Text

Proposed Inservice Inspection Alternative HNP-ISI-ALT-HDPE-01, Version 2.0, Final Stress Analysis for the Design of the Hdpe System
	ML15246A156
Person / Time
Site:	Hatch
Issue date:	09/03/2015
From:	Pierce C; Southern Co, Southern Nuclear Operating Co
To:	; Document Control Desk, Office of Nuclear Reactor Regulation
References
	NL-15-1628
	Download: ML15246A156 (112)
	v • d • e

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Charles R. Pierce Regulatory Affairs Director SEP 0 3 20\\5 Docket Nos.: 50-366 Southern Nuclear Operating Company, Inc.

40 Inverness Center Parkway Post Office Box 1295 Birmingham, AL 35242 Tel 205.992.7872 Fax 205.992.7601 U.S. Nuclear Regulatory Commission ATTN: Document Control Desk Washington, D. C. 20555-0001 Edwin I. Hatch Nuclear Plant-Unit 2 SOUTHERN<<\\

NUCLEAR A SOUTHERN COMPANY NL-15-1628 Proposed lnservice Inspection Alternative HNP-ISI-ALT-HDPE-01, Version 2.0 Final Stress Analysis for the Design of the HOPE System Ladies and Gentlemen:

By letter dated September 19, 2014, Southern Nuclear Operating Company (SNC) submitted for Nuclear Regulatory Commission (NRC) approval proposed inservice inspection (lSI) alternative HNP-ISI-AL T-HDPE-01, Version 2.0. In this submittal, SNC stated that the final stress analysis for the HOPE design would be available by the end of August, 2015. The letter also stated that, to allow completing the review, the stress analysis would be provided to the NRC.

The eight enclosures to this letter provide the agreed upon final deliverable.

This letter contains no NRC commitments. If you have any questions, please contact Ken McElroy at (205) 992-7369.

Rec.cty~ed, C. R. Pierce Regulatory Affairs Director CRP/OCV

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U.S. Nuclear Regulatory Commission NL-15-1628 Page 2

Enclosures:

1. Design Scope Description

2. P&ID and Isometric Drawings

3. Minimum Wall Thickness Calculation for Plant Service Water HOPE Piping

4. Summary of Hydraulic Evaluation Calculation of Plant Service Water HOPE Piping

5. Summary of Stress Analysis Calculation for PSW Buried HOPE Piping

6. Summary of Stress Analysis Calculation for New U2 Div II PSW Subgrade Vault Piping

7. Summary of Stress Analysis Calculation for Existing U2 Div II PSW Subgrade Vault Piping

8. 14 inch HOPE to Metallic Flanged Joint Analysis Calculation cc:

Southern Nuclear Operating Company Mr. S. E. Kuczynski, Chairman, President & CEO Mr. D. G. Bost, Executive Vice President & Chief Nuclear Officer Mr. D. R. Vineyard, Vice President-Hatch Mr. M.D. Meier, Vice President-Regulatory Affairs Mr. D. R. Madison, Vice President-Fleet Operations Mr. B. J. Adams, Vice President-Engineering Mr. G. L. Johnson, Regulatory Affairs Manager - Hatch RTYPE: CHA02.004 U.S. Nuclear Regulatory Commission Mr. V. M. McCree, Regional Administrator Mr. L. D. Wert, Regional Administrator (Acting)

Mr. R. E. Martin, NRR Senior Project Manager - Hatch Mr. D. H. Hardage, Senior Resident Inspector-Hatch

Edwin I. Hatch Nuclear Plant - Unit 2 Proposed In service Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0 Final Stress Analysis Design of the HOPE System Design Scope Description

Hatch Nuclear Plant-Unit 2 Plant Service Water Piping Replacement Using High Density Polyethylene (HOPE) Piping Scope Description General The scope of this design change is to replace approximately 1 000 feet of the buried 10 inch carbon steel Unit 2 Division II PSW (Plant Service Water) header with HOPE (High Density Polyethylene) Piping. This DCP replaces the 2P41-HBC piping between valve 2P41-F380B and the Unit 2 Reactor Building. Sections of the existing carbon steel piping not required to be removed to facilitate installation of the HOPE piping will be retired-in-place. The HOPE piping will be approximately 1,200 feet in length.

HOPE Piping & HOPE to Stainless Steel Transition

The IPS 14 DR7 HOPE Piping internal diameter is sized as close as practically possible to the existing carbon steel piping to provide equivalent or lower hydraulic resistance.

Metallic reducing slip-on flanges will be used at the transition from the NPS 10 stainless steel piping to the IPS 14 HOPE piping to maximize the HOPE flange adapter seating area and to maximize hydraulic performance.

Existing Unit 2 Service Water Valve Pit/Sub-Grade Vault 2B Existing penetration 2P41-SW-A826B is enlarged to facilitate the larger diameter of the IPS 14 DR7 HOPE Piping.

The HOPE piping is transitioned to stainless steel inside the sub-grade vault using a transition flange. The stainless steel (Pipe Class HAC) portion of piping upstream of the transition flange contains a flanged removable spool piece to allow for inspection of the HOPE piping without having to unbolt the transition flange. A spectacle blind flange and a drain valve are added to facilitate periodic pressure testing of the HOPE piping.

Valve 2P41-F380B and associated operator 2P41-F315 are relocated in the existing Unit 2 Division II vault closer to the 30 inch supply line to facilitate room for the HOPE to stainless steel transition. New stainless steel piping is added to replace the carbon steel piping from the Valve 2P41-F380B to the 30" supply line. This connection will be a flanged connection. An additional support will be added to the metallic piping in the vault to minimize the load on the HOPE.

Underground Piping The HOPE piping (Pipe Class LCF) exits the existing sub-grade vault and is routed underground to the a new Unit 2 Division II sub-grade vault near the Unit 2 Reactor Building. The piping route follows much of the Unit 1 PSW Division II piping route.

Page 1 of 2

Hatch Nuclear Plant - Unit 2 Plant Service Water Piping Replacement Using High Density Polyethylene (HOPE) Piping New Unit 2. Division II Service Water Sub-Grade Vault A new sub-grade vault is added in the yard area near the Unit 2 Reactor Building to allow for transition from HOPE to stainless steel (Pipe Class HAC). The stainless steel section of piping contains a tee which allows access to the interior of the piping without unbolting the transition flange or having to remove the tee. A vent valve is added on the piping tee and a spectacle blind flange is added downstream of the piping tee to facilitate periodic pressure testing of the HOPE.

The carbon steel piping (Pipe Class HBC) from the new sub-grade vault to the first flanged connections inside the Reactor Building is replaced with stainless steel (HAC).

A new stainless steel to carbon steel flange is added on the 2P41-HBC-4" piping within the Reactor Building.

Page 2 of 2

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Edwin I. Hatch Nuclear Plant-Unit 2 Proposed lnservice Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0 Final Stress Analysis Design of the HOPE System Minimum Wall Thickness Calculation for Plant Service Water HOPE Piping

SOUTHERN A COMPANY Plant:

Hatch Nuclear Plant

Title:

I Unit:

01 IEI2 Southern Nuclear Design Calculation Calculation Number:

SMNH-14-013 01&2 I Discipline:

Mech

Subject:

Minimum Wall Thickness for Plant Service Water HOPE Piping Pipe Wall Thickness Calculation Purpose I Objective:

The purpose of this calculation is to determine the minimum required wall thickness for the straight sections of buried High Density Polyethylene (HOPE) piping that is to replace the currently existing buried carbon steel piping as part of the Hatch Nuclear Plant Unit 2 Division II Plant Service Water (PSW) System under DCP SNC591628. This calculation does not apply to fittings and components.

System or Equipment Tag Numbers:

2P41 Division II Piping Contents Topic Page Attachments

of (Computer Printouts, Technrcal Papers, Pages Sketches, Correspondence)

Purpose of Calculation 1

Summary of Conclusions 1

DesiQn Inputs 1

Acceptance Criteria 2

Methodology 2

Assumptions 3

References 3

Body of Calculation 3

Total #of Pages including

~

cover sheet & Attachments :

Nuclear Quality level I IEJSafety-Related 0 Safety Significant 0 Non-Safety -Significant Version Record Version Originator Reviewer I Approval1 Approval2 No.

Description Printed Namo Ptlnltdllomo Printed Noma Printed Namo lnllloll Doto lnltloi/Dote lnati::aiiOouo lnlt~l/ Dale eo,., IZ11.J~IIl'~"cfr Cory

(!_,~, *..,

1 1

~Jared Dobbs Roy S Rosenfel Richardson Issued

'OIUJ'J1Lb;(:t!J~ CM.d~~.~~~ (.(,'() \\1i.r.:\\.:;c:t.~,..) 01'9tnh5

-':i,.n, I

.,,,,-,.s Notes: None NMP-ES-039-F01 NMP-ES-039-001

Plant:

Calculation Number:

Sheet:

Hatch Nuclear Plant SMNH-14-013 1

1.0 Purpose of Calculation:

The purpose of this calculation is to determine the minimum required wall thickness for the straight sections of buried High Density Polyethylene (HOPE) piping that is to replace the currently existing buried carbon steel piping as part of the Hatch Nuclear Plant Unit 2, Division II, Plant Service Water (PSW) System under DCP SNC591628 [Ref. 7.2.1]. This calculation does not apply to fittings and components.

2.0 Summary of

Conclusions:

The required minimum calculated wall thickness value provided in Table 2.0-1 below and Section 8.2 of this calculation apply to straight pipe only. It does not include erosion or other degradation allowances, but it includes a mechanical/installation allowance of 0.040". Nominal wall thickness values must account for manufacturing tolerances such that the as-manufactured minimum wall thickness values for the piping are greater than or equal to the required values.

Table 2.0 Minimum Pipe Wall Thickness IPS Nominal Pipe Size Min. Wall Thickness (in) 141nch 1.949" 3.0 Design Inputs:

3.1 Unit 2, Division II, Plant Service Water (PSW) Header HOPE Pipe Internal design pressure (Po) = 180 psig [Ref. 7.1.1]

Design temperature (To)= 123°F[Ref. 7.1.1]

Outside diameter of pipe (D)= 14.00" [Ref. 7.2.1]

Piping Material: PE 4710 with material properties of cell classification 445574C [Ref. 7.1.1]

3.2 Design Code The design is in accordance with SNC Proposed lnservice Inspection Alternative HNP-ISI-ALT-HDPE-01, Version 2.0, Enclosure 2,

"Proposed Alternative Technical Requirements to ASME Section XI Requirements for Replacement of Class 3 Buried Piping in Accordance with 10CFR50.55a(a)(3)(i).", [Ref. 7.1.1] and Design Specification HM-S-14-001

[Ref 7.2.2].

3.3 Erosion and Degradation Allowances No erosion or other degradation allowances (C) are considered in the minimum wall thickness calculation (See Assumption 6.1) [Ref. 7.2.3, p. 7].

NMP-ES-039-F02 NMP-ES-039-001

Plant:

Calculation Number:

Sheet:

Hatch Nuclear Plant SMNH-14-013 2

3.4 Mechanical/Installation Allowance The mechanical/installation allowance (C) for p1p1ng surface damage during installation is considered in the minimum wall thickness calculation to be the maximum allowable indentation of 0.040" [Ref. 7.1.1, Subsubarticle 4130]. Using this value in the minimum wall thickness calculation assures that the pipe thickness is adequate even with the maximum allowable indentation of 0.040".

3.5 Service Life The replacement buried piping shall be designed for a service life of 50 years under normal system operating conditions as stated in Reference 7.2.2, Section 5.1.

3.6 Allowable Stress The long-term allowable stress value at the Design Temperature (T 0 ) for HOPE piping is interpolated from Table 3131-1 of Reference 7.1.1 and is listed below in Table 3.6-1.

Table 3.6-1 -Allowable Stress Service Tern Load Duration Allowable Stress 570 4.0 Acceptance Criteria Not Applicable

5.0 Methodology

The required minimum wall thickness is determined for the piping identified in Section 1.0 of this calculation. The required minimum wall thickness (tdes*gn} is calculated in accordance with Paragraph 3131.1 of Reference 7.1.1 as follows:

where:

P0D tdesign -

+ C 2S+P0

'----v----'

tmin tdesign

= minimum required wall thickness (in) tmin

= minimum wall thickness for pressure (in)

C

= the sum of mechanical/installation allowance, erosion allowance, and other degradations allowance (in)

Po

= Piping system internal Design Pressure (gage) at the corresponding Design Temperature T 0. This pressure does not include the consideration of pressure spikes due to transients (psig)

D

= pipe outside diameter at the pipe section where the evaluation is conducted (in)

S

= allowable stress (psi)

NMP-ES-039-F02 NMP-ES-039-001

Plant:

Calculation Number:

Sheet:

Hatch Nuclear Plant SMNH-14-013 3

6.0 Assumptions

6.1 Assumption

No erosion or other degradation allowances (C) are considered in the minimum wall thickness calculation Justification:

From Reference 7.2.3, page 7, "PE pipe will not rust, rot, pit, corrode, tuberculate or support biological growth. It has superb chemical resistance and is the material of choice for many harsh chemical environments." Therefore, no erosion or other degradation allowances are required.

This is consistent with Section 5.2 of Reference 7.2.2.

7.0

References:

7.1 Codes 7.1.1 SNC Proposed lnservice Inspection Alternative HNP-ISI-ALT-HDPE-01, Version 2.0,, "Proposed Alternative Technical Requirements to ASME Section XI Requirements for Replacement of Class 3 Buried Piping in Accordance with 1 OCFR50.55a(a)(3)(i)."

7.2 Miscellaneous 7.2.1 Hatch Nuclear Plant Design Change Package SNC591628, "Unit 2 Plant Service Water Piping Replacement".

7.2.2 Hatch Nuclear Plant Design Specification HM-S-14-001, Ver.1, "Design Specification for High Density Polyethylene (HOPE) Piping for the Plant Service Water System "

7.2.3 The Plastics Pipe Institute, Inc. "Handbook of Polyethylene Pipe", Second Edition 8.0 Body of Calculation:

8.1 14" IPS HOPE Pipe Size-Minimum Calculated Wall Thicknesses Under an internal design pressure of 180 psig (Design Input 3.1 ), design temperature of 123°F (Design Input 3.1 ), the sum of mechanical/installation allowance, erosion allowance, and other degradations allowance of 0.040" (Design Input 3.3 and 3.4), allowable stress of 570 psi (Table 3.6-1) the PE 4710, 14.00" diameter, Unit 2, Division II, Plant Service Water (PSW) buried HOPE piping minimum required wall thickness for straight sections is:

tdesignl4 NMP-ES-039-F02 (180 psig X14.00") + 0.04011 = 1.94911 2(570 psi)+ 180 psig NMP-ES-039-001

Plant:

Calculation Number:

Sheet:

Hatch Nuclear Plant SMNH-14-013 4

8.2 Results The required minimum wall thickness for the straight sections of Hatch Nuclear Plant Unit 2, Division II, Plant Service Water (PSW) buried High Density Polyethylene (HOPE) piping is shown below in Table 8.2-1 :

T bl 8 2 1 M'.

a e 1m mum p*

W II Th' k Jpe a

1c ness IPS Nominal Pipe Size Min. Wall Thickness Jin) 141nch 1.949" NMP-ES-039-F02 NMP-ES-039-001

Edwin I. Hatch Nuclear Plant-Unit 2 Proposed lnservice Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0 Final Stress Analysis Design of the HOPE System Summary of Hydraulic Evaluation Calculation of Plant Service Water HOPE Piping

Summary Report On Hydraulic Performance Calculation of Plant Service Water HOPE Piping Hatch Nuclear Plant-Unit 2 1.0 Purpose of Summary Report:

This report summarizes the hydraulic calculation that has been prepared to support SNC's lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0. This lSI Alternative is needed to support the planned replacement of buried steel piping in Hatch Nuclear Plant Unit 2, Plant Service Water (PSW) system with High Density Polyethylene (HOPE) piping.

Computations from the calculation are not included in the summary. Calculation results and conclusions are shown in Section 2.0.

The calculation evaluates the hydraulic performance of the existing carbon steel Division II P!ant Service Water (PSW) piping compared to the HOPE (High Density Polyethylene) material piping to be installed by DCP SNC591628. The hydraulic resistance of the replacement piping must be less than or equivalent to the currently installed piping to ensure adequate cooling water flow is maintained. The piping to be replaced is from the Unit 2 Service Water Valve pit 2B to the Unit 2 Reactor Building.

2.0 Results and

Conclusions:

Table 2-1 below shows the summary of pressure losses for the existing carbon steel piping and for the HOPE piping to be installed. The pressure losses for these configurations are calculated for both the normal flow rate (8200 gal/min) and the minimum flow rate (4428 gal/min). Nominal HOPE piping sizes of 12 inches and 14 inches are calculated.

Based on results from Table 2-1, replacement HOPE with a nominal size of 14 inches (I.D of

9. 760) ensures the pressure loss through the piping is less than or equivalent to the currently installed piping for both the normal and minimum flow conditions.

T bl 2 1 S a e -

ummary o fP ressure L f

E. f p*.

osses or XIS lf!g tpmg an dW" h HOPEI It II d nsta e Configuration Pressure Loss@ 8200 gal/min Pressure Loss @ 4428 gal/min Existing CS Pipe 234.2 psi 74.9 psi 14" HOPE 198.8 psi 63.6 psi 12" HOPE 297.8 psi 95.2 psi 3.0 Design Inputs:

3.1 PSW Shutdown Cooling Minimum Flow Requirement, 4428 gal/min (Ref. 7.1) 3.2 Measured Normal Operation PSW Division II Flow Rate-8200 gal/min (Ref. 7.2, Attachment A, Page 5, Step 7.3.1.9) 3.3 The existing PSW piping is 10", schedule 40, carbon steel (Ref. 7.5) with an internal diameter of 10.02" (0.835 ft) (Ref. 7.6, Page B-17). The replacement stainless steel piping is sized to match this piping.

3.4 The existing section of PSW piping is 1,008' in length, contains nine (9) 45° elbows, and three (3) 90° elbows (Ref. 7.3, 7.7, 7.8, 7.9) (See Attachment A)

Pagel

3.5 The length of the HOPE replacement piping is 1, 186' 1" and contains nine (g) goo elbows, six (6) 45° elbows, twenty two (22) 22.5° elbow, one sudden enlargement, and one sudden contraction (Ref. 7.12) (Attachment B).

3.6 The internal diameter of the 12 inch HOPE piping is 8.88g inches (.7408 ft). This value is the average 10 for the 12 DR7 PE471 0 piping (Ref. 7.1 0, Pg. 18) 3.7 The internal diameter of the 14 inch HOPE piping is 9.760 inches (.8133 ft). This value is the average IDforthe 14 DR7 PE4710 piping (Ref. 7.10, Pg. 18) 3.8 The HOPE 90° elbows are considered to have 3 or more miters with a Representative Fittings Factor, K' of 24 (Ref. 7.11, Chapter 6 Table 2-2) with outside diameter reinforcement such that the interior diameter of the fittings is equal the interior diameter of the piping (Ref. 7.12).

3.g The HOPE 45° elbows are considered to have 2 or more miters with a Representative Fittings Factor, K' of 15 (Ref. 7.11, Chapter 6 Table 2-2) with outside diameter reinforcement such that the interior diameter of the fittings is equal the interior diameter of the piping (Ref. 7.12).

3.10 The HOPE 22.5° elbows are considered to be 2-segment elbows with a Representative Fittings Factor, K' of 8 (Ref. 7.11, Table 2-2) (Assumption 6.7) with outside diameter reinforcement such that the interior diameter of the fittings is equal the interior diameter of the piping (Ref. 7.12).

3.11 Elevation of the metallic to HOPE transition immediately downstream of P41-F380B is 120'7" (Ref. 7.12). The elevation of the metallic piping being replaced in this location is 120'6" (Ref. 7.3).

3.12 Elevation of Division II, PSW reactor building penetration is 125' 5" (Ref. 7.4, Penetration #10) 3.13 The section of stainless steel piping in the existing sub-grade vault downstream of P41-F380B contains one (1) spectacle flange. The section of stainless steel piping from the HOPE to stainless steel transition to the reactor building penetration contains two (2) goo elbows, one (1) full size straight piping tee, one (1) spectacle flange, and 51' 2" of piping (Ref. 7.12)

(Assumption 6.1 0).

3.14 The weight density of water is 62.371 lbs/ft3 (Ref. 7.6,Pg, A-6) (Assumption 6.1) 3.15 Acceleration of gravity, g is 32.2 ft/s2*

3.16 The dynamic viscosity of water at 60°F is 1.1 Centipoise (Ref. 7.6, Pg. A-3) (Assumption 6.1) 3.17 The Hazen-Williams Friction Faction, C is 150 for HOPE (Ref. 7.11, Pg. 175) 3.18 The HOPE absolute surface roughness design value is for that of "smooth pipe" (Ref 7.11, Pg.

172) (Assumption 6.6).

Page2

4.0 Acceptance Criteria The replacement HOPE piping shall have a pressure loss less than or equivalent to the existing carbon steel piping for both the normal and minimum required flow rates.

5.0 Methodology 5.1 Pressure Loss through Carbon Steel Piping The section of piping evaluated is the PSW Division II piping immediately downstream of P41-F380B (Ref. 7.13) to the reactor building piping penetration #1 0 (Ref. 7.4). The pressure loss of the flow through the existing carbon steel piping and fittings is determined from Bernoulli's theorem as depicted in (Ref. 7.6, Equation 1-3).

Z + 144P1 + v~ = Z + 144P2 + v~ + h 1

P1 2g 2

Pz 2g L

Where:

Z1-Elevation of Point 1-ft (Design Input 3.11)

Z2-Elevation of Point 2-ft (Design Input 3.12)

P1-Pressure at Point 1-psi P2-Pressure at Point 2-psi v1-flow velocity at Point 1-ft/s v2-flow velocity at Point 2-ft/s p1-Density of water at Point 1-lbs/ft3 (Design Input 3.14) p2-Density of water at Point 2-lbs/ft3(Design Input 3.14) g-Acceleration of Gravity-ft/s2 (Design Input 3.15) hL-Loss of static pressure head-ft Eq. 5-1 Because the density of water is unchanged from Point 1 to Point 2, p1 = p2 = p. Therefore Eq. 5-2 Because the pipe flow area and flow rates at Point 1 and Point 2 are equal, the pipe flow velocities at Point 1 and Point 2 are equal, v1 = v2. Therefore, Eq. 5-3

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Based on the equation above, the pressure loss through the section of piping,P1 - P2 is determined by knowing the density of the water, p and the static pressure head loss due to fluid flow, hL. The pressure loss due to the elevation change between Point 1 and Point 2, Z2 - Z1 is not considered in this calculation because both the existing piping and the piping configuration with HOPE installed are assumed to have equivalent Point 1 and Point 2 elevations (Assumption 6.11 ). The density of water and the elevation of Point 1 and Point 2 are input values. The static pressure head Joss due to fluid flow is the summation of pressure Joss due to the straight piping and the pressure loss through fittings.

Therefore, P1 -

P2 = 1~4 ( hL(pipe and fittings))

Eq. 5-4 The pressure loss through the piping and fittings is determined by the Hazen-Williams equation (Ref. 7.14, Page 27) below

(

100)1.85 Ql.BS hL(pipe and fittings) = 0.002083 (L) C (d4.B655)

Where:

Eq. 5-5 L-Length of pipe including equivalent length of pipe, Leq due to the loss through fittings-ft.

(Design Input 3.4)

C-Hazen-Williams Friction Factor-100 (Assumption 6.4)

Q-Flow of water, gpm d-Inside diameter of the pipe-inches (Design Input 3.3)

The equivalent length of pipe, Leq due to the pressure Joss through the goo and 45° elbows is obtained from (Ref. 7.14, Page 4g).

For 1 0" pipe, the friction Joss for a long radius goo elbow is equivalent to 11.0 feet of pipe.

For 1 0" pipe, the friction loss for a long radius 45° elbow is equivalent to 7.1 feet of pipe.

Page4

5.2 Pressure Loss with HOPE Piping The section of piping evaluated is the PSW Division II p1p1ng immediately downstream of P41-F3808 (Ref. 13) to the reactor building piping penetration #10 (Ref. 7.4). The metallic to HOPE piping transition occurs in the sub-grade vault immediately downstream of P41-F3808 (Assumption 6.12). The HOPE piping is transitioned back to new stainless steel piping in a new sub-grade vault just outside of the reactor building penetration. The carbon steel piping upstream of P41-F3808 and within the reactor building replaced with stainless steel will result in an increase in hydraulic efficiency and is conservatively not accounted for in this calculation.

The pressure loss with the replacement HOPE piping is the summation of the pressure loss through the HOPE piping using the characteristics of HOPE piping determined from (Ref. 7.11 ),

the pressure losses due to the sudden contraction in the transition from metallic to HOPE, the pressure loss due to sudden enlargement in the transition from HOPE to stainless steel (Design Input 3.5), and the pressure loss through the stainless steel piping and fittings. From Eq. 5.4, the equation for the pressure loss is shown below P1 -

Pz = 1: 4 ( hL(pipe.[ittings,contr,enlarg)) + SS!Jp Eq. 5-6 Where:

p-Density of water-lbstfe (Design Input 3.14) hL(pipe.fittings,contr,enlarg)- Pressure loss through the HOPE piping, fittings, sudden contraction, and sudden enlargement, in feet of head SStJr Pressure loss through the section of stainless steel piping and fittings, in psi The pressure loss through the HOPE piping, fittings, sudden contraction, and sudden enlargement is determined by the Hazen-Williams equation (Ref. 7.14, Page 27) below

(

100)1.85 Ql.SS hL(pipe,fittings,contr,enlarg) = 0.002083 (L) c (d4.a6ss)

Where:

Eq. 5-7 L-Length of pipe including equivalent length of pipe, Leq due to the loss through fittings, sudden contraction, and sudden enlargement-ft. (Design Input 3.5)

C-Hazen-Williams Friction Factor-150 (Design Input 3.17)

Q-Flow of water, gpm d-Inside diameter of the pipe-inches

PageS

The equivalent pipe length, Leq due to the pipe fittings is determined by using the equation below (Ref. 7.11, Eq 2-9)

Leq = K'D Where:

Leq = Equivalent pipe length due to the fittings, feet K'- Representative fittings factor from (Design Input 3.8, 3.9, 3.1 0)

D-Internal pipe diameter, feet Eq. 5-8 The equivalent pipe length, Leq due to the sudden contraction from stainless steel to HOPE is determined below (Ref. 7.6, Eq 2-4, 2-10)

Where:

f-Friction factor of piping (Based on Reynolds' Number) (Ref. 7.6, Pg. A-24)

D-HOPE Internal Pipe Diameter-ft.

d1-HOPE pipe diameter-in d2-stainless steel pipe diameter-in Eq. 5-9 The equivalent pipe length, Leq due to the sudden enlargement from HOPE to stainless steel is determined below (Ref. 7.6, Eq 2-4,2-9)

((1-~)

2

)co)

L

_ K D _ __,___d-=-2 --'--

eq-f --

f Where:

f-Friction factor of piping (Based on Reynolds' Number) (Ref. 7.6, Pg. A-24)

D-HOPE Internal Pipe Diameter-ft.

d1-HOPE pipe diameter-in dr steel pipe diameter-in Eq. 5-10 The pressure loss through the section of stainless steel piping from the HOPE to stainless steel transition to the reactor building penetration, ssAP is determined by SSAP = P1-Pz = 1: 4 ( hL(pipeandfittings))

Eq. 5-11

Page6

The pressure loss through the stainless steel piping and fittings is determined by the Hazen-Williams equation (Ref. 7.14, Page 27) below

(

100)1.85 Ql.85 hL(pipe and fittings) = 0.002083(L) C (d4.8655)

Where:

L-Length of stainless steel piping including equivalent length of fittings, Leq - ft.

(Design Input 3.13)

C-Hazen-Williams Friction Factor-100 (Assumption 6.4)

Q-Flow of water, gpm d-Inside diameter of the pipe-inches (Design Input 3.3)

Eq. 5-12 The equivalent length of pipe, Leq due to pressure loss through the full size straight tee is obtained from (Ref. 7.14, Page 4g). For 10" pipe, the friction loss through the run of a full size straight tee is equivalent to 16.g feet of pipe.

The equivalent length of pipe, Leq due to the pressure loss through the goo elbows is obtained from (Ref. 7.14, Page 4g). For 10" pipe, the friction loss for a long radius goo elbow is equivalent to 11.0 feet of pipe.

Page7

6.0 Assumptions

6.1 Assumption

The density and viscosity of water is evaluated at an assumed temperature of s*ooF.

Justification:

This temperature is the value for which the Hazen-Williams equation is based (Ref. 7.14, Page 27). Using this temperature through the calculation ensures consistency when comparing the existing piping to the new piping.

6.2 Assumption

The flow area of the existing carbon steel piping is assumed to have the same flow area as new piping.

Justification:

Due to the age of this piping and its raw water service, biological growth on the interior surface of this piping is likely to be present causing a reduction in flow area. The extent of flow area reduction due to biological grown is unknown. This reduction in flow area would cause an increased pressure loss in the piping. Therefore, this assumption is conservative for sizing the HDPE piping flow area.

6.3 Assumption

The fusion beads on the interior of the HDPE piping are assumed to have a negligible resistance on the system.

Justification:

The fusion beads have a negligible effect on fluid flow From (Ref. 7.11, Page 30) Also, The Hazen-Williams Friction Factor, C for HDPE piping was determined in a hydraulics laboratory with the fusion beads present (Ref. 7.11, Page 175). Therefore, any flow impact by the fusion beads has been incorporated through the Hazen-Williams Friction Factor.

6.4 Assumption

The Hazen-Williams Friction Factor, C is assumed to be 100 for the existing carbon steel piping and for the new stainless steel piping.

Justification:

Corrosion to the interior surface of the existing piping is known to be present and significant.

The average C value for clean, new pipe is 130 and the C value for corroded piping is 80 (Ref.

7.14 Pg.3-8). Using a C value of 100 for the existing carbon steel piping is not as restrictive as the value provided for corroded pipe is conservative and reasonable given the corrosion issues known to be present in the piping. Using a C value of 100 is conservative as the margin gained from using clean, new pipe is not credited.

6.5 Assumption

It is assumed the all of the metallic piping up to HDPE transition flange in the existing sub-grade vault is the existing carbon steel.

Justification:

A portion of this carbon steel piping will be replaced with stainless steel when the HDPE is installed. The pressure loss though the section of piping is greater by considering this section of piping to be the existing carbon steel and is conservative.

6.6 Assumption

No flow degradation over time is considered for the HDPE piping.

Justification:

From (Ref 7.6), page 7, "PE pipe will not rust, rot, pit, corrode, tuberculate or support biological growth." page 10, "Without corrosion, tuberculation, or biological growth PE pipe maintains its smooth interior wall and its flow capabilities indefinitely"

6.7 Assumption

HDPE 22.5° elbows are considered to be 2-segment elbows with a Representative Fittings Factor, K' of a 2-segment 30° elbow. (Design Input 3.1 0)

Justification:

A 30° elbow will provide more resistance than a 22.5° elbow. Therefore, this assumption is conservative.

PageS

6.8 Assumption

The PSW shutdown cooling flow of 4428 gal/min is all considered to be supplied through the Division II piping being replaced.

Justification:

Shutdown cooling of the plant requires only one PSW pump, delivering 4428 gal/min. The Division I and Division II headers are completely redundant to each other. Therefore, assuming the flow to be only through the Division II header is justified and conservative. Flow diverted for diesel generator cooling is not considered as it has, at most a negligible impact on the results of this calculation.

6.9 Assumption

The pressure losses through the two (2) stainless steel spectacle flanges (Ref.

7.12) are neglected.

Justification:

The spectacle flanges are full bore resulting in a negligible impact to the flow.

6.10 Assumption: An additional 5'1" is added to the length of stainless steel piping.

Justification:

The increased length of stainless steel piping added by this DCP increases the overall hydraulic resistance of the new piping and is conservative for this calculation.

6.11 Assumption: The elevation of the HOPE piping installed in the existing sub-grade vault by DCP SNC591628 is assumed to be at the same elevation as the currently installed carbon steel piping.

Justification:

The difference in the elevation between the HOPE piping to be installed and the existing piping is one inch. This difference is negligible. Additionally, the overall change in elevation from the 30 inch Division II header and the reactor building penetration is unchanged.

6.12 Assumption: The small stainless steel piping downstream of valve P41-F3808 before the transition to HOPE is negligible and not considered in this calculation.

Justification:

This small section of straight piping (approximately 5 feet) has a negligible impact in comparison to the resistance of the approximately 1200 feet overall length of the piping and is enveloped by the conservatism in the calculation.

Page9

7.0

References:

7.1 Hatch Calculation SMNH-08-011, Ver. 3.0- Minimum River Water Level Required to Meeting NPSH and Minimum Submergence Requirements for the Safety Related Pumps in the River Intake Structure During Safety Related Operation of the Pumps 7.2 Hatch Calculation SMNH-03-004, Ver. 5.0- Generate Unit 2 Plant Service Water (PSW)

PROTO-FLO Database for Latest Test Data 7.3 Hatch Drawing H21109, Ver. 9.0- Yard Piping Sheet 1 7.4 Hatch Drawing H26302, Ver. 12.0- Reactor Building Penetrations in Walls & Floors Below EL.

130'0" 7.5 Drawing A21000 Sh. 332, Rev. 5-Piping Class HBC 7.6 Crane Technical Paper. No. 410, 1988 edition 7.7 Hatch Drawing H11146, Ver. 35.0- Unit 1 Piping-Service Water Pump Structure To Building 7.8 Hatch Drawing H21110, Ver. 17.0- Yard Piping Sheet 2 7.9 Hatch Drawing H21111, Ver. 8.0- Yard Piping Sheet 3 7.10 ISCO Product Catalog, Version4.1, 2013, www.isco-pipe.com 7.11 Plastics Pipe Institute (PPI) Handbook of PE Pipe, Second Edition 7.12 DCP Worksheets SNC591628M201, Ver. 1.0 & SNC591628M202, Ver. 1.0 -Unit 2 PSW Division II Underground Piping Isometric Worksheets (Attachment B) 7.13 Hatch Drawing H21033, Ver. 59.0- Turbine Building Service Water System P&ID Sheet 1 7.14 Cameron Hydraulic Data, Fourteenth Edition 7.15 Hatch Drawing H26252, Ver. 10.0- Plant Service Water System-Reactor Building Piping Below El. 130'0" East

Page10 --------------------------------

Edwin I. Hatch Nuclear Plant-Unit 2 Proposed In service Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0 Final Stress Analysis Design of the HOPE System Summary of Stress Analysis Calculation for PSW Buried HOPE Piping

Summary Report on Stress Analysis for PSW Buried HOPE Piping Hatch Nuclear Plant-Unit 2 1.0 Purpose of Summary Report This report summarizes the stress analysis calculation, SMSH-14-011, that has been prepared to support SNC's lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0 (ATR). This lSI Alternative is needed to support the planned replacement of buried steel piping in Hatch Nuclear Plant Unit 2, Plant Service Water (PSW) system with High Density Polyethylene (HOPE) piping.

The piping to be replaced is the supply piping from the Unit 2 Service Water Valve Pit 2B to a new subgrade vault located outside of the Unit 2 Reactor Building. The stress analysis calculation evaluates the replacement piping to the design requirements in the ATR.

All computations from the calculation are not included in the summary. Calculation results and conclusions are shown in Section 2.0.

2.0 Results and Conclusions The replacement HOPE piping meets all of the acceptance criteria outlined in the lnservice Inspection (lSI) Alternative Request HNP-ISI-ALT-HDPE-01, Version 2.0.

Detailed results are shown in this section. All margin factor values are greater than 1.0 and are therefore acceptable. The controlling margin factor that was dependent on pipe loading (other than pressure) was 1.31 which was for ring deflection due to soil and surcharge loads (ISFSI cask and crawler surcharge loading controls). The stress analysis calculation also evaluated pipe floatation and concluded that the unanchored pipe will not float.

The piping has been analyzed to the following conditions per Section 1100 of Design Input 3.1. A design temperature value of 125°F was conservatively used which bounds the actual design temperature of 123°F.

Condition Temperature, °F Pressure, psig Normal Operating 95 140 Maximum Operating 97 190 Design 123 180 Allowable Service Level Spikes Due to Transient Pressures There are no credible fluid transients identified. Therefore, no fluid transient loads are considered.

Pressure Design of Joints and Fittings (Based on GSRs)

Component Pressure Rating Design Pressure Margin Factor (psi)

(psi)

Flange Adapters 187 180 1.04 Miter Elbows 207 180 1.15 The margin factors are greater than 1.0 for all components and are therefore acceptable.

Pressure Design of Miter Elbows (Based on Design Equations)

Pressure Design Pressure Margin Factor Rating (psi)

(psi) 218 180 1.21 The margin factor is greater than 1.0 and is therefore acceptable.

Paget

Ring Deflection due to Soil and Surcharge Loads Ring Deflection Max. Ring Margin Factor

(%)

Deflection (%)

2.13 2.8 1.31 The margin factor is greater than 1.0 at all locations for the design truck and the ISFSI crawler loads and is therefore acceptable.

Compression of Sidewalls Due to Soil and Surcharge Loads Circumferential Allowable Compressive Stress Compression Margin Factor (psi)

Stress (psi) 145 630 4.34 The margin factor is greater than 1.0 at all locations for the design truck and the ISFSI crawler loads and is therefore acceptable.

Buckling Due to External Pressure Case Total Pressure Pressure Limit Margin Factor on Pipe (psi) for Buckling (psi)

>4ft Cover 45.12 169.43 3.76

<4ft Cover 17.41 53.60 3.08 The margin factors are greater than 1.0 at all locations for the design truck and the ISFSI crawler loads and are therefore acceptable.

Effects of Negative Internal Pressure Per Section 3.1 0.1 of Enclosure 8 of design input 3.1, there are no negative internal pipe pressures anticipated.

Flotation Since the downward force acting on the pipe is greater than the upward force, the unanchored pipe will not float.

Longitudinal Stresses The seismic-induced stresses due to seismic wave passage were determined by following the methodology outlined in Section 4.5 of Ref. 7.3. The methodology involves calculating the controlling strain induced in the pipe due to seismic wave passage. The seismic wave induced strain is limited by breakaway between the pipe and soil. The controlling strain is converted into an equivalent differential temperature which is input into the SAP2000 Finite Element Models (FEM) for the seismic load cases. The axial soil strains caused by soil curvature were calculated as 1.357x1 011-8 and 2.544x1 011-8 for OBE and DBE seismic responses, respectively. The axial soil strains caused by wave passage were calculated as 1.6x1 011-4 and 3x1 011-4 for OBE and DBE seismic responses, respectively. Therefore, the controlling axial soil strains due to wave passage are input into the FEM as equivalent differential temperatures.

The relative displacements between the valve vaults and the surrounding soil are considered negligible since the valve vaults are relatively small and the surrounding soil has a high compaction level (minimum 95% of the maximum dry density compaction criterion per Design Input 3.7, Section 2.5.1.2.8). The HOPE piping is connected to the steel piping inside the new subgrade vault located Page2

outside of the Reactor Building. The steel piping continues into the Reactor Building. Since a portion of the Reactor Building seismic anchor movements could be distributed to the portion of the HOPE piping, the full Reactor Building seismic anchor movements were conservatively applied at the model termination point inside the new valve vault. The Reactor Building DBE seismic anchor movements were taken from Table 1 of Design Input 3.11 at Elevation 130' as 0.078-in and 0.064-in for the horizontal and vertical directions respectively. One-half of these values were considered for the OBE load case which is consistent with Design Input 3.12.

The loads resulting from applying the seismic anchor movements in each direction were combined by square root sum of the squares (SRSS). The resulting seismic anchor movement loads were then combined with the other seismic loads by SRSS (per Design Input 3.1, Enclosure 2, Section 3410).

Loads due to seismic soil movement are considered zero since the soil is not susceptible to liquefaction (per Design Input 3.7, Sections 2A.4.8 & 2A.5.2).

The longitudinal stresses tabulated on the following page consider the worst case axial loads in combination with the worst case bending moments. This is conservative since the worst case axial load is likely to not occur at the same location as the worst case bending moment. The table below summarizes the maximum loading that is used in developing the bounding stresses. These loads include seismic wave passage and seismic anchor movements which were extracted from the analysis computer program output.

Max. Axial Load, OBE 20591bf

-c

~Q) Max. Moment, OBE 10252 in-lbf c::

J

a.

0 ro *-

Max. Axial Load, DBE 4101 lbf Ill

!::;0..

1/)

(J)

Cl Max. Moment, DBE 20503 in-lbf Q) c::

a.*-

a.'-

Max. Axial Load, OBE 1490 lbf

Jc%

1/)

Cl

~

Max. Moment, OBE 5938 in-lbf c::

0

u;

.c Max. Axial Load, DBE 2795 lbf

J jjJ Max. Moment, DBE 11651 in-lbf Max. Axial Load, OBE 9881bf

-c "5,a>

Max. Moment, OBE 6802 in-lbf c::

J

a.

0 ro *-

Max. Axial Load, DBE 1853 lbf Ill

!::;0..

1/)

(J)

Cl Max. Moment, DBE 13532 in-lbf Q) c::

~ *c::

Max. Axial Load, OBE 9581bf 0 a.

....l(J) 1/)

Cl

~

Max. Moment, OBE 6600 in-lbf c::

_g

u; Max. Axial Load, DBE 1796 lbf

J jjJ Max. Moment, DBE 12375 in-lbf The allowable stress value was conservatively based on a temperature of 125°F instead of the design temperature of 123°F.

Page3

Service Spring Load Pipe Stress Stress Factor Margin Component Stress xAIIowable Level Case (psi)

Factor Stress (psi)

Factor Straight Pipe N/A 315 1

561 1.78 A

Miter Elbow N/A 331 1

561 1.69 Upper Bound 411 1.1 617 1.50 Straight Pipe Lower Bound 380 1.1 617 1.63 B

Miter Elbow Upper Bound 388 1.1 617 1.59 Lower Bound 385 1.1 617 1.61 Upper Bound 490 1.33 746 1.52 Straight Pipe Lower Bound 425 1.33 746 1.76 D

Upper Bound 424 1.33 746 1.76 Miter Elbow.

Lower Bound 415 1.33 746 1.80 The loads included in each Service Level are defined in Section 3.10 of Enclosure 8 of Design Input 3.1. Service Level A includes stresses due to the design pressure. Service Level B includes stresses due to the maximum operating pressure and forces and moments due to the effects of OBE seismic wave passage, OBE seismic soil movement, and OBE seismic anchor movements. Service Level D includes stresses due to the maximum operating pressure and forces and moments due to the effects of DBE seismic wave passage, DBE seismic soil movement, and DBE seismic anchor movements. As previously noted, loads due to seismic soil movement are considered zero since the soil is not susceptible to liquefaction. All margin factors are greater than 1.0 and are therefore acceptable.

Seismic-Induced Stresses See the Longitudinal Stresses Section for discussion on the development of the seismic load cases.

Similar to the longitudinal stresses evaluations, the seismic-induced stresses consider the worst case axial loads in combination with the worst case bending moments. The same forces and moments previously tabulated in the longitudinal stress section for the DBE case were used to calculate the stress values in the table below. Only the DBE case is evaluated as it bounds the OBE case.

Component Spring Load Pipe Stress (psi)

Allowable Stress Margin Factor Case Range (psi)

Straight Pipe Upper Bound 315 2032 6.46 Lower Bound 185 2032 10.99 Miter Elbow Upper Bound 139 2032 14.63 Lower Bound 128 2032 15.83 All margin factors are greater than 1.0 and are therefore acceptable.

Short Duration Longitudinal Applied Mechanical Loads There are no short duration longitudinal applied mechanical loads for this piping.

Design for Combined Thermal Expansion and Contraction The thermal stresses due to thermal contraction are based on a minimum water temperature of 32°F. The thermal stresses due to thermal expansion are based on a temperature of 125°F which bounds the design temperature of 123°F. This is conservative since the normal operating temperature and maximum operating temperature are only 95°F and 97°F respectively. Both cases utilize a ground temperature of 70°F. Additionally, the allowable stress value was based on a temperature of 125°F instead of the design temperature of 123°F which is conservative. The stress Page4

due to thermal contraction is added to the stress due to thermal expansion, which results in a thermal range stress.

Stress Due to Stress Due to Combined Allowable Stress Thermal Thermal Stress (psi)

Range (psi)

Margin Factor Contraction (psi)

Expansion (psi) 439 202 641 2032 3.17 The margin factor is greater than 1.0 and is therefore acceptable.

Alternative Thermal Expansion and Contraction Evaluation The temperature values of 32°F and 125°F discussed in the Design for Combined Thermal Expansion and Contraction Section were modeled in the FEM as two separate load cases. These were input into the FEM as changes in temperature of -38°F (= 32°F - 70°F) and +55°F (= 125°F -

70°F). The resulting loads for the two thermal load cases were then combined by absolute sum to obtain the thermal range response.

Similar to the longitudinal stresses evaluations, the thermal stresses consider the worst case axial loads in combination with the worst case bending moments. The table below summarizes the maximum loading that is used in developing the bounding stresses for the thermal expansion and contraction load cases. The loads in this table are for the thermal range and were extracted from the analysis computer output.

Using Straight Max. Axial Load 371271bf Upper Pipe Max. Moment 19109 in-lbf Bound Max. Axial Load 42341 lbf Springs Elbows Max. Moment 109810 in-lbf Using Straight Max. Axial Load 265141bf Lower Pipe Max. Moment 77739 in-lbf Bound Max. Axial Load 267191bf Springs Elbows Max. Moment 191582 in-lbf These loads were used to calculate the pipe stresses reported in the table below.

Spring Load Allowable Component Pipe Stress (psi)

Stress Range Margin Factor Case (psi)

Upper Bound 588 2032 3.45 Straight Pipe Lower Bound 742 2032 2.74 Miter Elbow Upper Bound 785 2032 2.59 Lower Bound 985 2032 2.06 All margin factors are greater than 1.0 and are therefore acceptable.

Non-Repeating Anchor Movements Calculation SCNH-15-029 has determined that there is no potential for vault settlement but vault heaving has been calculated to be less than 1/8" for the new valve vault. Therefore, a worst case upward displacement of 1/8" was considered at the new valve vault support location.

Similar to the longitudinal stresses evaluations, the non-repeating anchor movement stress evaluations consider the worst case axial loads in combination with the worst case bending moments. The following table summarizes the maximum loading that is used in developing the bounding stresses. The loads in this table were extracted from the analysis computer output.

PageS

Straight Max. Axial Load 0.21bf Pipe Max. Moment 24601.1 in-lbf Max. Axial Load 0.21bf Elbows Max. Moment 9398.7 in-lbf These loads were used to calculate the pipe stresses reported in the table below.

Component Pipe Stress (psi)

Long Term Allowable Margin Stress (psi)

Factor Straight Pipe 123 1122 9.09 Miter Elbow 38 1122 29.82 Both margin factors are greater than 1.0 and are therefore acceptable.

Other Design Considerations As noted in Enclosure 2 of Design Input 3.1, other design considerations will be addressed under SNC design procedures in accordance with the existing design and license basis for HNP.

Page6

3.0 Design Inputs 3.1 Edwin I. Hatch Nuclear Plant-Unit 2, Proposed lnservice Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0.

The acceptance criteria and general methodology was taken from this design input document.

3.2 Drawing SK-101, Version 1.0, "Edwin I. Hatch Plant Unit No.2 Yard Piping-Plant Service Water U2 DIV II Valve Vault to Reactor Building."

The routing of the piping is detailed in this design input document.

3.3 Drawing SK-102, Version 1.0, "Edwin I. Hatch Plant Unit No.2 Yard Piping-Plant Service Water U2 DIV II Valve Vault to Reactor Building."

The routing of the piping is detailed in this design input document.

3.4 Drawing SK-103, Version 1.0, "Edwin I. Hatch Plant Unit No.2 Yard Piping-Plant Service Water U2 DIV II Valve Vault to Reactor Building."

The routing of the piping is detailed in this design input document.

3.5 Drawing SK-104, Version 1.0, "Edwin I. Hatch Plant Unit No.2 Yard Piping-Plant Service Water U2 DIV II Valve Vault to Reactor Building."

The routing of the piping is detailed in this design input document.

3.6 Drawing SK-105, Version 1.0, "Edwin I. Hatch Plant Unit No. 2 Yard Piping-Plant Service Water U2 DIV II Valve Vault to Reactor Building."

The routing of the piping is detailed in this design input document.

3.7 HNP-2-FSAR-2, Rev. 26 (9/08)

Site specific details were taken from this design input document. Specifically, seismic wave velocity values, peak horizontal ground accelerations and site soil conditions.

3.8 ISCO Product Catalog, Version 4.1, 2013.

The HOPE pipe section dimensional properties and weight are taken from this design input document.

3.9 Flowable Fill Website, http://flowablefill.org/performance.html The typical maximum and minimum densities for flowable fill are taken from this design input document.

3.10 DCR00-35, "Addition of RR Pad and New Track, and Evaluation of Crawler Path," Rev. 1.

The crawler track dimensions for the ISFSI cask transporter and the total weight for the cask and crawler are taken from this design input document.

3.11 Calculation BHO-C-S08-V001-0003, Version 1, Edwin I. Hatch Nuclear Plant, Units 1 and 2, "The Stress Analysis of Underground Piping and Electric Ducts."

The DBE seismic anchor movements for the Reactor Building are taken from this design input document.

3.12 Calculation SMSH-12-020, Version 2.0, "Stress Analysis of Unit 1 Div II Buried Service Water Pipe."

The OBE seismic anchor movements for the Reactor Building are taken as half of the value of the DBE anchor movements. This is consistent with this design input document.

3.13 BH2-C-S23-V012-0001, "Final Seismic Analysis Reactor Building and Internals", April15, 1975, Volume 1.

The DBE & OBE seismic anchor movements were compared to the values reported in this design input document and were found to be of negligible difference.

Page7

3.14 BH2-C-S23-V013-0001, "Final Seismic Analysis Reactor Building and Internals", April15, 1975, Volume 2.

The DBE & OBE seismic anchor movements were compared to the values reported in this design input document and were found to be of negligible difference.

3.15 Drawing H-26051, Version 51.0, "Edwin I. Hatch Nuclear Plan Unit No. 2 Reactor Building -

Plant Service Water System P & I. D. Sht 2 of 2."

3.16 Calculation SCNH-15-029, Version 1, "Soil Bearing Evaluation for New Unit 2, Div. II PSW Transition Vault."

This calculation documents the non-repeated anchor movement for the new vault (heave).

3.17 Calculation SMSH-15-001, Version 1, "Stress Analysis for New U2 DIV II PSW Subgrade Vault Piping."

This analysis overlaps with the analysis for the piping discussed in this summary report.

3.18 Calculation SMSH-15-002, Version 2, "Stress Analysis for U2 DIV II PSW Subgrade Vault Piping."

This analysis overlaps with the analysis for the piping discussed in this summary report.

PageS

4.0 Acceptance Criteria*

The stress analysis calculation followed the acceptance criteria outlined in the lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0.

5.0 Methodology The stress analysis calculation followed the methodology outlined in the lnservice Inspection (lSI)

Alternative Request HNP-ISI-ALT-HDPE-01, Version 2.0 (ATR). In addition to the design truck in the ATR, surcharge loads for an ISFSI cask transporter were also considered.

In order to obtain the pipe axial forces and resultant moments for Sections 3223.1, 3311.4, 3312, &

341 0 of the ATR, several SAP2000 Finite Element Models (FEM) were produced. The stress analyses for the piping inside the valve vaults, SMSH-15-001 and SMSH-15-002, overlap with the buried HOPE piping discussed in this summary report and qualified in SMSH-14-011.

The models from SMSH-14-011 are terminated using fictitious anchors at the new supports inside both of the valve vaults. This is considered conservative since flexibility of the piping inside the vaults are ignored. This has been validated as being conservative by comparing the controlling loads from the soil supported piping analyzed in SMSH-14-011 to those resulting from the overlapped models within calculations SMSH-15-001 and SMSH-15-002. The controlling loads in SMSH-14-011 were larger than those in SMSH-15-001 and SMSH-15-002 with the exception of the loads resulting from the non-repeated anchor movement applied in the new vault (for vault heaving).

The larger non-repeated anchor movement loads from SMSH-15-001 have been deemed acceptable because the margin factor from SMSH-14-011 is large (9.088) and the difference in controlling loads was small (approximately 25%).

The piping is supported at the model termination points considering fictitious anchors inside the valve vaults as previously discussed and at discrete node points for the soil supported piping that were modeled using springs. The springs in the FEM consist of pipe ovaling springs and soil springs, which are dependent on the modulus of elasticity of the pipe and soil properties in series.

Detailed calculations and resulting values for the springs are shown in Attachment A.

The modulus of elasticity of the pipe is different for the thermal and seismic load cases. Since soil parameters may vary, lower bound and upper bound spring values were determined for both the thermal and seismic load cases using a range of potential soil and pipe properties. Then, best estimate spring values were determined by taking the average of the calculated lower and upper bound spring values. These best estimate values were adjusted downward and upward by a coefficient of variation (COV) to calculate best estimate lower and upper bound springs. A COV of 1 was conservatively used in accordance with Ref. 7.7, Section 11.4.c.iii. This resulted in factors of 0.5 and 2.0 for the lower and upper bound best estimate cases respectively. The pipe was qualified for the bounding range of calculated upper and lower cases and best estimate upper and lower cases.

The spring values and node point spacing for the FEM were determined by following the methodology outlined in Appendix B & Appendix E of Ref. 7.3.

The spring values have an upper limit which corresponds to break-away between the pipe and the soil. Therefore, links are used in the models instead of springs. The links were modeled with a linear force-displacement relationship equal to the spring value while having upper limits equal to the breakaway displacement.

Page9

6.0 Assumptions & Design Considerations 6.1 Design Consideration:

A smaller trench width produces larger and less favorable ring deflection. A trench width of 3-ft is considered to be a lower bound value which is used as a worst case.

Justification:

This design consideration does not require verification as it is being incorporated into the detailed design through DCP SNC591628.

6.2 Design Consideration:

The flowable fill is conservatively considered to be used for all of the backfill (all the way to grade).

Justification:

This design consideration does not require verification as the backfill properties incorporated into the detailed design through DCP SNC591628 will be bounded by the properties used in the evaluation.

7.0 References 7.1 PPI, "Handbook of Polyethylene Pipe", 2nd Ed. with 6/6/12 Errata.

7.2 Marohl, M. P., "Comparison of Numerical Methods for Calculation of Vertical Soil Pressures on Buried Piping Due to Truck Loading," Proceedings of the ASME 2014 Pressure Vessels &

Piping Conference, PVP2014-28467, July 20-24, 2014.

7.3 EPRI Report 1013549, "Nondestructive Evaluation: Seismic Design Criteria for Polyethylene Pipe Replacement Code Case," Technical Update, September 2006.

7.4 Das, Braja M., "Principles of Foundation Engineering," 6th Ed.

7.5 Amencan Lifelines Alliance, "Guidelines for the Design of Buried Steel Pipe," July 2001 with Addenda through February 2005.

7.6 McGrath, T. J. and Hoopes, R. J.,* " 'Bedding Factors and E' Values for Buried Pipe Installations Backfilled with Air-Modified CLSM," The Design and Application of Controlled Low-Strength Materials (Fiowable Fill), ASTM STP 1331, A. K. Howard and J. L. Hitch, Eds.,

American Society for Testing and Materials, 1998.

7.7 NUREG-0800, Section 3.7.2, "Seismic System Analysis."

Page 10

Attachment A - Development of Spring & Breakaway Force Values This attachment develops the spring and breakaway force values that are input into the SAP2000 models.

See Design Input 3.2 through 3.6 for the replacement piping layout. The ground surface elevation (grade elevation), highest elevation to the pipe centerline and the deepest elevation to the pipe centerline are per Design Input 3.2 through 3.6.

ELground := I 29* ft ELmax := 125.4167*fl ELm in:= 120.5833* fl Epipe.th := 46000*psi Epipe.seis := 82000*psi Grade Elevation Pipe Centerline Maximum Elevation Pipe Centerline Minimum Elevation Modulus of Elasticity of Pipe for Thermal Evaluations (1 000-hrs & 73°F, From Table 3210-3 of Ref. 7.1)

Modulus of Elasticity of Pipe for Seismic Evaluations (0.5-hrs& 73°F, FromTable3210-3ofRef. 7.1)

The modulus of soil reaction for the backfill (flowable fill) is taken from Table 4 of Ref. 7.6 considering a minimum age of 28 days. Note that this value is similar to that of Table 3-8 of Chapter 6 of Ref. 7.1 for coarse-grain soil at a depth of cover ranging from 5 to 10 feet. Most of the pipe has a cover of at least 5 feet for the entire routing. There is a small portion near the new vault above 5 feet but the range of 5 to 10 feet is more appropriate for the entire piping.

E' := 3000* psi Modulus of Soil Reaction for Backfill The straight pipe portions consist of IPS 14 HOPE piping produced with PE4710 material of cell classification 445574C per Section 1100 of Design Input 3.1.

o := 14.00*in Average Outside Diameter of Pipe (See 3.8)

DR := 7 Dimension Ratio of Pipe 1fabmin := 2*in Minimum Fabricated Pipe Wall Thickness (See 3.8)

Di := D-2*trabmin = IO*in Inside Diameter of Pipe Delbow := 16*in Miter Elbow Outside Diameter Inside diameter for the miter elbows is set to match the inside diameter of the straight pipe.

0 i.elbow := D-2*trabmin = IO*in Delbow - 0 i.elbow 1elbow :=

= 3*m 2

0elbow DRelbow := --- = 5.333 1elbow D4- 0*4 lp:='TI'*

1 =1394.867*in4 64 Miter Elbow Inside Diameter Miter Elbow Wall Thickness Dimension Ratio of Miter Elbow Moment of Inertia of Pipe Page A-1

Attachment A - Development of Spring & Breakaway Force Values The maximum and minimum soil densities are taken as the upper and lower bound values for the typical range of weight for flowable fill (See 3.9). The upper bound value used is considered to include any soil saturation affects. The flowable fill is conservatively considered to be used for all of the backfill (all the way to grade) (See 6.3).

Pmax := 145*pcf Pmin := 70*pcf K:= 0.1 Maximum Density of Soil Above Pipe Minimum Density of Soil Above Pipe Bedding Factor (See 3.1, Section 321 0)

The methodology outlined in Appendix B & Appendix E of Ref. 7.3 is followed for developing pipe ovaling spring values, soil spring values, and spring spacing.

The soil properties range from a lower bound to an upper bound (soil density for example).

Therefore, each load condition will have a lower bound evaluation and an upper bound evaluation.

This results in four sets of spring values: seismic load condition lower bound, seismic load condition upper bound, thermal load condition lower bound, and thermal load condition upper bound.

The pipe ovaling springs are dependent on the pipe inside and outside diameter. Therefore, spring values for both the straight pipe and the miter elbows are calculated.

Lower Bound Pipe Ovaling Springs:

The lower bound modulus of soil reaction value is used for the lower bound springs.

E' = 3000* psi Modulus of Soil Reaction for Backfill (Previously Defined)

Per Ref. 7. 3, Section 5.1 0, a lag factor of 1. 0 is recommended for short-term loads and 1. 5 for long-term loads. Therefore, 1.0 is used for seismic load condition and 1.5 is used for thermal load condition.

Lt:s := 1.0 Lr.th := 1.5 Lag Factor (Ref. 7.3, Section 5.10)

Stiffness due to Pipe Ovaling for Straight Pipe (Seismic Load Condition) 2*Epipe.seis *(--l-)3 + 0.061 *E' 2

3 DR-I Kpo.s.ls := [)."

K-L

D = 12210.42*psl 1

f.s (Ref. 7.3, Eq. B-1a)

Stiffness due to Pipe Ovaling for Straight Pipe (Thermal Load Condition)

?.E.

h (

I

)3

_-__;,p-'1p'--e_.t_, ---

+ 0.06 I

E' 2

3 DR-I Kpo.th.ls := -

D = 6066.206*ps1 Di K-Lt:th (Ref. 7.3, Eq. B-1a)

Stiffness due to Pipe Ovaling for Miter Elbows (Seismic Load Condition) 2*E.

(

I

)3 p1pe.se1s.

+ 0.061. E' 2

3 DRelbow - 1 Kpo.s.le :=

=:..----~----

Delbow = 27354.407*psi (Ref. 7.3, Eq. B-1a)

Di.elbow K*Lt:s Stiffness due to Pipe Ovaling for Miter Elbows (Thermal Load Condition)

?.E.

h (

I

)3 plpe.t *

+ 0.061 *E' 2

3 DRelbow - 1 Kpo.th.le :=

~---____:;----

Delbow = I 1944.055*psi 0 i.elbow K*Leth (Ref. 7.3, Eq. B-1a)

Page A-2

Attachment A - Development of Spring & Breakaway Force Values Lower Bound Transverse Soil Spring:

The flowable fill is considered to behave like dense sand for determination of soil springs.

The minimum ground cover is used which, as a worst case, produces the smallest stiffness.

ELP := ELm ax = 125.417. ti Pipe Elevation (To Pipe Centerline) 1-1 := ELground-ELP = 3.583*ti Height of Ground Cover (Depth to Pipe Centerline)

The angle of internal friction, ljl,is dependent upon the relative density of the soil. The flowable fill is considered to be a dense soil and the value for 4> ranges between 40 and 45 degrees per Table 1.8 of Ref. 7.4. 40 degrees is considered as a best estimate for the actual value.

=40°: ( H) (1-1)2 (1-1)3 (1-1)4 Nqh := 10.959 + 1.783* D + 0.045* D - 0.005425* D - 0.0001153* D = 16.692 The minimum soil density is used for all of the lower bound spring calculations since this produces the smallest spring values. lbf ft 1 := D*p * *HN h = 407.066*- . s mm q in Transverse Breakaway Force (Per Table B-1 of Ref. 7.3, Sand) The transverse displacement ranges between 2% and 10% of (H+D/2) per Table B-1 of Ref. 7.3. The larger the displacement, the smaller the spring value. Therefore, the largest displacement values from Table B-1 of Ref. 7.3 are considered for all of the lower bound spring values. dt := 10%{ 1-1 + ~) = 5*in f t.ls Kt.ls := d = 81.414*pSI t Displacement (Per Table B-1 of Ref. 7.3, Sand) Transverse Soil Spring for Straight Pipe (Per Table B-1 of Ref. 7.3) PageA-3 Attachment A-Development of Spring & Breakaway Force Values Calculations for Miter Elbow Springs: Horizontal Bearing Capacity Factor (Ref. 7.5, Section B.2) for ljl=40°: Nqh := 10.959 + 1.783*( H ) + 0.045 *( H ) 2 - 0.005425*(-l-_l -) 3 - 0.0001153 *(-H-) 4 Delbow 0 elbow 0 elbow 0 elbow Nqh = 15.964 lbf tt.le := Delbow* Pmin' H* Nqh = 444.931*-; Transverse Breakaway Force (Per Table B-1 of Ref. 7.3, Sand) ( Delbow) dt := 10%* H + -- 2 = 5.1*in f t.le Kue := d = 87.242*ps1 t Lower Bound Axial Soil Spring: Displacement (Per Table B-1 of Ref. 7.3, Sand) Transverse Soil Spring for Miter Elbows (Per Table B-1 of Ref. 7.3) The coefficient of soil pressure at rest is calculated using Jaky's simplified equation from Section 7.2 of Ref. 7.4. This provides a reasonable empirical approximation of the true value. K0 := I -sin(<!>)= 0.357 Coefficient of Soil Pressure at Rest (Ref. 7.4, Section 7.2) Per Ref. 7.3, page B-4, the friction angle pipe-soil, 5, ranges between 0.5 and 0.8 times the internal angle of friction for sand. The smaller the friction angle pipe-soil, the smaller the spring value. Therefore, 0.5 times the internal angle of friction is considered for the lower bound spring values li := 0.5*<!> = 20*deg Friction Angle Pipe-Soil Calculations for Straight Pipe Springs: . ( D) lbf f 1 := TI*-

p * *H*(I + K )*tan(li) = 18.923*-

a. s 2

mm o in Breakaway Axial Force (Per Table B-1 of Ref. 7.3, Sand) da := 0.2*in f a.ls Ka.ls := -d- = 94.613*ps1 a Calculations for Miter Elbow Springs: Displacement (Per Table B-1 of Ref. 7.3, Sand) Lower Bound Axial Soil Spring for Straight Pipe (Per Table B-1 of Ref. 7.3) ( Delbow) ( ) lbf f 1 := TI*

p * *H* I + K *tan(li) = 21.626*-

a. e 2

mm o in Breakaway Axial Force (Per Table B-1 of Ref. 7.3, Sand) da := 0.2*in ra.le Ka.le := -d- = I 08.129* ps1 a Displacement (Per Table B-1 of Ref. 7.3, Sand) Lower Bound Axial Soil Spring for Miter Elbows (Per Table B-1 of Ref. 7.3) PageA-4 Attachment A-Development of Spring & Breakaway Force Values Lower Bound Vertical Soil Spring: The Downward Bearing Capacity Factors 1 and 2, Nq & Nv respectively, can be determined by using Figure B-3 of Ref. 7.3. In-lieu of using this figure, the equations in Section B.4 of Ref. 7.5 are used to obtain more accurate values. Nq := exp(1T*tan(cp))*tan(45*deg + ~r = 64.195 N1 := ex{O.I8* d!g-2.5) = 109.947 Calculations for Straight Pipe Springs: Downward Bearing Capacity Factor 1 (Ref. 7.5, Section B.4) Downward Bearing Capacity Factor 2 (Ref. 7.5, Section B.4) 2 N" lbf fd 1 := p * *H*N *D + p * *D *- = 2001.967*- . s mm q mm 2 in Vertical Down Breakaway Force (Per Table B-1 of Ref. 7.3, Sand) dd := 15%*D = 2.1*in f d.ls Kd.ls*= -d- = 953.317*psl d Displacement (Per Table B-1 of Ref. 7.3, Sand) Vertical Down Soil Spring for Straight Pipe (Per Table B-1 of Ref. 7.3) The Vertical Uplift Factor, Nqv* can be determined by using Figure B-4 of Ref. 7.3. These same figures are developed in Ref. 7.5. In-lieu of using the figures, the equation in Section B.3 of Ref. 7.5 is used, which provides a more accurate value. . ( <P*H ) Nqv := mm ,Nq = 2.792 44*deg*D du := 1.5%*H = 0.645*in f u.ls Ku.ls:= -d- = 105.569*psl u Vertical Uplift Factor (Ref. 7.5, Section B.3) Vertical Upward Breakaway Force (Per Table B-1 of Ref. 7.3, Sand) Displacement (Per Table B-1 of Ref. 7.3, Sand) (Using Upper Bound Value) Vertical Up Soil Spring for Straight Pipe (Per Table B-1 of Ref. 7.3) As recommended in Ref. 7.3, Appendix B, the total vertical soil spring is taken as the average of the vertical up and down. Similarly, the breakaway force is taken as the average of the up and down. K Ku.ls + Kd.ls "29 443 v.ls := = :l

psi 2

Vertical Soil Spring for Straight Pipe fu.ls + fd.ls lbf t~.ls := = 1035.029*- 2 in Vertical Breakaway Force for Straight Pipe PageA-5 Attachment A - Development of Spring & Breakaway Force Values Calculations for Miter Elbow Springs: dd := 15%* Del bow = 2.4* in f d.le Kd.le := -d- = 983.0l *pS1 d Vertical Down Breakaway Force (Per Table B-1 of Ref. 7.3, Sand) Displacement (Per Table B-1 of Ref. 7.3, Sand) Vertical Down Soil Spring for Miter Elbows (Per Table B-1 of Ref. 7.3) . ( c!>*H ) Nqv := mm , Nq = 2.443 44*deg* Del bow Vertical Uplift Factor (Ref. 7.5, Section B.3) lbf 1~.le := Delbow'Pmin*H*Nqv = 68.09l*j; Vertical Upward Breakaway Force (Per Table B-1 of Ref. 7.3, Sand) du := 1.5%*H = 0.645*in r u.le Ku.le := -d- = 105.,69*ps1 u K I + Kd le K

u. e

544.289psi v.lc

2 1~.le + 1d.le

_ lbf 1~.le :=

= 1213.6)8*-

2 in Lower Bound Influence Length:

Displacement (Per Table B-1 of Ref. 7.3, Sand)

Vertical Up Soil Spring for Miter Elbows (Per Table B-1 of Ref. 7.3)

Vertical Soil Spring for Miter Elbows Vertical Breakaway Force for Elbows Since the influence lengths affect the straight pipe portions, only the straight pipe properties are used to determine the influence length. The soil modulus {K0 ) is taken as the pipe ovalization spring stiffness following the example in Ref. 7.3, Appendix E.

Kpo.s.ls = 1221 0.42* psi I

f3s := (

Kpo.s.ls

) 4 = 0.072-~

4*Epipe.seis'1P m

3*7T LA 5 := -- = 32.782*in JJ*

4* f3s Kpo.th.ls = 6066.206*psi I

4 f3th := (. ~o.th. ls )

= 0.06973*~

4*Epipe.th*lp m

3*7T Lf3.th := -- = 33.792*in 4*f3th Lf3.l := max(Lf3.s* Lf3.th} = 33.792*in Previously Calculated Seismic Pipe Ovalization Spring Stiffness Characteristic of the System (Ref. 7.3, Eq. B-6)

Influence Length (Seismic Load Condition) (Ref. 7.3, Eq. B-5)

Previously Calculated Thermal Pipe Ovalization Spring Stiffness Characteristic of the System (Ref. 7.3, Eq. B-6)

Influence Length (Thermal Load Condition) (Ref. 7.3, Eq. B-5)

Controlling Influence Length for Lower Bound Springs Page A-6

Attachment A - Development of Spring & Breakaway Force Values Upper Bound Pipe Ovaling Springs:

The larger the modulus of elasticity of the soil, the larger the spring value. Therefore, the modulus of elasticity for the backfill is taken from Table 3-9 of Chapter 6 of Ref. 7.1 as 20000-psi, which is the largest tabulated, realistic value.

E' := 20000*psi Upper Bound Modulus of Elasticity for Backfill Stiffness due to P1pe Ovaling for Straight Pipe (Seismic Load Condition) pipe.seJs.

+ 0.061. E' 2*E.

. (

I

)3 2

3 DR-I Kpo.s.us := -*

D = 41246.42*psl Di K*Lt:s Stiffness due to Pipe Ovaling for Straight Pipe (Thermal Load Condition) 2*E. h ( I )3 plpe.t *

+ 0.061

E' 2

3 DR-I Kpo.th.us := -*

D = 2::>423.539*psl Di K*Lt:th Stiffness due to Pipe Ovaling for Miter Elbows (Seismic Load Condition) 2*E.

. (

I

)3 plpe.seJs.

+ 0.061. E' 2

3 DRelbm' - I Kpo.s.ue :=

>.----~----*Delbow Di.elbow K* Lt:s Kpo.s.ue = 60538.407*psi Stiffness due to Pipe Ovaling for Miter Elbows (Thermal Load Condition) 2 Kpo.th.ue := -D--

i.elbow 2*E.

h (

I

)3 plpe.t *

+ 0.061

E' 3

DRelbow-1 K L.

Delbow

. l:th Kpo.th.ue = 34066.721 *psi Page A-7 (Ref. 7.3, Eq. B-1a)

(Ref. 7.3, Eq. B-1a)

Attachment A - Development of Spring & Breakaway Force Values Upper Bound Transverse Soil Spring:

The maximum ground cover is used which, as a worst case, produces the largest stiffness.

ELP := ELmin = 120.583*ft Pipe Elevation (To Pipe Centerline)

H := ELground - ELP = 8.417

ft Height of Ground Cover (Depth to Pipe Centerline)

The angle of internal friction, <jl,is dependent upon the relative density of the soil. The flowable fill is considered to be a dense soil and the value for <jl ranges between 40 and 45 degrees per Table 1.8 of Ref. 7.4. 40 degrees is considered as a best estimate for the actual value.

cj> := 40*deg Angle of Internal Friction (Ref. 7.4, Table 1.8)

Calculations for Straight Pipe Springs:

The Horizontal Bearing Capacity Factor, Nqh* can be determined by using Figures B-1 and B-2(a) of Ref. 7.3. These same figures are developed in Ref. 7.5. In-lieu of using the figures, the equation in Section B.2 of Ref. 7.5 is used, which provides a more accurate value.

Horizontal Bearing Capacity Factor (Ref. 7.5, Section B.2) for <jl=40°:

Nqh := 10.959 + 1.783{~) + 0.045{~r _ 0.005425-(~f _ 0.0001153{~r = 23.815 The maximum soil density is used for all of the upper bound spring calculations since this produces the largest spring values.

lbf

~'t.us := D*Pmax*H*Nqh = 2825.689*-.-

Transverse Breakaway Force (Per Table B-1 of Ref. 7.3, Sand) m The transverse displacement ranges between 2% and 10% of (H+D/2) per Table B-1 of Ref. 7.3.

The smaller the displacement, the larger the spring value. Therefore, the smallest displacement values from Table B-1 of Ref. 7.3 are considered for all of the upper bound spring values.

d1 := 2%{ H + ~) = 2.16*in r t.us Kt.us := -d- = 1308.184*psl t

Displacement (Per Table B-1 of Ref. 7.3, Sand)

Transverse Soil Spring for Straight Pipe (Per Table B-1 of Ref. 7.3)

Page A-8

Attachment A-Development of Spring & Breakaway Force Values Calculations for Miter Elbow Springs:

Horizontal Bearing Capacity Factor (Ref. 7.5, Section B.2) for <!>=40°:

Nqh := 10.959 + 1.783*(-H-) + 0.045*(-H-)

2

- 0.005425*(-H-)

3

- 0.0001153*(-H-)

4 0 elbow 0 elbow 0 elbow 0 elbow Nqh = 22.46 lbf tt.ue := Delbow*Pmax*H*Nqh = 3045.589*ir;"" Transverse Breakaway Force (Per Table B-1 of Ref. 7.3, Sand)

Upper Bound Axial Soil Spring:

Displacement (Per Table B-1 of Ref. 7.3, Sand)

Transverse Soil Spring for Miter Elbows (Per Table B-1 of Ref. 7.3)

The coefficient of soil pressure at rest is calculated using Jaky's simplified equation from Section 7.2 of Ref. 7.4. This provides a reasonable empirical approximation of the true value.

K0 := I -sin(<!>)= 0.357 Coefficient of Soil Pressure at Rest (Ref. 7.4, Section 7.2)

Per Ref. 7.3, page B-4, the friction angle pipe-soil, o, ranges between 0.5 and 0.8 times the internal angle of friction for sand. The larger the friction angle pipe-soil, the larger the spring value. Therefore, 0.8 times the internal angle of friction is considered for the upper bound spring values o := 0.8*<1> = 32*deg Friction Angle Pipe-Soil Calculations for Straight Pipe Springs:

( D) lbf fa.us := 1!*2 *Pmax*H*(l + K0}*tan(O) =. 1:>8.064*ir;""

Breakaway Axial Force (Per Table B-1 of Ref. 7.3, Sand) da := O.l *in f a.us Ka.us := -d- = 1580.638*psi a

Calculations for Miter Elbow Springs:

Lower Bound Displacement (Per Table B-1 of Ref. 7.3, Sand)

Upper Bound Axial Soil Spring for Straight Pipe (Per Table B-1 of Ref. 7.3)

(

Delbow)

(

)

lbf t~.ue :=

TI"*--

2

Pmax*H* I + K0 *tan(o) = 180.644*ir;""

Breakaway Axial Force (Per Table B-1 of Ref. 7.3, Sand) da :=O.l *in ra.ue Ka.ue := -d- = 1806.443*pSI a

Displacement (Per Table B-1 of Ref. 7.3, Sand)

Upper Bound Axial Soil Spring for Miter Elbows (Per Table B-1 of Ref. 7.3)

PageA-9

Attachment A - Development of Spring & Breakaway Force Values Upper Bound Vertical Soil Spring:

The Downward Bearing Capacity Factors 1 and 2, Nq and Nv respectively, can be determined by using Figure B-3 of Ref. 7.3. In-lieu of using this figure, the equations in Section B.4 of Ref. 7.5 are used to obtain more accurate values.

Nq := exp(11*tan(cj>))*tan(45*deg + :r = 64.195 N1 := exp(0.18*_1_- 2.5) = 109.947 deg Downward Bearing Capacity Factor 1 (Ref. 7.5, Section B.4)

Downward Bearing Capacity Factor 2 (Ref. 7.5, Section B.4)

Calculations for Straight Pipe Springs:

dd := IOo/o*D = 1.4*in r d.us Kd.us := -d- = 6086.452*psJ d

Vertical Down Breakaway Force (Per Table B-1 of Ref. 7.3, Sand)

Displacement (Per Table B-1 of Ref. 7.3, Sand)

Vertical Down Soil Spring for Straight Pipe (Per Table B-1 of Ref. 7.3)

The vertical uplift factor can be determined by using Figure B-4 of Ref. 7.3. These same figures are developed in Ref. 7.5. In-lieu of using the figures, the equation in Section B.3 of Ref. 7.5 is used which provides a more accurate value.

. (

cJ>*H

)

Nqv := mm

, Nq = 6.5,:,8 44*deg*D lbf t~.us := D* Pmax*l+ Nqv = 778.176*-.-

Jn du := 0.5%*H = 0.505*in f u.us 0

Ku.us := -d- = b40.936*psJ u

Vertical Uplift Factor (Ref. 7.5, Section B.3)

Vertical Upward Breakaway Force (Per Table B-1 of Ref. 7.3, Sand)

Displacement (Per Table B-1 of Ref. 7.3, Sand)

(Using Lower Bound Value)

Vertical Up Soil Spring for Straight Pipe (Per Table B-1 of Ref. 7.3)

As recommended in Ref. 7.3, Appendix B, the total vertical soil spring is taken as the average of the vertical up and down. Similary, the breakaway force is taken as the average of the up and down.

Ku.us + Kd.us Kv.us :=

= 3813.694*psl 2

Vertical Soil Spring for Straight Pipe f

+ fd lbf

t.

u.us

.us = 4649.604*-

v.us :=

2 in Vertical Breakaway Force for Straight Pipe Page A-10

Attachment A - Development of Spring & Breakaway Force Values Calculations for Miter Elbow Springs:

lbf fd ue = 9885.937*-.-

In dd := IO%* Delbow = 1.6*in f d.ue Kd.ue:= -d- = 6178.7ll *ps1 d

Nqv :=min(

<l>*H

,Nq) = 5.739 44*deg* Del bow Vertical Down Breakaway Force (Per Table B-1 of Ref. 7.3, Sand}

Displacement (Per Table B-1 of Ref. 7.3, Sand}

Vertical Down Soil Spring for Miter Elbows (Per Table B-1 of Ref. 7.3)

Vertical Uplift Factor (Ref. 7.5, Section B.3}

lbf lu.ue := Delbow'Pmax*H*Nqv = 778.176*-.-

Vertical Upward Breakaway Force du := 0.5%*H = 0.505*in f u.ue Ku.ue := -d- = 1.:>40.936*ps1 u

K

+ Kd Kv.ue :=

u.ue

.ue = 3859.824*psi 2

1~.ue + 1ct ue lbt' 1~.ue :=

= 5332.057*-

2 in m

(Per Table B-1 of Ref. 7.3, Sand}

Displacement (Per Table B-1 of Ref. 7.3, Sand}

Vertical Up Soil Spring for Miter Elbows (Per Table B-1 of Ref. 7.3}

Vertical Soil Spring for Miter Elbows Vertical Breakaway Force for Elbows Page A-11

Attachment A - Development of Spring & Breakaway Force Values Upper Bound Influence Length:

Since the influence lengths affect the straight pipe portions, only the straight pipe properties are used to determine the influence length. The soil modulus (K0 ) is taken as the pipe ovalization spring stiffness following the example in Ref. 7.3, Appendix E.

Kpo.s.us = 41246.42*psi 4

f\\ := (

~o.s.us J = 0.097 *~

4* Epipe.seis* 1 P m

3*71 LA := -- = 24.181*in 1-'*s 4*(\\

Kpo.th.us = 25423.539*psi I

13th := (

Kpo.th.us J 4 = 0.09976 *~

4*Epipe.th*lp m

3*71 L13.th := -- = 23.618*in 4* 13th Previously Calculated Seismic Pipe Ovalization Spring Stiffness Characteristic of the System (Ref. 7.3, Eq. B-6)

Influence Length (Seismic Load Condition) (Ref. 7.3, Eq. B-5)

Previously Calculated Thermal Pipe Ovalization Spring Stiffness Characteristic of the System (Ref. 7.3, Eq. B-6)

Influence Length (Thermal Load Condition) (Ref. 7.3, Eq. B-5)

Controlling Influence Length for Upper Bound Springs Page A-12

Attachment A-Development of Spring & Breakaway Force Values To simplify the modeling process, the enveloping spring for either the straight pipe properties or the miter bend properties is used.

Kpo.s.l := min(Kpo.s.ls*Kpo.s.le) = 12210.42*psi Kpo.th.l := min(Kpo.th.ls* Kpo.th.le) = 6066.206*psi Kt.l := min(Ku5,Kue) = 81.414*psi Ka.l := min(Ka.ls*Ka.le) = 94.613*psi Kv.l := min(Kv.ls*Kv.le) = 529.443 *psi Kpo.s.u := max(Kpo.s.us*Kpo.s.ue) = 60538.407*psi Kpo.th.u := max(Kpo.th.us* Kpo.th.ue) = 34066.721 *psi Kt.u := max(Kt.us*Kt.ue) = 1397.054*psi Ka.u := max( Ka.us, Ka.ue) = 1806.443 *psi Kv.u := max(Kv.us* Kv.ue) = 3859.824*psi Lower Bound Pipe Ovaling Spring for Seismic Lower Bound Pipe Ovaling Spring for Thermal Lower Bound Transverse Soil Spring Lower Bound Axial Soil Spring Lower Bound Vertical Soil Spring Upper Bound Pipe Ovaling Spring for Seismic Upper Bound Pipe Ovaling Spring for Thermal Upper Bound Transverse Soil Spring Upper Bound Axial Soil Spring Upper Bound Vertical Soil Spring Average of Lower and Upper Spring Values & Lower and Upper Axial Breakaway Forces:

The average values are considered best estimates.

Kpo.s.a := 0.5*(Kpo.s.l + Kpo.s.u) = 36374.413 *psi Kpo.th.a := 0.5*(Kpo.th.l + Kpo.th.u) = 20066.464*psi Kt.a := 0.5*(Ku + Kt.u) = 739.234*psi Ka.a := 0.5*(Ka.l + Ka.u) = 950.528*psi Kv.a := 0.5*(Kv.l + Kv.u) = 2194.633*psi lbf l(.as := 0.5*(ti.ls + t(.us) = 1616.377-1;

- (.

)

9 lbf 1a.as := O.::l* 1a.ls + 1a.us = 88.4 3*1; I.

-o*(

1*

)-284?317lbf v.as.-

.)* 1v.ls + v.us -

'1; PageA-13 lbf l(.ae := 0.5*(tt.le + t(.ue) = 1745.26*1; t~.ae := 0.5*(t~.le + t~.ue) = 101.135* ~~~*

l~*.ae := 0.5 * (t~ *. le + t~.ue) = 3272.857* ~~~*

Attachment A-Development of Spring & Breakaway Force Values 1/2 x Average Spring Values:

Kpo.s.al := 0.5*Kpo.s.a = 18187.207*psi Kpo.th.al := 0.5*Kpo.th.a = 10033.232-psi Kt.al := 0.5* Kt.a = 369.617 *psi Ka.al := 0.5*Ka.a = 475.264-psi Kv.al := 0.5*Kv.a = 1097.317-psi 1/2 x Average Axial Breakaway Forces:

lbf ft.ale := 0.5* tt.ae = 872.63*1; t*

0 8 89 lbf t.als.=.5* tt.as = 80.I

. 1; lbf ra.ale := 0.5* ~~.ae = 50.568*1; lbf fa.als := 0..)* fa. as = 44.24 7*1; lbf fv.ale := 0..)* fv.ae = 1636.429*1; 1~.als := 0.5* 1~.as = 1421.158* ~~;*

Page A-14 2 x Average Spring Values:

Kpo.s.au := 2*Kpo.s.a = 72748.827-psi Kpo.th.au := 2*Kpo.th.a = 40132.927-psi Kt.au := 2*Kt.a = 1478.468-psi Ka.au := 2*Ka.a = 1901.056*psi Kv.au := 2*Kv.a = 4389.267*psi 2 x Average Axial Breakaway Forces:

lbf fi.aue := 2*ti.ae = 3490.52*1; lbf ft.aus := 2*ft.as = 3232.755*1; lbf fa.aue := 2*fa.ae = 202.27*1; lbf 1~. aus := 2* ~~.as= 176.986*1; lbf 1v.aue := 2*1v.ae = 6.)4.)*714'1; lbf rv.aus := 2* ~~.as= 5684.633*1;

Attachment A-Development of Spring & Breakaway Force Values Controlling Lower Bound Spnngs & Axial Breakaway Forces for Models:

Kpo.s.bl := min(Kpo.s.I*Kpo.s.al) = 12210.42*psi Kpo.th.bl := min(Kpo.th.I*Kpo.th.al) = 6066.206-psi Ka.bl := min(Ka.I*Ka.al) = 94.613*psi Kv.bl := min(Kv.l* Kv.al) = 529.443*psi Kt.bl := min(Kt.I,Kt.al) = 81.414-psi

f.

0 (t" 0

)

lbf" a.bl.s := mm a.ls* 1a.als = l8.923*i; 0

0

(

0 0

)

6 lbf 1a.bl.e := m m 1a.le* fa ale = 21.62

1; lbf" fv.bl.s := min( ~~.Is* fv.als) = I 035.029*-.-m 1~.bl.e := min(rv.Ie* f~.ale) = 1213.658* l~f In 0

0

(

0 0

)

6 lbf" tt.bl.s := mm tt.ls* tt.als = 407.06 *1; 0

0

(

0 0

)

lbf ft.bl.e := mm tt.le* tt.ale = 444.931*1; Lower Bound Pipe Ovalinq Spring for Seismic Lower Bound Pipe Ovaling Spring for Thermal Lower Bound Axial Soil Spring Lower Bound Vertical Soil Spring Lower Bound Transverse Soil Spring Lower Bound Axial Breakaway Force for Straight Pipe Lower Bound Axial Breakaway Force for Elbows Lower Bound Vertical Breakaway Force for Straight Pipe Lower Bound Vertical Breakaway Force for Elbows Lower Bound Transverse Breakaway Force for Straight Pipe Lower Bound Transverse Breakaway Force for Elbows Note: The breakaway displacements for the models are determined by dividing the breakaway force value by its corresponding spring value. These values aren't shown here due to the simplicity of calculating them from the reported values documented herein.

Page A-15

Attachment A - Development of Spring & Breakaway Force Values Controlling Upper Bound Springs & Axial Breakaway Forces for Models:

Kpo.s.bu := max( ~o.s.u, Kpo.s.au) = 72748.827 *psi Kpo.th.bu := max(Kpo.th.u*Kpo.th.au) = 40132.927*psi Ka.bu := max(Ka.u*Ka.au) = 1901.056*psi Kv.bu := max(Kv.u* Kv.au) = 4389.267*psi Kt.bu := max(Kt.u*Kt.au) = 1478.468*psi

f.

(f'

)

lbf a.bu.s := max a. us

fa.aus = 176.986-i; t~.bu.e := max(f~.ue* t~.aue) = 202.27* ~~:*

f*

(.

f

)

lbf v.bu.s := max fv.us* v.aus = :l684.633*-.-

m t~'.bu.e := max(t~.ue*f~.aue) = 6545.714*

1

~:*

ft.bu.s := max(tt.us* t[.aus) = 3232.755* ~~:*

ft.bu.e := max( lt.ue* tt.aue) = 3490.52* ~~:

Upper Bound Pipe Ovaling Spring for Seismic Upper Bound Pipe Ovaling Spring for Thermal Upper Bound Axial Soil Spring Upper Bound Vertical Soil Spring Upper Bound Transverse Soil Spring Upper Bound Axial Breakaway Force for Straight Pipe Upper Bound Axial Breakaway Force for Elbows Upper Bound Vertical Breakaway Force for Straight Pipe Upper Bound Vertical Breakaway Force for Elbows Upper Bound Transverse Breakaway Force for Straight Pipe Upper Bound Transverse Breakaway Force for Elbows Note: The breakaway displacements for the models are determined by dividing the breakaway force value by its corresponding spring value. These values aren't shown here due to the simplicity of calculating them from the reported values documented herein.

Controlling Influence Length:

L~:= max(L~.I*L~.u) = 33.792*in The spring values and breakaway forces above are dependent on their effective lengths. Therefore, the value modeled will vary from node to node.

Page A-16

Edwin I. Hatch Nuclear Plant-Unit 2 Proposed In service Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0 Final Stress Analysis Design of the HOPE System Summary of Stress Analysis Calculation for New U2 Div II PSW Subgrade Vault Piping

Summary Report on Stress Analysis for New U2 DIV II PSW Subgrade Vault Piping (Near the Reactor Building)

Hatch Nuclear Plant-Unit 2 1.0 Purpose of Summary Report This report summarizes the stress analysis calculation, SMSH-15-001, that has been prepared to support SNC's lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0 (ATR). This lSI Alternative is needed to support the planned replacement of buried steel piping in Hatch Nuclear Plant Unit 2, Plant Service Water (PSW) system with High Density Polyethylene (HOPE) piping.

The piping to be replaced is the supply piping from the Unit 2 Service Water Valve Pit 28 to a new subgrade vault located outside of the Unit 2 Reactor Building. Stress analysis calculation SMSH 001 evaluates the replacement piping inside the new subgrade vault to the design requirements in the ATR. Note that only specific sections of the ATR apply for the piping evaluated by the stress analysis since the piping being qualified is inside the valve vault (i.e. no surcharge loads, no ring deflection, etc.). The stress analysis also evaluates the metallic piping inside the vault and captures the mutual influence between the replacement HOPE piping and the metallic piping.

All computations from the calculation are not included in the summary. Calculation results and conclusions are shown in Section 2.0.

2.0 Results and Conclusions Following the methodology outlined in Section 5.0, the replacement HOPE piping meets all of the acceptance criteria outlined in the lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0. The metallic piping meets all of the applicable requirements of paragraph 3.9.2.1 of Design Input 3.4 and all of the applicable requirements of Design Input 3.5. The valve acceleration values are less than 4.5g in each orthogonal direction and are therefore acceptable per Section 3.5.3 of Design Input 3.13.

Detailed results are shown in this section. All margin factor values are greater than 1.0 and are therefore acceptable. The controlling margin factor was determined to be 1.63 which was for the HOPE piping, Service Level 8 loading for the lower bound spring condition.

The loads on the flanged connection between the metallic piping and the HOPE piping are reported in this section. These loads are evaluated in Calculation SMNH-15-008 (Ref. 7.9).

The piping has been analyzed to the following conditions per Section 1100 of Design Input 3.1. A design temperature value of 125°F was conservatively used which bounds the actual design temperature of 123°F.

Condition Temperature, °F Pressure, psig Normal Operating 95 140 Maximum Operating 97 190 Design 123 180 Page 1

Summary of SAP2000 Output A summary of the controlling pipe member forces is provided below.

HOPE Vault 1 0" Stainless 1" Stainless HOPE Straight HOPE Elbows Model Load Case Piping Steel Vault Steel Vault Pipe Soil Soil Supported Straight Pipe Piping Piping Supported Piping p (lb) M (lb-in) p (lb) M (lb-in) p (lb) M (lb-in) p (lb)

M (lb-in) p (lb)

FULLOEAD 0

4269 240 9391 18 0

0 3747 0

Seismic OBE 80 4587 192 11916 1

35 296 1207 155 Lower Bound DBE 122 6968 278 19248 1

49 538 1750 251 HEAVE 0

22767 0

28470 0

0 0

4421 0

FULLDEAD 0

1812 240 10372 18 0

0 3746 0

Seismic OBE 337 1974 366 5245 1

33 925 1321 415 Upper DBE 617 2945 682 8443 2

58 1733 2236 777 Bound HEAVE 0

34129 0

46386 0

0 0

9884 0

Thermal FULLOEAD 0

3064 240 10271 18 0

0 3753 0

Lower THERMAL ABS 1033 17382 1033 45476 0

0 8209 24242 3319 Bound HEAVE 0

14759 0

18815 0

0 0

1730 0

Thermal FULLOEAD 0

1330 240 10764 18 0

0 3743 0

Upper THERMAL_ABS 5128 13356 5128 33084 0

0 25117 13934 11802 Bound HEAVE 0

21678 0

29927 0

0 0

7058 0

FULLDEAD 0

4269 240 10764 18 0

0 3753 0

OBE 337 4587 366 11916 1

35 925 1321 415 Bounding DBE 617 6968 682 19248 2

58 1733 2236 777 Values HEAVE 0

34129 0

46386 0

0 0

9884 0

THERMAL_ABS 5128 17382 5128 45476 0

0 25117 24242 11802 A summary of the valve accelerations for the 1" stainless steel valve is provided below.

Load Case U1 (g)

U2 (g)

U3 (g)

Max. (g)

OBE Lower Bound 0.255 0.214 0.038 0.255 DBE Lower Bound 0.361 0.305 0.071 0.361 OBE Upper Bound 0.329 0.172 0.072 0.329 DBE Upper Bound 0.585 0.323 0.134 0.585 OBE Bounding Values 0.329 0.214 0.072 0.329 OBE Bounding Values 0.585 0.323 0.134 0.585 In order to verify that an adequate amount of soil supported piping is modeled, fictitious boundary check loads are applied to the vault piping and the resulting loads at the soil supported piping termination point are verified to be negligible as discussed in Section 5.0. A summary of the joint forces resulting from the boundary check fictitious loads at the model termination point is provided below. All of the loads are either zero or negligible indicating adequate modeling of the soil supported piping.

Node 1190 Model (Model Termination Point)

FX (lb)

FY(Ib)

FZ(Ib)

MX (lb-in)

MY (lb-in)

MZ (lb-in)

Seismic Lower Bound 0.24 0.81 0.52 0.00 0.00 0.00 Seismic Upper Bound 0.01 0.00 0.13 0.00 0.00 0.00 Thermal Lower Bound 0.12 0.44 0.49 0.00 0.00 0.00 Thermal Upper Bound 0.00 0.00 0.04 0.00 0.00 0.00 Bounding Values 0.24 0.81 0.52 0.00 0.00 0.00 Page2 M (lb-in) 1340 2449 3571 1470 849 3185 5660 9884 1266 51250 1438 694 68954 7058 1340 3185 5660 9884 68954

A summary of the loads on the flanged interface between the HOPE piping and the metallic piping is provided in the following table. Note that the OBE and OBE load cases include the SAM_OBE and SAM_OBE loads, respectively. THERMAL_MIN is based on the minimum operating temperature of 32°F, THERMAL_MAXOP is based on the maximum operating temperature of 97°F, and THERMAL_MAX is based on the design temperature which is conservatively taken as 125°F.

Node 1100 Model Load Case (Flanged Connection Between HOPE & S.S.)

FX (lb)

FY (lb) FZ (lb)

MX (lb-in)

MY (lb-in)

MZ (lb-in)

FULLOEAO 0

0 427 4264 203 0

OBE 126 75 4

132 35 4585 Seismic OBE 195 116 7

246 56 6963 Lower SAM OBE 24 22 0

7 0

756 Bound SAM_OBE 47 44 0

15 0

1511 HEAVE 0

0 634 22763 413 0

FULLOEAO 0

0 361 1810 101 0

OBE 81 335 10 78 21 1973 Seismic OBE 126 611 19 147 35 2941 Upper SAM OBE 4

18 0

2 0

137 Bound SAM_OBE 8

35 0

3 0

273 HEAVE 0

0 1362 34117 902 0

FULLOEAO 0

0 398 3057 205 0

Thermal THERMAL MIN 120 416 0

0 0

6778 Lower THERMAL MAXOP 77 288 0

0 0

4373 Bound THERMAL MAX 188 616 0

0 0

10604 HEAVE 0

0 451 14754 398 0

FULLOEAO 0

0 345 1327 89 0

Thermal THERMAL MIN 157 2092 0

0 0

844 Upper THERMAL MAXOP 112 1487 0

0 0

600 Bound THERMAL MAX 230 3036 0

0 0

1289 HEAVE 0

0 917 21673 457 0

FULLOEAO 0

0 427 4264 205 0

OBE 126 335 10 132 35 4585 OBE 195 611 19 246 56 6963 SAM OBE 24 22 0

7 0

756 Bounding SAM OBE 47 44 0

15 0

1511 Values THERMAL MIN 157 2092 0

0 0

6778 THERMAL MAXOP 112 1487 0

0 0

4373 THERMAL MAX 230 3036 0

'0 0

10604 HEAVE 0

0 1362 34117 902 0

Page3

A summary of the loads and pipe movements on the vault support is provided in the following table.

Note that the OBE and DBE load cases include the SAM_OBE and SAM_DBE loads, respectively.

THERMAL_MIN is based on the minimum operating temperature of 32°F, THERMAL_MAXOP is based on the maximum operating temperature of 97°F, and THERMAL_MAX is based on the design temperature which is conservatively taken as 125°F.

Node 3400 Model Load Case (Support Inside Vault)

FZ (lb)

DX (in)

DY (in)

FULLDEAD 1270.69 0.0000 0.0000 OBE 17.36 0.0535 0.0152 Seismic DBE 32.27 0.0888 0.0222 Lower SAM OBE 2.15 0.0058 0.0022 Bound SAM DBE 4.30 0.0116 0.0043 HEAVE 1103.88 0.0000 0.0000 FULLDEAD 1150.82 0.0000 0.0000 OBE 28.29 0.0146 0.0026 Seismic DBE 53.22 0.0242 0.0048 Upper SAM OBE 0.71 0.0001 0.0003 Bound SAM DBE 1.42 0.0003 0.0007 HEAVE 2133.41 0.0000 0.0000 FULL DEAD 1226.44 0.0000 0.0000 Thermal THERMAL MIN 0.00 0.1772 0.0504 Lower THERMAL MAXOP 0.00 0.1217 0.0344 Bound THERMAL MAX 0.00 0.2641 0.0754 HEAVE 821.75 0.0000 0.0000 FULLDEAD 1127.78 0.0000 0.0000 Thermal THERMAL MIN 0.00 0.0222 0.0102 Upper THERMAL MAXOP 0.00 0.0158 0.0073 Bound THERMAL MAX 0.00 0.0319 0.0148 HEAVE 1510.12 0.0000 0.0000 FULLDEAD 1270.69 0.0000 0.0000 OBE 28.29 0.0535 0.0152 DBE 53.22 0.0888 0.0222 SAM OBE 2.15 0.0058 0.0022 Bounding SAM DBE 4.30 0.0116 0.0043 Values THERMAL MIN 0.00 0.1772 0.0504 THERMAL MAXOP 0.00 0.1217 0.0344 THERMAL MAX 0.00 0.2641 0.0754 HEAVE 2133.41 0.0000 0.0000 Page4

HOPE Vault Piping Stresses Service level stresses for the HOPE vault piping are summarized in the following table.

Service Spring Pipe Stress Stress Factor x Margin Stress Level Condition

{psi)

Factor Allowable Stress (psi)

Factor A

Lower Bound 336 1

561 1.67 Upper Bound 324 1

561 1.73 B

Lower Bound 378 1.1 617.1 1.63 Upper Bound 356 1.1 617.1 1.73 0

Lower Bound 391 1.33 746.13 1.91 Upper Bound 365 1.33 746.13 2.05 Seismic-induced stresses for the HOPE vault piping are summarized in the following table.

Spring Pipe Stress Allowable Stress Margin Condition

{psi)

Range (psi)

Factor Lower Bound 73 2032 27.77 Upper Bound 46 2032 44.26 Thermal expansion and contraction stresses for the HOPE vault piping are summarized in the following table.

Spring Pipe Stress Allowable Stress Margin Condition (psi)

Range (psi)

Factor Lower Bound 101 2032 20.13 Upper Bound 135 2032 15.05 Non-repeated anchor movement stresses for the HOPE vault piping are summarized in the following table.

Pipe Stress Long Term Allowable Margin Factor (psi)

Stress (psi) 171 1122 6.55 Stainless Steel Vault Piping Stresses Primary stresses for the stainless steel vault piping due to Normal Conditions are summarized in the following table.

Component Spring Load Pipe Stress Allowable Margin Case (psi)

Stre.ss (psi)

Factor 10" Steel Lower Bound 1929 16600 8.60 Piping Upper Bound 1961 16600 8.46 1" Steel Lower Bound 332 16600 49.93 Piping Upper Bound 332 16600 49.93 Primary stresses for the stainless steel vault piping due to Upset Conditions are summarized in the following table.

Component Spring Load Pipe Stress Allowable Margin Case (psi)

Stress (psi)

Factor 10" Steel Lower Bound 2708 19920 7.36 Piping Upper Bound 2304 19920 8.65 1" Steel Lower Bound 846 19920 23.56 Piping Upper Bound 812 19920 24.54 Page 5

Primary stresses for the stainless steel vault piping due to Faulted Conditions are summarized in the following table.

Component Spring Load Pipe Stress Allowable Margin Case (psi)

Stress (psi)

Factor 1 0" Steel Lower Bound 3187 39840 12.50 Piping Upper Bound 2513 39840 15.85 1" Steel Lower Bound 1058 39840 37.65 Piping Upper Bound 1182 39840 33.70 Secondary stresses for the stainless steel vault piping are summarized in the following table. There are two combinations that may be used per Section 3.9.2.1.B of Design Input 3.4. The thermal expansion plus OBE anchor displacement stresses option is considered. Since the OBE anchor displacements have already been accounted for in the primary stress evaluations, the OBE anchor displacement is omitted from the secondary stress evaluations as permitted by note "a" in Section 3.9.2.1.B of Design Input 3.4.

Loading Component Spring Load Pipe Stress Allowable Margin Case (psi)

Stress (psi)

Factor Lower Bound 3962 24900 6.29 1 0" Steel Piping Upper Bound 2882 24900 8.64 Thermal Lower Bound 0

24900 N/A 1" Steel Piping Upper Bound 0

24900 N/A Non-Lower Bound 2480 49800 20.08 1 0" Steel Piping Repeated Upper Bound 4041 49800 12.32 Anchor Lower Bound 0

49800 N/A Movement 1" Steel Piping Upper Bound 0

49800 N/A Valve Accelerations Valves are capable of withstanding 4.5g in all directions as discussed in Section 4.0. Valve accelerations are less than 4.5g in all directions and are therefore acceptable.

Model Termination & Overlap Reconciliation The model terminations as described in Section 5.0 is adequate since the resulting loads at the termination point due to the fictitious boundary check load are negligible.

The models in analysis SMSH-14-011 (Design Input 3.14) are terminated using a fictitious anchor at the vault support. This was theorized as being conservative as discussed in Attachment A of SMSH-14-011. The models used in SMSH-15-001 overlap with the SMSH-14-011 models. All controlling loads for the soil supported HPDE piping in SMSH-14-011 are larger than those resulting in SMSH-15-001 except for the HEAVE load case for the upper bound spring models which has a slight difference (approximately 25% difference). Therefore, all of the evaluations within SMSH-14-011 for the soil supported HOPE piping are acceptable with the exception of the evaluation of non-repeated anchor movements (HEAVE load case). The non-repeated anchor movements remain acceptable for the larger loads determined in SMSH-15-001 since there is a large margin factor of 9.088 in SMSH-14-011 and the load difference is small.

Other Design Considerations As noted in Enclosure 2 of Design Input 3.1, other design considerations will be addressed under SNC design procedures in accordance with the existing design and license basis for HNP.

Page6

3.0 Design Inputs 3.1 Edwin I. Hatch Nuclear Plant-Unit 2, Proposed lnservice Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0.

The acceptance criteria and general methodology was taken from this design input document.

3.2 HNP-1-FSAR-12, Rev. 19 (7/01).

This design input document is used to develop the response spectrum for the vault piping.

3.3 HNP-2-FSAR-2, Rev. 26 (9/08).

This design input document is used to develop the response spectrum for the vault piping.

3.4 HNP-2-FSAR-3, Rev. 30 (9/12).

This design input document is used to develop the response spectrum for the vault piping.

3.5 ASME B&PV Code, Section Ill, 1971 Edition.

This design input document defines the acceptance criteria and provides material properties for the metallic piping.

3.6 USAS B31.1.0-1967 and ASME B31.1-2012.

These design input documents provides material properties for the metallic piping.

3.7 ISCO Product Catalog, Version 4.1, 2013.

The HOPE pipe section dimensional properties and weight are taken from this document.

3.8 Flowable Fill Website, http://flowablefill.org/performance.html The typical maximum and minimum densities for flowable fill are taken from this design input document.

3.9 Work Procedure ES-ESST-001, Version 2.0, "Piping Stress Analysis."

This design input document defines the applicable code for the metallic piping.

3.10 Drawing H-26051, Version 51.0, "Edwin I. Hatch Nuclear Plan Unit No. 2 Reactor Building-Plant Service Water System P & I. D. Sht 2 of 2."

3.11 Calculation SCNH-15-029, Version 1, "Soil Bearing Evaluation for New Unit 2, Div. II PSW Transition Vault."

This calculation documents the non-repeated anchor movement for the new vault (heave).

3.12 Drawing H-21111, "Edwin I. Hatch Nuclear Plant Unit No.2 Yard Piping Sheet 3."

This document shows the piping layout.

3.13 Specification No. SN9501, Version 6.0, "Specification for ASME Section Ill, Gate, Globe, Check, Ball, and Butterfly Valves and Replacement Valve Parts."

The valve acceleration limits are taken from this design input document.

3.14 Calculation SMSH-14-011, Version 1, "Stress Analysis of PSW Buried HDPE Piping."

This analysis overlaps with the analysis for the piping discussed in this summary report.

3.15 Calculation SCNH-15-029, Version 1, "Soil Bearing Evaluation for New Unit 2, Div. II PSW Transition Vault."

This design input document determines the potential for settlement or heave of the new vault.

3.16 Drawing H-53444, "Edwin I. Hatch Nuclear Plant Unit No.2 Plant Service Water System Div II Piping Isometric from Yard to Transition Vault."

This document shows the piping layout.

Page7

3.17 Calculation BHO-C-S08-V001-0003, Version 1, Edwin I. Hatch Nuclear Plant, Units 1 and 2, "The Stress Analysis of Underground Piping and Electric Ducts."

This design input document is used to determine the reactor building seismic anchor movements.

3.18 Calculation SMSH-12-020, Version 2.0, "Stress Analysis of Unit 1 Div II Buried Service Water Pipe."

This design input document is used to determine the reactor building seismic anchor movements.

3.19 BH2-C-S23-V012-0001, "Final Seismic Analysis Reactor Building and Internals," April 15, 1975, Volume 1.

This design input document is used to determine the reactor building seismic anchor movements.

3.20 BH2-C-S23-V013-0001, "Final Seismic Analysis Reactor Building and Internals," April 15, 1975, Volume 2.

This design input document is used to determine the reactor building seismic anchor movements.

4.0 Acceptance Criteria The stress analysis calculation followed the acceptance criteria outlined in the lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0 for the HOPE piping. Per Drawing H-21 033, the stainless steel piping has piping class designation HAC. Per Table 3.2-4 of Design Input 3.4, this corresponds to 150# ANSI stainless steel Nuclear Power Piping, ASME Section Ill, Class 3.

Table 3.9-29 of Design Input 3.4 indicates that the design criteria for ASME Class 3 piping is per paragraph 3.9.2.1 of Design Input 3.4. Per Section 6.1 of Design Input 3.9, ASME Section Ill, 1971 applies for Hatch Unit 2 ASME piping (Design Input 3.5). The stainless steel piping is qualified to the requirements of paragraph 3.9.2.1 of Design Input 3.4 and all of the applicable requirements of Design Input 3.5.

Per Section 3.5.3 of Design Input 3.13, all ASME Section Ill gate, globe, check, ball, and butterfly valves shall be demonstrated capable of withstanding a seismic acceleration of 4.5g applied simultaneously in all three orthogonal directions. Therefore, the valves are considered acceptable for acceleration values less than 4.5g.

PageS

5.0 Methodology In order to obtain the required data to meet the acceptance criteria discussed in Section 4.0, several SAP2000 Finite Element Models (FEM) were produced. The models include a portion of soil supported piping in order to sufficiently capture the interaction between the piping inside the vault and the soil supported piping outside of the vault. The soil supported piping is modeled using links which supports the piping while capturing the flexibility of the soil and piping. These links are based on soil and pipe ovaling springs as well as pipe-soil breakaway forces. Both of these parameters dependent on several pipe and soil properties. The modulus of elasticity for the pipe is different for the thermal and seismic load cases and several soil parameters fall within a range. Because of these variations, upper bound and lower bound soil springs and breakaway forces are determined for both the thermal and seismic load cases. The average of the calculated lower and upper bound values is considered as a best estimate. The best estimate is adjusted upward and downward by a coefficient of variation (COV) to calculate lower and upper bound best estimate spring and breakaway force values. A COV of 1 is conservatively used in accordance with Section 11.4.c.iii of Ref. 7.7. This results in factors of 0.5 and 2.0 for the lower and upper bound best estimate cases respectively. The calculated lower bound values are screened with the best estimate lower bound values to determine worst case lower bound values. Similarly, the calculated upper bound values are screened with the best estimate upper bound values to determine worst case upper bound values. This results in four different models; lower bound thermal, upper bound thermal, lower bound seismic, and upper bound seismic. Detailed calculations and resulting values for the springs and break-away forces are shown in Attachment A of Enclosure 5 for the HOPE piping and Attachment A of this enclosure for the stainless steel piping.

Design Input 3.11 has determined that there is no potential for vault settlement but vault heaving has been calculated to be less than 1/8". Therefore, a worst case upward displacement of 1/8" is considered at the vault support location. Since the vault heave is a long-term load condition, it is most appropriate to capture the movement in the thermal load case models. However, it is conservatively considered in all four models and the worst case resulting loads are considered. The loads resulting from the heave load case are secondary loads and are qualified as such.

The FEM include a portion of soil supported piping (buried) in order to sufficiently capture the interaction between the vault piping and the soil supported piping. The soil supported piping on one side of the vault is modeled up to the reactor building where it is considered anchored. The other side is modeled with two changes in direction. This introduces enough flexibility into the analysis to estimate loading from the thermal load cases without being excessively conservative. In order to verify the adequacy of this termination point for seismic loading, fictitious loads are applied to the piping inside the vault and the resulting reactions at the termination point inside the soil are verified as being negligible. The fictitious loads are applied in all directions at each of the transitions between soil supported piping and piping inside the vault. Loads of 5000-lbf have been logically chosen which bound the dead weight of the vault piping, its components, and any seismic response.

Since the loads are applied in all directions, two load cases were created; one for all positive directional loading and one for all negative directional loading. These load cases were screened.

The node point spacing does not define the finite element meshing parameters. These are determined internally by SAP2000. The node point spacing for the soil supported piping is chosen based on discrete soil springs while the node point spacing for the piping that is not soil supported is based on important locations such as support points, lumped weights, branch lines, pipe section changes, and other points of interest for data extraction like valve accelerations.

The node points for the soil supported HOPE piping are spaced in accordance with Appendix B of EPRI Report 1 013549 (Ref. 7 2). The node points for the soil supported stainless steel piping are spaced in accordance with Non-Mandatory Appendix VII to ASME B31.1-2012 (Design Input 3.6).

There are two seismic load cases considered; the operating basis earthquake (OBE) and the design basis earthquake (DBE). The design spectrum for horizontal ground motion is shown in HNP FSAR-3 (Design Input 3.4) Figures 3.7A-1 & 3.7A-20 for the OBE and Figures 3.7A-2 & 3.7A-21 for the DBE. Points are chosen from these figures to create bounding Response Spectrum functions in SAP2000. Per Sections 2 5.2.1 0 & 2 5.2.11 of HNP-2-FSAR-2 (Design Input 3.3), the vertical accelerations are two-thirds of the horizontal accelerations. Therefore, each vertical response is taken as two-thirds of its horizontal response. The modal response was combined using the NRC Page9

Ten-Percent Method which is in accordance with Section 3.7A.2.1.1 of HNP-2-FSAR-3. The responses in each direction were combined by the SRSS Method which is in accordance with Section 3.7A.3.7 of HNP-2-FSAR-3. All modes with frequencies lower than 33-Hz were considered in the models. This is consistent with Section 3.7A.2.1.1 of HNP-2-FSAR-3.

The seismic response on the soil supported portion of the piping that is included in the models is captured by modeling an equivalent differential temperature. This is the same methodology outlined in Section 4.5 of EPRI Report 1013549 (Ref. 7.2). This methodology involves calculating the controlling strain induced on the pipe due to seismic wave passage. The seismic wave induced strain is limited by breakaway between the pipe and the soil. The controlling strain is converted into an equivalent differential temperature which is input into SAP2000 for the seismic loading.

SAP2000 does not have the capability to limit which parts of the model the response spectrum load cases are applied. Because of this, the response spectrum load cases are applied to the entire piping system; including the soil supported piping. This is conservative since the soil supported piping seismic response is already accounted for by using the equivalent differential temperatures previously discussed. The additional seismic loading should not be excessively conservative as the soil supported portion of the piping is expected to be in the rigid range providing minimal contribution to the modal response. The seismic response from the soil supported piping is absolutely added to the design spectrum response to produce a conservative response for the entire system.

The models are terminated at the Reactor Building wall where the Reactor Building seismic anchor movements are applied. The seismic anchor movement response is taken as the square root sum of the squares (SRSS) of the response in all three directions. The seismic response from the equivalent temperatures and the response spectrum is combined with the response due to the seismic anchor movements by SRSS.

The relative displacements between the valve vault and the surrounding soil are considered negligible since the valve vault is relatively small and the surrounding soil has a high compaction level (minimum 95% of the maximum dry density compaction criterion per HNP-2-FSAR-2, Section 2.5.1.2.8).

There are three thermal load cases considered; pipe at 32°F, pipe at 97°F, and pipe at 125°F.

These are modeled as temperature differentials from 70°F; that is a change of -38°F, 27°F, and 55°F. The resulting loads for the 32°F and 125°F cases (minimum operating temperature and design temperature rounded up from 123°F) are combined by absolute sum in order to capture the stress range which is used to qualify pipe stresses. Using the design temperature for the stress range instead of the maximum operating temperature of 97°F is conservative. The 97°F load case (maximum operating temperature) is used for support and flange qualifications only.

6.0 Assumptions & Design Considerations 6.1 Design Consideration:

The flowable fill is conservatively considered to be used for all of the backfill (all the way to grade).

Justification:

This design consideration does not require verification as the backfill properties incorporated into the detailed design through DCP SNC591628 will be bounded by the properties used in the evaluation.

Page 10

7.0 References 7.1 PPI, "Handbook of Polyethylene Pipe", 2nd Ed. with 6/6/12 Errata.

7.2 EPRI Report 1013549, "Nondestructive Evaluation: Seismic Design Criteria for Polyethylene Pipe Replacement Code Case," Technical Update, September 2006.

7.3 American Lifelines Alliance, "Guidelines for the Design of Buried Steel Pipe," July 2001 with Addenda through February 2005.

7.4 McGrath, T. J. and Hoopes, R. J.," 'Bedding Factors and E' Values for Buried Pipe Installations Backfilled with Air-Modified CLSM," The Design and Application of Controlled Low-Strength Materials (Fiowable Fill), ASTM STP 1331, A. K. Howard and J. L. Hitch, Eds.,

American Society for Testing and Materials, 1998.

7.5 Das, Braja M., "Principles of Foundation Engineering," 6th Ed.

7.6 CSI Knowledge Base Website, "Damping in Response-Spectrum Analysis,"

https://wiki.csiamerica.com/display/kb/Damping+in+response-spectrum+analysis 7.7 NUREG-0800, Section 3.7.2, "Seismic System Analysis."

7.8 Calculation SMSH-15-002, Version 1, "Stress Analysis for U2 DIV II PSW Subgrade Vault Piping."

7.9 Calculation SMNH-15-008, Version 1, "14" HOPE to Metallic Transition Piping Flanged Joint Analysis."

Page 11

Attachment A - Development of Spring & Breakaway Force Values This attachment develops the spring and breakaway force values for the stainless steel piping that are input into the SAP2000 models. The spring and breakaway force calculations and values for the HOPE piping are shown in Attachment A of Enclosure 5.

The piping layout is shown in Design Input 3.12.

ELground := 129-ft ELmax:= 125.4167ft ELmin:= 120.5833ti Ess.seis := 28300 ksi Ess.th := 28300 ksi 0 55 := 10.75in t55 := 0.365 in Di.ss := Dss - 2*tss = 10.02:in Grade Elevation Pipe Centerline Maximum Elevation Pipe Centerline Minimum Elevation Modulus of Elasticity of Pipe for Seismic Evaluations (70°F)

Modulus of Elasticity of Pipe for Thermal Evaluations (70°F}

Outside Diameter of Pipe Pipe Wall Thickness Inside Diameter of 10" Stainless Steel Pipe 4

4 Dss - 0 i.ss

. 4 lr.ss := 71*

= 160.734m Moment of Inertia of 10" Stainless Steel Pipe 64 The maximum and minimum soil densities are taken as the upper and lower bound values for the typical range of weight for flowable fill (Design Input 3.8). The upper bound value used is considered to include any soil saturation affects. The flowable fill is conservatively considered to be used for all of the backfill (all the way to grade) (See 6.2).

Pmax := 145*pcf Pmin := 70*pcf Maximum Density of Soil Above Pipe Minimum Density of Soil Above Pipe The modulus of soil reaction for the backfill (flowable fill) is taken from Table 4 of Ref. 7.4 considering a minimum age of 28 days. Note that this value is similar to that of Table 3-8 of Ref. 7.1 for coarse-grain soil at a depth of cover ranging from 5 to 10 feet. The pipe has a cover of at least 5 feet for the entire routing.

E' := 3000* psi Modulus of Soil Reaction for Backfill The methodology outlined in Appendix 8 of Ref. 7.3 is followed for developing soil spring values and spring spacing.

The soil properties range from a lower bound to an upper bound (soil density for example).

Therefore, each load condition will have a lower bound evaluation and an upper bound evaluation.

This results in four sets of spring values: seismic load condition lower bound, seismic load condition upper bound, thermal load condition lower bound, and thermal load condition upper bound.

The pipe ovaling springs are considered negligible for the stainless steel piping due to its stiffness.

Page A-1

Attachment A - Development of Spring & Breakaway Force Values Lower Bound Axial Soil Spring:

The flowable fill is considered to behave like dense sand for determination of soil springs.

The minimum ground cover is used which, as a worst case, produces the smallest stiffness.

ELP := ELm ax = 125.417

ft H := ELground-ELP = 3.583*ft

<P := 40*deg Pipe Elevation (To Pipe Centerline)

Height of Ground Cover (Depth to Pipe Centerline)

Best Estimate Angle of Internal Friction (Ref. 7.5, Table 1.8)

The coefficient of soil pressure at rest is calculated using Jaky's simplified equation from Section 7.2 of Ref. 7.5. This provides a reasonable empirical approximation of the true value.

K0 := I - sin( <P) = 0.357 Coefficient of Soil Pressure at Rest (Ref. 7.5, Section 7.2)

Per Table B.1 of Ref. 7.3, the friction angle pipe-soil, o, ranges between 0.7 and 0.8 times the internal angle of friction for steel pipe coatings. The smaller the friction angle pipe-soil, the smaller the spring value. Therefore, 0.7 times the internal angle of friction is considered for the lower bound spring values o := 0.7*<P = 28*deg The minimum soil density is used for all of the lower bound spring calculations since this produces the smallest spring values.

I + Ko lbf T 1 := 'Tr*D *H*p **--*tan(&)= 21.226*-

u.

ss mm 2

in Max. Axial Soil Force (Breakaway Force)

(Per Eq. B-1 of Ref. 7.3, Sand: c=O)

The axial displacement that corresponds to the maximum soil force is defined in Section B.1 of Ref. 7.3.

Dense sand behavior is considered for the flowable fill.

~~ := O.l *in Displacement at Max. Soil Force Tu.l 6

Ka 1 :=- = 212.2 l *ps1

~~

Axial Soil Spring Page A-2

Attachment A-Development of Spring & Breakaway Force Values Lower Bound Transverse Soil Spring:

Horizontal Bearing Capacity Factor (Ref. 7.3, Section B.2) for ljl=40°:

Nqh := 10.959 + 1.783*(...!:!._) + 0.045*(...!:!._)

2

- 0.005425*(...!:!._)

3

- 0.0001153*(...!:!._)

4

= 18.434 0ss 0ss 0ss 0ss Max. Soil Force (Per Eq. B-2 of Ref. 7.3, Sand: c = 0)

The transverse displacement that corresponds to the maximum soil force is defined in Section B.2 of Ref.

7.3 as being 0.04(H+D/2) but less than 0.10D to 0.15D. The larger the displacement, the smaller the spring value. Therefore, the 0.15D is considered as an upward limit.

~p := mi{0.04{ H + D; 5).0.15*D55] = 1.613*in Pu.l Kt 1 :=- = 214.068*psl 0

~p Lower Bound Vertical Soil Spring:

2 Nq := exp(7T*tan(<j>))*tan(45*deg + ~) = 64.195 N1 := exp(O.I8*_!_- 2.5) = 109.947 deg

~qd := 0.1* 0 55 = 1.075 *in Qd.l Kd 1 := -- = 1357.601*psl 6-qd

. (

<f>*H

)

Nqv := mm

, Nq = 3.636 44*deg*D55 lbf Ou.l := Nqv'Pmin*H*D55 = 68.091*"Tr; Displacement at Max. Soil Force Transverse Soil Spring Downward Bearing Capacity Factor 1 (Ref. 7.3, Section B.4)

Downward Bearing Capacity Factor 2 (Ref. 7.3, Section B.4)

Max. Downward Soil Force (Eq. B-4 of Ref. 7.3)

Displacement at Max. Downward Soil Force (Ref. 7.3, Section B.4, Granular Soils)

Vertical Down Soil Spring Vertical Uplift Factor (Ref. 7.3, Section B.3)

Max. Upward Soil Force (Eq. B-3 or Ref. 7.3)

The upward displacement that corresponds to the maximum upward soil force is defined in Section B.3 of Ref. 7.3 as ranging between 0.01 Hand 0.02H but less than 0.1 D for sands. The larger the displacement, the smaller the spring value. Therefore, the 0.02H is considered as an upward limit.

~qu := min( 0.02* H, 0.1* 0 55) = 0.86* in Displacement at Max. Upward Soil Force Ou.l 6

Ku 1 := -- = 79.17 *psi

~qu Vertical Up Soil Spring Page A-3

Attachment A - Development of Spring & Breakaway Force Values As recommended in Ref. 7.2, Appendix B, the total vertical soil spring is taken as the average of the vertical up and down. Similarly, the breakaway force is taken as the average of the vertical up and down.

Ku.l + Kd.l Kv.l :=

= 718.389*psl 2

Vertical Soil Spring Oct.l + Ou.l lbf Ov 1 :=

= 763.756*-.-

2 In Vertical Breakaway Force Upper Bound Axial Soil Spring:

The maximum ground cover is used which, as a worst case, produces the largest stiffness.

ELP := ELmin = 120.583*ft Pipe Elevation (To Pipe Centerline)

H := ELground - ELP = 8.417

tl Height of Ground Cover (Depth to Pipe Centerline)

Per Table B.1 of Ref. 7.3, the friction angle pipe-soil, iS, ranges between 0.7 and 0.8 times the internal angle of friction for steel pipe coatings. The larger the friction angle pipe-soil, the larger the spring value. Therefore, 0.8 times the internal angle of friction is considered for the lower bound spring values o := 0.8* 4> = 32* deg The maxium soil density is used for all of the lower bound spring calculations since this produces the largest spring values.

I + K0 lbf Tu.u := 71'*055 H*Pmax* *tan(o) = 121.37*ir;"

Max. Axial Soil Force (Breakaway Force)

(Per Eq. B-1 of Ref. 7.3, Sand: c=O)

The axial displacement that corresponds to the maximum soil force is defined in Section B.1 of Ref. 7.3.

Dense sand behavior is considered for the flowable fill.

~t

-=O.I

in Displacement at Max. Soil Force Tu.u Ka u := -- = 1213.704*psl 0

~~

Axial Soil Spring Upper Bound Transverse Soil Spring:

Horizontal Bearing Capacity Factor (Ref. 7.3, Section B.2) for cjl=40:

Nqh := 10.959 + 1.783*(_!:!...] + 0.045*(_!:!...JZ- 0.005425*(_!:!...]

3

- 0.0001153*(_!:!...J 4

= 26.286 0 ss 0 ss 0 ss 0 ss lbf Pu.u := Nqh' Pmax*H*D55 = 2394.82*ir;"

Max. Soil Force (Per Eq. B-2 of Ref. 7.3, Sand: c = 0)

The transverse displacement that corresponds to the maximum soil force is defined in Section B.2 of Ref.

7.3 as being 0.04(H+D/2) but less than 0.1 OD to 0.15D. The smaller the displacement, the larger the spring value. Therefore, the 0.10D is considered as an upward limit.

~P := mi{o.o4{H +

0 5J.o.IO*D55] = 1.075*in Displacement at Max. Soil Force p

K1 u := ~

= 2227.739*psi 0

~p Transverse Soil Spring Page A-4

Attachment A-Development of Spring & Breakaway Force Values Upper Bound Vertical Soil Spring:

'J Nq := exp('IT*tan(<jJ))*tan(45*deg + ~r = 64.195 N1 := exp(O.I8*_!_- 2.5) = 109.947 deg Downward Bearing Capacity Factor 1 (Ref. 7.3, Section B.4)

Downward Bearing Capacity Factor 2 (Ref. 7.3, Section B.4) 2 0ss lbf Qd.u := Nq*Pmax*H*D55 + N,(Pmax*-

2

= 6381.771 *1; Max. Downward Soil Force (Eq. B-4 of Ref. 7.3) 6qd := 0.1*055 = 1.075*in K

Od.u 3

d U := -- = 5936.5 l *pSl 6qd

. (

<P*H

)

Nqv := mm

, Nq = 8.)41 44*deg*D55 lbf Ou.u := Nqv*Pmax*H*D55 = 778.176*-.-

m Displacement at Max. Downward Soil Force (Ref. 7.3, Section B.4, Granular Soils)

Vertical Down Soil Spring Vertical Uplift Factor (Ref. 7.3, Section B.3)

Max. Upward Soil Force (Eq. B-3 or Ref. 7.3)

The upward displacement that corresponds to the maximum upward soil force is defined in Section B.3 of Ref. 7.3 as ranging between 0.01 Hand 0.02H but less than 0.10 for sands. The smaller the displacement, the larger the spring value. Therefore, the 0.01 His considered as an upward limit.

6qu := min(O.OI*II,O.I*D55) = I.OI*in Displacement at Max. Upward Soil Force Ou.u K

= -- = 770.468*psl u.u 6qu Vertical Up Soil Spring As recommended in Ref. 7.2, Appendix B, the total vertical soil spring is taken as the average of the vertical up and down. Similarly, the breakaway force is taken as the average of the vertical up and down.

K

+ Kd K

u.u

.u 33.

V.U :=

=

).)* pSI 2

Vertical Soil Spring Q

._ Od.u + Ou.u _ 3 _ 9 9

~

vu *-

)7. 73*.

2 In Vertical Breakaway Force Page A-5

Attachment A - Development of Spring & Breakaway Force Values Influence Length:

The influence length for the buried steel piping is based on the methodology outlined in Non-Mandatory Appendix VII to ASME B31.1-2012.

H Nh := 0.28:.-,*- + 4.3 = 6.978 0ss Seismic Loading (Lower Bound):

[

k

]0.25 hi I

f11s :=

= 0.0 11*-:-

4* Ess.seis' 1P.ss m

3 *'IT L1.15 := -- = 219.159*in 4 *~1s Seismic Loading (Upper Bound):

[

k

]0.25 hu I

f1us :=

= 0.013*-:-

4* Ess.seis' 1P.ss m

1.71 Ll.us := ---- = 182.68*in 4 *~us Horizontal Force Factor Horizontal Stiffness Factor (Considered Compacted Soil)

Lower Bound Horizontal Modulus of Subgrade Thermal Loading (Lower Bound):

f31th := [

khl

]0.25 = 0.011*~

4*Ess.th' 1P.ss m

3*71 Ll.lth := -- = 219.159*in 4* (31th Upper Bound Horizontal Modulus of Subgrade Thermal Loading (Upper Bound):

f1uth := [

khu

]0.25 = 0.013 *~

4*Ess.th'1P.ss m

3*71 Ll.uth := -- = 182.68*in 4* f1uth Controlling Influence Length (Stainless Steel)

Page A-6

Attachment A-Development of Spring & Breakaway Force Values Average of Lower and Upper Spring Values & Lower and Upper Breakaway Forces:

Kt.a := 0.5*{Kt.J + Kt.u) = 1220.904*psi Ka.a := 0.5*{Ka.l + Ka.u) = 712.982*psi Kv.a := 0.5*{Kv.l + Kv.u) = 2035.944*psi 1/2 x Average Spring Values:

Kt.al := 0.5*Kt.a = 610.452*psi Ka.al := 0.5*Ka.a = 356.491*psi Kv.al := 0.5*Kv.a = 1017.972*psi 1/2 x Average Breakaway Forces:

lbf ft. a I := O.:l* ft.a = 68;,.00 l*in""

9 lbf fa al := O.:l*fa a= h64 *-.-

In lbf fv al := O.:l*lv a = I 08;,.932*-.-

In lbf ft.a := 0.5*{Pu.l + Pu.u) = 1370.002*in""

lbf fa.a := 0.5*{T u.l + T u.u) = 71.298*in lbf fv.a := 0.5*{ Ov.l + Ov.u) = 2171.865*in 2 x Average Spring Values:

Kt.au := 2* Kt.a = 2441.808* psi Ka.au := 2*Ka.a = 1425.964*psi Kv.au := 2*Kv.a = 4071.888*psi 2 x Average Breakaway Forces:

f.

- ?,.

- 2 40 oo* ~

t.au.- -* t.a -

7

. :l* in fa au:= 2* t~ a= 142.596* l~f In t~.au := 2* f~.a = 4343. 729* ~~:

Page A-7

Attachment A - Development of Spring & Breakaway Force Values Controlling Lower Bound Springs & Breakaway Forces for Models (for Stainless Steel Piping):

t.

. (

)

6 lbf a.bl := mm Tu.l* 1a.al = 21.22 'i; t~.bl := min(Ov.l* rv.al) = 763.756* ~~:*

. (

. )

- lbf 1t bl := mm pu I* tt al = 34:>.18:>*-.-

In Lower Bound Axial Soil Spring Lower Bound Vertical Soil Spring Lower Bound Transverse Soil Spring Lower Bound Axial Breakaway Force Lower Bound Vertical Breakaway Force Lower Bound Transverse Breakaway Force Note: The breakaway displacements for the models are determined by dividing the breakaway force value by its corresponding spring value. These values aren't shown here due to the simplicity of calculating them from the reported values documented herein.

Controlling Upper Bound Springs & Breakaway Forces for Models (for Stainless Steel Piping):

Ka.bu := max(Ka.u*Ka.au) = 1425.964*psi Kv.bu := max(Kv.u*Kv.au) = 4071.888*psi lbf t~.bu := max(T u.u.fa. au) = 142.596*i; lbf

~'v.bu := max(Ov.u*t~.au) = 4343.729*i;

(

)

- lbf tt.bu :=max Pu.u*'t.au = 2740.00:>*i; Upper Bound Axial Soil Spring Upper Bound Vertical Soil Spring Upper Bound Transverse Soil Spring Upper Bound Axial Breakaway Force Upper Bound Vertical Breakaway Force Upper Bound Transverse Breakaway Force Note: The breakaway displacements for the models are determined by dividing the breakaway force value by its corresponding spring value. These values aren't shown here due to the simplicity of calculating them from the reported values documented herein.

Controlling Influence Length:

Lf3.ss = 219.159*in The spring values and breakaway forces above are dependent on their effective lengths. Therefore, the value modeled will vary from node to node.

Page A-8

Edwin I. Hatch Nuclear Plant-Unit 2 Proposed In service Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0 Final Stress Analysis Design of the HOPE System Summary of Stress Analysis Calculation for Existing U2 Div II PSW Subgrade Vault Piping

Summary Report on Stress Analysis for Existing U2 DIV II PSW Subgrade Vault Piping Hatch Nuclear Plant-Unit 2 1.0 Purpose of Summary Report This report summarizes the stress analysis calculation, SMSH-15-002, that has been prepared to support SNC's lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0 (ATR). This lSI Alternative is needed to support the planned replacement of buried steel piping in Hatch Nuclear Plant Unit 2, Plant Service Water (PSW) system with High Density Polyethylene (HOPE) piping.

The piping to be replaced is the supply piping from the Unit 2 Service Water Valve Pit 28 to a new subgrade vault located outside of the Unit 2 Reactor Building. Stress analysis calculation SMSH 002 evaluates the replacement piping inside the existing subgrade vault to the design requirements in the A TR. Note that only specific sections of the A TR apply for the piping evaluated by the stress analysis since the piping being qualified is inside the valve vault (i.e. no surcharge loads, no ring deflection, etc.). The stress analysis also evaluates the metallic piping inside the vault and captures the mutual influence between the replacement HOPE piping and the metallic piping.

All computations from the calculation are not included in the summary. Calculation results and conclusions are shown in Section 2.0.

2.0 Results and Conclusions Following the methodology outlined in Section 5.0, the replacement HOPE piping meets all of the acceptance criteria outlined in the lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0. The metallic piping meets all of the applicable requirements of paragraph 3.9.2.1 of Design Input 3.4 and all of the applicable requirements of Design Input 3.5. The valve acceleration values are less than 4.5g in each orthogonal direction and are therefore acceptable per Section 3.5.3 of Design Input 3.13.

Detailed results are shown in this section. All margin factor values are greater than 1.0 and are therefore acceptable. The controlling margin factor was determined to be 1.32 which was for the 1 0" carbon steel piping, secondary stresses for the upper bound spring condition.

The loads on the flanged connection between the metallic piping and the HOPE piping are reported in this section. These loads are evaluated in Calculation SMNH-15-008 (Ref. 7.9).

The piping has been analyzed to the following conditions per Section 11 00 of Design Input 3.1. A design temperature value of 125°F was conservatively used which bounds the actual design temperature of 123°F.

Condition Temperature, °F Pressure, psig Normal Operating 95 140 Maximum Operating 97 190 Design 123 180 Page 1

Summary of SAP2000 Output A summary of the controlling pipe member forces for the piping inside the vault is provided below.

HOPE Vault 1 0" Carbon Steel 1 0" Stainless Steel 2" Stainless Steel Model Load Case Piping Vault Piping Vault Piping Vault Piping Straight Pipe p (lb) M (lb-in) p (lb)

M (lb-in) p (lb)

M (lb-in)

FULLOEAO 3

3429 10 60332 726 20682 57 0

Seismic OBE 111 3274 276 26999 277 22061 1

84 Lower Bound OBE 187 4694 384 40467 444 31383 2

120 FULLOEAO 12 2325 20 60841 744 20530 57 0

Seismic OBE 265 2626 195 19569 246 13615 2

69 Upper Bound OBE 475 3816 282 33346 450 21347 4

114 Thermal FULLOEAO 2

2592 9

61261 741 21052 57 0

Lower Bound THERMAL ABS 3332 5734 21 211728 3332 145127 0

0 Thermal FULLOEAO 11 1711 18 61692 755 20932 57 0

Upper Bound THERMAL ABS 10039 12022 494 574955 10039 275935 0

0 FULLOEAO 12 3429 20 61692 755 21052 57 0

Bounding OBE 265 3274 276 26999 277 22061 2

84 Values OBE 475 4694 384 40467 450 31383 4

120 THERMAL ABS 10039 12022 494 574955 10039 275935 0

0 A summary of the controlling pipe member forces for the soil supported piping IS provided below.

HOPE Straight Pipe HOPE Elbows Soil Model Load Case Soil Supported Supported Piping p (lb)

M (lb-in) p (lb)

M (lb-in)

FULLOEAO 21 3545 18 1277 Seismic OBE 254 1345 132 2945 Lower Bound OBE 473 2069 242 4811 FULLDEAD 19 3487 11 838 Seismic OBE 917 1184 426 3523 Upper Bound OBE 1720 1883 799 6603 Thermal FULLDEAO 20 3519 15 1165 Lower Bound THERMAL ABS 5788 12416 2811 49975 Thermal FULLDEAD 19 3471 13 735 Upper Bound THERMAL ABS 17474 8724 8768 60492 FULLDEAD 21 3545 18 1277 Bounding OBE 917 1345 426 3523 Values OBE 1720 2069 799 6603 THERMAL ABS 17474 12416 8768 60492 A summary of the valve accelerations for the 1 0" stainless steel valve and the 2" stainless steel valve is provided below.

10" Valve Accelerations 2" Valve Accelerations Load Case U2 (g)

U3 (g)

Max. (g)

U1 (g)

U2 (g)

U3 (g)

Max (g)

U1 (g)

OBE Lower Bound 0.223 0.164 0.085 0.223 0.205 0.14 0.018 0.205 OBE Lower Bound 0.313 0.243 0.128 0.313 0.284 0.212 0.027 0.284 OBE Upper Bound 0.217 0.118 0.13 0.217 0.18 0.099 0.042 0.18 DBE Upper Bound 0.324 0.208 0.227 0.324 0.291 0.182 0.077 0.291 OBE Bounding Values 0.223 0.164 0 13 0.223 0.205 0.14 0.042 0.205 DBE Bounding Values 0.324 0.243 0.227 0.324 0.291 0.212 0.077 0.291 Page2

In order to verify that an adequate amount of soil supported piping is modeled, fictitious boundary check loads are applied to the vault piping and the resulting loads at the soil supported piping termination point are verified to be negligible as discussed in Section 5.0. A summary of the joint forces resulting from the boundary check fictitious loads at the model termination point is provided below. All of the loads are either zero or negligible indicating adequate modeling of the soil supported piping.

Node 1120 Model (Model Termination Point)

FX (lb)

FY (lb)

FZ (lb)

MX (lb-in)

MY (lb-in)

MZ (lb-in)

Seismic Lower Bound 0.00 0.29 0.00 0.00 0.00 0.00 Seismic Upper Bound 0.00 0 00 0.00 0.00 0.00 0.00 Thermal Lower Bound 0.00 0.07 0.00 0.00 0.00 0.00 Thermal Upper Bound 0.00 0.00 0.00 0.00 0.00 0.00 Bounding Values 0.00 0.29 0.00 0.00 0.00 0.00 A summary of the loads on the flanged interface between the HOPE piping and the metallic piping is provided in the following table. THERMAL_MIN is based on the minimum operating temperature of 32°F, THERMAL_MAXOP is based on the maximum operating temperature of 97°F, and THERMAL_MAX is based on the design temperature which is conservatively taken as 125°F.

Node 1010 Model Load Case (Flanged Connection Between HOPE & S.S.)

FX (lb)

FY (lb)

FZ (lb)

MX (lb-in)

MY (lb-in)

MZ (lb-in)

Seismic FULLOEAO 10 3

417 3420 64 243 Lower OBE 59 131 7

270 194 3257 Bound OBE 83 220 11 392 282 4669 Seismic FULLOEAO 20 12 385 2286 189 378 Upper OBE 97 265 12 356 148 2598 Bound OBE 143 475 18 552 219 3769 FULLOEAO 9

2 396 2582 8

232 Thermal THERMAL MIN 12 1948 40 1162 293 2721 Lower THERMAL MAXOP 9

1384 29 826 208 1934 Bound THERMAL MAX 13 2644 58 1657 426 4242 FULLOEAO 18 11 364 1670 91 360 Thermal THERMAL MIN 289 5869 76 2299 195 6639 Upper THERMAL MAXOP 205 4170 54 1633 138 4717 Bound THERMAL MAX 419 8497 111 3329 283 9611 FULLOEAO 20 12 417 3420 189 378 OBE 97 265 12 356 194 3257 Bounding OBE 143 475 18 552 282 4669 Values THERMAL MIN 289 5869 76 2299 293 6639 THERMAL MAXOP 205 4170 54 1633 208 4717 THERMAL_MAX 419 8497 111 3329 426 9611 Page3

A summary of the loads and pipe movements on the vault support is provided in the following table.

THERMAL_MIN is based on the minimum operating temperature of 32°F, THERMAL_MAXOP is based on the maximum operating temperature of 97°F, and THERMAL_MAX is based on the design temperature which is conservatively taken as 125°F.

Node3110 Model Load Case (Support Inside Vault)

FZ (lb)

OX (in)

DY (in)

Seismic FULLDEAD 1268.11 0.0082 0.0006 Lower OBE 134.50 0.0299 0.0062 Bound DBE 212.24 0.0424 0.0095 Seismic FULLDEAD 1218.71 0.0058 0.0002 Upper OBE 129.31 0.0125 0.0043 Bound DBE 227.22 0.0179 0.0076 FULLDEAD 1232.73 0.0085 0.0006 Thermal THERMAL MIN 820.20 0.0555 0.0026 Lower THERMAL_MAXOP 582.76 0.0394 0.0018 Bound THERMAL MAX 1126.64 0.0788 0.0009 FULLDEAD 1185.74 0.0065 0.0003 Thermal THERMAL_MIN 2227.58 0.0507 0.0630 Upper THERMAL MAXOP 1582.76 0.0360 0.0447 Bound THERMAL MAX 3224.93 0.0733 0.0912 FULLDEAD 1268.11 0.0085 0.0006 OBE 134.50 0.0299 0.0062 Bounding DBE 227.22 0.0424 0.0095 Values THERMAL MIN 2227.58 0.0555 0.0630 THERMAL_MAXOP 1582.76 0.0394 0.0447 THERMAL_MAX 3224.93 0.0788 0.0912 Page4

HOPE Vault Piping Stresses Service level stresses for the HOPE vault piping are summarized in the following table.

Service Spring Pipe Stress Stress Factor x Margin Stress Level Condition (psi)

Factor Allowable Stress (psi)

Factor A

Lower Bound 332 1

561 1.69 Upper Bound 327 1

561 1.72 B

Lower Bound 368 1.1 617.1 1.68 Upper Bound 361 1.1 617.1 1.71 0

Lower Bound 376 1.33 746.13 1.99 Upper Bound 370 1.33 746.13 2.02 Seismic-induced stresses for the HOPE vault piping are summarized in the following table.

Spring Pipe Stress Allowable Stress Margin Condition (psi)

Range (psi)

Factor Lower Bound 52 2032 39.02 Upper Bound 51 2032 39.93 Thermal expansion and contraction stresses for the HOPE vault piping are summarized in the following table.

Spring Pipe Stress Allowable Stress Margin Condition (psi)

Range (psi)

Factor Lower Bound 73 2032 27.85 Upper Bound 193 2032 10.50 Metallic Vault Piping Stresses Primary stresses for the metallic vault piping due to Normal Conditions are summarized in the following table.

Component Spring Load Case Pipe Stress Allowable Margin (psi)

Stress (psi)

Factor 10" Carbon Lower Bound 5261 15000 2.85 Steel Piping Upper Bound 5289 15000 2.84 1 0" Stainless Lower Bound 2634 16600 6.30 Steel Piping Upper Bound 2626 16600 6.32 2" Stainless Lower Bound 593 16600 27.98 Steel Piping Upper Bound 593 16600 27.98 Primary stresses for the metallic vault piping due to Upset Conditions are summarized in the following table.

Component Spring Load Case Pipe Stress Allowable Margin (psi)

Stress (psi)

Factor 10" Carbon Lower Bound 7025 18000 2.56 Steel Piping Upper Bound 6568 18000 2.74 1 0" Stainless Lower Bound 4075 19920 4.89 Steel Piping Upper Bound 3515 19920 5.67 2" Stainless Lower Bound 885 19920 22.51 Steel Piping Upper Bound 835 19920 23.86 Page 5

Primary stresses for the metallic vault piping due to Faulted Conditions are summarized in the following table.

Component Spring Load Case Pipe Stress Allowable Margin (psi)

Stress (psi)

Factor 10" Carbon Lower Bound 7905 36000 4.55 Steel Piping Upper Bound 7468 36000 4.82 1 0" Stainless Lower Bound 4684 39840 8.51 Steel Piping Upper Bound 4021 39840 9.91 2" Stainless Lower Bound 1010 39840 39.44 Steel Piping Upper Bound 991 39840 40.22 Secondary stresses for the metallic vault piping are summarized in the following table. There are two combinations that may be used per Section 3.9.2.1.B of Design Input 3.4. The design pressure plus weight plus thermal expansion and OBE anchor displacement stresses option is considered.

There are no OBE anchor displacement stresses as discussed in Section 5.0.

Component Spring Load Case Pipe Stress Allowable Margin (psi)

Stress (psi)

Factor 10" Carbon Lower Bound 23707 37500 1.58 Steel Piping Upper Bound 28316 37500 1.32 1 0" Stainless Lower Bound 15277 41500 2.72 Steel Piping Upper Bound 26665 41500 1.56 2" Stainless Lower Bound 593 41500 69.94 Steel Piping Upper Bound 593 41500 69.94 Valve Accelerations Valves are capable of withstanding 4.5g in all directions as discussed in Section 4.0. Valve accelerations for both valves are less than 4.5g in all directions and are therefore acceptable.

Model Termination & Overlap Reconciliation The model terminations as described in Section 5.0 is adequate since the resulting loads at the termination point due to the fictitious boundary check load are negligible.

The models in analysis SMSH-14-011 (Design Input 3.14) are terminated using a fictitious anchor at the vault support. This was theorized as being conservative as discussed in Attachment A of SMSH-14-011. Loads evaluated in SMSH-14-011 are bounding when compared to the loads from SMSH-15-002 (summarized in this section). This confirms the model termination in SMSH-14-011 as being conservative.

Other Design Considerations As noted in Enclosure 2 of Design Input 3.1, other design considerations will be addressed under SNC design procedures in accordance with the existing design and license basis for HNP.

Page6

3.0 Design Inputs 3.1 Edwin I. Hatch Nuclear Plant-Unit 2, Proposed lnservice Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0.

The acceptance criteria and general methodology was taken from this design input document.

3.2 HNP-1-FSAR-12, Rev. 19 (7/01).

This design input document is used to develop the response spectrum for the vault piping.

3.3 HNP-2-FSAR-2, Rev. 26 (9/08).

This design input document is used to develop the response spectrum for the vault piping.

3.4 HNP-2-FSAR-3, Rev. 30 (9/12).

This design input document is used to develop the response spectrum for the vault piping.

3.5 ASME B&PV Code, Section Ill, 1971 Edition.

This design input document defines the acceptance criteria and provides material properties for the metallic piping.

3.6 USAS B31.1.0-1967 & ASME B31.1-2012.

These design input documents provides material properties for the metallic piping.

3.7 ISCO Product Catalog, Version 4.1, 2013.

The HOPE pipe section dimensional properties and weight are taken from this document.

3.8 Flowable Fill Website, http://flowablefill.org/performance.html The typical maximum and minimum densities for flowable fill are taken from this design input document.

3.9 Work Procedure ES-ESST-001, Version 2.0, "Piping Stress Analysis."

This design input document defines the applicable code for the metallic piping.

3.10 Drawing H-26051, Version 51.0, "Edwin I. Hatch Nuclear Plan Unit No. 2 Reactor Building-Plant Service Water System P & I. D. Sht 2 of 2."

3.11 Calculation SCNH-15-029, Version 1, "Soil Bearing Evaluation for New Unit 2, Div. II PSW Transition Vault."

This calculation documents the non-repeated anchor movement for the new vault (heave).

3.12 Drawing H-11146, "Edwin I. Hatch Nuclear Plant Unit No. 1 Yard Piping-Service Water Pump Structure to Building."

This document shows the piping layout.

3.13 Specification No. SN9501, Version 6.0, "Specification for ASME Section Ill, Gate, Globe, Check, Ball, and Butterfly Valves and Replacement Valve Parts."

The valve acceleration limits are taken from this design input document.

3.14 Calculation SMSH-14-011, Version 1, "Stress Analysis of PSW Buned HOPE Piping."

This analysis overlaps with the analysis for the piping discussed in this summary report.

3.15 Calculation SCNH-15-029, Version 1, "Soil Bearing Evaluation for New Unit 2, Div. II PSW Transition Vault."

This design input document determines the potential for settlement or heave of the new vault.

3.16 Drawing H-53442, "Edwin I. Hatch Nuclear Plant Unit No.2 Plant Service Water System Div II Piping Isometric from PSW Valve Vault to Yard."

This document shows the piping layout.

Page 7

3.17 Calculation BHO-C-S08-V001-0003, Version 1, Edwin I. Hatch Nuclear Plant, Units 1 and 2,

The Stress Analysis of Underground Piping and Electric Ducts."

This design input document is used to determine the reactor building seismic anchor movements.

3.18 Calculation SMSH-12-020, Version 2.0, "Stress Analysis of Unit 1 Div II Buried Service Water Pipe."

This design input document is used to determine the reactor building seismic anchor movements.

3.19 BH2-C-S23-V012-0001, "Final Seismic Analysis Reactor Building and Internals," April 15, 1975, Volume 1.

This design input document is used to determine the reactor building seismic anchor movements.

3.20 BH2-C-S23-V013-0001, "Final Seismic Analysis Reactor Building and Internals," April15, 1975, Volume 2.

This design input document is used to determine the reactor building seismic anchor movements.

4.0 Acceptance Criteria The stress analysis calculation followed the acceptance criteria outlined in the lnservice Inspection (lSI) Alternative Request HNP-ISI-AL T-HDPE-01, Version 2.0 for the HOPE piping. Per Drawing H-21 033, the steel piping has piping class designation HBC & HAC. Per Table 3.2-4 of Design Input 3.4, HAC corresponds to 150# ANSI stainless steel Nuclear Power Piping, ASME Section Ill, Class 3 and HBC corresponds to 150# ANSI carbon steel Nuclear Power Piping, ASME Section Ill, Class 3.

Table 3.9-29 of Design Input 3.4 indicates that the design criteria for ASME Class 3 piping is per paragraph 3.9.2.1 of Design Input 3.4. Per Section 6.1 of Design Input 3.9, ASME Section Ill, 1971 applies for Hatch Unit 2 ASME piping (Design Input 3.5). The steel piping is qualified to the requirements of paragraph 3.9.2.1 of Design Input 3.4 and all of the applicable requirements of Design Input 3.5.

Per Section 3.5.3 of Design Input 3.13, all ASME Section Ill gate, globe, check, ball, and butterfly valves shall be demonstrated capable of withstanding a seismic acceleration of 4.5g applied simultaneously in all three orthogonal directions. Therefore, the valves are considered acceptable for acceleration values less than 4.5g.

PageS

5.0 Methodology In order to obtain the required data to meet the acceptance criteria discussed in Section 4.0, several SAP2000 Finite Element Models (FEM) were produced. The models include a portion of soil supported HOPE piping in order to sufficiently capture the interaction between the piping inside the vault and the soil supported piping outside of the vault. The soil supported piping is modeled using links which supports the piping while capturing the flexibility of the soil and piping. These links are based on soil and pipe ovaling springs as well as pipe-soil breakaway forces. Both of these parameters dependent on several pipe and soil properties. The modulus of elasticity for the pipe is different for the thermal and seismic load cases and several soil parameters fall within a range.

Because of these variations, upper bound and lower bound soil springs and breakaway forces are determined for both the thermal and seismic load cases. The average of the calculated lower and upper bound values is considered as a best estimate. The best estimate is adjusted upward and downward by a coefficient of variation (COV) to calculate lower and upper bound best estimate spring and breakaway force values. A COV of 1 is conservatively used in accordance with Section 11.4.c.iii of Ref. 7.7. This results in factors of 0.5 and 2.0 for the lower and upper bound best estimate cases respectively. The calculated lower bound values are screened with the best estimate lower bound values to determine worst case lower bound values. Similarly, the calculated upper bound values are screened with the best estimate upper bound values to determine worst case upper bound values. This results in four different models; lower bound thermal, upper bound thermal, lower bound seismic, and upper bound seismic. Detailed calculations and resulting values for the springs and break-away forces are shown in Attachment A of Enclosure 5.

The valve vault is an existing structure. Any settlement is considered to have already occurred.

Therefore, no settlement load cases are considered.

The 1 0" steel piping is decoupled from the 30" header pipe. Per sheet 34 and 39 of BHO-C-S08-V001-0003, the 1 0" steel pipinp has a moment of inertia of 178.1 in4 and the 30" header pipe has a moment of inertia of 3976.1 in. The ratio of the 30" header moment of inertia to the 1 0" pipe moment of inertia is 22.3. The intent of this calculation is to qualify the 1 0" steel piping, 2" steel piping, and HOPE piping branching off of the 30" header. WRC Bulletin 300, Section 2.2.2 recommends that if the ratio of run to branch pipe moment of inertia is 25 to 1, or more, the branch pipe may be decoupled from the run pipe. The use of a lower ratio of run to branch pipe moment of inertia is also judged to be acceptable providing that adequate technical justification is given. The moment of inertia of 22.3 is close to 25 indicating that there will be little influence on the 30" piping from the 1 0" piping. The 30" piping inside the vault is a straight 12-ft long span. This piping is considered to have rigid behavior based on the length of the short straight span in relation to the size of the pipe.

Therefore, decoupling the 1 0" piping from the 30" header is considered acceptable. The model considers a fictitious anchor at this location. This approach to model termination is considered conservative for thermal load cases because any flexibility of the header piping is ignored. Thermal movements from the short straight 30" header are considered negligible since the maximum operating temperature is only 97°F. The differences for seismic load cases are theorized to be negligible based on the 30" header piping being considered rigid.

The FEM include a portion of soil supported piping (buried) in order to sufficiently capture the interaction between the vault piping and the soil supported piping. The soil supported piping is modeled with two changes in direction. This introduces enough flexibility into the analysis to estimate loading from the thermal load cases without being excessively conservative. In order to verify the adequacy of this termination point inside the soil for seismic loading, fictitious loads are applied to the piping inside the vault and the resulting reactions at the termination point are verified as being negligible. The fictitious loads are applied in all directions near the transition between soil supported piping and piping inside the vault. Loads of 2000-lbf have been logically chosen which bound half of the dead weight of the vault piping, its components, and any seismic response (the other half is considered to be carried by the 30" line). Since the loads are applied in all directions, two load cases were created; one for all positive directional loading and one for all negative directional loading. These load cases were screened.

The node point spacing does not define the finite element meshing parameters. These are determined internally by SAP2000. The node point spacing for the soil supported piping is chosen based on discrete soil springs while the node point spacing for the piping that is not soil supported is Page9

based on important locations such as support points, lumped weights, branch lines, pipe section changes, and other points of interest for data extraction like valve accelerations.

The node points for the soil supported HOPE piping are spaced in accordance with Appendix 8 of EPRI Report 1013549 (Ref. 7.2).

There are two seismic load cases considered; the operating basis earthquake (OBE) and the design basis earthquake (DBE). The design spectrum for horizontal ground motion is shown in HNP FSAR-3 (Design Input 3.4) Figures 3.7A-1 & 3.7A-20 for the OBE and Figures 3.7A-2 & 3.7A-21 for the DBE. Points are chosen from these figures to create bounding Response Spectrum functions in SAP2000. Per Sections 2.5.2.1 0 & 2.5.2.11 of HNP-2-FSAR-2 (Design Input 3.3), the vertical accelerations are two-thirds of the horizontal accelerations. Therefore, each vertical response is taken as two-thirds of its horizontal response. The modal response was combined using the NRC Ten-Percent Method which is in accordance with Section 3.7 A.2.1.1 of HNP-2-FSAR-3. The responses in each direction were combined by the SRSS Method which is in accordance with Section 3.7A.3.7 of HNP-2-FSAR-3. All modes with frequencies lower than 33-Hz were considered in the models. This is consistent with Section 3.7A.2.1.1 of HNP-2-FSAR-3.

The seismic response on the soil supported portion of the piping that is included in the models is captured by modeling an equivalent differential temperature. This is the same methodology outlined in Section 4.5 of EPRI Report 1013549 (Ref. 7.2). This methodology involves calculating the controlling strain induced on the pipe due to seismic wave passage. The seismic wave induced strain is limited by breakaway between the pipe and the soil. The controlling strain is converted into an equivalent differential temperature which is input into SAP2000 for the seismic loading.

SAP2000 does not have the capability to limit which parts of the model the response spectrum load cases are applied. Because of this, the response spectrum load cases are applied to the entire piping system; including the soil supported piping. This is conservative since the soil supported piping seismic response is already accounted for by using the equivalent differential temperatures previously discussed. The additional seismic loading should not be excessively conservative as the soil supported portion of the piping is expected to be in the rigid range providing minimal contribution to the modal response. The seismic response from the soil supported piping is absolutely added to the design spectrum response to produce a conservative response for the entire system.

The relative displacements between the valve vault and the surrounding soil are considered negligible since the valve vault is relatively small and the surrounding soil has a high compaction level (minimum 95% of the maximum dry density compaction criterion per Design Input 3.3, Section 2.5.1.2.8).

There are three thermal load cases considered; pipe at 32°F, pipe at 97°F, and pipe at 125°F.

These are modeled as temperature differentials from 70°F; that is a change of -38°F, 27°F, and 55°F. The resulting loads for the 32°F and 125°F cases (minimum operating temperature and design temperature rounded up from 123°F) are combined by absolute sum in order to capture the stress range which is used to qualify pipe stresses. Using the design temperature for the stress range instead of the maximum operating temperature of 97°F is conservative. The 97°F load case (maximum operating temperature) is used for support and flange qualifications only.

6.0 Assumptions & Design Considerations 6 1 Design Consideration:

The flowable fill is conservatively considered to be used for all of the backfill (all the way to grade).

Justification:

This design consideration does not require verification as the backfill properties incorporated into the detailed design through DCP SNC591628 will be bounded by the properties used in the evaluation.

Page 10

7.0 References 7.1 PPI, "Handbook of Polyethylene Pipe", 2nd Ed. with 6/6/12 Errata.

7.2 EPRI Report 1013549, "Nondestructive Evaluation: Seismic Design Criteria for Polyethylene Pipe Replacement Code Case," Technical Update, September 2006.

7.3 American Lifelines Alliance, "Guidelines for the Design of Buried Steel Pipe," July 2001 with Addenda through February 2005.

7.4 McGrath, T. J. and Hoopes, R. J.," 'Bedding Factors and E' Values for Buried Pipe Installations Backfilled with Air-Modified CLSM," The Design and Application of Controlled Low-Strength Materials (Fiowable Fill), ASTM STP 1331, A. K. Howard and J. L. Hitch, Eds.,

American Society for Testing and Materials, 1998.

7.5 Das, Braja M., "Principles of Foundation Engineering," 6th Ed.

7.6 CSI Knowledge Base Website, "Damping in Response-Spectrum Analysis,"

https://wiki.csiamerica.com/display/kb/Damping+in+response-spectrum+analysis 7.7 NUREG-0800, Section 3.7.2, "Seismic System Analysis."

7.8 Calculation SMSH-15-001, Version 1, "Stress Analysis for New U2 DIV II PSW Subgrade Vault Piping."

7.9 Calculation SMNH-15-008, Version 1, "14" HOPE to Metallic Transition Piping Flanged Joint Analysis."

7.1 0 WRC Bulletin 300.

Page 11

Edwin I. Hatch Nuclear Plant-Unit 2 Proposed lnservice Inspection Alternative HNP-ISI-AL T-HDPE-01, Version 2.0 Final Stress Analysis Design of the HOPE System 141nch HOPE to Metallic Flanged Joint Analysis Calculation

SOUTHERN A COMPANY

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Plant:

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Title:

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01 002 Southern Nuclear Design Calculation Calculation Number:

SMNH-15-008 01&2 I Discipline:

Stress 14" HOPE to Metallic Flanged Joint Analysis I

Subject:

PSW HOPE Piping Purpose I Objective:

To qualify the 14" x 10" HOPE to Metallic Transition Piping Flanged Joint.

System or Equipment Tag Numbers:

PSWSystem Contents Topic Page Attachments

of (Computer Printouts, Technical Papers, Pages Sketches, Correspondence)

Purpose of Calculation 2

Att. A - Excerpts from Selected References 12 Summary of Conclusions 2

Design inputs 2

Acceptance Criteria 6

Methodology 6

Assumptions 6

References 7

Analysis 7

Total # of Pages Including 23 cover sheet & Attachments :

OSafet Si niflcant 0 Non* Safet Version Record Version No.

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Hatch Table of Contents 1.0 Purpose of Calculation 2.0 Summary of Conclusions 3.0 Design Input 3.1 3.2 3.3 Flange Loads Flange Data Flange Bolting 4.0 Acceptance Criteria 5.0 Methodology Calculation Number:

SMNH-15-008 6.0 Assumptions and Design Considerations 7.0 References 8.0 Analysis 8.1 8.2 Slip-on Flange with Hydrostatic Pressure -(Initial Condition Only)

Slip-on Flange with 190 psi (Maximum Pressure)- Long Term Condition Attachment A - Excerpts from Select References Sheet: 1 of 10 Sheet 2

2 2

3 5

5 6

6 6

7 7

7 9

A-1

Plant:

Calculation Number:

Sheet: 2 of 1 0 Hatch SMNH-15-008 1.0 Purpose of Calculation This calculation determines the seating stress and bolt torque requirements for the transition flange assembly between the steel and High Density Polyethylene (HOPE) piping. This piping is part of the Plant Service Water (PSW) piping for the Hatch Nuclear Plant Unit 2 Division II. A section of the existing buried 10" Carbon Steel (CS) piping will be replaced by a IPS 14 DR 7 HOPE piping. The HOPE piping inside diameter is 10.0" and matches the inside diameter (10.02") of the replaced 10" SCH 40 steel pipe segment Thus, the flow area is not compromised.

The steel/ HOPE piping transition joint includes a custom-made 14" 150# 816.5 Stainless Steel (SS) flange with slip-on flange dimensions, except Its inside diameter matching a 10" 150#

816.5 slip-on flange. The flanged joint typically consists of a metallic (SS) flange, a PE flange adapter, a metallic (SS) backing ring and high strength bolting. The SS transition piece will be welded back to the existing buried CS pipe, whereas the PE adapter will be joined with the PE piping.

The method used in this calculation is a combination of the guidance from the Plastic Pipe Institute (PPI) Specification TN-38 (Ref. 7.2), the ASME Subsection ND-3600 and the ASME Ill Appendix XI for metallic flanges. Per the PPI TN-38, the HOPE material is visco-elastic which will relax its Initial seating stress over a period of time.

2.0 Summary of Conclusions Following is the summary of results from Paragraph 8.0 of this calculation.

2.1 The 14" x 150# 816.5 slip-on (raised face) flange dimensions have been used for this calculation, except the inside diameter, which matches the 10" x 150# slip-on flange dimension.

2.2 If a gasket is used (between the metallic flange and the HOPE flange adapter), its m value must be limited to 2.0 and they value limited to 1800 psi. These values are used in the analysis.

2.3 The maximum operating pressure of 190.0 psi instead of the hydrostatic pressure (180x1.5 = 270 psi) has been used as the long-term pressure for the qualification of this flanged joint.

3.0 Design Input The following design data have been obtained from PPI TN-38 (Ref. 7.2)

Maximum Allowable HOPE Gasket Seating Stress y Nut Factor K Long Term Creep Factor c1 Gasket Factor m

= 1800 psi.

= 0.16 (Lightly Grease Lubricated)

= 0.35

= 2.0 (Maximum)

The following HOPE Flange Adapter data have been obtained from ISCO Product Catalog Flange Outside Diameter Outside Diameter Inside Diameter Thickness Flange Thickness

= 17.5"

= 14.0"

= 10.0" (DR 7)

= 2.0"

= 1.5" Minimum

Plant:

Calculation Number:

Sheet: 3 of 10 Hatch SMNH-15-008 P1 := 190.0*psi Maximum Operating Pressure P2 := 180.0*psi Design Pressure P3 := P2*1.5 P3 = 270*psi (Hydrostatic Pressure)

P := max(P1, P3)

P = 270*psi T 0 := 123.0*deg

F Design Temperature Flange= ANSI 816.5 *150 # 14" x 10" Special Transition Flange Flange Material: SA182 F316L (or Equivalent) sh = 15700.0 psi (Ref. 7.3 Appendix 1.7-2)

Bolt Material: SA193 87 S1 = 25000.0 psi (Ref. 7.4)

The bolt material used in this analysis has the following properties:

Bolt size Bolt Material Sy

= 1" UN (16 TPI)ASME SA193 B7 bolt material (high strength Chrom-Moly Steel)

= ASME SA 193 B7 Chrom-Moly Steel

= 105,000 psi Minimum Yield Strength Su

= 125,000 psi Minimum Tensile Strength sh

= 25,000 psi (Allowable Stress up to 1100 OF)

Following alternative bolt materials (with higher strength) can be used:

Bolt Material Sy Su sh

= ASME SA540 B21/ B22 Class 4 Chrom-Moly-Vanadium Steel

= 120,000 psi Minimum Yield Strength

= 135,000 psi Minimum Tensile Strength

= 27,000 psi (Allowable Stress up to 850 °F)

Nut Material: Heavy Hex (To match Bolt properties)

Flange Adapter: ISCO Style 3545 (Ref. 7.1) a 5 = 1800.0 psi-Max Seating Stress (Ref. 7.2) 3.1 Flange Loads (Ref. 7.5)

The following load sets represent 14" flange assembly nodepoints 1100 and 1010. Loads from node 1010 have been used in this calculation, which governs both flange locations.

Axial Force Fdw := 12.0*1bf Fobe := 264.B*Ibf Fhve := O.O*Ibf Due to Deadweight Fth := 5869.2*1bf Due to OBE Earthquake F dbe := 474.9*1bf Due to Heave Due to Max. Thermal Due to DBE Earthquake

Plant:

Calculation Number:

Sheet: 4 of 10 Hatch SMNH-15-008 Torsional Moment Mrdw := 188.6*in-lbf Due to Deadweight Mnh := 292.6* in *lbf Due to Max. Thermal Mrobe := 193.7-in*lbf MThve := 0.0-in-lbf Due toOBE Mrdbe := 281.7 *in *lbf Due to DBE Due to Heave Bending Moment M2dw:= 3419.7-in*lbf M2obe := 356.5-in*lbf Due to Deadweight M2th := 2298.9 *in *lbf Due to OBE M2dbe :=:: 551.5-in*lbf Due to Heave M2hve := 0.0-in-lbf M3dw := 377.8-in *lbf M3obe:= 3257.1-in*lbf M3hve := O.O*in*lbf Due to Deadweight M3th := 6638.7 *in *lbf DuetoOBE Due to Heave Combined Loads Acting @Flanged Joint Fs := Fdw + Fth Fd := Fdw + Fth + Fdbe + Fhve M2s := {M2dw + M2th)

M3s := {M3dw + M3th)

M2d := {M2dw + M2th + M2dbe + M2hve)

M3d := {M3dw + M3th + M3dbe + M3hve)

Mrs := ( Mrdw + Mnh)

Mrd := {Mrdw + Mnh + MTdbe + Mrhve)

(

2 2)0.5 Mtsb := M2s + M3s

(

2 2)0.5 Mtdb := M2d + M3d M3dbe := 4668.9-in*lbf Fs = 5881.2-lbf Fd = 6356.1-lbf M2s = 5719-in-lbf M3s = 7017-in*lbf M2d = 6270-in-lbf M3d = 11685-in*lbf Mrs= 481-in-lbf Mrd = 763-in*lbf Mtsb = 9052-in *lbf Mfdb = 13261-in-lbf Mts = 9052-in*lbf Mtd = 13261-in *lbf Due to Max. Thermal Due to DBE Due to Max. Thermal Due to DBE

Plant:

Calculation Number:

Hatch SMNH-15-008 3.2 FLANGE DATA (Ref. 7.3 Appendix Xl-3130 & Ref. 7.10)

A:= 21.00-in B := 10.88*in C := 18.75*in x := 12.00*in x - B 91 := --

2 go:= g1 t:= 1.3125*in 0 := 10.00-in s0 := max(B, D)

G0 := 17.50-in R := 16.25*in R-Bo 14" x 150 #Flange Outside Diameter 10" x 150 #Flange Inside Diameter 14" x 150 #Bolt Circle Diameter 10" x 150 #Hub Cia.@ Base 91 = 0.56-in Hub Thickness@ Large End Hub Thickness@ small End 14" x 150 #Flange Thickness HOPE Pipe Inside Cia.

Bo = 10.88-in Maximum Gasket Contact Face Inside Cia.

Flange Adapter (Gasket) Outside Cia.

14" x 150 # Flange Raised Face Diameter Sheet: 5 of 1 0 N:=--

2 N

bo:=-2 N = 2.685*in Gasket Contact Width (Ref. 7.3 Table Xl-3221.1-2) b0 = 1.3425 *in Basic Gasket Seat Width (b. )0.5 o*m b :=....:..___2....:,__

b = 0.5793 *In Effective Gasket Seat Width (Since b0 > 0.25: b = (b00*5)1 2)

G := R- (2-b)

G = 15.091. in Diameter@ Location of Gasket Load Reaction m := 2.0 Gasket Factor y := 1800-psi Maximum Gasket Compressive Seating Stress (Ref. 7.2) 3.3 Flange Bolting

SA-193 87 Bolt Size= 1" Thread Type= 16 x UN di := 0.932*in Bolt Minor Cia.

n := 12 Number of Bolts 2

Ak = 0.682*1n (Ref. 7.4)

Root Area of 1" x 16 UN Bolt Total Bolt Root Area

Plant:

Hatch Sa:= 25000*psi sb := 25000 *psi ua := min(1.0*Sa, 1.0*Sb) 4.0 Acceptance Criteria Calculation Number:

SMNH-15-008 Allowable Bolt Stress @ 100 °F Allowable Bolt Stress@ 123 °F ua = 25000*psi 4.1 The bolt stresses should be limited to 1.0 S8 4.2 The gasket seating stress (HOPE flanger adapter) should be less than 1800.0 psi Sheet: 6 of 10 4.3 If a gasket is used (between the metallic flange and the HOPE flange adapter), its m value must be limited to 2.0 and they value limited to 1800 psi.

5.0 Methodology The flange analysis is based on the calculation steps delineated in ASME Appendix XI. The equivalent pressure term P eq is derived from ASME Section Ill ND-3658-1. The bolt torque requirements and limits are based on the Plastic Pipe Institute (PPI) TN-38 guidelines "Bolt Torque guidelines for Polyethylene Flanged Joints"- July 2011.

This Stainless Steel special flange assembly is custom-fabricated with the 14" 150# 816.5 slip-on flange parameters, except its inside flange diameter, which matches a 10" 150# 816.5 slip-on flange. The 14" IPS DR7 HOPE Flange Adapter dimensions are obtained from Ref. 7.1 Design Information Transmittal (DIT). The 1" bolt material properties used in this analysis are for Low Alloy Crome-Moly Steel A193 87 bolt, from ASME Section Ill Appendices Table 1-7.3.

A few alternate bolt materials that can be used with this flange assembly (with equal or higher Sh allowables) are listed in Section 3.0 of this calculation.

This calculation is performed by using MATHCAD program version 14.

6.0 Assumptions and Design Considerations 6.1 The special flange fabricated matches the outside dimensions of a 14" B16.5 #150 flange, except its inside dimensions matching a 10" 816.5 #150 flange.

6.2 If a gasket (optional) is used between the metallic flange and the HOPE flange adapter, its design factors matches as follows: y = 1800 psi (maximum); m = 2.0 (maximum) 6.3 Following alternative flange bolting materials to A 193 87 for this assembly are used:

1) A540 821

2) A540 822 6.4 Two separate conditions are analysed. The first condition is to qualify the flange assembly to withstand the hydrostatic pressure of 270 psi for initial gasket seating. In this analysis, the creep factor cf = 1.0 is used. The second condition analyzed is by using long-term operating (maximum) pressure of 190.0 psi, and the creep factor cf= 0.35. Use of the maximum operating pressure of 190.0 psi for long-term condition is justified, because the full relaxation to 35% of its initial seating is not expected concurrently with the hydrostatic test pressure.

Plant:

Calculation Number:

Sheet: 7 of 10 Hatch SMNH-15-008 6.5 The flange bolts must be re-torqued to the calculated torque value of 202.0 ft-lb, per Paragraph 8.2 of this calculation, before any subsequent hydrostatic test.

7.0 References 7.1 DIT No 12695-004-01: HOPE Flange Joint Calculation Input.

7.2 Plastic Pipe Institute (PPI) Tech. Note (TN) 38: Bolt Torque for Polyethylene Flanged ints.

7.3 ASME Boiler and Pressure Vessel Code Section Ill-ND & Appendices -1974 7.4 ASA B1.1-1960: USA Standard - Unified Screw Threads 7.5 Southern Company Calculation SMSH-15-001.

7.6 Southern Company Procedure: NMP-ES-050-F01 Ver. 2.0 7.7 Southern Company Drawing: H21033 (P & ID) 7.8 ISCO Product Catalog: IPS HOPE Fittings -14" Flange Adapter 7.9 GARLOCK Catalog: GYLON Gasketing 7.10 TUBE-TURN Catalog 311: Welding Fittings, Flanges 7.11 Southern Company Procedure: NMP-ES-039-001 Ver. 5.1-Calculations-Preparation and Revision 7.12 HNP-2-FSAR-3 Revision 27 (1 0/09) 7.13 Southern Company Calculation SMSH-15-002.

8.0 Analysis A 14" x 10" 150# special slip-on flange dimensions (14" x 150# B16.5) are used in this analysis, except the inside flange diameter, which matches the 10" 150# B16.5 flange. Dimensions are from the TUBE-TURN Catalog.

The 1" UN 16 Threads Per Inch (TPI) bolt minor diameter value is used.

Following Two bounding options are calculated:

1) Apply Hydrostatic Pressure (1.5 x Design Pressure)- For initial installation only

2) Apply 190.0 psi (Maximum Pressure)- For long term condition 8.1 Slip-on Flange w1th Hydrostatic Pressure- (Initial Condition Only)

This evaluation using a creep factor c1 = 1.0 provides assurance that the gasket is initially leak-tight.

Plant:

Calculation Number:

Sheet: 8 of 10 Hatch SMNH-15-008 Equivalent Pressure- (Ref. 7.3 ND-3658/ Appendix XI))

16-Mfs 4*F6 Peq1 :=

3 + --

'1\\"*G

'n"*G2 8Mfd 4*Fd P eq2 := --3 + --2

'1\\"*G

'n"*G P eq := max(P eq1 'P eq2)

Flange Load Computation

'1\\"

2 H := -*G.pfd 4

Wm2 := '1\\"*b*G*y Wm1 Am1 := --

sb Wm2 Am2:= --

Sa Am:= max(Am1,Am2)

Am

-=0.446 Ab W := max(wm1, wm2)

P eq1 = 46.3*psi P eq2 = 45.4*psi Peq = 46.3*psi Pfd = 316.3*psi H = 56576.2 -lbf Hp = 34749.8*1bf wm1 = 91326*1bf wm2 = 49439.9-lbf Am1 = 3.653*in2 Am2 = 1.978*in2 Am

-!>1.0 Ab w = 91326*1bf w = 91326*1bf Required Bolt Load and Torgue cf := 1.0 Creep factor (Ref. 7.2)

Equivalent Pressure on the Flanged Joint due to External Load (ND-3658)

P =Hydrostatic pressure= 270 psi Total End Force Total Joint contact Surface Compression Load Minimum Required Bolt Load for Design Conditions -ASME Xl-3221.1)

Total Bolt Load for Gasket Seating Flange Design Bolt Load (OR) w Fbolt := -

Fbolt = 91326*1bf Required Initial Bolt Load Cf

Plant:

Hatch Fbolt Fi:=--

n Fi = 7610*1bf Calculation Number:

SMNH-15-008 Load per Bolt F'

F's:= ~

Ak F's = 11156*psi Bolt Stress s

Sa = 25000*pst- = 0.446 ua Required Bolt Torgue (Ref. 7.2)

K:= 0.16 Nut Factor-Lightly greased Bolts and Nuts Tbolt = 95*ft*lbf Bolt Torque Seating Stress in Flange Adapter 71'(2

2) 2 As := 4* R - B0 A5 = 114.423*in Seating Area of Slip-on Flange Fbolt Ss:= --

As S8 = 798*psi Seating Stress in Flange Adapter Ss

= 0.443 us 8.2 Slip-on Flange with 190 psi (Maximum Pressure)- Long-term Condtion This evaluation uses a long-term creep factor c, = 0.35.

Equivalent Pressure- (Ref. 7.3 ND-3658/ Appendix XI))

16*Mfs 4*F8 Peq1 :=

3 + --2 71'*G 71'*G P eq1 = 46.3 *psi 8Mfd 4*Fd P eq2 := --3 + --2 71'*G 71'*G Peq2 = 45.4*psi F's

-~1.0 ua Sheet: 9 of 10 (OK)

P eq = 46.3*psi Equivalent Pressure on the Flanged Joint due to External Load (ND-3658)

Flange Load Computation 71' 2

H := -*G.pfd 4

Pfd = 236.3*psi H = 42266.3 *1bf Hp = 25960.5 *lbf wm1 = 68226.8*1bf wm2 = 49439.9*1bf P1 = 190 psi (Maximum Operating Pressure)

Total End Force Total Joint contact Surface Compression Load Minimum Required Bolt Load for Design Conditions -ASME Xl-3221.1)

Total Bolt Load for Gasket Seating

Plant:

Hatch Wm1 Am1 := --

Sb Wm2 Am2:= --

Sa Am:= max(Am1,Am2)

Am

-=0.333 Ab Calculation Number:

SMNH-15-008 Am2 = 1.978*in2 2

Am= 2.729*1n Am

-$1.0 Ab W := max(wm1, wm2) w = 68226.8*1bf Flange Design Bolt Load (OR) w = 68226.8 *lbf Required Bolt Load and Torque Cf := 0.35 w

Fbolt :=-

Cf Fbolt Fi:=--

n Fi Long term creep factor (Ref. 7.2)

Fbolt = 194934*1bf Required Initial Bolt Load F1 = 16244*1bf Load per Bolt F's F's :=-

Ak F's = 23811*psi Bolt Stress Sa= 25000*psi

= 0.952

<Ta Required Bolt Torgue (Ref. 7.2)

K:= 0.16 Nut Factor-Lightly greased bolts & Nuts Tbolt = 202*ft*lbf Bolt Torque Seating Stress In Flange Adapter As:= ;.( R2 - s0

2)

As= 114.423*in2 Seating Area of SUp-on Flange Fbolt Ss:= --

As S8 = 1704*psi Seating Stress in Flange Adapter Sheet: 10 of 10 Ss s1.0 (OK)

<rs

Plant Hatch Calculation Number SMNH-15-008 Sheet: A-1 ATTACHMENTS FOR ENCLOSURE 8

Plant:

Hatch Southern Nuclear Design Calculations X

Calculation Number:

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SMSH~5~08 A~

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Southern Nuclear Design Calculations Plant:

Calculation Number:

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Plant:

Hatch Southern Nuclear Design Calculations Calculation Number:

Sheet:

SMSH-15-008 A-4

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Calculation Number:

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Plant:

Hatch Southern Nuclear Design Calculations Calculation Number:

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SMSH-15-008 A-7

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B.2 Summary of Output (Continued)

Joint Forces for Flanged Connection Between HOPE & Stainless Steel Piping (Absolute Values)*

Node 1010 Comments Model Load Case (Flanged Connection Between HOPE & S.S.l FX {lb)

FY(Ib}

FZ (lb)

MX(Ib-in) MY (I b-in)

MZ(Ib-in)

Se1sm1c FULLDEAD 9.93 2.91 416.77 3419 70 64 25 242.75 Lower OSE 59.27 130.72 7.03 269.73 193.69 3257 13 Sound DBE 83.19 219.80 11.06 391.68 281.67 4668.87 Seismoc FULLDEAD 19.86 12.00 385.20 2286.40 188.56 377.79 Upper OSE 97.19 264.80 11.51 356.49 147.74 2597.72 Bound DBE 143.12 474.92 18.02 551.52 219.45 3769.21 Thermal FULLDEAD 9.08 2.10 396.14 2581.71 8.27 232.00 Lower THERMAL MIN 12.43 1947.72 40.47 1161.99 292.64 2721.42 Minimum Operating Temperature (32 deg F)

THERMAL MAXOP 8.79 1383.91 28.75 825.50 207.93 1933.95 MaXImum Operatmg Temperature {97 deg F)

Bound THERMAL MAX 13.06 2644.38 58.25 1657.06 426.31 4241.68 Based on Design Temperature (125 deg F)

Thermal 1 FULLDEAD 17.59 10.80 363.97 1670.02 91.31 360.42 I

TrlERMAL MIN 288.79 5869 18 76.44 2298.85 194.71 6638.66 Monimum Operating Temperature (32 deg F)

Upper THERMAL MAXOP 205 19 4170.21 54.31 1633.40 138.35 4716.94 Ma~imum Operatmg Temperature (97 deg F)

Bound I THERMAL MAX 418.51 8496.91 110.69 3328.54 283.04 9611.10 Based on Design Temperature (125 deg F)

FULLDEAD 19.86 12.00 416.77 3419.70 188.56 377.79 OBE 97.19 264.80 11.51 356 49 193.69 3257.13 Boundmg DBE 143.12 474.92 18.02 551.52 281.67 4668.87 Values THERMAL MIN 288.79 5869.18 76.44 2298.85 292 64 6638 66 Mommum Operating Temperature (32 deg F)

THERMAL MAXOP 205.19 4170.21 54.31 1633.40 207.93 4716.94 Max1mum Opera tong Temperature (97 deg F)

THERMAL MAX 418.51 8496.91 110.69 3328.54 426.31 9611.10 Based on Design Temperature (125 deg F)

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Joint Forces and Dosplacements for Support Inside Valve Vault (Absolute Values)

Node 3110 Model Load Case (Sup crt Inside Vault)

Comments FZ (lb)

DX(in)

DV(in)

Se1smoc FULL DEAD 1268.11 0.0082 0.0006 Lower OBE 134.50 0.0299 0.0062 Bound DBE 212.24 0.0424 0.0095 Se1smoc FULLDEAD 1218.71 0.0058 0.0002 Upper OBE 129.31 0.0125 0.0043 Bound DBE 227.22 0.0179 0.0076 Thermal FULLDEAD 1232.73 0.0085 0.0006 Lower THERMAL MIN 820.20 0.0555 0.0026 Minimum Operating Temperature (32 deg F)

THERMAL MAXOP 582.76 0.0394 0.0018 Maximum Operatmg Temperature (97 deg F)

Bound THERMAL MAX 1126.64 0.0788 0.0009 Based on Design Temperature (125 deg F)

Thermal FULLDEAD 1185.74 0.0065 0.0003 Upper THERMAL_ MIN 2227.58 0.0507 0.0630 Minimum Operating Temperature {32 deg F)

THERMAL MAXOP 1582.76 0.0360 0.0447 Max1mum Operating Temperature (97 deg F)

Bound THERMAL MAX 3224.93 0.0733 0.0912 I Based on Design Temperature (125 deg F)

FULLDEAD I 1268.11 0.0085 o.ooo6 I OBE I 134.50 0.0299 0.0062 I I Bounding DBE 227.22 0.0424 00095 I Values THERMAL_ MIN 2227.58 0.0555 0.0630 ! Minimum Operating Temperature (32 deg F)

THERMAL MAXOP 1582.76 0.0394 0.0447 I Maximum Operatmg Temperature (97 deg F)

THERMAL_ MAX 3224.93 0.0788 o.0912 I Based on Des1gn Temperature (125 deg F)

Note:

There are only support loads in the X d~rectoon and there are no movements in the Z direction.

C"

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T~

B.2 Summary of Output (Continued)

Jotnt Forces for Boundary Check Load Case (Absolute Values)*

Node 1190 Model (Model Termination Pomt)

FX (lb)

FY (lb)

FZ(Ib)

MX(Ib-in) MY(Ib-in) MZ(Ib-in)

Seismic Lower Bound 0.24 0.81 0.52 0.00 0.00 0.00 Seismic Upper Bound 0.01 0.00 0.13 0.00 0.00 0.00 Thermal Lower Bound 0.12 0.44 0.49 0.00 0.00 o.oo Thermal Upper Bound 0.00 0.00 0.04 0.00 0.00 0.00 Boundmg Values 0.24_ L_ _ _Q.8l O.S2 0.00 -

0.00 0.00 Joint Forces for Flanged Connection Between HOPE & Stainless Steel Piping (Absolute Values)*

Node 1100 Comments Model Load Case (Flanged Connection Between HOPE & S.S.)

FX(Ib)

FY(Ib)

FZ(Ib)

MX(Ib-in) MY(Ib-in) MZ(Ib-in)

FULLDEAD 0.00 0.00 427.14 4264.36 202.78 0.00 Seismic OBE 125.63 7S.48 3.86 132.Z7 34.87 4584.77 Includes SAM OBE Lower OBE 195.29 116.04 7.24 246.45 55.69 6963.19 Includes SAM OBE SAM OBE 23.66 22.14 0.18 7.49 0.05 755.70 Bound SAM_DBE 47.31 44.28 0.35 14.98 0.11 1511.40 HEAVE 0.00 0.00 633.79 22762.94 413.14 0.00 FULLDEAD 0.00 0.00 361.10 1809.65 101.28 0.00 Seismic OBE 81.33 334.76 9.95 78.30 20.98 1972.78 Includes SAM OBE Upper DBE 126.03 611.00 18.73 147.25 34.56 2940.81 Includes SAM DBE SAM OBE 3.92 17.60 0.05 1.55 0.05 136.67 Bound SAM DBE 7.85 35.20 0.10 3.10 0.09 273.34 HEAVE 0.00 0.00 1362.28 34117.04 901.98 0.00 FULLDEAD 0.00 0.00 397.97 3057.00 204.64 0.00 Thermal THERMAL MIN 120.27 416.41 0.00 0.00 0.00 6778.21 Minimum Operating Temperature (32 deg F)

Lower THERMAL MAXOP 77.38 288.33 0.00 0.00 0.00 4373.34 Maximum Operating Temperature (97 deg F)

Bound THERMAL MAX 188.46 616.12 0.00 0.00 0.00 10603.84 Based on Design Temperature (125 deg F)

HEAVE 0.00 0.00 450.71 14754.10 398.27 o.oo FULLDEAD 0.00 0.00 344.86 1327.01 89.22 0.00 Thermal THERMAL MIN 157.00 2092.20 0.00 0.00 0.00 844.26 Minimum Operating Temperature (32 deg F)

Upper THERMAL MAXOP 111.55 1486.56 0.00 0.00 0.00 599.87 Maximum Operating Temperature (97 deg F)

Bound THERMAL MAX 230.00 3035.71 0.00 0.00 0.00 1288.81 Based on Design Temperature (125 deg F)

HEAVE 0.00 0.00 916.78 21672.82 457.08 0.00 FULLDEAD 0.00 0.00 427.14 4264.36 204.64 0.00 OBE 125.63 334.76 9.95 132.27 34.87 4584.77 Includes SAM OBE DBE 195.29 611.00 18.73 246.45 55.69 6963.19 Includes SAM DBE Bounding SAM OBE 23.66 22.14 0.18 7.49 0.05 755.70 SAM DBE 47.31 44.28 0.35 14.98 0.11 1511.40 Values THERMAL MIN 157.00 2092.20 0.00 0.00 0.00 6778.21 Minimum Operatmg Temperature (32 deg F)

THERMAL_MAXOP 111.55 1486.56 0.00 0.00 0.00 4373.34 Maximum Operating Temperature (97 deg F)

THERMAL_ MAX 230.00 3035.71 0.00 0.00 0.00 10603.84 Based on Design Temperature (125 deg F)

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1-----IIO_J This drawing is the propOfty of Gec~g Fischer Central PlosUcs LLC !Georg Fischer) and is wbjectto copyright and other intenecluol property protection. It Is not allowed to copy. reproduce. or tronsmil this drawing to lhfd parties without Pfiof wtitten permission of Georg Fischer.

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Initiation Request, Request Acceptance... Administration Questions Answers Results Approval Other:
MONTHYEAR2015-12-21 [Table View]

NL-15-1628, Proposed Inservice Inspection Alternative HNP-ISI-ALT-HDPE-01, Version 2.0, Final Stress Analysis for the Design of the Hdpe System

Text

Enclosures:

Title:

Subject:

Conclusions:

5.0 Methodology

6.0 Assumptions

6.1 Assumption

References:

Conclusions:

6.0 Assumptions

6.1 Assumption

6.2 Assumption

6.3 Assumption

6.4 Assumption

6.5 Assumption

6.6 Assumption

6.7 Assumption

6.8 Assumption

6.9 Assumption

References:

544.289psi v.lc

Title:

Subject:

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NL-15-1628, Proposed Inservice Inspection Alternative HNP-ISI-ALT-HDPE-01, Version 2.0, Final Stress Analysis for the Design of the Hdpe System

Text

Enclosures:

Title:

Subject:

Conclusions:

5.0 Methodology

6.0 Assumptions

6.1 Assumption

References:

Conclusions:

6.0 Assumptions

6.1 Assumption

6.2 Assumption

6.3 Assumption

6.4 Assumption

6.5 Assumption

6.6 Assumption

6.7 Assumption

6.8 Assumption

6.9 Assumption

References:

544.289*psi v.lc *

Title:

Subject:

Navigation menu

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544.289psi v.lc