB Combustion Engineering Nuclear Power

ABB C-PENG-CALC-020, Rev. 00 Page 25 of 47 C7 = 1.4149 L3= 0.0265 L, = 0.2923 Solving for the stiffness of the top plate:

2w*

K., ,a, = = 7500000 lb a'

D (C,L'C, -LJ 5.2.3.2.3 Equivalent Stiffness Tie Rods-Top Plate System i

The equivalent stiffness of tie rod-top plate system is determined as follows:

I K,, =

+

Ka K,, e l

Therefore:

1 K,, 1630,000 lb I m 2.085E6 7.5E6 5.2. 3.3 TotalDeflection, At The base equation developed previously is used to determine Ax:

1 2

-K,, Ar - F, Ar - F, Gap = 0

=> Ax = [F t (F.2 + 2K,q F, Gap)") / K.q Given:

F = 1250lbs K., = 1,630,000 lb / in Gap = 0.019 in.

=> Ax = 0.0062 in.

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k naher C-PENG-CALC-020, Rev. Ou l

Page 26 of 47 l 5.2.4 Impact Force The impact force, F.,,,u, on the top plate and tie rods is then ca'~ ted:

F,,,,, = K,, Ar Therefore:

F,,, = 1,630,000 S(0.0062in) = 10.1 ikips m

The value of F, ,,, = 11,000 lb will be used in subsequent analysis of the MNSA components.

s l

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7"EIfIF C-PENG-CALC-020. Rev. 00 Page 27 of 47 5.3 StructuralStress Analysis ofthe MNSA Components The Design Loads for the various MNSA components will be a function of either bolt preload, the impact load, and/or thermal expansion loads, depending upon the effects of the source load upon a particular component.

5.3.1 Tie Rod 5.3.1.1 Design Bolt LoadStresses The design bolt load for the tie rod is considered to be the hydrostatic load which results from Design Pressure only, since the tie rod is not used for gasket-joint purposes.

Section 5.4.2.1 determines the service stress in the tie rod for a pressure which bounds the Design Pressure; this stress value is compared to the design bolt load stress allowable:

6.45 ksi < 26.9 ksi (S. at 650*F)

The hydrostatic load stress is below the stress allowable, which indicates that the actual bolt area (A3) exceeds the minimum required bolt area (A.).

5.3.1.2 StressDue to ThermalExpansion The differential thermal expansion between the tie rod and the upper flange (or top plate) will create an additional tensile load on the tie rod. For the analysis of the tie rod, this additional load is assumed to be completely taken up by deformation of the tie rod.

The stress effect of this differential thermal expansion is determined below.

From Reference 6.22:

o t = E a. A T The at to be used is the differential in coeflicients for the materials of the tie rod and flanges, Aa. Therefore:

ot.. = E A a A T Given:

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I A ItIt l HEBEF C-PENG-CALC-020, Rev. 00 Page 28 of 47 E = 25.25 E6 psi at 61l'F l

Aa = an , - %,,a = (9.55 E 8.95 E-6) = 0.6 E-6 (a at 61l'F) '

AT = 611 - 70 = 541 *F ct.. = 8.2 ksi The effective tensile force due to this thermal expansion is determined by:

2 P = ct.. A, = (8.2 ksi)(0.0775 in ) = 0.636 kips l t3.1.3 Preload The tie rod and nuts are being preloaded to 75 in-lbs. To determine the load in each tie rod, the following equation is used (Reference 6.15):

T = 0.2 F d

=> F =T / 0.2 d 1 Given:

T = the applied torque = 75 in-lbs d = is the nominal major tie rod diameter = 0.375 in.

l F = (75 in-lbs) / (0.20) (0.375 in) = 1.00 kips.

5.3.1.4 Impact Load Since the impact load is distributed evenly to each of 4 tie rods, the impact load on the tie ro 11.0/4 = 2.75 kips 5.3.1.5 Maximum Tie RodLoad The load on the tie rod will be the greater of the load due to preload plus thermal expansion and the load due to the irr. pact:

l Preload + thermal expansion = 1.00 + 0.636 = 1.636 kips I l

Impact = 2.75 kips (> 1.636 kips) l ABB Combustion Engineering Nuclear Power i

A ItIt M IDIF C-PENG-CALC-020 Rev. 00 Page 29 of 47 Therefore, the maximum tie rod load is impact load P = 2.75 kips.

5.3.1.6 Average Stress,ar (NB-3232.1)

The average (axial) stress (oi) in the tie rod is due to the maximum tie rod load:

ci = P/A i 2

Ai = 0.0775 in P = 2.75 kips 2

at = (2.75 kips /0.0775 in ) = 35.48 ksi < 2 Sm = 53.8 ksi (S. at 650*F) 5.3.1.7 Maximum Stress (NB-3232.2)

The maximum stress in the tie rod is essentially a stress intensity due to a combination of the average stress, bending stress from the OBE (Operating Basis Earthquake) event, and the torsional shear stress due to residual torque (from preload).

5.3.1.7.1 Seismic Bending Stress Prior to any (complete) weld failure, a seismic event will cause accelerations of the MNSA. Most components will experience little or very little effects from these seismic accelerations. However, because of motions associated with the top plate, the inboard end of each tie rod will be subjected to bending stress. This bending stress (o6) will be conservatively added to the average and torsional shear stress for determining the maximum stress in the tie rod.

The bending stress at each of the tie rods is determined by applying the rnaximum acceleration occurring at OBE event to the top plate, following with the even distribution of the resulting force to each of 4 tie rods, and then calculating the stress at the tie rod inboard end.

on = 1/4 Mc/I = 1/4 (16.61 x 0.149) / 3.82 x 10" = 1.62 ksi, where M=FU2 = bending moment (Reference 6.11, Table 3, Case Ib)

M=FU2 = (8.14 x 4.08)/2 = 16.61 in-lbs - bending moment F = W(1+a) = 4.4 (1+0.85) = 8.14 lbs - acting force; ABB Combustion Engineering Nuclear Power

l M IDIF C-PENG-CALC-020, Rev. 00 Page 30 of 47 W = pV = 0.29 x 15.12 = 4.4 lbs - weight of the top plate 2 2 a = ]G,2 +G,2 = ]0.6 + 0.6 = 0.85g - maximum acceleration at OBE, according to Reference 6.14, accelerations in any horizontal and vertical direction shall be applied simultaneously.

G, = G y= 0.6g - conservatively assumed in both horizontal directions and in the vertical direction OBE acceleration values, taken from Reference 6.17.

p = 0.29 lb/m' - density of the stainless steel 2 2 V = n/4 x (4.81 - 1.12 ) x 0.88 = 15.12 in' - volume of the top plate (conservative)

L = (4.75 - 1. I 15 + 0.88/2) = 4.08 in - length of the tie rod from upper flange to the center ofgravity of top plate I = nd*/64 = 3.82 x 10" in' - moment ofinertia of the tie rod d = 0.297 in - basic minor diameter of the tie rod c = d/2 = 0.297/2 = 0.149 in 5.3.1.7.2 Residual Torque, Ta The residual torque due to preload may be calculated using the following equation for standard threads (Reference 6.15, Equation 6):

Ta = 0.5625 T = 0.5625 (75) = 0.042 in-kips where:

0.5625 = multiplier based upon a coefficient of friction of 0.15 and standard bolt dimensions T = 75 in-lb = applied torque 5.3.1.7.3 Torsional Shear Stress,17 The torsional shear stress is determined using the following equation derived from Equation 7 of Reference 6.15; tr = 16Ta / nd' =16 (0.042) / n (0.32)' = 6.53 ksi where:

d = average of basic pitch diameter (0.3344 in ) and minor diameter (0.297 in.) =

0.32 in.

l ABB Combustion Engineering Nuclear Power

b l

l FL DIF C-PENG-CALC-020, Rev. 00 Page 31 of 47 5.3.1.7.4 Maximum Stress, a.x The maximum stress intensity (c. ) is determined using the following equation (Reference Equation 8):

c.x = 2 ( #'

2 ')*+(tr) = 39.33 ksi < 2.7 Sm = 72.6 ksi (S. at 6507)

< 3.0 Sm = 80.7 ksi (S. at 6507)

S.3.1.8 Shear Stress (r) - Threads At Top Plate (hex nuts)

The tie rods pass through the top plate and are held in place with hex nuts at the top and at bottom. The impact load, in directly loading the top plate and top nut, will create stresses in th section o~f the tie rod which are in addition to the tie rod / nut preload stresses. The nuts are o same material as the rods. Therefore, the parameters associated with the external threads of the r are used (i.e., because of the smaller shear area).

From Reference 6.18:

AS = x n Le Kn max [(1/2n) + 0.57735 (Es min - Kn 2max)] = 0.288 in where:

n = numb,;r of threads per inch = 16 Le = the length of engagement (nut thickness) = 0.5 in (Ref. 6.8)

Kn max = maximum minor diameter ofinternal thread = 0.321 in Es min = minimum pitch diameter ofexternal thread = 0.3287 in P = 2.75 kips 2

t = 2.75 kips / 0.288 in = 9.55 ksi < 0.6 Sm = 16.14 ksi At Uoper Flange On the other side, the tie rods thread into the Upper Flange. The lower strength Uppe threads are evaluated below. (The external tie rod threads in the Upper Flange hav same stress as the external tie rod threads in the top plate region, which were evaluated p From Reference 6.18:

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M IDIF C-PENG CALC-020, Rev. 00 l

Page 32 of 47 AS = x n Le Ds min [(1/2n) + 0.57735(Ds min - En max)]2 = 0.414 in where:

n = number of threads per inch = 16 Le = the length of engagement. Assume equal to 0.5 in En max = maximum pitch diameter ofinternal thread = 0.3401 in Ds min = minimum major diameter of external thread = 0.3643 in P = 2.75 kips 2

t = 2.75 kips / 0.414 in = 6.64 ksi < 0.6 Sm = 9.72 ksi The minimum allowable length of engagement of the tie rod into the Upper Flange may be ca as a simple proportion:

Le

(Shear Stress / Allowable Stress) x Assumed Length ofEngagement

= (6.64/9.72) x 0.5 = 0.34 in.

5.3.2 Hex Head Bolt l

S.3.2.1 Design Bolt LoadStresses Design Bolt Load for the Design Pressure, W.i .

W.i= H + H, H = 1.953 kips (from Section 5.2.3.1.1)- hydrostatic end force H, = 2b x 3.14GmP - compression load to ensure a tight joint b = 0.25 in (width of seal)

G = 1.24 in '(average diameter of seal, Reference 6.8) m = 1.3 (from Reference 6.21, p. 47)

P = 2.500 ksi

=> H, = 6.33 kips

=> W.i =1,953 + 6.33 = 8.28 kips Stress due to W.i = W.i / A.

= 8.28 kips / 4(0.1599 in2 )

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A It11

1. 1If17 C-PENG-CALC-020, Rev. 00 Page 33 of 47

= 12 95 ksi < 26.9 ksi (S. at 650*F)

Design Minimum Initial Bolt Load, W.2 W 2 is taken as the total preload. Bolt stress due to preload only (ci.g) is calculated in Section 5.4.3.1:

a .g = 22.51 ksi < 28.3 ksi (S. at 100*F)

The stress due to W.i and W.2 are below their respective allowables, which indicates that the actual bolt area (AS) exceeds the minimum required area (A.)

5.3.2.2 StressDue to ThermalExpansion The differential thermal expansion between the hex head bolt and the upper flange and compression collar will create an additional tensile load on the bolt. For the analysis of the bolt, this additional load is assumed to be completely taken up by deformation of the bolt.

The stress effect of this differential thermal expansion is considered to be equivalent to that of the tie rod since the respective tie rod - top plate and bolt-flange materials are the same:

ct.. = S.2 ksi I l

The effective tensile force due to this thermal expansion is determined by:

2 P = ct.. A. = (8.2 ksi)(0.1599 in ) = 1.31 kips 3.3.2.3 Preload The bolts are being preloaded to 30 ft-lb. To determine the load in each bolt, the following equati is used (Reference 6.15) l T=0.2Fd

=> F =T / 0.2 d i Given:

T = the applied torque = 360 in-lbs d = is the nominal major bolt diameter = 0.50 in.

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Ak It!k i

l M IDID C-PENG-CALC-020, Rev. 00 Page 34 of 47 F = (360 in-lbs) / (0.20) (0.50 in) = 3.600 kips.

S.3.2.4 Marimum Bolt Load As it was described in Section 2.3.2, the impact from ejection of the nozzle will increase the load on the bolts.

The total load on the bolt can be expressed as follows: l l

e s  :

y'd F, . = Preload + F,,,,,  ;

< Ksa + k p,a <

In the above expression, Kn, is considered to be the equivalent stiffness of the components whic are put.in compression due to the torquing / tightening of the hex head bolts.

The stiffness of the components in the flanged connection between the MNSA and the Hot L l calculated below.

Stiffness of Hex Head Bolts:

The stiffness of the bolts is calculated using the same methods described for the tie rods in Section 5.2.3.2.1. Dimensions are taken from Reference 6.8.  !

\

2 g

Kw,, = 4 =4 (0.172in )(25.0X10' k*f )

= 10,269,000 I 1.675m. m 2

where: A = 0.172 in , cross-sectional area of the bolt, based on the basic pitch diameter 0.4675 in.

E = 25.0 x 10' psi (at 650*F) 1 = effective length of bolt, assuming 0.5 in of thread engagement

= thread engagement + lower flange + upper flange + washer

= 0.5 + 0.365 + .75 + 0.06 = 1.675 inch l

1

[

ABB Combustion Engineering Nuclear Power

A It R M IDIN C-PENG-CALC-020, Rev. 00 Page 35 of 47 Sligness of Overall Flanne:

The Hot Leg RTD MNSA has two components which represent the flanged connection to the Hot l Leg, the upper flange and the compression collar. The stiffness of each of these components is calculated with the use ofReference 6.11, Table 24, Case Ia. '

l w a' C, L, E t'

-L 3 ); D =

y = D ( C, 12(1 - y',)

l where Ci, C7, L9, and L3 are constants, and are calculated using the equations of Reference 6.11. Since neither the upper flange, nor the compression collar have a rectangular cross section, the dimensions are selected to produce the lowest component stiffness.

Upper flange:

All dimensions are taken from Reference 6.8. 3 FIGURE 2. UPPER Ft.ANGE h 2.0w =l y w jf"f I @~2 I mares gy j oy. ^

[MJ,Vfj uwt 0.750" l I h ggs! l , . .. ,,O,X. y .

0.2085"

> <ar =

a = outer radius (point of the load reaction),1.906 in b = inner radius,1.048 in

r. = radius of applied load,1.048 in t = thickness,0.75 in y = Poisson's ratio,0.3 E = elastic modulus,25.0 X 10' psi Ci = 0.4358 C2 = 0.5773 L3 = 0.0ll2 L, = 0.2809 ABB Combustion Engineering Nuclear Power

A ItIt M IDIF C-PENG-CALC-020, Rev. 00 Page 36 of 47 25.0X 10' k'" (0.75)'in' 3

Et l

D= = = 965,831 in -lb l 12(1 - y',) 12(1- 03 )

2 SoMng for the stiffness of the upper flange:

K*=b= y a ' C,L, _

= 4,572,500 '

in D C, Compression Collar:

All dimensions are taken from Reference 6.8.

FIGURE 3. COMPRESSION COLLAR

" noer  ;;

sur l -

- oie

-- imr = J ,l

iar

2 4$r  ;

a = 1.047 in b = 0.846 in r = 0.623 in t = 0.75 in (conservatively) y = 0.3 E = 25.0 X 10' psi Ci = 0. lr'l C, - 0.1955 L3 = 0.0078 L = 0.2636 Et 3 25.0X 10' #" , ' I (0.75)'in' D= = = 965,831/n -lb 12(1-y ) 2 12 (1- 03 )

2 ABB Combustion Engineering Nuclear Power

M IDIF C-PENG-CALC-020, Rev. 00 Page 37 of 47 2xr* = 13,274,000Ib K,,n, = W = [ (C,L, _

y in D C, Determination of equivalent flange stiffness:

The effective stiffness of the component is calculated below.

I Ky,, - - 3,401,000 lb 4 m K,,,, K,,iw Therefore, the maximum bolt load is I '

0,269,000 11 F" = 3.6 + .

= 5.66 kips s 10,269,000 + 3,401,000j 4 5.3.2.5 Average Stress, m (NB-3232.1)

The average (axial) stress (ci) in the bolt is due to a combination of stresses from the maximum bolt load and from differential thermal expansion:

ci= P/A + c c..

i A, = 0.1599 in2 P = 5.66 kips ot,. = 8.2 ksi 2

c = (5.66 kips /0.1599 in ) + 8.2 = 43.6 ksi < 2 Sm = 53.8 ksi (at 650*F) 5.3.2.6 Maximum Stress (NB-3232.2)

The maximum stress in the bolt is essentially a stress intensity due to a combination of the avera stress and the torsional shear stress due to residual torque (from preload).

l l

ABB Combustion Engineering Nuclear Power

i A ItIt l FRIFIF C-PENG-CALC-020, Rev. 00 l

Page 38 of 47 5.3.2.6.1 Residual Torque, Ta l

The residual torque due to preload may be calculated using the following equation for standard threads (Reference 6.15, Equation 6):

Ta = 0.5625 T = 0.5625 (360) = 0.203 in-kips where:

0.5625 = multiplier based upon a coefficient of friction of 0.15 and standard bolt dimensions T = 360 in-Ib = applied torque 5.3.2.6.2 Torsional Shear Stress, tr l l

1 The torsional shear stress is determined using the following equation derived from Equation 7 of Reference 6.15: '

tr = 16Ta / xd' =16 (0.203) / x (0.4525)' = 11.16 ksi where:

d = average of basic pitch diameter (0.4675 in.) and minor diameter (0.4374 in.) =

0.4525 in.

5.3.2.6.3 Maximum Stress, c.,

l The maximum stress intensity (c.) is determined using the following equation (Reference 6.15,  !

Equation 8):

l o.x = 2 ( )2 + (rr)2 = 48.98 ksi < 2.7 Sm = 72.6 ksi (S. at 6507)

< 3.0 Sm = 80.7 ksi (S. at 650T)

5. 3.2. 7 ThreadShearStress, r From Reference 6.18:

2 AS. = x n Le Kn max [(1/2n) + 0.57735(Es min - Kn max)) = 0.400 in where:

n= number of threads per inch = 20 ABB Combustion Engineering Nuclear Power

i i

kk M IfIF C-PENG-CALC-020 Rev. 00 Page 39 of 47 Le = the length of engagement. Assume equal to 0.5 in ,

Kn max = maximum minor diameter ofintemal thread = 0.457 in Es min = minimum pitch diameter of external thread = 0.4619 in l

P = 5.66 kips 1

1 2

i = 5.66 kips / 0.400 in = 14.15 ksi < 0.6 Sm = 16.14 ksi The minimum allowable length of engagement of the hex head bolt into the Hot Leg pipe may be calculated as a simple proportion, based on the bolt threads.

Le , = (Shear Stress / Allowable Stress) x Assumed Length of Engagement =

1

= (14.15/16.14) x 0.5 = 0.44 in. I 5.3.3 Top Plate 3.3.3.1 Shear Stress, r z

A. = x (D) t =x (1.325 in) (0.88 in) = 3.663 in where:  !

D = the maximum diameter of the thermowell = 1.325 in (Reference 6.4)  ;

t = the thickness of the top plate = 0.88 in l

P = 11.0 kips 2

t = 11.0 kips / 3.663 in = 3.0 ksi < 0.6 Sm = 9.72 ksi 5.3.3.2 Bendingstress, ob l The impact load is distributed over the area of the top plate in contact with the thennowell. 1 Conservatively, the impact load is applied at the location of the outer radius of the thermowell. From Reference 6.11, Table 24, Case Ia (dimensions are taken from Reference 6.8 and 6.4):

r, = 0.663 in a = 1.906 in b = 0.56 in t = 0.88 in w=F%a / [(2 n) r.]= 11.0 kips / [(2 x) 0.663] in = 2.641 kips / in l b /a= 0.2938 ABB Combustion Engineering Nuclear Power

M IDIF C-PENG-CALC-020, Rev. 00 Page 40 of 47 K. (b/a) = 0.6122 M = K. w a = 0.6122 (2.641 kips /in)(l.906 in) = 3.082 kips-in 2

o = 6M/t = 6 (3.082 kips-in) / (0.88)2 in2 = 23.88 ksi c6 = 23.88 ksi < l.5 Sm = 24.3 ksi 5.3.4 Compression Collar 5.3.4.1 Shear Stress, r 2

A.= (x)(D)(t) = (x)(2.093 in) (0.312 in) = 2.052 in P = 3.6 kips / bolt x 4 bolts = total preload from bolts

= 14.4 kips l

t = 14.4 kips / 2.052 in2 = 7.02 ksi < 0.6 Sm = 9.72 ksi 5.3.4.2 Bearingstress, ob 2 2 A6 = (x/4)(D co.,wu,oo - d eo.p.ou-m) = (x/4)(l.49'- 1.002 2) in = 0.955 in2 l P = 14.4 kips i

2 c6 = 14.4 kips / 0.955 in = 15.08 ksi < Sy = 17.9 ksi 5.3.5 Upper Flange Shear stress, t i

z A, = (x)(D) (t) = (n)(2.513 in) (0.438 in) = 3.458 in P = 14.4 kips i t = 14.4 kips / 3.458 in 2= 4.16 ksi < 0.6 Sm = 9.72 ksi i

l Due to the proximity of the bolts and support surface, bending stresses are considered to be small and are neglected.  !

ABB Combustion Engineering Nuclear Power

A R It M IDIF C-PENG-CALC-020, Rev. 00 Page 41 of 47 14 Fatigue Analysis The fatigue analysis of the components will conservatively consider loads which may exist on the components after weld or nozzle failure has occurred. Prior to failure, components will be subjecte to loads due mainly to preload and thermal expansion. After failure, and assuming that the nozzle / valve is free to move, certain components will be additionally stressed because of the internal pressure forcing the nozzle / valve up against the top plate. The load on these components would be cyclical, given the change in pressure and temperature that occurs as the plant heats up and th cools down.

The critical components for fatigue analysis purposes are the tie rod and hex head bolt, on the basis of:

e i preload tensile stresses e th rmal expansion tensile stresses e i stress concentrations in the threaded sections, and e

for the levels of stresses involved, a more restrictive number of allowable cycles (versus the stainless steel MNSA components; see Table I-9.1 ofReference 6.9)

It is noted that, in the fatigue analyses below, the stresses produced by the one-time applicati impact load are not considered since the contribution to fatigue from this one occurrence is not significant.

5.4.1 Normal Operating Pressure Force The effect of the force acting on *he MNSA components due to internal pressure is similar to that o the impact load, only of a smaller magnitude, and it is a function of the internal pressure and the a of the nozzle. The pressure used to determine the force is based on the maximum internal p for all Normal and Upset conditions from Reference 6.3, which is 2350 psia (for the 10% Ste Increase transient). The normal operating force is determined as a function of the Design Press force calculated in Section 5.2.3.1.1:

F, = 1.953 kips (2350 psia / 2500 psia) = 1.836 kips Use F, = 2.00 kips This force will be used to calculate normal operating tensile stresses in the tie rod and bolt as the Peak Stress Intensity calculation.

ABB Cornbustion Engineering Nuclea Power

kk!k

1. "tIf17 C-PENG-CALC-020, Rev. 00 Page 42 of 47 5.4.2 Tie Rods 5.4.2.1 Peak Stress The maximum Peak Stress in the tie rod is calculated, as fo!!ows:

Peak Stress = fstf*(om.)

where:

fstf = fatigue strength reduction factor = 4.0 (from NB-3232.3) o, . = maximum stress using the load from normal operating pressure instead of the impact load:

o,.x. = 2 (#' "' )2 + (r7 )'

where:

ci = ci, + oij+ ot.. (conservatively), and ci_, = tensile stress due to pressure

= (2.00 kips /4)] / 0.0775 in z

= 6.45 ksi ciw = tensile stress due to preload (preload from Section 5.3.1.3)

= 1.0 kips / 0. 0775 in 2

= 12.9 ksi ot . = tensile stress due to thermal expansion (stress from Section 5.3.1.2)

= (8.2 ksi)*(30E6/25.25E6) (E/E per NB-3232.3(d))

= 9.74 ksi on

= 1.62 ksi(stress from Section 5.3.1.7.1)

Tr = 6.53 ksi(stress from Section 5.3.1.7.3)

=> o.,x.= 33.3"i ksi

=> Peak Stress = 4 (33.3 7) = 133.48 ksi Based upon a minimum stress value of 0.0 ksi (this is a conservative approach, since preload never goes away), the maximum Peak Stress Intensity Range (S,) in the tie rod is:

S, = 133.48 - 0 = 133.48 ksi ABB Combustion Engineering Nuclear Power

M IDIF C-PENG-CALC-020, Rev. 00 Page 43 of 47 5.4.2.2 Usage Factor The calculation of the usage factor for the tie rod is based upon the maximum Peak Stress In Range of 133.48 ksi:

S. = S,/2 = 66.74 ksi This range is considered to occur a total number of 700 cycles, which is the sum of the numbers of cycles for Heatup/Cooldown (500) and Plant Leak Test (200), p r Reference 6.20 (seismic loads are conservatively included in a total number of cycles).

For a S. = Sp/2 = 66.74 ksi, the allowable number of cycles per Table I-9.1 (Reference 6.9) is approximately 1800. The usage factor (U)is:

U = 700/1950 = 0.359 There are other Normal and Upset transients which are defined for the Piping (per Reference 6 but their contribution to fatigue in the tie rod is not significant.

5.4.3 Hex Head Bolt 5.4.3.1 PeakStress i

The maximum Peak Stress in the bolt is calculated, as follows: '

Peak Stress = fstf*(e -)

l where:

fstf = fatigue strength reduction factor = 4.0 (from NB-3232.3) o

m. = maximum stress using the load from normal operating pressure instead of the impact load:

om.=2( )2+(tr )*

where:

c, = ci., + oi.pi+ c .a(conservatively), and ABB Combustion Engineering Nuclear Power

A ItIt C-PENG-CALC-020, Rev. 00 Page 44 of 47 ci., = tensile stress due to pressure

= (2.00 kips /4) / 0.1599 in 2

= 3.13 ksi i

oi.pi = tensile stress due to preload (preload from Section 5.3.2.3)

= 3.6 kips / 0.1599 in'

= 22.51 ksi c1.. = tensile stress due to thermal expansion (stress from Section 5.3.2.2)

= (8.2 ksi)*(30E6/25.25E6) (E/E per NB-3232.3(d))

= 9.74 ksi i

TT = 11.16 ksi (stress from Section 5.3.2.6.2) l

=> o e = 41.83 ksi

=> Peak Stress = 4 (41.83) = 167.32 ksi l

Based upon a minimum stress value of 0.0 ksi (this is a conservative approach, since preload never goes away), the maximum Peak Stress Intensity Range (Sp ) in the bolt l

S, = 167.32 - 0.0 = 167.32 ksi S.4.3.2 Usage Factor The calculation of the usage factor for the hex head bolt is based upon the maximum Peak Stress .

Intensity Range of167.32 ksi:

S. - S pt2 = 83.7 ksi This range is considered to occur a total number of 700 cycles, which is the sum of the numbers of cycles for Heatup/Cooldown (500) and Plant Leak Test (200), per Reference 6.20.

For a S. = S/2 = 83.7 ksi, the allowable number of cycles per Table I-9.1 (Reference 6.9) is approximately 1400. The usage factor (U)is:

U = 700/1400 = 0.5 There are other Normal and Upset transients which are defined for the Piping (per Reference 6.20),

but their cont;ibution to fatigue in the bolt is not significant.

ABB Combustion Engineering Nu lear Power

i

' k M IDIF C-PENG-CALC-020. Rev. 00 Page 45 of 47 5.5 Consideration ofHydrostatic Test Pressure Conditions Per Paragraph 1.3.1 of Reference 6.7, the deliverable MNSA hardware is not required to be hydrostatically tested. However, it is noted that the pressure load created by the hydrostatic p of 3125 psi (Reference 6.20)is less than the impact load. Therefore, stresses resulting from hydrostatic testing would be acceptable.

5.6 Consideration ofFaulted Conditions Reference 6.20 lists the Faulted Conditions which are identical to the Design Condition except they also include a Design Basis Earthquake and Pipe Rupture events. Pipe Rupture event has no effect on the MNSA components. An assessment is made of the effect of Faulted Conditions reviewing the maximum stress results for the tie rod (Sections 5.3.1.7.1), which is the compon most signincantly affected by an earthquake event (either OBE or Maximum /DBE). Cons doubling the OBE bending stresses to simulate the effects of the Maximum Earthquake event res in stresses which meet the 3S. allowable for Maximum stress. Therefore, the stresses res Faulted Conditions are acceptable.

ABB Combustion Engineering Nuclear Power

A ItIt l 2"EIFlE C-PENG-CALC-020, Rev. 00 Page 46 of 47 l

6 REFERENCES 6.1. ABB Project Plan No. C-NOME-IPQP-0263, Revision 0, "Waterford Mechanical Nozzle Seal Anemblies", March 1999.

6.2. ABB Combustion Engineering Nuclear Power Quality Procedures Manual QPM-101, Revision 03.

6.3.

" Analytical Report for Waterford Unit No. 3 Piping," Report No. CENC-1444, May 1981.

6.4.

• Rosemaunt Engineering Company Drawing H33760-1201, Rev. C, " Sensor, Temperature" 6.5.

ABB CE Drawing 74470-771-003, Revision 02, " Primary Pipe Assembly".

6.6.

ABB CE Drawing 74470-772-001, Revision 04, " Instrument Nozzles Waterford III Piping" 6.7.

" Design Specification for Mechanical Nozzle Seal Assembly (MNSA) Waterford Unit 3",

Specification No. C-NOME-SP-0067, Revision 01.

6.8. ABB Drawing No.

6.8.1. E-MNSAWFD-228-001, Revision 02, " Hot Leg RTD Mechanical Nozzle Seal Assembly Flat" 6.8.2. E-MNSA-228-004, Revision 05, " Mechanical Nozzle Seal Assembly Details" 6.9.

American SocienfMcchanical Engineers Boiler and Pressure Vessel Code,Section III,1989 Edition (No Addenda).

6.10.

" Test Report for MNSA Hydrostatic and Thermal Cycle Tests," Test Report No. TR-PENG-042, Rev.00.

6 11. "Roark's Formulas for Stress and Strain," Warren C. Young, Sixth Edition,1989, McGraw-Hill.

6.12. " Heat Transfer A Basic Approach", M. Necati Ozisik, D85, McGraw-Hill.

6.13. Inter-Office Correspondence from J. T. McGany to K. H. Haslinger, "Waterford MNSA Stress Analysis", Letter No. NOME-99-C-0122, Revision 01, dated March 18,1999.

6.14.

" General Specification for Reactor Coolant Pipe and Fittings", Specification No. 00000-PE-140, Rev. 02, July 1973.

ABB Combustion Engineering Nuclear Power

F1IFIF C-PENG-CALC-020, Rev. 00 Page 47 of 47 6.15. "How to Calculate and Desiga for Stress in Preloaded Bolts", A.G. Hopper and G.V. Thompson, Product Engineering,1964.

6.16. Engineering Report No. C-NOME-ER-0120, Revision 00, "Desige Evaluation of MNSA for Various Applications at Waterford Unit 3", March 1999.

6.17. Calculation No. C-PENG. CALC-021, Revision 00, "Detemunation of Waterford 3 Hot Leg Seismic l

Response Spectra & Accelerations for Use in the Analyses ofMNSAs", March 1999.

6.18. ANSI Standards for Threads, Appendix B, Bl.1,1982.

6.19. " Addendum to CENC-1365 and CENC-1507 Analytical Rept " Southern California Edison San Onofre Units 2 and 3 Piping", Design Report No. S-PENG-DR-wa, Rev. 01.

l 6.20. " Project Specification for Reactor Coolant Pipe and Fittings for Entergy Operations, Inc. Waterford, Unit 3", Specification No. 09270-PE-140, Rev. 07, December 1993.

6.21.

Union Carbide Grafoil, "Engmeenng Design Manual, " Volume One, Sheet and Laminated Products, by R.A. Howard,1987.

6.22. " Strength ofMaterials", F. L. Singer, Second Edition, Harper & Row, New York,1%2.

6.23. "Engmeenng Mechanics: Statics and Dynamics", F.L. Singer, Third Edition, Harper & Row, New York, 1975.

6.24. " Mechanical Engineers' Handbook", M. Kutz, ed., John Wd' ey & Sons, Inc.,1986.

• This Reference is used per Reference 6.13 ABB Combustion Engineering Nuclear Power

. - . -- . .- . . . . _ _ _ . . . - _ _ _ .__ _=., .. _. . . _ _ .

3 A It R C-PENG-CALC-020, Rev. 00 5 19 EF 33 Page Al of A2 APPENDIX A ASSEMBLY DRAWING 1

ABB Combustion Engineering Nuclear Power

sw l I

9 to

\ 3

\ g \ ,

W SECTION A~ A Y SECTION B-B ji

,iFE S A ,

l l

____h @ '

._., M; ';i i

-(

@ COMPRESSION COLLAR e.993 - -- y - -

. _ . 1, f )

8..VE W///#

a q,

C

\

\

'N ,

@ HOT LEG RTD MECM i

n 1 8 8 7 I 6 '

o=~~ cr

Paae A'2 o f A 2.

~%~ w::.a,,

1IIIIT

_ . _ . -. I ..,.. .,

s ,

,m o us m 2 2 5 R ,

%D

. m-- w,.U. . N 'wT y

. . m- - -- -

.,,,.r,,, ,n g 4.-.Ju6,r., , . , , ,- ,.

,,u ,...t.,.-

C,,G,) .

., ,. . wi

. ., _ , . r h..r,.d.r wad.. w-

. r. m,-nent-m ,,v 5

% ,,, v id h1- .

S I , _

/

( I

_j_ _. .

I c.

i i

w- I -

.b "r=.2. .. .

l is.l.iC.

. .L*. ." Nun I Y VIEW C--C r HOT LEG PIPE M ACHINING p

A -

1 cm

,i -.

l ,

s Ll L E

,- .......i~

- n w- , n.cm . . .. . .,, .

7 __./

ir i

. , , = . . . . . . . . . ,

.4 y '/,. _. _ .. . . _ . c. m ou ... -

/ I A -in -a ,2 o ' ."."'"'; ",c.U.a'J,*;*RM M.,7,MMYo !".'.7c! "#

• i c os l -

ff El

. m.,. .. . .. ,_, ._ .. ... ~ - . _ _m. ..,

,......_-..._....e.,. .. . . . . . .. . . . .......... . . . . _... -,_...m....._.

,. ., y2 ... . . . _

. . . . . . . n, s

il > --s -

....c........

- o C . ,.

wv ....

s ,' < ,,.,....s..... .. . . . . . . . . . . . . . . .

$ /

j G tt.Cse.dR t C*sL O.et Co e go .ete 7 e, us s , a is. s. sea. sa.or s.Must y j o Lggg me 9 A Pc== n g.0. - JRE A s xs . . . ... , . .c_ , . . _ . _. .,, . . . . . . , , ,

CAR 0, p-'"g NA' N N ~ _ . . . . . . , . . ,

t j N ,.....,,..,,,,,....s

...omo...,,,,.. ,ee,.-.,......

Also Avallatio on l cu ce., .. Aperturo r:ard l ~

l LNIC AL N0ZZLE SE AL ASSEMBtY FL AT lE- W C- M T L2d n

- I -

CERTIFIED FOR l s CONSTRUCTION 6 f- t.Mt2P -

l

.. . ., . ,,.a...... -, m

.nr_m e .., s....

==qr::pe,, -

• - . lG 3

y".....t,. r -

- - .Raii $

l gg,d==".,  %"

• II If l .

• ~

".0 .'"+.'3'r =' l"GilEf

's"-- diO"

, ev = w4. . . _ _ , , - ~

, A s ..w-se. u., oa Q2cm3 .",, *, .* **g,,, k,qg, 4 0 g .rlg.==i5' c':.*:*!C:p.,.f ll.

. st.e".

ut

..E.I ".,Ern.e o .ssivet rt p -.

=. -.i. = . .. . .n3-- Ts - .r - i.-- ...- 'E .M'N5HtD.-DB.004, J

' ' 4 ' 3 1 0.T8.% ...

2 ' D't:I L- - - - _ _ _ _ - _ . . - - _ _ _ - _ - _ _ - - . - _ _ . . _ . . __n

A ItIt f! C-PENG-CALC-020, Rev. 00 Page B1 ofB2 '

i i l l

l l

i APPENDIX 8

CALCULATION OF THE TIE ROD AVERAGE TEMPERATURE l J

i 1

i i

b ABB Combustion Engineering Nuclear Power

r . _ . _ _ - _ . .

A ItIt gggg C-PENG-CALC-020, Rev. 00 Page B2 of B2 Calculation of average temperature due to axial conduction Infinite Fin (" Heat Transfer - a Basic Approach", M.N. Ozisik, McGraw-Hill Inc.,1985)

Reference 6.12 T, - T, ' "5 Ambient Temperature T,inifinite 120 *F Base Temperature T,o 500 'F Nominal dimension tie rod OD 0.375 in Nominal dimension te rod OD 0.03125 ft Tie Rod area A 0.000767 ft2 Tie Rod perimeter P 0.098175 ft Tie Rod therm cond (Ref 6.9,300*F) k 8.8 btu /hr-ft-F generic outside film coefficient h 1.7 btu /hr-ft2-F calulcated coefficient m 4.972652 Length x T 0 0 500 0.6 0.0500 416 l 1.2 0.1000 351 1.8 0.1500 300 i

2.4 0.2000 261 3.0 l 0.2500 230 l 3.6 0.3000 205l 323l 4.2 0.3500 187 4.8 0.4000 172 5.4 0.4500 161 6.0 0.5000 152 Average temperature at 3.6 inches is 323'F l

l l

l ABB Combustion Engineering Nuclear Power

A IkIk gggg C.PENG-CALC-020, Rev. 00 Page Cl ofC6 l

APPENDIX C i

QUALITY ASSURANCE FORMS ABB Combustion Engineering Nuclear Power

gg C-PENG-CALC-020, Rev. 00 Page C2 of C6 <

Design Analysis Verification Checklist Instructions:

If a major topic area (generally unnumbered, bold face type such as Use of Computer Software) is not

' applicabic, be left blank. then N/A (not applicable) next to the topic may be checked and the check boxes for all items un Where there is no check box under N/A for a numbered item, such a response is generally inappropriate. If N/A is checked in such a situation, document the basis at the e'd of this checklist in the Comments section.

Author IR Overall Assessment Yes l N/A Concur.

1. Are the results/ conclusions correct and appropriate for theirintended use?

@ I  !

2. Are alllimitations and contingencies on the results/ conclusions documented?

@ ,2 Assignment of Cognizant Engineers, Independent Reviewers and Mentors 1.

tf there are multiple Cognizant Engineers, has their scope been documented?

2. If there are muluple t h Rei ..is, has theirscope been d~"--d?

[ [

3.

If there wilt be muluple Management Approvers, has their scope been d~~-d?

@ [

4

] @ [ )

tf an Independent Reviewer is the supervisor or Project Manager, has authonzahon as an Independent Reviewer been documented? O E 5.

tf there is a Mentor, has their scope and responsibilities been adequately d~'-a-a d?

Use of Computer S "ivvare

{

For software wtuch has been valid-wunder QP 3.13:

1. Is the software hsted on an Approved QC 1 Software Last?

g

2. Is the software apphcable for tius analysts? O O for Code-1As Cormtructs vahdated under QP 3.I4. O 1.

Is the Code-take Construct hsted on an Approved QC-1 Software List? b 2.

Is the Code Lake Construct apphcable for tha analysis?

3. Was the Code IAe construct used t..diy a the coraround 6..?

No O O

- If No stiove, a the copy idetsacalto the L;.~u a the careroDed .' .am? (leeve blank afnot apphcable.)

- If changes were made to the Code-Like Cw-u ct to meet wJic analyus needa, were such changes documented as non-validated software following paragraph 3.3.37 (Leave blank if not applicable. Complete the next section of this Checklist if'Yes".)

For software wtuch has not been vahdated under QP 3.13 or QP 3.14:

1. Is the i - N type, program name and revision identificatson da'-ted?

g 2.

Is a copy of the software included a the Design Analysis?

3.

Have tests been documented wtuch are ad~;'.'ete to /---- _ e e correct operapon for the software's mtetuled use?

4 is the output thnn the tests mcluded m the Design Analysis?

5. Has the Cogruzant Engmeer ^=

correctiv for its intended use?

Jed the results of the tests and the basis for concludmg the software is operstmg 6.

Did the software, as used in this analysts, give correct results?

ABB Combustion Engineering Nuclear Power

F If p C-PENG-CALC-020, Rev. 00 Page C3 ofC6 Design Analysis Verification Checklist Author IR Use of Cornputer Software (continued) 1 Yes l N/A Concur.

Were spreadsheets used in tius Design Analysis m any way . data display plottmg, computauons, etc.?

. If data display 2Bl% (no computauons or plotung ), check "Yes" and slup remauung quesuons.

. If used for computations:

. us o,e coniputauoos adequaieiy --#

+ Are the results correct?

0

- liused for plotmg:

e is the data to be plotted correct?

e Are the plots -i in other respects? (titles, scales, labels, etc.)

8.

Is a copy of the spred e included in the Design Analysis? ( A copy of the fde may be included or sufficient detad included in the analvsis docuneion to pemut recnatmg the .v.--' ' 1) -

Objective of the Design Analysis

1. Has information necessary to defme the task been included or nfennad?

2. Have the objecuves been enumerated?

@ [

3. Has the applicability and intended use of the resuhs been ?---

@ [

"?

Assessment of Significant Design Changes

[

1.

Have signiricant daign.nlated changes that might impact this analysis been considend?

2.

If any such changes have been identified, have they been adequately addressed?

[

Analytical Techniques (Methods)

@ [

l. Are the analytse4 techniques (methods) d c.. bed in sufficiera detad tojudge their appropnateness?

2.

Are the analytical techniques used or their apphcation governed by an NRC issued SER7 if yes, have the applicable SER: been documented? No If yes has the basis for concludmg the analysis is in confonnance been documented?

]

3. Have analyucal Wques .,,.r.

ed by ref.aw to genenc analysca, lead plant analyses or previous cycle analyses O

been previously venfsed?

4.

An any ---Mcanons or Api.rmres frorn previously approved analyucal techniques or Convenuonal or Automated Procedures hem and justified?

3. -

If sugi- 4.4 a.n ..ed analyucal techruques or engmeenng procedures are used, as their usejustified and approved?

6.

Does the tasus date ofreferenced approved Convenuonal or Automated Procedures predate their une in tius analysis?

Schion of DesignInputs

1. Are the design inputs h--H?

2.

Are the design inputs correctly selected and traceable to their source?

3.

Are the bases for selecuon of all design inputs h--d?

4.

Is previously unverified design i wit uined in this analysis? No

[

If Yes, is it treated in accordance with QP 3.2, paragraph 3.4 for use of unvenfied desing information?

] ]

ABB Combustion Engineering Nuclear Power

t pggg C-PENG-CALC-020, Rev. 00 Page C4 of C6 l Design Analysis Verification Checklist Author IR Selection of Design Inputs (continued)

Yes l N/A Concur.

5.

Is the venficahon status of design inputs transrrutted from customers or CENP Nuclear Systems appropnate and A.-,w.A?

[

6.

Is the use of customer. controlled sources such as Tech Specs, LTSARs etc. authonzed, and does the authonzation specify amendment level, revision number, etc.? [

l Assumptions

1. If there are no assumpuans, is this documente.d?

2. Are local asumptions documented, fuuyjustdied and venfied?

] [

3.

]

Are Intemal and External Assumptions which must be cleared by CENP or the customer hsted on a Contagencies and Amumphons fann?

4.

la the 'roject Managerresponsible for cleanng the Assumphons identdied on the form?

Results/ Conclusions 1.

Are all resuhs contained m or referenced in the Resuha/ Conclusion section? (Where feasible, in the enumerated order '

of the objectives.) @ [

2.

Are all Imutations on the reeuks/ conclusions and their applicabibty documented in this esction?

3. Are allw

_smgencies _ on,_there,-,

resuhs that must be cleared listed in the Results/ Conclusion section or the Contagencies

] @ ,

4.

Is the Project Manager responsible for cleanng the Assumptions or Contingencies identified on the form?

Other Elements

] @

1.

Has a He-differences of the explained? results with those of a previous cycle or stnular analysis been documented and signdicant 3.

Have appbcable Codes (e.g., ASME Code) and statu1ards beenately .n,., referenced and apphed?

3.

Is the infonnauon from relevant htcrature searcN tackground data adequately documented and referenced?

4 Are hand calculations s.4 and .n m.ately * ^

_?

5. Is all apphcable a- ,-p- W and input mcludut?

6.

Is all wmputer software used idenufied by narne and revision identificanon?

References

1. Are all refaw used to perform the analysis listed?

@ ] 1'

2. Are the references as direct as possible and appropnate to the source?

> m @

e_ __ -c ~ __. _ __ -,c, ~ w,n. _ -

id sificnuan of the locanon ofthe sfonnanon m the Mannce, such as page, table or paragraph W a g Independent Reviewer's Statement of Verification Activities:

t= ' r- 's Reviewer to desenbe details of venfication activities beyond the obvious on this checklist ,

includmg but not linuted t th o e review of new methods, previousiv use ofinputs, unvenfied software etc. under paragraph 3.3.3, spreadsheet use, assessment of design and rnethodology changes, engmeenn l

I ABB Combustion Engineering Nuclear Power

F C PhhG-C ALC-020, Rev. 00 Page C5 ofC6 Design Analysis VeriGcation Checidist The Form and Independent Format section of the Checklist below may be completed by a Checker under the direction of the\

Reviewer. '

Author IR Fonn/Fornist Yes N/A Coxur.

I.

Is the *- -

• legible, ,m Wacible and in a form suitable for filmg and retneving as a Quahty Record?

2.

Except as pemuned by 3.1.3.a. are all pages ident6ed with the document number, includmg revnion number?

@ [

3.

Except as pernutted by 3.1.3.a. do a!! pages have a unique page number?

[

4.

@ }

Have all changes been ah*=H by the irutials and date of the Quality Records Controller?

S.

Are all files on CD-ROM identified by the path name?

)( g 6.

Are all -qwter disks identified with the analysis number?

[ @

7.

Are any unvenfied sections of an otherwise verded analysis clearly ind- w?

] [

For a " Memorandum Revaion* to a completed Design Analysis:

1.

Have the trtie and document n2 der been preserved without change?

@ [

2.

Q Q Does tius revision meet the enteria for a " simple revision"?

3.

]

Are the Author, fad a ad

• Reviewer and Management Approver and theirroles identded?

r and "Page ha=e Package" formats:

] ]

For a revision to a -v;eted analysis in the

• Complete Revision" 1.

Where practical, have changes and addrtions been identified by ==M

[

=- such as vertical lees, etc.?

2.

Where practical, have deletions been identded by u :- such as strike outs, etc.?

3.

Have indications of change in previous revisions been removed?

] l 4.

O O i Does the distribution of the revision include those on the distribution of the previous revision?

For a " Complete Revision":

[ '

1.

g Have the title and document number been preserved without diange?

2.

O ]

Has the revision number been incremented by one?

For a *Page Change Package"-

Q ] 1 1.

Are pages numbered in accordance with the anginal analysis?

] [

Are instructions provided for the insertion and deletion of revisied pages?

Has a new Title Page been prepared?

4.

O [

Does the Package Contents Page reDect the change package contents?

O O Form / Format section completed by the Independent Reviewer.

O Form /Fonnat section completed by the Checker identified below:

Checker Name: Signature:

ABB Combustion Engineering Nuclear Power

A IkIk gggg C-PENG-CALC-020, Rev. 00 Page C6 ofC6 Reviewer's Comment Form

Title:

Analysis of Waterford 3 Hot Leg RTD MNSA Document Number: C-PENG-CALC-020 Revision Number: 00 Comment Reviewer's Comment Response Author's Response Response Number Required? Accepted?

I .f5.21-Il could be charer as k Wo A /E AS uki Me dahm 4 for fi, 2 [63.17 - 08E is xor 1)ssisd VES hluno. ed lesh "

Ewwaurz- r s OnearN4 8 Asis 3 Rwe 35 - Raurz .2 - uk) ih *a" Ye.s Mm' %intof l V SNds denniban. Jan'ooohh2k m 1

l ABB Combustion Engineering Nuclear Power

l C-PENG-DRM Rev. 01 Page El of E29 i

ATTACHMENT E  !

i A-WATERFD-9449-1213, Revision 01,

" Evaluation of Attachment Locations for Mechanical Nozzle Seal Assembly on Waterford Unit 3  ;

l Partial Penetration Welded Nozzles in the Hot Leg Piping" (29 pages including cover) i

SUMM.ARY OF CONTENTS l Calculation L Pages Appendices O Pages Anachtnents 9 Pages Diskene Attached _Yes X_No EVALUATION OF ATTACIIMENT LOCATIONS FOR MECIIANICAL NOZZLE SEAL ASSEMBLY ON WATERFORD UNIT 3 PARTIAL PENETRATION WELDED NOZZLES IN THE HOT LEG PIPING A-WATERFD-9449-1213 REV. 01 Quality Class: QC-1 (Safety-Related)

Contingencies: None PURPOSE: To evaluate the structural integrity of the attachment locations for the mechanical nozzle seal assemblies about the partial penetration welded nozzles in the hot leg piping.

This Design Analysis is complete and verified. Management authorizes the use ofits results.

PREPARED BY: P. L. Anderson ,

. _ /. sd,_ .raf DATE: $ ([96 .

VERIFICATION STATUS: COMPLETE The Safety-Related design mformatico contamed in this document has been venfied to be correct by means of Design Review using the Checkhst m QP-3.4 of QPM-101.

Name R. E. Johnson Independent Rev=wer Signatwe 6/

w Date: F/MW '

APPROVED BY: D. P. Siska DATE: f-##

ABB COMBUSTION ENGINEERING CHATTANOOGA, TENNESSEE TOTAL P.01

. - . _ _ _ _ . ~ . . . _ . . . _ . . _ ._ ._ . _ _ _ . _ _ _ _ _ . . _

A-WATERFD-9449-1213. Rev. 01 Page 2 of 25 ABB - CENP PROPRIETARY RECORD OF REVISIONS i

PARAGRAPH (s) PREPARED INDEPENDENT APPROVED NUMBER DATE INVOLVED BY REVIEWER BY 0 Original Issue B. A. Bell R. E. Johnson D. P. Siska l

l l

l 1 3/19/99 Sections 1.0,5.0, P. L. Anderson R. E. Johnson D. P. Siska 6.1.1, 6.1.2 & 7.1 I

I I

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1 CSE-99-046

A-WATERFD.9449-1213. Rev 01 Page 3 of 25 ABB - CENP PROPRIETARY TABLE OF CONTENTS P_ags a

1.0 OBJECTIVE OF THE DESIGN ANALYSIS. .4 2.0 ASSESSMENT OF SIGNIFICANT DESIGN CHANGES. . .4 3.0 ANALYTICAL TECHNIQUES. . . .

.5 4.0 SELECTION OF DESIGN INPUTS.. .. ... . .5 5.0 ASSUMPTIONS. . . . . . .. .. .. .

. .6 6.0 DETAILED ANALYSIS 6.1 Seal Assembly Bolted Connection to Piping Evaluation.... .. . . .7 6.2 Re-evaluation ofPenetration Reinforcement Area for Standard MNSA. . . . .9 6.3 Re-evaluation of Penetration Reinforcement Area for Sampling Nozzle MNSA..13 6.4 Range of Stress Intensity Evaluation of 1/2" Tapped Holes. . . . . . . . . . 17 6.5 Fatigue Evaluation of Tapped Holes . . . . . . . . . . . . . . . . . . . .21 i

7.0 RESULTS / CONCLUSIONS.

. . . . . . . . . . . . . .. .. . . . . . .24

! 8.0 REFERENCE S . . . . . . . . . . . . . .. . . . . . . .. . . . . . . . . . . . . . . . . . . . . . .25 Attachment 1: Design Analysis Verification Checklist & Reviewer's Comment Form (For QA Record only) l l Attachment 2: References l

CSE-99-046

A-WATERFD-9449-1213, Rev. OI Page 4 of 25 ABB - CENP PROP!UETARY 1.0 OBJECTWE OF THE DESIGN ANALYSIS As a result of three (3) leaking partial penetration nozzles on the Hot Leg Primary Piping at the  !

Entergy Waterford Unit 3 Plant, ABB-CE designed Mechanical Nozzle Sealing Assemblies (MNSA) to seal these partial penetration welded nozzles on the hot leg piping. The MNSAs are described in References 1, 2,14,15 and 16. The partial penetration nozzles included in this report include the RTD, the PDT (Pressure Measurement), and Sanipling nozzles. A MNSA is attached to the outside surface of a pipe by four 1/2 inch bolts. The bolt size, pattern and depth of attachment bolt holes in the Hot Leg are the same for the RTD, the PDT and Sampling nozzle j MNSAs. The four objectives of this analysis relate to the machined tapped holes made in the '

pipe to receive the attachment bolts. The objectives are as follows:

(1) Establish the maximum permissible bolt load on the tapped holes considering the minimum engagement length (including tolerance stack-up) for the bolt threads.

l (2) Show that the ASME Code (Reference 7) area of reinforcement requirement for the nozzle i penetration remams in compliance, when the areas removed by the tapped holes and defective weld including all of the buttering area are considered. This area of reinforcement area is re-evaluated for both the standard MNSA design as well as the special case for a sampling nozzle MNSA design.

(3) Calculate the range of stress intensity for the tapped holes to demonstrate that the ASME Code allowable of 3 Sm for primary plus secondary stresses is not exceeded.

(4) Perform a fatigue evaluation for the base of the tapped holes to demonstrate that the ASME Code total usage factor (U) requirement of 1.0 is not exceeded.

The results of this analysis will demonstrate that the use of a MNSA in the hot leg pipe will comply with the ASME Code requirements. This analysis can be used for future repair work on any other leaking partial penetration nozzles provided their criteria are the same as those used in this report.

2.0 ASSESSMENT OF SIGNIFICANT DESIGN CHANGES The "as-designed" configurations for the Unit 3 hot leg pipes defmed in References 3, 4, 5, and 13 are considered. The design conditions and operating transients (TNS) as presented in Reference 6 are not changed by the results of this analysis.

CSE-99-046

A-WATERFD-9449-1213. Rev 01 Page 5 of 25 l ABB - CENP PROPRIETARY 3.0 ANALYTICAL TECHNIOUES The ASME Code design force capacity of the thread shear area is determined. The analytical  ;

expressions used in this determination are classical to the structural discipline.

The tapped hole areas as well as the defective weld including all of the buttering areas are subtracted from the available reinforcement area about a penetration following the rules of the 1 ASME Code, Reference 7, and conforming to the reinforcement limits stated in the piping design report, Reference 6.

The structural evaluation for the base of a tapped hole is accomplished using acknowledged stress expressions for the outside surface of a cylindrical shell. The range of stress intensity is  !

evaluated first followed by the fatigue analysis. In the fatigue evaluation, the longitudinal (c.)

and circumferential (ae) stresses determined in the range investigation are multiplied by the maximum concentration factor of five (5) per Paragraph NB-3222.4 (e) (2) of the ASME Code. The peak stress intensities are then calculated, and the fatigue usage values determined i using the applicable ASME Code Sa/N curves.

The analysis applies to all the partial penetration nozzle locations including RTD, the PDT (Pressure Measurement), and Sampling nozzles. The bolt size, pattern and depth of attachment j bolt holes in the hot leg are the same for the RTD, the PDT and Sampling nozzle MNSAs, and  !

therefore, the attachment bolt load and thread calculations are the same.

l 4.0 SELECTION OF DESIGN INPUTS The design conditions for the Waterford Unit 3 hot leg piping are from Reference 6, page 4, and are as follows:

Design Pressure: 2500 psia Design Temperature: 650*F The transient conditions for the hot leg piping are also from Reference 6, pages 4, 5, A-22, A-23, and B-15.

i l

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CSE-99-046 '

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A-WATERFD-9449-1213. Rev 01 Page 6 of 25

- ASB - CENP PROPRIETARY 5.0 ASSUMPTIONS The assumptions included in this design analysis are as follows:

(1) In the determination of the maximum allowable load based on thread shear, the minimum length of engagement used is 0.525 in. in lieu of the 0.68 in. minimum thread depth given in Reference 1. The lower value for the length of engagement results from the stack-up of tolerances on the bolt and mating compression parts and the neglect of incomplete threads. The use of a lower length of engagement results in a lower value for the allowable load.

(2) In the determination of the bolt installation axial force, a mean coefficient of friction of 0.15 is conservatively used in lieu of the 0.18 given in Reference 11. For a known installation torque, use of a lower value results in a higher axial force.

(3) In the re-evaluation of the reinforcement area, the minimum design pipe wall thickness of 3.75 in. is used for the standard MNSA design and an actual pipe wall thickness of 3.85 in. for the special case of the sampling nozzle MNSA design. The 3.85 in.

dimension comes from the subtraction of the maximum clad thickness of 0.25 in. from 4.10 in., which is less than the thicknesses from Reference 18 where the range was 4.106 in. to 4.207 in. for the RTD, Sample Tap, and PDT nozzles on Hot Leg Two.

(4) The moments, stated for the hot leg pipe on Page A-22 of Reference 6, were considered the most critical with respect to the seismic and expansion loads that exist on the hot leg piping.

(5) The pipe moments are assumed to be transverse section loads; therefore, o, = Mc/I is applicable.

(6) The tapped holes are positioned on the pipe section at the location of the maximum bending stress; therefore, shear stresses are not considered in calculating stress intensities.

(7) Longitudinal stresses (o) are conservatively calculated for the pipe outside surface rather than at the base of the tapped holes.

(8) The instantaneous coefficient of thermal expansion (ot)i is used, rather than the usual ,

mean value, for calculating thermal stress; therefore, the thermal stresses are I approximately 15 percent higher.

(9) The thennal stress imposed on the outside surface of the pipe by the temperature difference between the linear and non-linear temperatures is conservatively omitted.

(10) The ASME Code maxunum recommended stress concentration factor of five (5) per i Paragraph NB-3222.4 (e) (2) is used to determine the peak stress intensities.

CSE-99-046 i

A-WATERFD-9449-1213. Rev 01 '

Page 7 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.1 SEAL ASSEMBLY BOLTED CONNECTION TO PIPING EVALUATION 6.1.1 DETERMINATION MAXIMUM ALLOWABLE LOAD BASED ON THREAD SHEAR Reference 7; Paragraph NB- 3227.2(a) and Reference 8, page 59 ASn = x n Le Ds [1/(2 n) + .57735(Ds, - En )] = .568 in 2 ass = x n Le Kn (1/(2 n) + .57735(Es - Kn )] = .419 in 2

Where: Threads = .500-20 UNF-2B; Reference 1 ASn = Internal Thread Shear Area (Pipe) '

ass = External Thread Shear Area (Bolt) n = Threads per Inch = 20 Le = Length of Thread Engagement = .525 in.(min), Refs. 1, 2, 14, 15 & 16)

Ds = Min. Major Diameter ofExternal Thread = 0.4906 in.; Reference 9, page 27 l !

i En = Max. Pitch Diameter ofInternal Thread = 0.4731 in., Reference 9, page 27 Kn = Max. Minor Diameter ofInternal Thread = 0.457 in., Reference 9, page 27 Es = Min. Pitch Diameter ofExternal Thread = 0.4619 in., Reference 9, page 27 t =0.6S. = Thread Shear Stress Allowable (Reference 7; Paragraph NB- 3227.2(a))

For Pipe: t, = .6(18.4) = 11.04 ksi S. = 18.4 ksi @650*F; Ref. 7, Pg. 389 (C-Mn-Si)

Pipe Material is SA-516, Gr.70; Refs. 3 & 13 For Bolt: T = .6(26.9) = 16.14 ksi S. = 26.9 ksi @ 650'F; Reference 7, Page 401 Bolt Materialis SA453, Gr.660; Reference 1 The maximum allowable bolt load based on the available thread area is:

F,=t, ASn F, = 6.27 kips F.=t ass F, = 6.76 kips Therefore, the 6.27 kips value is the maximum allowable bolt load. The maximum applied bolt load during operation is calculated in Reference 17 and compared to this allowable value.

CSE-99-046 l

l 1

A-WATERFD-9449-1213, Rev 01 Page 8 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.1 SEAL ASSEMBLY BOLTED CONNECTION TO PIPING EVALUATION (Cont'd) 6.1.2 EVALUATION OF INSTALLATION TORQUE (PRELOAD) 1 From Reference 12 the equation for the relation between torque and the axial load F is: I T = (F/24) [D v + (d, Tan ( p + 4 )) / Cos(a)]

Substituting, using the below values and solving for F F= 3987 lb. for Preload Condition Where: T = Maximum Torque = 32 FT - LBS , Reference 1 D = Average Diameter of Bolt Head = .5( .75 + .5 ) = .625 in.

Lubricant = Neolube or approved equivalent, Reference 10 v = Coefficient of Friction Beneath Bolt Head,0.18, Ref.1I (Conservatively, use 0.15) d, = Pitch Diameter of Bolt Threads = .4619 in., Minimum value per Reference 9 n = Threads per Inch = 20 p = Helix Angle of Threads = Tan( 1/n )/( xd, ) =~ 1.9734' p = Coefficient ofFriction @ Thread Interface = 0.15

& = Thread Friction Angle = Tanp = 8.53l' j

ot = 1/2 Profile Angle ofThreads = 30' ~

l The applied preload of 3.99 kips is less than the allowable bolt load of 6.27 kips as calculated on the preceding sheet.

i CSE-99-046

A-WATERFD-9449-1213. Rev 01 1 Page 9 of 25 ABB - CENP PROPRLETARY 6.0 DETAILED ANALYSIS 6.2 RE-EVALUATION OF PENETRATION REINFORCEMENT AREA FOR STANDARD MNSA The reinforcement : ea calculation for the standard MNSA design is accomplished by using the ASME code rules from Paragraphs NB-3643.3 as well as NB-3332.2. Paragraph NB-3630

{

gives the alternative rules for using other sections outside of the piping design. I 6.2.1 DETERMINATION OF EXCESS PIPE WALL THICKNESS ( to )

The minimum wall thickness, t., of the hot leg pipe required for design pressure is calculated below using equation (3) from Reference 7, Paragraph NB-3641.1.

t. = [ P d + 2 a( S. + P y ) ] / 2( S. + P y - P ) = ( P R, ) / ( L .6P )

t. = 3.144 in.

Then, the excess pipe wall thickness, t , is

t. = ta - t. = 3.75 - 3.144 = 0.606 in.

Where: P = Design Pressure = 2.5 ksia ID = Clad Surface Inside Pipe Diameter = 42 in. Maximum; Reference 3 R, = ( ID/2 ) + tu = 21.25 in.

a = Corrosion Allowance = 0; Cladded Pipe Pipe Material = SA-516, Gr 70; Reference 3 S. = Allowable Stress = 18.4 ksi @ 650*F; Reference 7, page 389 ( C-Mn-Si )

y = 0.4; Reference 7, Paragraph NB-3641.1

. ta = Minimum Design Pipe Wall Thickness = 3.75 in.; Reference 3 The MNSA clamp holes will protrude into the minimum wall thickness, t., by 1.125 - 0.606 or 0.519 inches as shown in Figure 6.2-1. Thus, the remainmg wall thickness will be less than the minimum wall thickness in these local regions. This condition is acceptable provided the reinforcement requirements are met in Paragraphs NB-3643.3 and NB-3332.2.

6.2.2 REQUIRED AREA OF REINFORCEMENT (Ref. 7, Paragraphs NB-3643.3 & NB-3332.2)

In the required area of reinforcement calculation, two geometry planes are considered. The first geometry plane is where the MNSA clamp holes are located in relation to the nozzle penetration. This geometry plane is 45 degrees from the longitudinal axis for the standard MNSA clamp hole design. The second geometry plane is on the longitudinal axis where only the nozzle penetration is located. In both geometry planes, the area of the defective weld including all of the buttering is considered to be part of the removed area requiring reinforcement.

l CSE-99-046 u _. __

A-WATERFD-94491213. Rev 01 Page 10 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.2 RE-EVALUATION OF PENETRATION REINFORCEMENT AREA FOR STANDARD MNSA 6.2.2 REQUIRED AREA OF REINFORCEMENT (continued)

A. Geometry Plane where both the MNSA clamp holes and nozzle penetrations are located 1

A, = Required Area of Reinforcement (See Figure 6.2-1) f.

4 A$

= [D, ( t. ) + 2 da-(h - to) + A w] F = 3.458 in'

Where : D, = Max. Diameter of Penetration = 1.00 in., Reference 3

t. = Required Mmimum Thickness of Pipe = 3.144 in. a h = Maximum Depth of Hole = 1.125 in., Reference I da = Diameter of Hole = 0.5 in., Reference 1 p**

t = Excess Pipe Wall Thickness = 0.606 in. ,Y ~

F = Correction Factor which compensates for the variation in Pressure Stresses on different planes with respect to the longitudinal axis of the component

= 0.75 for 45' angle of plane with the longitudinal axis per Ref. 7 Fig. NB-3332.2-1 A w = Area ofdefective weld = 2 (A i + A2 + A3 + A4) = 0.947 in Where:

t. = weld + butter thickness = 15/16 in. (Reference 13) tu = clad thickness = 1/8 in. (conservatively, use minimum value, Ref 3)

r. = weld + butter radius = 1/2 in. (Reference 13)

A i = (t. - tu)(l.25 - D, )/2 = 0.1016 inz -

2 A = 0.5 (t. - tu - r.) [2 r./cos(l5')

2

+ (t. - tu - r.) tan (15')] = 0. I 748 in' A3 = x (75/360) r.2 = 0.1636 in A4=0.5r.2 tan (15') = 0.0335 in' A. = Area Available for Reinforcement (See Figure 6.2-1) 2

= 2 L A( Im )

• D r.e( tm )

• 2( de )*Im = 3.939 in Where: La = Limit of Reinforcement on Each Side of Penetration L 4= ta + D, /2 = 4.25 in.; Reference 7, Paragraph NB-3643.3(c)(1)(b) and NB-3334.1 ta = 3.75 in. = Minimum Pipe Thickness

t. = Excess Pipe Thickness = 0.606 in.

A = 3.939 in' > A, = 3.453 ig: l Therefore, the reinforcement area requirement is acceptable in this geometry plane.

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l CSE-99-046 )

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A-WATERFD-9449-1213. Rev, 01 Page11of25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.2 RE-EVALUATION OF PENETRATION REINFORCEMENT AREA FOR STANDARD MNSA 6.2.2 REQUIRED AREA OF REINFORCEMENT (continued)

B. Geometry Plane on the Longitudinal Axis where only the nozzle penetration is located A, = Required Area of Reinforcement

= (D, ( t. ) + A w) F = 4.091 in' Where : D,,,, = Max. Diameter of Penetration = 1.00 in., Reference 3 i t. = Required Minimum Thickness of Pipe = 3.144 in.

F = Correction Factor which compensates for the variation in Pressure Stresses on different planes with respect to the longitudinal axis of the component

= 1.0 for the longitudinal axis per Ref. 7 Fig. NB-3332.2-1 2

A w = Area of defective weld = 2 (Ai + A2 + A3 + A4) = 0.947 in l A. = Area Available for Reinforcement

= 2 La( t ) - D,,,,( t ) = 4.242 in2 Where : LA = Limit of Reinforcement on Each Side ofPenetration in the longitudinal axis, which is limited to 4.00 in because of the 8 inch spacing between nozzles on the lonnitudinal axis oniv. as shown in References 4 and 13.

t = Excess Pipe Thickness = 0.606 in.

A. = 4.242 in' > A, = 4.091 in 2 Therefore, the reinforcement area requirement is acceptable in this geometry plane.

l l

l Cf E-99-046  ;

i

A WATERFD-9449-1213. Rev. 01 Page 12 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.2 RE-EVALUATION OF PENETRATION REINFORCEMENT AREA FOR STANDARD MNSA 6.2.2 REQUIRED AREA OF REINFORCEMENT (continued)

Limits of Reinforcement: Reference 7, Paragraph NB-3643.3(c)(1)(b) and NB-3332. l(b) 2

/ 3of the required reinforcement must be within Li of the centerline of the nozzle opening.

Lx' = D, /2 +.5-[(R, +.5(t)} t]4 = 5.16 in.

where: D, , = 1.00 in., R,, = 21.25 in., t = 3.75 in. (Minimum Pipe Thickness)

Since Li = 5.16 in. > tr = 4.25 or 4.00 in. depending upon the geometry plane, all the area furnished is within Li and therefore the 2/ 3requirement is satisfied.

6.2.3 REINFORCEMENT OVERLAP CONSIDERATION (Ref. 7, NB-3643.3(d)(3) and NB-3335(f))

Metal available for reinforcement shall not be considered as applying to more than one opening.

For the partial penetration nozzles and the neighboring Surge and Shutdown Cooling Outlet Nozzles the minimum spacing is 20.73 in.; therefore, there is no overlap and this requirement is  !

met. ,

)

i Fieure 6.2 1 Reinforcement of Partial Penetration Nozzle Opening i b^ = t + DJ Pi ing L 4= 14.438" 1.906" (Ref.1)

L ,

t. =.606" j

I I l A. I i i i i u i  %!A , % P6* Fd H j

, os ., g :

i 4 h=1.125" ' #] ,

l

/

d w 4 5" k A, I

}

{ p 7

t. = 3.144"

/I I Adjacent Large t, = t. + t. = 3.75" Nozzle Opening D.,,,, = 1.00" CSE-99-046

_ _ _ . ~ _ _ . _ . _ . _ _ _ _ _ _ . _ _ _ . _ - _ _ _ _ _ . . . _ _ _ ___ _._. _ ....._

A-WATERFD-9449-1213, Rev. OI 1 Page 13 af 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.3 RE-EVALUATION OF PENETRATION REINFORCEMENT AREA FOR SAMPLING NOZZLE MMiA l

This reinforcement area calculation is for the special case of the sampling nozzle MNSA design where the MNSA clamp holes cannot be tapped on the geometry plane 45' to the longitudinal axis. Instead its geometry plane is 20' to the longitudinal axis. An actual pipe thickness of 3.85 in. (Assumption (3) in Section 5.0) is used in place of the minimum design pipe wall thickness of 3.75 in. The reinforcement area calculation is accomplished by using the ASME code rules from Paragraphs NB-3643.3 as well as NB-3332.2. Paragraph NB-3630 gives the alternative rules for using other sections outside of the piping design.

6.3.1 DETERMINATION OF EXCESS PIPE WALL THICKNESS ( to )

The minimum wall thickness, t., of the hot leg pipe required for design pressure is calculated below using equation (3) from Reference 7, Paragraph NB-3641.1.

t. = [ P d + 2 a( S.

• P y ) ) / 2( S. + P y - P ) = ( P R, ) / ( S. .6P )

t. = 3.144 in. I Then, the excess pipe wall thickness, to, is to = 14- t. = 3.75 - 3.144 = 0.706 in.

Where: P = Design Pressure = 2.5 k5ia ID = Clad Surface Inside Pipe Diameter = 42 in. Maximum; Reference 3 i R, = ( ID/2 ) + ta.4 = 21.25 in.

a = Corrosion Allowance = 0; Cladded Pipe Pipe Material = SA-516, Gr 70; Reference 3 S. = Allowable Stress = 18.4 ksi @ 650*F; Reference 7, page 389 ( C-Mn-Si )

y = 0.4; Reference 7, Paragraph NB-3641.1 ta = Actual Design Pipe Wd Thickness = 3.85 in.; (Assu.nption (3) in Section 5.0)

The MNSA clamp holes will protrude into the minimum wd thickness, t., by 1.125 - 0.706 or 0.419 inches as shown in Figure 6.3-1. Thus, the remaimng wd thickness will be less than the muumum wall thickness in these local regions. This condition is acceptable provided the reinforcement requirements are met in Paragraphs NB-3643.3 and NB-3332.2.

6.3.2 REQUIRED AREA OF REINFORCEMENT (Ref. 7, Paragraphs NB-3643.3 & NB-3332.2)

In the required area of reinforcement calculation, two geometry planes are considered. The first geometry plane is where the MNSA clamp holes are located in relatic.;i to the nozzle penetration. This geometry plane is 20 degrees from the longitudinal axis for the special case of the sampling nozzle MNSA clamp hole design. The second geometry plane is on the longitudinal axis where only the nozzle penetration is located. In both geometry planes, the area of the defective weld including all of the buttering is considered to be part of the removed area requiring reinforcement.

CSE-99-046

._ . _ _ _ . _ __ _ __ __ __ __. ._. ~.

l A WATEkFD-9449-1213. Rev 01 Page 14 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.3 MEla 11E-EVALUATION OF PENSI. RATION REINFORCEMENT AREA FOR SAMPLING NOZ 6.3.2 REQUIRED AREA OF REINFORCEMENT (continued)

A. Ceometry Plane where both the MNSA clamp holes and nozzle penetrations are located A, = Required Area of Reinforcement (See Figure 6.3-1) , - #=

43

, = [D,.,( t. ) + 2 dw,-(h - t ) + A w] F = 4.262 in' ,

' t,gg a, 4, ,

Where : D, = Max. Diameter of Penetration = 1.00 in., Reference 3 l t. = Required Minimum Thickness of Pipe = 3.144 ir "

h = Maximum Depth of Hole = 1.125 in., Reference 1 dw. = Diameter of Hole = 0.5 in., Reference 1 Op 2-

_ p *, )

1

t. = Excess Pipe Wall Thickness = 0.706 in. ,y %  !

F = Correction Factor which compensates for the variation in Pressure Stresses on different planes with respect to the longitudinal axis of the component

= 0.945 for 20' angle of plane with the longitudinal axis per Ref. 7 Fig. NB-3332.2-1

{ A.,,,w = Area of defective weld = 2 (Ai + A2 + A3 + A4) = 0.947 in2

, Where:

i t. = weld + butter thickness = 15/16 in. (Reference 13)

u = clad thickness = 1/8 in. (conservatively, use nummum value, Ref. 3)

r. = weld + butter radius = 1/2 in. (Reference 13)

Ai = (t. - tu) (1.25 - D, )/2 = 0.1016 in' A2 = 0.5 (t. - tu - r.) [2 r.,/cos(15') + (t. - tu - r.) tan (15')] = 0.1748 inz 2

A3 = x (75n60) r.2 = 0.1636 in 2

A4 = 0.5 r.' tan (15') = 0.0335 in j

A = Area Available for Reinforcement (See Figure 6.3-1)

= 2 Li ( t. ) - Dy( t. ) - 2( dw. ) t = 4.730 inz Where: L A= Limit of Reinforcement on Each Side of Penetration La = ta + D,,,,/2 = 4.35 in.; Reference 7, Paragraph NB-3643.3(c)(1)(b) and NB-3334.1 ta = 3.85 in. = Actual Pipe Thickness to = Excess Pipe Thickness = 0.706 in.

A. = 4.730 in* > A, = 4.262 in' Therefore, the reinforcement area requirement is acceptable in this geometry plane.

CSE-99-046

A-WATERFD-9449-1213. Rev 01  !

Page 15 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.3

' RE-EVALUATION OF PENETRATION REINFORCEMENT .REA FOR SAMPLING NOZZLE M!f16 6.3.2 REQUIRED AREA OF REINFORCEMENT (continued)

B. Geometry Plane on the Longitudinal Axis where only the nozzle penetration is located A, = Required Area of Reinforcement

= (D,,,,( t. ) + A u) F = 4.091 in' Where : D,,,, = Max. Diameter of Penetration = 1.00 in., Reference 3

t. = Required Mmimum Thickness of Pipe = 3.144 in. '

F = Correction Factor which compensates for the variation in Pressure Stresses on different planes with respect to the longitudinal axis of the component

= 1.0 for the longitudinal axis per Ref. 7 Fig. NB-3332.2-1 2

A u = Area of defective weld = 2 (A i + A2 + A 3+ A4) = 0.947 in A, = Area Available for Reinforcement

= 2 La( t. ) - D, ( t. ) = 4.942 in' '

Where : La = Limit of Reinforcement on Each Side of Penetration in the longitudinal axis, which is limited to 4.00 in because of the 8 inch spacing between nozzles on the lonnitudinal axis only, as shown in References 4 and 13.

to = Excess Pipe Thickness = 0.706 in.

4, = 4.942 in' > A, = 4.091 in' Therefore, the reinforcement area requirement is acceptable in this geometry plane.

CSE-99-046

A-WATERFD-9449-1213, Rev 01 Page 16 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALM 6.3 RE-EVALUATION OF PENETRATION REINFORCEMENT AREA FOR SAMPLING NO771_E M.tfliA 6.3.2 REQUIRED AREA OF REINFORCEMENT (continued)

Limits of Reinforcement: Reference 7, Paragraph NB-3643.3(c)(1)(b) and NB-3334.1(b) 2

/ 3of the epred reinforcement must be within LA' of the centerline of the nozzle opening.

L A' = D,,,,/2 +.5-[{R,,,,,+.5(t)} t]' = 5.22 in.

where: D,,,, = 1.00 in., R,,,,, = 21.25 in., t = 3.85 in. (Actual Pipe Thickness)

Since L4' = 5.22 in. > L =A 4.35 or 4.00 in. depending upon the geometry plane, all the area furnished is within LA' and therefore the 2/ 3requirement is satisfied.

6.3.3 REINFORCEMENT OVERLAP CONSIDERATION (Ref. 7, NB-3643.3(d)(3) and NB-3335(f))

Metal available for reinforcement shall not be considered as applying to more than one opening.

For the sampling nozzles and the neighboring Surge and Shutdown Cooling Outlet Nozzles the minimum spacing is 25.57 in.; therefore, there is no overlap and this requirement is met.

Fieure 6.3-1 Reinforcement of Partial Penetration Nozzle Opening

^~

Pi ing L 4= 14.438" 1.906" (Ref.1) l t ,

m

t. = .706" I

I l A. I

^ i i i i u

?A , 46 G;E h!

' u 9fd x]

)

\ [ h=1.125" k A, l

~

dw&S" ~

{ p y t. = 3.144"

/I I  !

Adjacent Large t = t. + t. = 3.85" Nozzle Opening D , = 1.00" CSE-99-046

A-WATERFD-9449-1213. Rev. Or Page 17 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.4 RANGE OF STRESS INTENSITY EVALUATION OF 1/2" TAPPED HOLES The stresses at the location of the tapped holes for attachment of the MNSA to the hot leg pipe are calculated 4

by classical methods of strength of materials. In order to address the specine location of these holes (i.e. the outside surface of the pipe), the stress calculations are performed according to the provisions of Subsection NB-3200 of ASME Section III, Reference 7.

The drilled holes are considered to be local structural discontinuities that effect the peak stresses and fatigue usage factor in the pipe. Consenatively, a stress concentration factor of 5, the mammum value required per NB 3222.4(e)C) of Reference 7, is used to represent the local structural discontinuity effect of the holes. The effect of these holes on the prunary plus mandary stresses and the range of stress intensity is considered to be negligible. De previous analysis for the hot leg pipe on pages A-21 through A-26 of CENC-1444, Reference 6, was performed by the stress index method for piping as defined in Subsection NB-3600 of Reference 7. 'Ihis method is a simplified approach that conservatively envelops the stresses for all locations and applied loads on the pipe; consequently, it often yields overly conservative results for a specific location. For this reason it is not appropriate for the analysis of the MNSA =*=Amant holes.

6.4.I OPERATING TRANSIENT CONDITIONS CONSIDERED Reference 6 (2) (1) (2) (3) (4) (4) l SYMBOL' TITLE CYCLES ;P

,1 i M x 10( DI: ATi  :

?ni KSIA IN-KIP -

• F. 8 F

Al Heatup 500 2.25 64.042 536 10 B1 Cooldown 500 0.1 1.469 79 -10 B2* Ambient 500 0.0 1.469 70 0 Cl Plant Loading 15000 2.29 64.042 595 19 DI Plant Unloading 15000 2.21 64.042 571 -19 E1 10% Step Inc. 10' 2.35 64.042 611 0 F1 10% Step Dec. 10' 2.15 64.042 611 0 G4 Reactor Trip (100 sec)" 480 1.65 64.042 597 -44 G5 Reactor Trip (125 sec)" 480 1.69 64.042 592 -52 G6 Reactor Trip (400 sec)** 480 1.75 64.042 564 -30 Il Leak Test Heatup 200 2.25 64.042 391 10 12 Leak Test Cooldown 200 2.25 1.469 109 -10 M1 Hydrostatic Test 10 3.125 1.469 400 0 S1 100% SS + OBE 200 2.25

~

72.250 611 0 S2 100% SS - OBE 200 2.25 55.834 611 0

- (1) Reference 6, pages 4 & 5 (2) Reference 6, page A-23 (3) Reference 6, page A-22 (Moments for transient cases are combinction of thermal + dead weight)

(4) Reference 6, page B-15: Where a positive AT i indicates that the outside surface of the pipe is cooler than the inside surface; thereby, inducing an outside surface tension stress.

• - Condition for the end of the Cooldown cycle

" - Includes Loss of Flow and Loss of Load CSE-90-046

,- - .. - .= .- _. ._- . - . . - . - _ . - .. - . - - - . .. - -.

A-WATERFD-9449-1213. Rev. 01 Page 18 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.4 RANGE OF STRESS INTENSITY EVALUATION OF 1/2" TAPPED HOLES (cont'd) 6.4.2 UNCONCENTRATED STRESS EQUATIONS Longitudinal or Axial Stress Component in Pipe (ksi)

Ic x= (b' P)/2Rst + Mro /I + Eni/2(l. )[AT ] or Io, = 2.5427 P + 0.1705x10M + (Eai)/1.4[AT ] (radius to OD conservatively. used)

Circumferential or Hoop Stress Component in Pipe (ksi)

Ice = (b P)/t + Eai/2(1-p)[ATi] or Ice = 5.6 P + (Eai/1.4)(ATi]

Radial Stress Component in Pipe (ksi) e, = - (h/t) P = - 0.300 P (Stress varies linearly from -P on ID to 0 on OD) i Where:b = Clad Surface Inside Radius = 21 in. (maximum value conservatively used, Reference 3)

P = Pressue, ksia R. = Mean Radius of Pipe Base Metal = 23.125 in. (Reference 3) t = Pipe Base Metal Wall Thickness = 3.75 in (minimum value conservatively used, Reference 3)

.M = Pipe Mechanical Moment, in-kip r, = Outside Radius of Pipe = 25 in. (maximum value on page A-21 of Reference 6) j I = Pipe Moment ofInertia = 146,647 in'(page A-21 ofReference 6) '

E = Modulus ofElasticity of Pipe Base Metal @ Mean Temperature of Pipe (TM) I ai = Instantaneous Coefficient of Thermal Expansion @ TM

' = Poisson's Ratio = 0.3 AT i= Linear Thermal Gradient Through the Pipe Base Metal h = 1.125 in. = Maximum Permitted Tapped Hole Depth (Reference 1)

The above stress equations are based on the classical methods of strength of materials. The resulting stress components are then used to calculate stress intensities that can be compared to the allowables given in Subsection NB-3200 of Reference 7. The stress intensities resulting from these equations are comparable to the values calculated from the equation for Sn on page A-23 of Reference 7.

After the application of a stress concentration factor to the axial and hoop stress components, the resulting stress intensities in Section 6.5.1 are comparable to the values calculated from the equation for Sp on page A-25 of Reference 7.

CSE-99-046

A-WATERFD-9449-1213. Rev 01 Page 19 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.4 RANGE OF STRESS INTENSITY EVALUATION OF 1/2" TAPPED HOLES (cont'd) 6.4.3 Eai THERMAL STRESS PARAMETERS (5) (6)

TNS* TM" . E x 10' . mix 104 Eai SYMBOL ~ 'F KSI - infin/*F. Kip /in2,op Al 536 '26.148 8.1040 0.2119 B1 79 27.886 6.1090 0.1704 B2 70 - -

N/A Cl 595 25.735 8.3310 0.2144 D1 571 25.903 8.2398 0.2134 El 611 - -

N/A F1 611 - -

N/A G4 597 25.721 8.3386 0.2145 __

G5 592 25.756 8.3196 0.2143 G6 564 25.952 8.2132 0.2131 Il 391 27.036 7.5022 0.2028 I2 109 27.840 6.2432 0.1738 M1 400 - -

N/A S1 611 - -

N/A S2 611 - -

N/A N/A = Not/ Applicable, ATi = 0 (5) Reference 7; page 414, Carbon Steels with C 5 0.30%

(6) Reference 7; page 413, Carbon Steel Group

- Reference 6, page A-23

• * - Reference 6, page B-15 The material properties shown in th'e above table can be compared to the values from page A-24 in Reference 6:

E = .027910' ksi at 70 F 4

a = 6.0710 in/in/ F at 70 F i

CSE-99-046 I

A-WATERFD-9449-1213. Rev 01 Page 20 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.4 RANGE OF STRESS INTENSITY EVALUATION OF 1/2" TAPPED HOLES (cont'd) 6.4.4 RANGE OF STRESS INTENSITY Stress Intensities TNS* Io, Ice - e, Io, . Ice Io, - o, Ice - o, l SYMBOL KS1 KSI KSI KSI KSI KSI Al 18.154 14.114 -0.675 4.040 18.829 14.789 BI -0.712 -0.657 -0.030

-0.055 -0.682 -0.627 B2 0.250 0.000 0.000 0.250 0.250 0.000 Cl 19.652 15.734 -0.687 3.918 20.339 16.421 D1 13.642 9.479 -0.663 4.163 14.305 10.142  !

El 16.895 13.160 -0.705 3.735 17.600 13.865 F1 16.386 12.040 -0.645 l

4.346 17.031 12.685 G4 8.374 2.499 -0.495 5.875 8.869 2.994 G5 7.257 1.505 -0.507 5.752 7.764 2.012 G6 10.801 5.233 -0.525 5.569 11.326 5.758 11 18.089 14.049 -0.675 4.040 18.764 14.724 12 4.730 11.358 -0.675 -6.628 5.405 12.033 M1 8.1 % 17.500 -0.938 -9.304 9.134 18.438 S1 18.040 12.600 -0.675 5.440 18.715 13.275 ,

S2 15.241 12.600 -0.675 2.641 15.916 13.275

'{

• Reference 6, page A-23 RANGE Io, - Ice = 5.875 - (- 9.304) = 15.178 ksi Io, - o =r 20.339 -(- 0.682) = 21.021 ksi < 3Sm = 3( l8.4 ) = 55.2 ksi @ 650*F Io - o, = 18.438 - (- 0.627) = 19.064 ksi The resulting range of primary plus secondary stress intensity of 21.02 ksi as calculated above can be compared to the value of Sn = 36.33 ksi from page A-24 of Reference 6. As noted earlier, the previous result was calculated by a simplified, overly conservative approach that envelops any location and all applied loads for the hot leg pipe. The present result is calculated by a more 1 accurate method and is representative only of the speci6c locations of the MNSA attachment bolts.

The previous result still applies for other locations in the hot leg pipe.

1 CSE-99-046

A-WATERFD-9449-1213, Rev. 01 Page 21 of 25 ABB - CENP PROPRIETARY 6.0 )

DETAHED ANALYSIS 1 6.5 FATIGUE EVALUATION OF TAPPED HOLES Reference 7; Paragraph NB-3222.4 6.5.1 PEAK STRESS INTENSITIES Njgig Per Reference 7, Paragraph NB-3222.4 (e) (2), a maximum stress concentration factor of 5.0 is used on the longitudinal and circumferential stresses.

(7) (8) (8) Peak Stress Intensides TNS 5Io.i; 5Io. o, SIo - SIo.  : SIo.- e, SIo. - o, SYMBOL KSI7 :KSI - KSI- KSI : KSI KSI Al 90.769 70.568 -0.675 20.201 91.444 71.243 B1 -3.560 -3.284 -0.030 -0.276 -3.530 -3.254 B2 1.252 0.000 0.000 1.252 1.252 0.000 Cl 93.258 78.668 -0.687 19.590 98.945 79.355 D1 68.210 47.397 -0.663 20.813 68.873 48.%0 El ,,84.473 65.800 -0.705 18.673 85.178 66.505 F1 81.930 60.200 -0.645 21.730 82.575 60.845

G4 41.870 12.4 % -0.495 29.373 42.365 12.991 G5 36.287 7.525 -0.507 28.762 36.794 8.032 G6 54.007 26.163 -0.525 27.844 54.532 26.688 Il 90.445 70.244 -0.675 20.201 91.120 70.919  !

12 23.650 56.792 -0.675 -33.142 24.325 57.467 I S1 90.199 63.000 -0.675 27.199 90.874 63.675 S2 76.204 63.000 -0.675 13.204 76.879 63.675 (7) Per Reference 7, the 10 cycles ofHydrostatic Test (TNS II) are omitted.

l (8) The thermal stress imposed on the outside surface of the pipe by the temperature difference  ;

between the linear and non-linear temperatures is omitted, because it would decrease the o and o. stresses.

The peak stren intensities calculated above can be compared to the tabulated values of Sp on pages A-25 and A-26 in Reference 6. The value Sp is the stress index equivalent of a peak stress intensity.  ;

Two of the three peak stress intensities tabulated in the table above are more conservative than the i

previous result because of the assumption of a stress concentration factor of 5.0 to represent the local discontinuity effect of the tapped hole.

CSE-99-046

A-WATERFD-9449-1213. Rev 01 Page 22 of 25 ABB - CENP PROPRIETARY 6.0 DETAILED ANALYSIS 6.5 FATIGUE EVALUATION OF TAPPED HOLES (continued) 6.5.2 ACCUMULATED FATIGUE DAMAGE (IU) 6.5.2.1 5Io,- SIce PEAK STRESS INTENSITY (SI)

Maximum Minimum (9) (10)

TNS SI- TNS SI. .Sa .Sa* N n U = n/N KSI" KSI ' KSI KSI .

G4 29.373 I2 -33.142 31.258 37.138 10813 200 0.01850 G4 29.373 B1 -0.276 14.825 17.614 158066 280 0.00177 S1 27.199 B1 -0.276 13.737 16.322 210174 200 0.00095 F1 21.730 B1 -0.276 11.003 13.073 667878 20 0.00003 F1 21.730 S2 13.204 4.263 5.065 e 200 0.0 (9) Per Reference 7, Paragraph NB-3222.4(e)(4), the following:

Sa* = E n /E65o( Sa ) = 1.195a Where : E.. = 30 x 10' psi ; Ref. 7, Figure I-9-1, page 433 E63o = 25.25 x 10' psi ; Ref. 7, Table I-6.0, page 414, Carbon Steels with C 5 0.30%

(10) Reference 7; Figure I-9-1, page 433, UTS 5 80 ksi EU = 0.0212 < Allowable = 1.0 -

6.5.2.2 SIo,- o, PEAK STRESS INTENSITY (SI)

Comment: Refer to notes (9) and (10) in report Section 6.5.2.1.

Maximum Minimum (9) (10)

TNS: -

?TNSg -

iN n U = n/N C1 98.945 B1 -3.520 51.238 60.877 2346 500 0.21317,,

Cl 98.945 12 24.325 37.310 44.329 6331 200 0.031 fs Cl 98.945 G5 36.794 31.076 36.922 11030 480 0.N352 Cl 98.945 D1 68.873 15.036 17.865 150195 13820 0.09201 Al 91.444 D1 68.873 11.286 13.409 531441 500 0.00094 Il 91.120 D1 68.873 11.124 13.216 605369 200 0.00033 S1 90.874 D1 68.873 11.000 13.070 669294 200 0.00030 El 85.178 D1 68.873 8.153 9.686 e 280 0.0 IU = 0.3819 < Allowable = 1.0 CSE-99-046

A-WATERFD 9449-1213. Rev 01 Page 23 of 25 ABB - CENP PROFRIETARY 6.0 DETAILED ANALYSIS 6.5 FATIGUE EVALUATION OF TAPPED HOLES (continued) 6.5.2 ACCUMULATED FATIGUE DAMAGE (IU)(continued) 6.5.2.3 SIce - o, PEAK STRESS INTENSITY (SI)

Comment: Refer to notes (9) and (10) in report Section 6.5.2.1.

Maximum Minimum (9) (10)

TNS SI' tnt: SI Sa. Sa*- N n U = n/N KSI: ' KSI KSI KSI Cl 79.355 B1 -3.254 41.305 49.075 4659 500 0.10731 Cl 79.355 G5 8.032 35.662 42.370 7240 480 0.06630 Cl 79.355 D1 48.060 15.648 18.591 130101 14020 0.10776 Al 71.243 D1 48.060 11.592 13.772 456449 500 0.00110 Il 70.919 D1 48.060 11.430 13.580 486752 200 0.00041 El 66.505 D1 48.060 9.223 10.958 m 280 0.0 IU = 0.2829 < Allowable = 1.0 The fatigue usage factors calculated in Sections 6.5.2.1 through 6.5.2.3 can be compared to the previous results for the hot leg pipe given in Reference 6, page A-26. The current maximum result is 0.3819 versus the earlier value of 0.0123. The increase is caused by the conservative stress concentration factor of 5.0 that has been used in this calculation. This assumption was made to demonstrate that the usage factors would be acceptable even though the maximum stress concentration factor required by ASME Section III, Reference 7, is assumed.

CSE-99-046 i

i

~ A-WATERFD-9449-1213. Rev. 01 Page 24 of 25 ABB - CENP PROPRIETARY 7.0 RESULTS / CONCLUSIONS 7.1 RESULTS The objectives and their resolutions are presented below.

Objective (1): Establish the maximum permissible bolt load.

Resolution: The allowable maximum load, including preload during MNSA impacting, is 6.27 kips. The maximum applied bolt load is calculated and shown to meet this allowable in Reference 17. -

Objective (2): Considering the tapped holes and defective weld with all of the buttering area,  !

show that the ASME Code area of reinforcement requirement for the RTD l penetration is maintained.

Resolution: l For the standard MNSA design with minimum pipe thickness of 3.75 m.

Area Available = 3.939 in' > Area Required = 3.458 in' for the geometry plane at 45' to longitudin Area Available = 4.242 in: > Area Required = 4.091 in' for the longitudinal axis geometry For :he special case of a sampling nozzle MNSA design with actual pipe thickness of 3.85 in.

Area Available = 4.730 in' > Area Required = 4.262 in' for the geometry plane at 20' to longitudinal Area Available = 4.942 in* > Area Required = 4.091 in' for the longitudinal axis geometry Objective (3): Demonstrate that the ASME Code primary plus secondary stress intensity .

allowable of 3 Sm is not exceeded.

4 i

Resolution: For minimum pipe thickrmss of 3.75 in.

SIm = 21.02 ksi < Allowable 3Sm = 55.2 ksi Objective (4): Demonstrate that the ASME Code fatigue usage factor allowable of 1.0 is not exceeded.

Resolution: For minimum pipe thickness of 3.75 in.

IUm = 0.382 < Allowsble = 1.0

7.2 CONCLUSION

Acknowledging the assumptions sta'ed in report Section 5.0 and following limits stated above, the MNSAs can be attached to a hot leg pipe at all the partial penetration welded nozzle locations and meet the ASME Code stress requirements.

CSE-99-046

A-WATERFD-9449-1213. Rev 01 Page 25 of 25 ABB - CENP PROPRIETARY s.o BEEERENCES l

1. ABB/CE Drawing No. E-MNSAWFD-228-001, Rev. 02, " Hot Leg RTD Mechanical Nozzle Seal 1 Assembly Flat"

2. ABB/CE Drawing No. E-MNSA-228-004, Rev.05, Sheets 1 to 4, " Mechanical Nozzle Seal Assembly Details"

3. ABB/CE Drawing No. E-74470-722-001, Rev. 04, " Primary Pipes, Waterford III Piping"

4. ABB/CE Drawing No. E-74470-771-003, Rev. 03, " Primary Piping Assembly, Waterford III Piping"

5. ABB/CE Drawing No. D-74470-772-001, Rev. 04, " Instrument Nozzles, Waterford III Piping"

6. ABB/CE Report No. CENC-1444, " Analytical Report for Wateford Unit No. 3 Piping", May 1981.

7. ASME Boiler and Pressure Vessel Code,Section III, Division 1 - Subsection NB, " Rules for Construction ofNuclear Power Plant Components",1971 Edition thru Wi' iter 1971 Addenda.

8. USA Standard ASA-Bl.1-1960," Unified Screw Threads",1960.

9. FED-STD-H28/2, " Federal Standard Screw-Thread Standards for Federal Services Section 2 Uni 6ed Thread Form and Thread Series for Bolts, Screws, Nuts, Tapped Holes and General Applications",31 March 1978.

10. ABB/CE Report No. TR-PENG-012, Rev. 00, " Test Report for Verification Testing of RTD Nozzle Seal Assembly", February 1,1995. -

11. ABB/CE Report No. C" 9448-CSE91-1101, Rev.1, " Acceptance Criteria for Florida Power and Light St. Lucie #1 and #2 Manway and Handhole Stud Hole Threads", Appendix C " Good Bolting Practices", Volume 1 issued by EPRI - 3412 Hillview Avenue Palo Alto, Califomia 94304.

12. "Preloading of Bolts", by Bernie J. Cobb; Product Engineering, August 19,1963.

13. ABB/CE Drawing No. D-74470-771-002, Rev. 03, "Pnmary Pipe Assembly, Waterford III Piping"

14. ABB/CE Drawing No. E-MNSAWFD-228-002, Rev. 02, " Hot Leg PDT MNSA"

15. ABB/CE Drawing No. E-MNSAWFD-228-003, Rev. 02, " Hot Leg Sampling MNSA"

16. ABB/CE Drawing No. E-MNSA-228-013, Rev. 07, " Mechanical Nozzle Seal Assembly Details"

17. ABB/CE Report No. C-PENG-CALC-020, Rev. 00," Analysis of Waterford Unit 3 Hot Leg RTD MNSA"

18. ABB/CE E-mail to K. M. Rajan, et. al. from G. P. Bundick on "H.L. Pipe Thickness" (copy in Attachment 2).

CSE-99-046

A-WATERFD-9449-1213, Rev. 01 Page 2-1 of 2-3 1 ABB . CENP PROPRETARY l ATTACHMENT 2 REFERENCES

18. ABB/CE E-mail to K. M. Rajan, et. al. from G. P. Bundick on "H.L. Pipe Thickness", dated 3/12/99.

CSE-99-046

A-WATERFD-94491:13. Rev 01 Page - of 2 3 ABB - CENP PROPRIETARY Gary P. Bundick/CENOlUSNUS/ASB 03/12/99 09:18 AM (pnene: 1.sosner1e)

To:

Kneh M. RasenNSIMS/A880A88_USSEV_IMS se:

Bruce A. SeWCENONSNUS/A880A88.USSEV.lMS. Donald P.

Seeka/CENONSNUS/A880A88_US8EV_IMS, John T.

McGany/CENOMSNUS/A880A88.US8EV_IMS Subject Re: H.L Pies TNehnese p '

1 Here is the inspecton report. They identWied them sonording to the steem Generator, so from top to bosom, the irWormodon is Ihr the RTD, Sergio Top and the PDT on hot log two.

Gary 1

UT of Piping.p Krish M. Rejen/USIMSIA88 1

Krleh M.Rafen/USWAStA38 os/12/se 07:ss Au cenene:.5 4:s.m-ases To: Gary P. GurutWCENOM4NUS/ASSGASS.USSEV_ ass es: 8eune A. newCENONSNUS/ASSGA88.USSEY_Est, DoneM P.

• SlekeCENOM8NuS/ ASS $A80 ug8EY,ast,JohnT.

uscanymencr i" :3_uesav_ mas Subjoet: H.L Pipe Thidmese Please send the documentaton on the H.L pipe field thicknese measuremonte. We cannot do any 1.R i ofHeial ceice vnthout this into. Also what le tie task teleted to justWying buner se rewWorcement for 7 i

I CSE-99-046 i;

4 l

A-WATERFD-94491213. Rev 01 Page 2-3 of:-3 ABB - CENP PROPRETARY ULTRASONIC TECKNESS VAING8 OF N'e

• A??M ON IK3 P! PING NOS A!! thiebuss medhss are is lesbes.

SG 81 Nossis 8 0900 W t ed.Js i 4.1!$

4.!!3 4.lM 4.078 4J06 4.183 4.lM 4.143 .

m ul r--+ a 1500 - 3'ed=4 4.1 41 4.143 4.181 4.147 4.181 4.!G 4.1$1 4.134 30 C Weeds a 1800:

4M 4.239 4.200 AIN 4.1N 4.206 4.2M 4.189 4.182 JJ i 0LWhlbKJ mi evi n am .t-049 -

CSE-99-046

- .s

• 9 :37 l ABB March 23,1999 NOME.99-C-0134 l

Entergy Operations, Inc.

Waterford Steam Electric Station Unit 3 P.O. Box B l Killona, LA 70066  ;

SUBJECT:

Mechanical Naszie Seal Assemblies (MNSA)

Entergy Contract No. NWC00385 Proprietary Information

Dear Mr. Pmetor:

The following documentation has been submitted to Entergy for the subject contract

1. A-WATERFD-9449-1213 Rev 01. Evaluation of An-% Locanons for Mechanical %zzle Seal Assembly on Waterford 3 Partial Penetation Welded Nonlesin the Hot Leg Piping.

l

2. C-PENG-CALC-021 Rev 00, Determination of Waterford 3 Hot Les Seismic Response Spectra and Accelerations for Usein Analysis ofMNSAs

3. Design Report No. C-PENG-DR-006 Rev 01 Addendum to Stress Report No. CENC-1444 Analytical Report for Waterford Unit No. 3 Piping Item 1 is an Attachment to Item 3. Pages 2 through 25 ofitem 1 are marked "ABB -

CENP Pmprietary" Please disregard this propnetary designation. ABB does not  ;

consider this proprietary item 2 includes an Appendix A which is stamped " Proprietary Information". Please disregard this propnetary designation. ABB no longer considers this information y.sg,&.it.

If there are any questions on the above, please call me.

Project xc: G. Bundick ABB Combustion Erigineering Nuclear Operations 00%440iLa%* tomo WW PO sem :01 Tw W urs 18co)rds.t9tt i W C r few E N W504 i

a a a_ m _. c..s. m a A. A,. , . . .d;_- J w,,--wa 4 s_ 4E._a J&&M__t&- a-u4g.as-w.J. .am J_-aM3.,.de_m.LA._.m_JJm4_

_a4.. ._ ,4.4 4 m A _ a hd J -- O Me f.2 WATERFORD 3 SES SQ-MN-347 MECHANICAL NOZZLE SEAL ASSEMBLY (MNSA)

SECTION 2 REVISION RECORD l l

REVISION PAGES DESCRIPTION OF CHANGE PREPARED BY/

NUMBER DATE AFFECTED REVIEWED BY 0 3/23/99 ALL Initial Issue u,m D

v 1

1 i

1 4

t

WATERFORD 3 EQUIPMENT SEISMIC QUALIFICATION Th[

SUMMARY

DATA SHEET SQRT No.: SQ-MN-347  ; Section 6

1. GENERAL INFORMATION 1.0 Component

Description:

Mechanic Nozzle Seal Assembly 2.0 Component UNID No(s): See Section 3 3.0 Vendor and Model No.: ABB CENP - RCS Hot Leg Nozzles 4.0 Size or Range:  %" to 1" diameter nozzles i

5.0 Purchase Order No.: NWC00385 s

6.0 Specification No.: 09270-PE-0140 and C-NOME-SP-0067 7.0 Physical Description of Component:

• Appearance: See drawings: 5817-12141,5817-12142, and 5817-12143

. Dimensions: See drawings: 5817-12141,5817-12142, and 5817-12143

• Weight: See drawings: 5817-12141,5817-12142, and 5817-12143 8.0 Location Where Component is To Be Installed:

• Building: Reactor Containment Building

. Elevation: +15' to +17' elevation

. Cabinet or Panel (if applicable) N/A

. System in Which it is Located: RCS

. Functional Description of Component: ,

e is the Equipment Required For.

() Hot Standby () Cold Shutdown (x) Both () Other (specify) f Page 1

I l

&g O

WATERFORD 3 EQUIPMENT SEISMIC QUALIFICATION

SUMMARY

DATA SHEET SQRT No.: SQ-MN-347  ; Section 6 8.0 Location Where Component is To Be Installed (Cont.):

• How is Component to be Mounted in the Field:

( x) Bolted / Screwed a No. of Bolts See dwgs 5817-12141 thru 12143

• Size of Bolts See dwgs 5817-12141 thru 12143

• Torque Values: See dwgs 5817-12141 thru 12143

() Welded

• Weld Size

• Weld Configuration (See Attachment No. )

• Mounting Orientation: __

See TR-PENG-033 9.0 Qualification Documentation:

• Qualification Report (No., Title, Date):

See the attached list of design reports / calculations (page 7)

• Company That Prepared Report: ABB - CENP

• Company That Reviewed Report: ABB - CENP /

Entergy Operations, Inc.

• Where Report is Filed or Available: SQRT File: SQ-MN-347

• Qualification Method:

() Test (Complete Section 11 Only)

() Analysis (Complete Section ill Only)

(x ) Combination of Test and Analysis (11 & lil)

Prepared By: Ken P. Wilson ((g Date: g.fg py

'/

Reviewed By: John P. BurkeQg hg Date: 3.g3, qq Page 2 \

l

WATERFORD 3 EQUIPMENT SEISMIC QUALIFICATION C-h[

SUMMARY

DATA SHEET SQRT No.: SQ-MN-347  ; Section 6

11. TEST INFORMATION 1.0 Test Frequency Content

() Single Frequency *

(x ) Multi-Frequency 2.0 Test input Motion

( x) Random *

() Sine Beat

() Other (Specify): ,

j 3.0 Test Type j

() Single Axis *

( x)_ Bi-axial

() Tri-axial l

4.0 Number of Qualification Tests l

• 5 OBE

• 1 SSE l

• Other (Specify):

5.0 Test Frequency Range: 1 - 50 Hertz 6.0 Natural Frequencies:

1

• S/S: 11.8 Hertz (C-PENG-CALC-021)

• F/B:

• Vert.:

7.0 Method of Determining Natural Frequencies:

() Lab Test *

() In-Situ Test

• (x ) Analysis 8.0 For Multi-Frequency Tests, TRS Envelopes RRS:

• (x ) YES * () NO e () Not Applicable 9.0 For Single Frequency Tests / input Accelerations Exceed Requirements:

• () YES * () NO + ( x) Not Applicable l

l Page 3

1' I'

(O )

WATERFORD 3 EQUIPMENT SEISMIC QUALIFICATION l

'O [

SUMMARY

DATA SHEET SQRT No.: SQ-MN-347  ; Section 6 )

10.0 Describe Test Mounting and Orientation:

( x) Bolted / Screwed )

• No. of Bolts See TR-PENG-033 j

= Size of Bolts

• Torque Values:

() Welded a Weld Size

• Weld Configuration (See Attachment No. )

. Mounting Orientation on Shake Table: See TR-PENG-033 11.0 Functional Operability Verified: (x) YES () NO () N/A 12.0 Brief Description of Test Results including Any Modifications Made:

No observance of leaks or significant pressure drops 13.0 Brief Description and Resolution of Test Anomalies:

None 14.0 Other Tests Performed:

() Aging (service life) *

() Radiation Exposure

() Mechanical Aging *

(x ) Other See remarks 15.0 Remarks: In addition to the seismic tests performed, the MNSA Qualification also included Hydrostatic and thermal cycle tests.

(TR-PENG-042)

Prepared By: Ken P. Wilson [ Date: 3.fr g Reviewed By: John P. Burke g >Rd Date: 3 2g,qt U Page 4

h WATERFORD 3 EQUIPMENT SEISMIC QUAllFICATION

~ h[

SUMMARY

DATA SHEET SQRT No.: SQ-MN-347  ; Section 6 111. ANALYSIS 1.0 Method of Analysis

• () Static Coefficient Analysis (gp eak X1.5)

• (x) Equivalent Static Analysis (rigid component)

See the attached list of applicable qualification calculations /

reports

• () Response Spectrum Analysis (flexible component)

• () Time History Analysis (flexible component) 2.0 Natural Frequencies in each Direction:

• S/S: 11.8 Hz e F/B:

• Vert:

3.0 Method of Combining Loads:

() Absolute Sum *

( x) SRSS

() Other (Specify):

4.0 Model Used to Represent Component:

( x) 3-D *

() 2-D +

() 1-D

=

() Finite Element *

() Closed-formed Solution 5.0 Computer Software Code (name): None Frequency Range and Number of Modes Considered (if computer code used):

N/A 6.0 Method of Combining Dynamic Response (stress, etc.):

() Absolute Sum *

( x) SRSS

() Other (Specify):

7.0 Damping Values Used:

• OBE 1% Basis for Damping Used Specification 09270-PE-0140

• SSE 1% Basis for Damping Used Specification 09270-PE-0140 Page 5

WATERFORD 3 EQUIPMENT SEISMIC QUALIFICATION Oh[

SUMMARY

DATA SHEET SQRT No.: SQ-MN 347  ; Section 6 8.0 Support Consideration in Analysis:

See the attached list of applicable design reports / calculations 9.0 Critical Structural Elements

. Stress Governing Load or Dynamic Allowable Idei. .. son Location Response Combination Stiesa Stress

. Deflections Maximum Critical Location Maximum Allowable Deflection to Assure Functional Deflections Operability

. Failure Mode:

See the attached list of design reports / calculations i

Prepared By: Ken P. Wilson

/

~

[ Date: psy.ff Reviewed By: John P. Burke bQ ~

Date: 343.q('

O Page 6

WATERFORD 3 8[

-~ - h EQUIPMENT SEISMIC QUALIFICATION SQRT No.:

SUMMARY

DATA SHEET SQ-MN-347  ; Section 6 4

The following documents were prepared for and used as the basis for seismic 4

qualification of the MNSA at Waterford SES Unit-3:

4 C-NOME-ER-0120 Design Evaluation of MNSA for Various Applications at Waterford Unit 3 dated 3/11/99.

C-NOME-CALC-0107 Seismic Qualification of Hot Leg RTD, Hot Leg PDT, and Hot Leg Sampling MNSA Hardware dated 3/11/99.

S-PENG-CALC-008 Nozzle Loads for which SONG Bottom Mounted PZR MNSA was Qualified dated 2/3/99.

C-PENG-CALC-021 Analysis of Waterford 3 Hot Leg Seismic Response Spectra and Accelerations dated 3/10/99.

In addition to the above documents, several other design reports / testing performed for San Onofre, Units 2 and 3 (SONGS) are used as input for the

qualification:

TR-PENG-033 Test Report- Seismic Qualification of the San Onofre,

Units 2 and 3 MNSA Clamps for Pressurizer Instrument

! Nozzles and RTD Hot Leg Nozzles dated 713/97.

S2-NOME-CALC-0085 Seismic Qualification of the Steam Generator PDT, Hot Leg PDT, and Hot Leg Sampling MNSA Hardware dated 2/10/98.

TR-PENG-042 Test Report for MNSA Hydrostatic and Thermal Cycle Test dated 7/3/97.

l 4

't J

J e . o Reviewed By: John P. Burke ghk Date: 3 2 S. Q q U Page 7

._ .