ML20008E874

Amend 56 to PSAR

ML20008E874
Person / Time
Site:	Allens Creek File:Houston Lighting and Power Company icon.png
Issue date:	03/31/1981
From:	HOUSTON LIGHTING & POWER CO.
To:
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ML20008E873	List: ML20008E872 ML20008E874
References
NUDOCS 8103100209
Download: ML20008E874 (475)

Text

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l ACNG S- PS A R l 110USTON LIGilf1NG & POWER C(UPANY

! ALLENS CREEK NUCLEAR GENERATING STATION - UNIT NO. 1 l PRELIM'N\RY SAFETY ANALYSIS RE POR'I i

AMENDMENT No. 56 l l INSTRUCTION Slu:ET i

j j

This amendment contains infc,rmation pertaining to New Fuel and PSAR Update.

l Each revised page bears the notation Am. No. 56, (3/81) at the bottom of the ,

i nage. Vertical bars with the number 56 representing Amendment No. 56 have i I

been used in the margins of the revised pages to indicate the location of the i revision on the page.

The following pa >e removals and insertions should be made to inco rpe ra te Amendment No. 56 into the PSAR.

t REMOVE INSERT (EXIISTING PAGES) (AMENDMENT No. 56 PAGES) i Chapter 1_ Chapter 1 l l

, 1* 1* l j 2a* 2a*

1.2-4 1.2-4 l 1.2-6 1.2-6 1.2-7 1.2-7 1.2-12 1.2-12 1.6-2 1.6-2 i

1. 6- 3 1. 6- 3 1.6-4 1. 6- 4 1.6-5 1. 6- 5 1.6-Sa 1. 6- Sa 1.6-5b 1. 6- 5b 1.6-6 1.6-6 [

1.6-7 1. 6- 7 ,

Chapter 2 Chapter 2 r

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2.2.A-1 2.2.A-1 2.2.A-13 2.2,A-13 Chapter 3 Chapter 3 1* 1*

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Chapter 3 (Cont'd) Chapter 3 (Cont'd) i xxviii xxviii l l

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- xxxix 3.1-2 3.1-2 3.1-9 3.1-9 3.1-12 3.1- 12 3.1-19 3.1-19 .

3.1-21 3.1-21 3.1-22 3.1-22 3.1-23 3.1-23 f; 3.1-24 3.1-24 3.1-26 3.1-26 i 3.1-27 3.1-27 3.1-28 3.1-28 3.1-29 3.1-29  !

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I' REMOVE II;S E RT (EXISTING PAGES) (AMENDMENT SO. 56 PAGES) l Chapter 3 (cont'd) Chapter 3 (Cont'd)

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ACNGS- PSAR INSERT REMOVE (AMENDMENT NO. 56 PAGES)

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Chapter 6 (Cont'dl Chapter 6 (Cont'd) 6.3-26, 6,1-27 6.3-29 6.3-29 6.3-30 6.3-30 6. 3- 31 a 6.3-31a 6.3-32 6.3-32 6.3-33 6.3-33, 6.3-34 6.3-34 6.3-35 6.3-35 6.3-36 6.3-37 6.3-38 6.3-39 F6.3-3 F6.3-3 F6.3-Sa F6.3-Sa F6. 3- 6 F6.3-6, F6.3-7 F6.3-7 F6.3-9 F6.3-9 F6.3-10 F6. 3- 10 F6.3-11 F6. 3- 11 F6. 3- 12 F6. 3- 12 F6.3-13 F6.3-13 F6.3-13a, b, c, d F6. 3- 13a F6.3-13b ~

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@ F6.3-38, 39, 40 F6.3-38 F6.3-39 F6.3-40 F6.3-41 F6.3-41 IS-5 Am. No. 56, (3/81)

ACNCS- PS AR REMOVE INSERT (EXISTING PAGES) (AMENDMENT No. 56 PAGES)

Chapter 6 (Cont'd) Chapter 6 (Cont'd)

F6.3-42, 43, 44, 45a , 45b, 45c F6. 3- 42

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j Chapter 7 Chapter 7 1* 1*

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7.2-7 7.2-7 i

j 7.3-14 7.3-14 7.4-21 7.4-21 1

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7 . 7- 9 7. 7- 9 Chapter 15 Chapter 15 1* 1*

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d REMOVE INSERT l (EXISTING PAGES) (AMENDMENT NO. 56 PAGES)

Chapter 15 (Cont'd) Chapter 15 (Cont'd) i

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j Chapter 16 Chapter 16 i

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l Appendix C Appendix C l

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C1.44-1 C1.44-1

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Enclosed are fili 3 tabs l for Appendices F through 0 These Appendices can be  !

j 9 found at the back of the PSAR under Resolution of Staff Questions Appendix 3 l

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LIST OF EFFECTIVE PAGES i CHAPTER 1 j INTRODUCTION AND GENERAL DESCR_IPTION OF_ PLANT Page Amendment 1* 56 i 2* 46 l 2a* 56 i 3* 37 j 4* 39 i

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ACNGS-PSAR I LIST OF EFFECTIVE PAGES (Cont'd)  !

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2a Am. No. 56, (3/81) ,

s ACNGS-PSAR 1.2.2.2 Nuclear System The nuclear system includes a single-cycle, forced circulation, General Electric boiling water reactor that produces steam for direct use in the steam turbine ( see NED0-10569A) . A heat balance showing the major parameters of the nuclear system for the rated power condition is shown in Figure 1.2-28, 1.2.2.2.1 Reactor Core and Control Rods Fuel for the reactor coro consists of slightly enriched uranium dioxide pellets sealed in Zircaloy-2 tubes. These tubes (or fuel rods) are assembled into individual fuel assemblies. Gross control of the core is achieved by movable, bottom-entry control rods. The control rods are cruciform in shape and are dispersed throughout the lattice of fuel assemblies. The control rods are positioned by individual control rod drives.

Each fuel assembly has several fuel rods with gadolinia (Gd23 0 ) mixed in solid solution with the UO .2 The Gd 023 is a burnable poison which diminishes the reactivity of the fresh fuel. It is depleted as the fuel reaches the end of its first c yc le.

A conservative limit of plastic strain is the design criterion used for fuel rod cl dd.ng failure. The peak linear heat generation for steady-state operatius is well below the damage limit even late in life. Experience has shown that the control rods are not susceptible to distortion and have an O average life expectancy many times the residence time of a fuel loading.

details see Sections 4.2 and 4.3.

For 1.2.2.2.2 Reactor Vessel and Internals The Reactor Vessel contains the core and supporting structures; the steam separators and dryers; the jet pumps; the control rod guide tubes; the distribution lines for the feedwater, core sprays, and liquid control; the incore instrumentations; and other components. The main connections to the vessel include steam lines, coolant recirculation lines feedwater lines, control rod drive and other in-core housings, core spray lines, liquid control lines, dif ferential pressure line, jet pump and fluid level instrumentation, fI 56 and control rod drive hydraulic return lines (capped).

The Reactor Vessel is designed and fabricated in accordance with applicable codes for a pressure of 1250 psig. The nominal operating pressure in the steam space above the separators is 1025 psig. The vessel is fabricated of low alloy steel and is clad internally with stainless steel (except for the top head and those nozzles which connect to carbon steel pipes).

The reactor core is cooled by demineralized water that enters the lower portion of the core and boils as it flows upward around the fuel rods. The steam leaving the core is dried by steam separators and dryers located in the upper portion of the Reactor Vessel. The steam is then directed to the turbine through the main steam lines. Each steam line is provided with two isolation valves in series, one on each side of the Containment barrier. For details see Section 5.5.5.

i Os 56 1.2-4 Am. No. 56, (3/81) l- . _ -

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V' a) Main steam lines b) Reactor Water Cleanup (RWCU) System c) Residual Heat Removal (RHR) System d) Reactor Core Isolation Cooling (RCIC) System e) Instrument li ne s Small leaks generally are detected by temperature and pressure changes, fillup rate of drain sumps, and fission product concentration inside the Containment. Large leaks are also detected by changes in reactor water level and changes in flow rates in process lines. For details see Section 5.2.7.

1.2.2.3 Nuclear Safety Systems and Engineered Safety Features 1.2.2.3.1 Reactor Protection System The Reactor Protection System initiates a rapid, automatic shutdown (scram) of the reactor. It acts in time to prevent excessive fuel cladding damage and any nuclear system process barrier damage following abnormal operational transients. The Reactor Protection System overrides all operator actions and process controls and is based on a fail safe design philosophy tha* allows f- , appropriate protective action even if a single failure occurs. For details see Section 7.2.

(v) 1.2.2.3.2 Neutron Monitoring System Although not ali portions of the Neutron Monitoring System qualify as a nuclear sa fety system, those that provide high neutron flux signals to the Reactor Protection System do. The Intermediate Range Monitors (IRM) and Average Power Range Monitors ( APRM), which monitor neutron flux via in core detectors, signal the Reactor Protection System to scram in time to prevent excessive fuel cladding damage as a result of overpower transients. For details see Section 7.6.

1.2.2.3.3 Control Rod Drive System When a scram is initiated by the Reactor Protection System, the Control Rod Drive System inserts the negative reactivity necessary to shut down the

~

reactor. Each control rod i. controlled individually by a hydraulic control unit. When a scram signal is received, high pressure water stored in an accumulator in the hydraulic control unit forces its control rod into the core. For details see Section 4.6.1.1.

l 56 1.2.2.3.4 Control Rod Drive Housing Supports Control rod drive housing supports are located underneath the reactor vessel near the control rod housings. The supports limit the travel o f a control rod in the event that a control rod housing is ruptured. The supports prevent a (f-^3)'

nuclear excursion as a result of a housing failure and thus protect the fuel

's / barrier. For details see Sections 4.5.3 and 4.6.1.2. 56

1. 2- 6 Am. No. 56, (3/81)

ACNGS-PSAR 1.2.2.3.5 Control Rod Velocity Limiter A control rod velocity limiter is attached to each control rod to limit the l velocity at which a control rod can fall out of the core should it become detached from its control rod drive. This action limits the rate of teactivity insertion resulting from a rod drop accident. The limiters contain no moving parts. For details see Section 4. 6. 2. 3.1.9.

l 56 1.2.2.3.6 Nuclear System Pressure Relief System A Pressure Relief System consisting of safety / relief valves mounted on the main steam lines is provided to prevent excessive pressure inside the nuclear system following either abnormal operational transients or accident s. For details see Section 5.5.13.  ;

1.2.2.3.7 Reactor Core Isolation Cooling System The Reactor Core Isolation Cooling System (RCICS) provides makeup water to the Reactor Vessel when the vessel is isolated. The RCICS uses a steam-driven turbine pump unit and operates automatically in time and with sufficient ,

coolant flow to maintain adequate water level in the Reactor Vessel. For [

details see Section 5.5.6.

l.2.2.3.8 Emergency Core Cooling Systems Four Emergency Core Cooling Systems are provided to maintain fuel cladding below fragmentation temperature in the event of a loss of reactor coolant.

The systems are:

High Pressure Core Spray (HPCS) System Automatic Depressurization System (ADS)  :

Low Pressure Core Spray (LPCS) System Low Pressure Coolant Injection (LPCI), an operating mode of the Residual Heat Removal System a) High Pressure Core Spray System

( The HPCS System provides and maintains an adequate coolant inventory L inside the Reactor Vessel to maintain fuel cladding temperatures below fragmentation teraperature in the event of breaks in the Reactor Coolant Pressure Boundary. The system is initiated by either high pressure in the drywell or low water level in the Reactor Vessel. It operates independently of all other systems over the entire range of pressure

! dif ferences from greater than normal operating pressure to zero. The HPCS cooling can decrease Reactor Vessel pressure to enable the Low Pressure Cooling System to function. The HPCS system pump motor is powered by a diesel generator if auxiliary power is not available. The

system may also be used as a backup for the RCIC System.

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, matica11y restarted on the Standby Power System. See Section 6.2.3 for system details.

1.2.2.3.17 ECCS Area Filtered Exhaust System The ELCS Area Filtered Exhaust System will establish and maintain a nega-i tive pressure in the ECCS area within the Auxiliary Building following a LOCA and ensure that all airborne leakage from the ECCh area is filtered through IIEPA and charcoal filters before release to the environment. The ECCS area includes that portion of the Auxiliary building which bouses the i ECCS pumps, RilR heat exchangers and the penetration areas. See Section 6.5 for system details.

1.2.2.3.18 Onsite Standby AC Power Supply l The onsite standby power supply will consist of three diesel generator sets, each one serving its respective safety related loads (Divisions 1, 2, 37(U) or 3), their attendant air starting and fuel supply system and automatic control circuitry. Any two of the three diesel generatora and their asso-ciated load groups will be capable of providing a safe and orderly shutdown i of the plant and/or mitigating the consequences of a LOCA. The Standby AC Power System will be seismic Category I. See Section 8.3 for details.

! 1.2.2.3.19 DC hs der Supply

There will be six electrically independent and separate 125 v de load 37(U)

groups. These groups will supply power for plant control and instrumenta-tion and for operation of small de motor operated equipment in their re-spective non-safety related load groups and safety related Divisions 1, 2, <

3, and 4. See Section 8.3 for details.

37(U) 1.2.2.3.20 Standby Liquid Control System Although not intended to provide prompt reactor shutdown, like the control rods, the Standby Liquid Control System provides a redundant, independent, and different way to bring the reactor to subcriticality and to maintain subcriticality as the reactor als. The system makes possible an orderly

+ and safe shutdown in the co Lhat not enough control rods can be inserted into the reactor core to accomplish shutdown in the normal manner. The system is sized to counteract the positive reactivity effect over the range i

of rated power to the cold shutdown condition. See Section 4.6.6 for de-tails.

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! 1.2.2.3.21 Safe Shutdown from Outside the Control Room l

l Auxiliary control panels located within the Control Building but outside 37(U) r the Control Room will be available to operate those items of equipment re-quired to bring the units to and maintain them in a hot standby condition.

it will also be possible to reach a cold shutdown through the use of the auxiliary control panels and other local controls.

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l (U)-Update 1.2-12 Am. No. 56, (3/81)

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I TABLE 1.6-1 REFERENCED REPORTS A. General Electric Company Reports >

i SAR Sections Report where Number Title Re ferenced

< APED-4827 Maximum Two-Phase Vessel Blowdown 5.2 from Pipes ( April 1965)

APED-5286 Design Basis for Critical Heat Flux in 16.1 Boiling Water Reactors (September 1966)

APED-5454 Metal Water Reactions - Ef fects on Core 1.5 Cooling and Containment (March 1968)

APED-5458 Ef fectiveness of Core Standby Cooling 1.5, 5.5 Systems for General Electric Boiling Water Reactors (March 1968) lj [

APED-5460 De sign and Performance of General 5.2

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Electric Boiling Water Reactor Jet Pumps (September 1968)

APED-5529 Core Sprav and Core Flooding Heat 6.3 Transfer Effectiveness in a Full-Scale Boiling Water Reactor Bundle (June 1968) 1 APED-5555 Impact Testing on Collet Assembly for 4.6 56 Control Rod Drive Mechanism 7RDB144A (November 1967)

APED-5608 General Electric Company Analytical and 1.5 Experimental Programs for Resolution of ACRS Safety Concerns (April 1968) 56 APED-5640 Xenon Considerations in Design of Large 4.1, 4.3 Boiling Water Reactors (June 1968)

APED-5654 Considerations Pertaining to Containment 1.5 Inerting ( August 1968)

I APED-5698 Summary of Results Obtained From a Typi- 1.5, 16.2 cal Startup and Power Test Program for a ps General Electric Boiling Water Reactor (February 1969)

)

. APED-5706 In-Core Neutron Monitoring System for 7.6, 16.4 General Electric Boiling Water Reactors (November 1968; revised April 1969)

1.6-2 Am. No. 56, (3/81)

\ ACNGS-PSAR TABLE 1.6-1 (Cont'd)

SAR Sections Report where Number Title Re ferenced APED-4986 Consequences of Operating Zircaloy-2 Clad 4.2 Fuel Rods Above the Critical Heat Flux 56 (October 1965)

APED-5640 Xenon Considerations in Design of Boiling 4.1 Water Reactors (June 1968)

APED-5736 Guidelines for Determining Sa fe Test Appendix B Intervals and Repair Times for Engineered Safeguards (April 1969)

APED-5750 Design and Performance of General 1.5, 5.5 Electric Boiling Water Reactor Main Steam Line Isolation Valves (March 1969) p APED-5756 Analytical Methods for Evaluating the 1. 5 , l 'i .1 Radiological Aspects of the General

(~ Electric Boiling Water Reactor (March 1969)

GEAP-4059 Vibration in Fuel Rods in Parallel Flow 3.9, 4.2 (July 1962)

GEAP-13112 Thermal Response and Cladding Performance 4.2 of an Internally Pressurized, Zircaloy Clad, Simulated BWR Fuel Bundle Cooled by Spray Under Loss-of-Coolant Conditions (April 1971) 56 GEAP-4616 Two-Phase Pressure Drop in Straight Pipes 4.4 and Channels; Water-Steam Mixtures at 600 to 1400 psia (May 1964)

! GEAP-4966 Vibration of SEFOR Fuel Rods in Parallel 3.9 Flow (September 1965)

GEAP-5620 Failure Behavior in ASTM A106B Pipes 5.2 Containing Axial Through-Wall Flows (April 1968)

GEAP-10117 Response of a Simulated BWR Fuel Bundle 6.3 Cooled by Flooding Under Loss-of-Coolant (December 1969)

GEAP-13197 Emergency Cooling in BWR's Under Simulated 6.3 Lo s s-o f-Coolan t (BWR FLECHT Final Report)

(June 1971)

1. 6- 3 Am. No. 56, (3/81)

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TABLE 1.6-1 (Cont'd)

SM fections Report 2*are Number Title Re fr renced i

GEAP-1054 Theory Report for Creep-Plast Computer 4.1 Program, (January, 1972) 56 i

NEDO-10029 An Analytical Study on Britt? Fracture 1.5 of GE-CBWR Vessel Subject to the Design Basis Accident (July 1969)

I

NEDO-10045 Consequences of a Steam Line Break in 1.5 a General Electric Boiling Water Reactor (July 1969) 4 NEDO-10139 Compliance of Protection Systems to 7.2 Industry Criteria

General Electric BWR

Nuclear Steam Supply System (June 1970) i NEDO-10173 Current State of Knowledge, High Per- 1.5, 4.2* 56 l formance BWR Zircaloy-Clad UO2 Fuel 11.1 j

(May 1970)

NEDO-10174 Consequences of a Postulated Flow 1.5, 4.2 56 Blockage Incident in A Boiling Water q

Reactor (May 1970)

NEDO-10179 Ef fects of Cladding Temperature and 1.5, 6.3 Material on ECCS Performance (June 1970)

NEDO-10189 An Analysis of Functional Common-Mode 1.5 Failures in GE BWR Protection and Control Instrumentation (July 1970)

NEDO-10208 Effects of Fuel Rod Failure on ECCS 1.5 Performance (August 1970) l56 NEDO-10320 The General' Electric Pressure Sup- 6.2 pression Containment Analytical Model (April 1971), Supplement 1 (May 1971)

J NED0-10329 Loss-o f-Coolant Accident and Emergency 1.5, 5.2, Core Cooling Models for General 6.2, 6.3 Electric Boiling Water Reactors (April 1971) Supplement 1 (April 1971) Addenda (May 1971)

O NEDO-10349 Analysis of Anticipated transients 1.5, 15.5 Without Scram (March 1971) 1.6-4 Am. No. 56, (3/81)

- _ - . . _ _ _ _ _ _ - _ _ . , _ _ - . _ _ _ . ~ . . . _ . . _ _ , _ . . . - _ . _ - . _ _ . . __

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l ACNGS-PSAR TABLE 1.6-1 (Cont'd)

SAR Sections Report where Number Title Re ferenced NEDD-10505 2xperience with BWR Fuel Through 4.2, 11.1 September 1971 (May 1972)

I NEDO-10527 Rod Drop Accident Analysis for Large 4.3, 15.2, 4 Boiling Water Reactors (March 1972) 15.4 56 Supplement 1 (July 1972), Supplement 2 (January 1973)

NEDO-10569A BWR/6 Nuclear System from General Electric - 1.2.2.2 A Performance Description (April, 1972) i NEDO-10571 General Electric Boiling Water Reactor 6.2 Mark III Containment Concept (April 1972)

NEDO-10585 Behavior of iodite in Reactor Water 15.4 i During Plant Shutdown and Startup (August 1972) l 56 4

NEDO-10677 Analysis of Recirculation Pump Overspeed 1.5

()g g in a Typical GE-BWR I NEDO-10678 Seismic Qualification of Class I Electric 7, 3.10, Equipment 3.11

! NEDO-10698 Environmental Qualification of Class I 7, 3.10, Control and Instrumentation Equipment 3.11 4

NEDO-10179 Effects of Cladding Temperature and 1.5, 6.3 i Material on ECCS Performance (June 1970)

16 NEDD-10189 An Analysis of Functional Common-Mode 1.5 Failures in GE BWR Protection and Control i

Instrumentation (July 1970)

! NEDO-10208 Effects of Fuel Rod Failure on ECCS 1.5

! Performance ( August 1970) 56 NED0-10320 The General Electric Pressure Suppression 6.2 Containment Analytical Mode ( April 1971).

Supplement 1 (May 1971)

, NEDO-10329 Loss-of-Coolant Accident and Emergency 1.5, 5.2,

! Core Cooling Models for General Electric 6.2, 6.3 Boiling Water Reactors ( April 1971)

/'~'\ Supplement 1 (April 1971) Addenda (May 1971)

NEDO-10349 Analysis of Anticipated Transients Without 1.5, 15.5

.. Scram (March 1971)

1. 6- 5 Am. No. 56, (3/81)

ACNGS-PSAR TABLE 1.6-1 (Cont'd)

SAR Sections Re po rt where Number Title Re ferenced Experience with BWR Fuel Through September 11.1 56 NEDO-10505 1971 (May 1972)

NEDO-10527 Rod Drop Accident Analysis for Large 15.2, 15.4 Boiling Water Reactors (March 1972) l56 NEDO-10585 Behavior of Iodine in Reactor Water 15.4 During Plant Shutdown and Startup (August 1972)

NEDO-10734 A Ceneral Justification for Classification 11.2, 11.3 of Effluent Treatment System Equipment as Group D (February 1973)

NEDO-10751 Experimental and Operational Confirmation 11.3 of Of fgas System Design Parameters (January 1973) (Company Proprietary) 16 NEDO-10801 Modeling the BWR/6 Loss-of-Coolant Accident: 1.5 Core Spray and Bottom Flooding Heat Trans-fer Effectiveness (March 1973)

NEDO-10846 BWR Core Spray Distribution (April 1973) 1.5 NEDO-11013 An Analytical Procedure for the conserva- 1.5 tive calculation of Core Metal-Water Reaction following a Design Basis Loss-of-Coolant Accident NEDM-10735 Densification Considerations in BWR Fuel 4.2 Design and Performance (December 1972)

I NEDO-10802 Analytical Methods of Plant Transient 4.4 Evaluations for General Electric Boiling Water Reactor, (February 1973)

) NEDO-10958 General Electric Thermal Analysis Bases 4.2 (CETAB): Data, Correlation, and Design l Application, General Electric Co., 56 November 1973 NED0-20340 Process Computer Performance Evaluation 4.3 Accuracy (June 1974) f\ NEDE-20606 Creep Collapse Analysis of BWR Fuel Using Safe Collapse Model (August 1974) 4.2 SQ NEDO-20605 l NEDO-20377 8x8 Fuel Bundle Development Support 4.2 (February 1975)

1. 6- Sa Am No. 56, (3/81)

ACNGS-PSAR TABLE 1.6-1 (Cont'd)

SAR Sections Report where Number Title Re ferenced NEDO-20939 Lattice Physics Methods Verification 4.3 (August 1975)

NEDO-20964 Generation of Void and Doppler Reactivity 4.3 Feedback For Application to BWR Plant Transient Analysis: (August 1975)

NED0-20994 Peach Bottom Atomic Power Station Units 4.4 2 and 3, Safety Analysis Report for Plant Modifications to Eliminate Significant In-Core Vibration (September 1975)

NED0-21156 Supplemental Information for Plant Modifi- 4.4 56 cation to Eliminate Significant In-Core Vibracion (January 1976)

NEDO-20566 General Electric Company Model for Loss-of- 4.3 p) f,

%d Coolant Accident Analysis In Accordarce with 10CFR50 Appendix K, (January 19 6)

NEDO-20360 General Electric Boiling Water Reactor 4.2 Generic Reload Application for 8x8 Fuel (March 1976)

NEDO-20360-1P General Electric Boiling Water Reactor 4.2 Generic Reload Application for 8x8 Fuel (March 25, 1976)

NEDO-20953 3D BhR Core Simulator (May 1976) 4.3 NEDO-20948-P BWR/6 Fuel Design (June 1976) 4.2 NEDO-10722A Core Flow Distribution in a Modern 4.4 Boiling Water Reactor as Measured in Monticello (August 1976)

NEDE-21354-P BWR Fuel Channel Mechanical Design and 4.2 NEDO-21354 Deflection (September 1976)

NEDO-21231 Banked Position Withdrawal Sequence 4.3 (September 1976)

NEDE-21175-P BWR/6 Fuel Assembly Evaluation Combined 4.2

'SSE and LOCA Loadings (November 1976)

\ NEDE-20943 Urania-Gadolinia Nuclear Fuel Physical and 4.2 NED0-20943 Irradiation Characteristics and Material Properties (January 1977)

1. 6- Sb Am. No. 56, (3/81)

ACNGS-PSAR FA BLF. 1.6-1 (Cont'd)

SAR Sectione Report where Number Title Referenced NEDO-2tt'844-1 HWR/4 and BWR/S Fuel Design Am. ndment 4.2 Gen.ral Electric Co., (January 1977)

N E DO- 10'3 "> 8 A Con ral Electric Thermal Analysis Basis 4.3, 4.4 (GE rAB ): Data, Corr 'lation, and Design Application (January 1977)

NEDO-21506 Stability and Dynamic Performance of the 4.1, 4.4 Ceneral Electric Bailing Water Reactor (January 1977)

NED0-20946A BWR Simulator Methods Verification 4.3 (January 1977)

NED0-20913A Lattice Physics Methods (February 1977) 4.3 56 NEDE-23542-P Fuel Assembly Evaluation of Shipping and 4.2 NE00-23542 Handling Loadings (March 1977)

(-

NEDO- 23786- 1 Fuel Rod Prepressurization Amendment 1 4.2, 4.4 (May 1978)

NED0-23786-1-P Fuel Rod Prepressurization (March 1978) 4.2, 4.4 NED0-24154 Qualification of the One-dimensional Core 4.3 Transient Model for Boiling Water Reactors (October 1978)

Bailly Generating Station Nuclear 1, PSAR 1.5 Amendment 13 Browns Ferry Nuclear Power Station, Units 1 1.5 2 and 3, PSAR Amendment 15 Brunswick Steam Generating Plant, Units 1 1.5 and 2, PSAR, Supplements 3, 4 and 6.

Summary Memorandum on Excursion Analysis 15.1

Uncertainties, Dresden Nuclear Power l Station, Unit 3, Plant Design Analysis Report, Amendment 3 Dresden Nuclear Power Station, Unit 3, 1.5 Special Report No. 14 O

h Dresden Nuclear Power Station, Units 2 and 3, PSAR Amendments 7, 8, 14 and 15 1.5 33(U)

(U)-Upda te

, 1. 6- 6 Am. No. 56, (3/81) 1 I , . , .- _ _ . _.___ _ _

- _ _ - = _ , - . -

ACNGS-PSAR TABLE 1.6-1 (Cont'd)

SAR Sections Report where Number Title Re ferenced Dresden Nuclear Power Station, Units 2 and 6.3 3, PSAR Amendments 7 and 8 Hatch Nuclear Plant, Unit 1, FSAR Amendment 15.1 56 10, Appendix L Oyster Creek Nuclear Power Station Unit 1 1.5 FSAR Amendment 10 Millstone Nuclear Power Station, 6.3 PSAR Amendment 14 Pilgrim Nuclear Power Station, 6.3 PSAR Amendment 14 i

B. Other Reports l 56 BRH/ DER 70-1 Radiological Surveillance Studies at 11.1 a Boiling Water Nuclear Power Reactor

\ (March 1970)

CF 59-6-47 Removal of Fission Product Cases from 11.3 (ORNL) Reactor Of f-Cas Streams by Adsorption (June 11, 1959)

UCRL-50451 Improving Availability and Readiness 16.3 of Field Equipment Through Periodic Inspection, p. 10 (July 16, 1968)

WACP-6065 Melting Point of Irradiated Uranium 4.2 Dioxide (February 1965)

WAPD-TM-283 Effects of High Burnup on Zircaloy- 4.2 1 56 Clad, Bulk UO2 Plate Fuel Element Samples (September 1962)

WAPD-TM-629 Irradiation Behavior of Zircaloy-Clad 4.2 Fuel Rods Containing Dished End UO2 Pellets (July 1967)

ETR-1001 Ebasco Nuclear Quality Assurance 17.0 33(U)

Program Manual

-pae

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ACNCS-PSAR APPENDIX 2.2-A PIPELINE BREAK EVALUATION 1.0 _ BREAK OF 6" LPG LINE FILLED WITH PROPANE The consequences of a complete severance of the 6 inch LPG line, assumed to l be pipine propane have been evaluated on the basis of the following assump-tions:

a) Double ended rupture of the line occurs instantaneously and at the a closest point to plant Category I st ructures ( 7,000 feet).

b) "Ihe released petroleum liquid-gas mixture escapes from the break at the critical velocity for two phase flow, and at the design pressure i of the line, 1,000 psig (a conservative assumption since the oper-atina pressure is only 750 psi).

c) The temperature of the atmosphere is assumed to be 72 F. Higher temperatures would lead to higher vaporization of escaping propane, 41

. but the flow rate would be less due to the higher quality at the Q 312.6 exit plane.

d) Five percentile meteorology is assumed, which is equivalent to a Pasquill F inversion with wind speed of 0.8 mps in the direction I O 1.1 of the plant structures.

CALCULATION OF FLOW RATE OUT OF THE BREAK The propane in the line will, upon the instant of the break, deumpress isenthalpically to a saturation pressure of 125 psia immediately because of the very large speed of sound in the liquid. A decompression wave will travel very rapidly away from the break leaving the fluid behind at the saturation pressure. Since propane would issue - from the break at 72 F, approximately 1/3 of it would quickly vaporize, cooling the remainder to its boiling point of about - 44 F. Hence the process of decompression >

is described by the throttling process shown in Figure 1.1 . From that figure the exit plane quality, x, of the fluid can be estimated from:

v=v +xy l48(U) f f8 '56 3 3 3 where v=2.4 f t /lb, v =.0275 ft /lb, v =6.6 ft /lb, v fg

=v y -v f a f

.l Hence x = 0.36 To estimate the flow rate out of the break, Fauske's equation 1/ for critical two-phase mass velocity is used:

Cg = (-a/(k dvg /dp+k dx/dp+k 2 3 d"f

/d P))

1 A

(U)-Update 2.2.A-1 Am. No. 56, (3/81)

! ACNGS-PSAR g'

In reality the upstream pressure will decrease with time with a correspond-inn decrease in the flow rate out of the break. Furt t.ar because of friction effects, the maxi um flow rat e out of the break will be lower, hence the model assumed is known to be conservative. The followine figure compares the model used with what is expected in reality.

3.2 CALCULATION OF DETONABLE CLOl'D SIZE ne dimension of the detonable plume downwind of the break have been evaluated for a Category F stability, and a constant, invariant wind speed of 2.6 ft/sec.

For purposer of conservation buoyancy ef fect s have been ienored in this section. Buoyancy and the jer momentum are considered in Subsection 3.5.

The centerline (directly downwind) concentration of the methane (excludine buoyancy) is determined by:

• c1 =_

Off-centerline concentrations are determined by 41 2 2 Q X=y Exp/ +

n c1 I

' _1,,_{(y/,y) 2 (t/az)) 312.5

! b V Since the flow out of the break varies with time it is necessary to discuss which flow out of the break is chosen for subsequent analysis.

Durine the initial 2.5 minutes of the event, flow out of the break is essentially sonic. Althouch a large portion of the total mass of cas is released during this time period, (i.e. 70 percent of the total quan-tity of gas emitted from the break in the 2.5 hrs, estimated to be required to terminate flow is released during the initial decompression 56 period), the velocity at the break, coupled with the buoyancy of the , gas will propel it away from the surface and part low atmospheric inversions so that less than 500 lbs. of this initial mass is calculated to fall within the fla nmable limits in the vicinity of the surface, but most of it will be dispersed in the upper levels.

To establish therefore the configuration of a credible, low lyine detonable cloud in the vicinity of the plant (ie, cloud travelline toward the plant under worst meteorological conditions), it is assumed that the flow out of the break is a constant 108 lb/sec, corresponding to the condition existine af ter the initial 2.5 minutes transient. Buoyancy was neelected in this calculation.

Re potential cloud configuration (neglectine buoyancy) is plotted in Figure 3.3a. The dashed portion represent the fraction of the cloud which f alls within the flammable limits (4.8 and 14.0 volume percent).

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EFFECTIVE FIGURE LIST

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CHAPTER 3  !

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Figure No. Amendment No.

3.8-5 34 I 3.8-6 -

3.8-7 -

3.8-8 -

3.8-9 -

1 3.8-11 54 li 3.8-12 26 l 3.8 26 i

3.8-14 26 l I

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!' 3.8-16 26 j 3.8-17 35  !

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3.9-3 56 i 3.9-4 56 i

! 3.9- 5 56  ;

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56 l 3.9-9 56 l l 3.9-10 56 ,

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3.9-12 56 3.9-13 56 3.11-1 17 '

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• , All Figures whether labelled " Unit 1" or " Units 1 & 2" are to be con-sidered applicable to Unit No. 1.

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18 Am. No. ~ 56, (3/81)

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ACNGS-PSAR TABLE OF CONTENTS

}

V CHAPTER 3 (Cont'd)

Section Title Page 3.9.3 COMPONENTS NOT COVERED BY ASME CODE 3.9-6 3.9.3.1 General 1.9-6 3.9.3.2 Fuel Mechanical Design and Analytical Procedures 3.9-7 3.9.3.3 Control Rod Drive Operability and Control Rod Insertability Under LOCA and Seismic Loadings 3.9-7 3.9.4 REACTOR CORE SUPPORT STRUCTURES AND INTERNALS 3.9-9 MECHANICAL DESIGN 3.9.4.1 Design Bases 1.9-9 3.9.4.1.1 General Design Bases 3.s-9 3.9.4.1.1.1 Safety Design Bases 3.9-9 3.9.4.1.1.2 Power Generation Design Bases 3.9-9 O)

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3.9.4.1.2 Specific Design Characteristics 3.9-9 3.9.4.1.2.1 Design Loading Combinations 3.9-9 3.9.4.1.2.2 Stress, Deformation and Fatigue Limits for Reactor Internals 3.9-9 56 3.9.4.1.2.3 Stress, Deformation and Fatigue Limits for Core Support Structures 3.9-10 3.9.4.1.2.4 . Fuel Assembly Restraints 3.9-10 3.9.4.1.2.5 Material Selection 3.9-10 3.9.4.1.2.6 Radiation Ef fects 3.9-10 i

3.9.4.1.2.7 Shock Loads 3.9-10 3.9.4.2 Description 3.9-10a 3.9.4.2.1 Core Support Structure 3.9-10a i

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TABLE OF CONTENTS CHAPTER 3 (Cont 'd)

Section Title Pajgt 3.9.4.2.1.1 Core Shroud 3.9-10b 3.9.4.2.1.2 Shroud Head and Steam Assembly 3.9-10b 3.9.4.2.1.3 Core Support Plate 3.9-10b f 3.9.4.2.1.4 Top Guide 1.9-10b 3.9.4.2.2 Fuel Support 3.9-10c i 3.9.4.2.3 Control Rod Guide Tubes 3.9-10e 3.9.4.2.4 Jet Pump Assemblies 3.9-10e i

3.9.4.2.5 Steam Dryers 3.9-10d 56 3.9.4.2.6 Feedwater Spargers 3.9-10d t 3.9.4.2.7 Core Spray Lines 3.9-10d s

3.9.4.2.8 Vessel Head Cooling Spray Nozzle 3.9-10e

, 3.9.4.2.9 Dif ferential Pressure and Liquid Control Line 3.9-10e 3.9.4.2.10 In-Core Flux Monitor Guide Tubes 3.9-10e .

t 3.9.4.2.11 Surveillance Sample Holders 3.9-10e 3.9.4.2.12 Low Pressure Coolant Injection Lines 3.9-10f 3.9.4.3 Safety Evaluation 3. 9-10 f 3.9.4.3.1 Evaluation Methods 3.9-10f 3.9.4.3.1.1 Input for Safety Evaluation 3.9-10f 3.9.4.3.1.2 Events to be Evaluated 3.9-10f 3.9.4.3.1.3 Pressure Dif ferential During Rapid Depressurization 3.9-10g 3.9.4.3.2 Recirculation Line and Steam Line Break 3.9-10g i

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i ACNGS-PSAR TABLE OF CONTENTS l CHAPTER 3 (Cont'd)

Section Title Page 3.9.4.3.2.1 Accident Definition 3.9-10g

) 3.9.4.3.2.2 Effects of Initial Reactor Power and Core Flow 3.9-10h  !

l 3.9.4.3.2.3 -Break Size Spectrum Analysis 3.9-10i 1 3.9.4.3.2.4 Conclusions 1.9-10i 3.9.4.3.2.5 Response of Structures Within the Reactor Vessel to Pressure Difference 3.9-10i i 56 3 3.9.4.3.3 Earthquake 3.9-10k

i 3.9.4.3.4 Conclusions 3.9-10m 3.9.4.4 Inspection and Testing 3.9-10n

3.9 REFERENCES

3.9-39 APPENDIX 3.9. A 3.9A-1 a

APPENDIX 3.9.B 3.9B-1 1

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ACNGS-PSAR CHAPTER 3 LIST OF TABLES (Cont 'd)

Table Title Page 3.7.A-7C Comparison of Shears and Moments Containment 4g Spring -- a vs Flush - a 3.7.A-16 (U) 3.7.A-7D Comparison of Shears and Moments Drywell Spring - a vs Flush - a 3.7.A-17 3.7.A-8 Deconvoluted Zero Period Acceleration 3.7.A-18 3.8-1 Stress Limits for Containment Vessel 3.8-79 46 3.8-2 Buckling Criteria for Containment Vessel 3.8-81 (C) 3.8-3 Load Combinations and Load Factors 3.8-82 3.8-4 Load Combinations for Cuard Pipes - Deleted 3.8-83 3.9-1 ASME Code Class 2 and 3 Components 3.9-11 s-- ) 3.9-2 Design Loading Combinations for ASME Code Class 2 and 3 Components 3.9-16 3.9-3 Allowable Stresses for ASME Class 2. and 3 Components 3.9-17 3.9-4 Safety 'Related Components Not Covered by ASME Code 3.9-19 3.9-5 Applicable Codes and Standards for Heating, Ventilating and Air Conditioning Systems I and Components 3.9-22a 3.9-6 Fracture Toughness Requirements for Code Class 2 and 3 Components 3.9-22b 3.9-7 Design Loading Combinations for Class 1,2,3, MC Linear Type Supports (EBASCO design) 3.9-22g 41 4

3.9-8 Stress Categories and Stress Limit Factors for Class 1,2,3. and MC Linear Type Supports Designed By Elastic Analysis 3.9.22h

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(U)-Upda te (C)-Consistency xxxitsa Ama No. 56, (3/81)

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i ACNGS-PSAR

[) CHAPTER 3 LIST OF TABLES (Cont'd)

Table Title Page

3.4-9 Deformation Limits 3.9-23 3.9-10 Primary Stress Limits 3.9-24 3.9-11 Buckling Stability Limit 3.9-26 3.9-12 Fatigue Limit 3.9-27 3.9-13 Core Support Structures Stress Categories and Limits 56 i of Stress Intensity for Normal and Upset Conditions 3.9-28 3.9-14 Core Support Structures Stress Categoties and Limits of Stress Intensity for Emergency Conditions 3.9-31 ,

3.9-15 Core Support Structures Stress Categories and Limits i,

of Stress Intensity for Faulted Conditions 3.9-34 3.9-16 Design Loading Conditions and Combinations 3.9-37 3.9-17 Pressure Differentials Across Reactor Vessel Internals 3.9-38 3.10-1 Typical Vendor Supplied Class IE Devices 3.10-5 3.11-1 Accident Environment - Containment (Inside Drywell) 3.11-8 3.11-2 Accident Environment - Containment

(Outside Drywell) 3.11-10 l 3.11-3 Environmental Conditions - Normal Plant l Operation and Post-Accident Radiation 4

Exposures 3.11-14 3.11-4 Accident Basis Envelope 3.11-19 i

3.11-5 NOTES FOR TABLES 3.11-2 and 3.11-3 3.11-21

3.11-6 Environmental Conditions Upon Loss of Air Conditioning 3.11-23 f

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l j ACNGS-PSAR CHAPTER 3 1 LIST OF FIGURES (Cont'd)

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l Figure Title j 3.9-3 Reactor Vessel Cutaway c

i 3.9-4 Reactor Internals Flow Paths i 3.9-5 Steam Separator  !

, c

4. 3.9-6 Fuel Support Pieces -

i 3.9-7 Jet Pump 56 i

3.9-8 Steam Dryer t i

3.9-9 Pressure Nodes Used for Depressurization Analysis I

3.9-10 Power Flow Map

! 3.9-11 Maximum Reactor Internal Pressure Loais as a Function of Steam Line Break Area  ;

i 3.9-12 Transient Pressure Differentials Following a Steam Line Break at the 105% Rated Steam Flow Condition 3

3.9-13 Seismic Mathematical Model for the Reactor Pressure Vessel and Internals i

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xxxix ,im, No. 56, (3/81)

ACMGS-PSAR assurance program is to assure sound engineering in all phases of design

('~')j and construction through conformity to regulatory requirements and design

( bases described in the license application. In addition, the program assures adherence to specified standards of workmanship and implementation of recognized codes and standards in fabrication and construction. It also includes the observance of proper preoperational and operational testing and maintenance procedures as well as the documentation of the f o'regoing by keeping appropriate records. The total quality assurance program of the applicant and its principal contractors is responsive to and satisfies the quality-related requirements of Title 10CFR50, including Appendix 3.

S t ruc t ure s , systems, and components are first classified in Chapter 3 with respect to their location and service and their relationship to the safety f unction to be pe rf ormed . Recognized codes and standards are applied to the equipment in these classifications as necessary to assure a quality produc t in keeping with the required safety function. In cases where codes are not available or the existing code must be modified, a rigorous expla-nation is provided.

Documents are maintained which demonstrate that all the requirements of the quality assurance program are being satisfied. This documentation shows that appropriate codes, standards and regulatory requirements are observed, specified materials are used, correct procedures are utilized, qualified personnel are provided and that the finished parts and components meet the applicable specifications for safe and reliable operation. These records are available so that any desired item of information is retrievable for 7~ reference. These records will be maintained during the life of the oper-

/ ) ating licenses.

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The detailed quality assurance program developed by the applicant and its contractors satisfies the requirements of Criterion 1.

For further discussion, see the following sections:

a. Principal Design Criteria 1.2

b. Plant Description 1.2

c. Design of Structures, Components, and Systems 3.0

d. Fuel Mechanical Design 4.2 l 56

e. Reactivity Control System 4.6

f. Nuclear Design 4.3

g. Thermal and Hydraulic Design 4.4 l35(C)

h. Control Rod Drive Housing Supports 4.5

1. Overpressurization Protection 5.2

j. Reactor Vessel and Appurtenances 5.4

k. Reactor Recirculation System 5.5

1. Main Steam Line Flow Restrictor 5.5

m. Main Steam Line Isolation Valves 5.5

n. Reactor Core Isolation Cooling System 5.5

o. Residual Heat Removal System 5.5

p. Containment Systems 6.2

q. Emergency Core Cooling Systems 6.3 f ~3 r. Reactor Protection System 7.2 k )

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(C)-Consistency

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3.1-2 Am. No. 56,"(3/8'l)

l ACNGS-PSAR a) Good load-following with a well-damped behavior and little under-s hoot or overshoot in the heat transfer response.

b) Load-following with recirculation flow control.

c) Strong damping of spatial power disturbances.

l The Reactor Protection System design provides protection from excessive fuel cladding temperatures and protects the Reactor Coolant Pressure Boun-dary f rom excessive pressures which threaten the integrity of the system.

Local abnormalities are sensed, and, if protection system limits are reach-ed, corrective action is initiated through an automatic scam. High in-tegrity of the protection system is achieved through the combination of logic arrangement, trip channel redundancy, power supply redundancy, and physical separation.

The reactor core and associated coolant, control and protection systems are designed to suppress any power oscillations which could result in ex-ceeding fuel design limits. These systems assure that Criterion 12 is met.

For further discussion see the following sections:

a. Principal Design Criteria 1.2

b. Reactivity Control System 4.6 l 56

c. Nuclear Design 4.3

d. Thermal and Hydraulic Design 4.4 35(C)

, e. Nuclear System Stability Analysis 4.4

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Overpressurization Protection Reactor protection System 5.2 7.2

h. Reactor Manual Control System 7.7

1. Accident Analysis 15.0 3.1.2.2.4 Criterion 13 - Instrumentation and Control Instrumentation shall be provided to monitor variables and systems over their anticipated ranges for normal operation, for anticipated operational occurrences, and for accident conditions as appropriate to assure adequate safe ty, including those variables and systems that can affect the fission process, the integrity of the reactor core, the Reactor Coolant Pressure Boundary, and the Containment and its associated systems. Appropriate con-trols shall be provided to maintain these variables and systems within pre-scribed operating ranges.

l 3.1.2.2.4.1 Evaluation Against Criterion 13 1

The fission process is monitored and controlled for all conditions from source range through power operating range. The intermediate and power ranges of the Neutron Monitoring System detect core conditions that threat-en the overall integrity of the fuel barrier due to excess power generation and provide a signal to the Reactor Protection System. Fission detectors, located in the core, are used for neutron detection. The detectors are i located to provide optimum monitoring in the intermediate and power ranges.

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(C)-Consistency 3.1-9 Am. No. 56, (3/81)

ACNGS- PS AR For further discussion. see the following sections:

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\ j a Principal Design Criteria 12 56

b. Reactivity Control System 4.6 c Reactor Coolant Pressure Boundary Leakage Detection System 5.2 and 7.6 d Main Steam Line Isolation Valves 5. 5

e. Containment Systems 6.2

f. Reactor Protection System 7.2

g. ESF Systems 7.3

h. Safety Related Display Instrumentation 7.5

i. Neutron Monitoring System 7.6

j. Reactor Vessel - Instrumentation and Control 7.6

k. Process Computer System 7.6

1. Reactor Manual Control System 7.7

m. Recirculation Flow Control System 7.7 3.1.2.2.5 Oriterion 14 - Reactor Coolant Pressure B ounda ry The Reactor Coolant Pressure Boundary shall he designed, f abricated, erect-ed, and tested so as to have an extremely low probability of abnormal leakage, of rapidly propagating failure, and of gros s rupture.

3.1.2.2.5.1 Evaluation Against Criterion 14 0)

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The piping and equipment pressure parts within the Reactor Coolant Pressure Boundary through the outer isolation valve (s) are designed, fabricated, erected, and tested to provide a high degree of integrity throughout the plant li fe t i ce . Chapter 3 classifies and discusses systems and components within the Reactor Coolant Pressure Boundary. The design requirements and codes and standards applied to the Reactor Coolant Pressure Boundary ensure a quality product in keeping with the safety functions to be performed.

In order to minimize the possibility of brittle fracture within the Reactor Coolant Pressure Boundary, the fracture toughness properties and the oper-ating temperature of ferritic materials are controlled to ensure adequate t ou ghne s s. Section 5.2.3, " General Material Considerations," describes the l 35 methods utilized to control toughness properties. Materials are to be (U) impact tested in accordance with ASME Boiler and Pressure Pessel Code,Section III. Where Reactor Coolant Pressure Boundary piping penetrates the Containment, the fracture toughness temperature requirements of the Reactor Coolant Pressure Boundary materials apply.

Piping and equipment pressure parts of the Reactor Coolant Pressure Bound-ary are assembled and erected by welding unless applicable codes permit flanged or screwed joints. Welding procedures are employed which produce welds of cociplete fusion and free of unacceptable defects. All welding procedures , welde rs , and welding machine operators are qualified in accord-ance with the requirements of Section IX of the ASME Boiler and Pressure Vessel Code for the materials to be welded. Qualification records, includ-O ing the results of procedure and performance qualification tests and iden-tification symbols assigned to each welder are maintained.

(U)-Update 3.1.}2 Am. No. 56, (3/81')

! ACNGS-PS AR

! system is prompt and the total scram time is short. Control rod scram I motion starts in about 170 milliseconds af ter the high flux set po int is exceeded A fully withdrawn control rod will traverse 75% of its full stroke in l sufficient time to assure that acceptable fuel design limits are not exceeded.

In addition tn the reactor protection system which provides for automatic l'

shutdown of the reactor to prevent fuel damage, protection systems are provided to sense accident conditions and initiate automatically the

operation of other systems and components important to safety. Systems j such as the emergency core cooling system are initiated automatically to i limit the extent of fuel damage following a loss-of-coolant accident.

! Other systems automatically isolate the reactor vessel or the containment i to prevent the release of significant amounts of radioactive materials from the fuel and t he reactor coolant pressure boundary. The controls and inst rumentation for the emergency core cooling systems and the isolation systems are initiated automatically when monitored variables exceed pre- 35(G) 4 selected operational limits.

The design of the protection system satisfies the functional requirements as specified in Criterion 20.

l

.i For further discussion, see the following sections:

. a) Principal Design Criteria 1.2 56 b) Reactivity Control Mechanical Design 4.6 c) Control Rod Drive Housing Supports 4.5

, d) Ove rpressurization Protection 5.2 e) Main Steam Line Isolation valves 5.5 f) Emergency Core Cooling System 6.3

) g) Reactor Protection System 7.2 i

h) Containment and Reactor Vessel Isolation Control 7.3 l i S ys t em

, i) Emergency Core Cooling Systems - Instrumentation 7.3 and Control I

j j) Neutron Monitoring System 7.6 k) Process Radiation Monito-ing System 7.6

1) Leak Detection System 7.6 l

4 m) Accident Analysis 15.0 i

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(G)-GESSAR 3.1- 19 Am. No. 56, '3 ( 81')

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i ACNGS-PS AR control room i ns t rument at ion. More importantly, the hydraulic control unit scram accumulator and the scram discharge volume level are continuously monitored.

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The main steam line isolation valves may be tested during full reactor operation. Individually, they can be closed to 90% of full open position i

without af fect ing the reactor operation. If reactor power is reduced sufficiently, the isolation valves may be fully closed. Provisions are provided to evaluate valve stem leakage during reactor shutdown. During 56 refueling operation, valve leakage rates can be determined.

Residual heat removal system testing can be performed during normal operation. Main system pumps can be evaluated by taking suction from the suppression pool and discharging through test lines back to the suppression pool. System design and operating procedures also permit testing the discharge valves to the reactor recirculation loops. The low pressure 35(G) coolant injection mode can be tested after reactor shutdown. Each active l component of the emergency core cooling systems provided to operate in a j design basis accident is designed to be operable for test purposes during normal operation of the nuclear system.

The high functional reliability, redundancy, and inservice testability of the protection system satisfy the requirements specified in Criterion 21.

For further discussion, see the following sections:

a) Principal Design Criteria 1.2 N

b) Reactivity Control System 4.6 c) Main Steam Line Isolation Valves 5.5 d) Residual Heat Removal System 5.5 e) Containment Systems 0.2 i

f) Emergency Core Cooling Systems 6.3 i g) Reactor Protection System 7.2 i

h) Containment and Reactor Vessel Isolation 7.3 l Control System i) Emergency Core Cooling Systems - Instrumentation and Control 7.3 j) Neutron Monitoring System 7.6 k) Process Radiation Monitoring System 7.6 l 1) Leak Detection System 7.6 l

m) Accident Analysis 15.0

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(G)-GESSAR 3.1-21 Am. No. 56,'(3/81)

ACNGS-PS AR 3.1.2.3.3 Criterion 22 - Protection System Independence

) The protection system shall be designed to assure that the effects of natural phenomena, and of normal operating, maintenance, testing, and postulated accident conditions on redundant channels do not result in loss of the protection function, or shall be demonstrated to be acceptable on some other defined basis. Design techniaues, such as functional diversity or diversity in component design and principles of operation, shall be used to the extent practical to prevent loss of the protection function.

3.1.2.3.3.1 Evaluation Against Criterion 22 The components of protect ita systems are designed so that the mechanical 139(U and thermal environment resulting from any emergency situation in which the components are reoaired to function will not interfere with the operation of that function. Wiring for the reactor protection system outside of the control room enclos ares is run in rigid metallic wireways. No other wiring is run in these wireways. The wires from duplicate sensors on a common process tap are run in sepsrate wireways. The system sensors are electri-cally and physically separated. Only circuits of the sam division may be 35(G) run in the same wireway.

The reactor pr atection system is designed to permit maintenance and diagnostic work while the reactor is operating without restricting the plant operation or hindering the output of their safety functions. The flexibility in design af forded the protection system allows operational system testing by the use of an independent input for each actuator logic.

[ } When an essen ial monitored variable exceeds its scram trip point, it is

(,/ sensed by four independent sensors. An intentional bypass, maintenance operation, calibration operation, or test will result in a single channel trip. This leaves three channels per monitored variable, each of which can in this condition initiate a scram. Only two actuator logics must trip to initiate a scram. Thus, the two-out-of-four arrangement assures that a scram will occur as a monitored variable exceeds its scram setting.

The protection system meets the design requirements for functional and physical independence as specified in Criterion 22.

For further di scussion, see the following sections:

a) Principal Design Criteria 1.2 56 b) Reactivity Control System 4.6 c) Main Steam Line Isolation Valves 5.5 d) Residual Heat Removal System 5.5 e) Emergency Core Cooling Systems 6.3 f) Reactor Protection System 7.2 A

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v' (U)-Update (G)-GESSAR 3.1-22 Am. No. 56, (3/81)

ACNGS-PSAR 3.1. 2 . 3. 5 Criterion 24 - Separation of Protection and Control Systems The protection system shall be separated from control systems to the extent that f ailure of any single control system component or channel, or failure or removal from service of any single protection system component or chan-nel which is common to the <ontrol and protection systems leaves intact a system satisfying all relirbility reaundancy, and indepcoilence re<luirements of the protection system. Interconnection of the protection and control systems shall be limited so as to assure that safety is not significantly impaired.

3.1.2.3.5.1 Evaluation Against Criterion 24 There is separation between the Reactor Protection System and the process control systems. Sensors, trip channels, and trip logics of the Reactor Protection System are not used directly for automatic control of process systems. Therefore, failure in the controls and instrumentation of pro-cess systems cannot induce failure of any portion of the protection sys-tem. iiigh scram reliability is designed into the Reactor Protection System and hydraulic control unit for the control rod drive. The scram signal and mode of operation overrides all other signals.

The Containment and Reactor Vessel 1 solation Control Systems are designed so that any one f ailure, maintenance operation, calibrction operation, or test to verify operational availability will not impair the functional ability of the isolation control system to respond to essential variables.

O Process radiation monitoring is provided on the main steam lines. Four v instrumentation channels are used to prevent an inadvertent scram and iso-lation as a result of instrumentation malfunctions. The output trip sig-nals from each channel are combined in two out of four logic that two chan-nels must signal high radiation to initiate scram and main steam isolation.

The protection system is separated from control systems as required in Criterion 24.

For further discussion, see the following sections:

a. Principal Design Criteria 1.2

b. Reactivity Control System 4.6 56

c. Emergency Core Cooling System 6.3

d. Reactor Protection System 7.2 e- Containment and Reactor Vessel Isola-tion Control System 7. 3

f. Emergency Core Cooling System -

Instrumentation and Control 7. 3

g. Neutron Monitoring System 7. 6

h. Process Radiation Monitoring 7. 6

1. Leak Detection System 7. 6

j. Reactor Manual Control System 7. 7 3.1-23 Am. No. 56, (3/81) .

4 ACNGS-PSAR i

! 3.1.2.3.6 Criterion 25 - Protection System Requirements for Reactivity Control Halfunctions The protection system shall be designed to assure that specified acceptable fuel design limits are not exceeded for any single malfunction of the re-activity control systems, such as accidental withdrawal (not ejection or dropout) of control rods.

3.1. 2 . 3. 6.1 Evaluation Against Criterion 25 The Reactor Protection System provides protection against the onset and consequences of conditions that threaten the integrity of the fuel barrier i

and the Reactar Coolant Pressure Boundary. Any monitored variablet which I exceeds the scram set point will initiate an automatic scram and not im-pair the remaining variables from being monitored, and if one channel fails the remaining portions of the Reactor Protection System shall function.

The Reactor Manual Control System is designed so that no single failure can j negate the effectiveness of a reactor scram. The circuitry for the Reactor Manual Control System is completely independent of the circuitry control-ling the scrua valves. This separation of the scram and normal rod control functions prevents failures in the reactor manual control circuitry from affecting the scram circuitry. Because each control rod is controlled as an individual unit, a failure that results in energizing any of the insert or withdraw solenoid valves can affect only one control rod. The effec-tiveness of a reactor scram is not impaired by the malfunctioning of any one control rod.

l35(G)

, The design of the protection system assures that specified acceptable fuel design limits are.not exceeded for any single malfunction of the reactivity I control systems as specified in Criterion 25.

i 4 For further discussion, see the following sections:

a. Principal Design Criteria 1.2

b. Reactivity Control System 4.6 Nuclear Design 4.3 l 56 c.

i d. Thermal and Hydraulic Design - 4.4

e. Reactor Scram System 7. 2

f. Reactor Manual Control System 7. 7

g. Accident Analysis 15.0

3.1.2.3.7 Criterion 26 - Reactivity Control System Redundancy and Ca-

! pability Two independent reactivity control systems of different design principles shall be provided. One of the systems shall use control rods, preferably including a positive means for inserting the rods, and shall be capable of

! reliably controlling reactivity changes to assure that under conditions of normal operation, including anticipated operational occurrences, and with appropriate margin for malfunctions such as stuck rods, specified acceptable fuel design limits are not exceeded. The second reactivity l control system shall be capable of reliably controlling the rate of reacti-j.

v j vity changes resulting from planned, normal power changes (including. xenon (G)-GESSAR (C)-Consistency, 3.1-24 Am. No. 56, (3/81)

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ACNGS-PSAR 7

The redundancy and capabilities of the reactivity control systems for the BWR satisfy the requirements of Criterion 26.

For further discussion, see the following sections

a. Principal Design Criteria 1.2 l 35((

b. Reactivity Control System 4.6 l56

c. Standby Liquid Control System -

Instrumentation and Control 7.4

d. Reactor Manual Control System 7. 2 l 35(C 3.1.2.3.8 Criterion 2 7 - Combined Reactivity Control Systems Capability The reactivity control systems shall be designed to have a combined capa-bility in conjunction with poison addition by the Emergency Core Cooling System, of reliably controlling reactivity changes to assure that under postulated accident conditions and with appropriate margin for stuck rods the capability to cool the core is maintained.

I 3.1.2.3.8.1 Evaluation Against Criterion 2 7 i

There is no credible event applicable to the BWR which requires combined capability of the Control Rod System and poison additions by the Emergency Core Cooling network. The primary reactivity control system for the BWR i during postulated accident conditions is the Control Rod System. Abnor-p malities are sensed, and, if protection system limits are reached, correc-i tive action is initiated through an automatic scram. High integrity of the protection system is achieved through the combination of logic arrangement, trip channel redundancy, power supply redundancy, and physical separation.

High reliability of reactor scram is further achieved by separation of scram and manual control circuitry. individual control units for each con-trol rod, and fail-safe design features built into the rod drive system.

Response by the Reactor Protection System is prompt and the total scram I

time is short.

In operating the reactor there is a spectrum of possible control rod worths, depending on the reactor state and on the control rod pattern chosen for operation. Control rod withdrawal sequences and patterns are selected to achieve optimum core performance and low individual rod worths.

The rod pattern control system prevents rod withdrawal other than by the ,

pre-selected rod withdrawal pattern. The rod pattern control system 35(G) assists the operator with an effective backup control rod monitoring routine, that enforces adherence to established startup, shutdown, and low

povar level operations. As a result of this carefully planned procedure, prompt shutdown of the reactor can be achieved with scram insertion of less than half of the many independent control rods. If accident conditions re-i quire a reactor scram this can be accomplished rapidly with appropriate margin for the unlikely occurrence of malfunctions such as stuck rods.

The reactor core design assists in maintaining the stability of the core under accident conditions as well as during power operation. Reactivity f'*5g coefficients in the power range that contribute to system stability are:

\ l q) 1 (G)-GESSAR (C)-Consistency 3.1-26 'Am.'No. 56, (3/81) d

,,,,-,,--~w.m.-,- ,.y ,-

--v -,y- e. - _ _ - , , - - ,

._,e .c-wc.,--%ry,,..wcv- w.,,.ry- ,w,, ,,y,- <- ,y-,,,,,,-,.p -

3~w,,v.+*w v- = vv +vr v --m--- -c - w

AC'iGS-PSAR a) fuel temperature or Doppler coefficient; b) moderator soid coefficient; c) moderator temperature coefficient.

The overall power reactivity coef ficient is negative and provides a strong negative reactivity feedback under severe power transient conditions.

The design of the reactivity control systems assures reliable control of reactivity under postulated accident conditions with appropriate margin for stuck rods. The capability to cool the core is caintained unde; all postulated accident conditions; thus, Criterion 2 7 is satisfied.

For further discussion, see the following sections:

a. Principal Design Criteria 1.2 l 35(t

b. Reactivity Control System 4.6 l56

c. Nuclear Design 4.3

d. Thermal and Hydraulic Design 4.4

e. Reactor Protection System 7. 2

f. Reactor Manual Control System 7. 7

g. Accident Analysis 15.0 l35(C 3.1.2.3.9 Criterion 28 - Reactivity Limits The reactivity control systems shall be designed with appropriate limits on the potential amount and rate of reactivity increase to assure that the ef fects of postulated reactivity accidents can neither (1) result in damage to the Reactor Coolant Pressure Eoundary greater than li=ited local yielding, nor (2) sufficiently disturb the core, its support structures or other reactor pressure vessel internals to impair significantly the capa-bility to cool the core. These postulated reactivity accidents shall in-clude consideration of rod ejection (unless prevented by positive means),

rod dropout, steam line rupture, changes in reactor coolant tempe rature and pressure, and cold water addition.

3.1.2.3.9.1 Evaluation Against Criterion 28 i The Control Rod System design incorporates appropriate limits on the poten-l tial amount and rate of reactivity increase. Control rod withdrawal se-i quences and patterns are selected to achieve optimum core performance and l low individual rod worths. The rod pattern control systeu prevents with-35

, drawal other than by the preselected rod withdrawal pattern. The rod pat-

tern control system assists the operator with an effective backup control

rod monitoring routine that enforces adherence to established startup, j shutdown, and low power level operations control rod procedures.

l The control rod mechanical design incorporates a hydraulic velocity limiter j in the' control rod which prevents rapid rod ejection. This engineered

safeguard protects against a'high reactivity insertion rate by limiting the

! control rod velocity to less than 5 ft/sec. Normal rod movement is limited j \ to 6 in. increcents and the rod withdrawal rate is limited through the j j hydraulic valve to 3 in./sec.

(C)-Consistency i

(G)-GESSAR j 3.1-27 Am. No. 56, (3/81)

[ .

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

i 1l ACNCS-PSAR l

The accident analysis (Chapter 15) evaluates the postulated reactivity ac-cidents as well as abnormal operational transients in detail. Analyses are t included for rod dropout, steam line rupture, changes in reactor coolant

] temperature and pressure, and cold water addition. The initial conditions, j assumptions, calculational models, sequences of events, and anticipated re-sults of each postulated occurrence are covered in detail. The results of these analyses indicate that none of the postulated reactivity transients i or accidents result in damage to the Reactor Coolant Pressure boundary. 6

! In addition, the integrity of the core, its support structures or other reactor pressure vessel internals are maintained so that the capability to cool the core is not impaired for any of the postulated reactivity acci-dents described in the accident analysis.

a The design features of the reactivity control systems which limit the potential amount and rate of reactivity increase ensure that criterion 28 is satisfied for all postulated reactivity accidents.

For further discussion, see the following sections:

a. Principal Design Criteria 1.2 l35

b. Design Criteria - Structures, Components j Equipment and Systems 3.0 i c. Reactor Core Support Structures and Internals Mechanical Design 3.9.4 56

d. Reactivity Control System 4.6  !

, c. Nuclear Design 4.3

[5g

f. Control Rod Drive Ilousing Supports 4.5 V g.

h.

Overpressurization Protection Reactor Vessel and Appurtenances 5.2.2 5.4

! 1. Main Steam Line Flow Restrictor 5.5 J. Main Steam Line Isolation Valves 5.5

k. Accident Analysis 15.0 5

' 3.1.2.3.10 Criterion 29 - Protection Against Anticipated Operational

. Occurrences I

i The protection and reactivity control systems shall be designed to assure

! an extremely high probability of accomplishing their safety functions in l the event of anticipated operational occurrences.

3.1.2.3.10.1 Evaluation Against Criterion 29 l

l The high functional reliability of the protection and reactivity control systems is achieved through the combination of logic arrangement, re-dundancy, physical and electrical independence, functional separation, fail-safe design, and inservice testability. These design features are discussed in detail in Criteria 21, 22, 23, 24, and 26.

l An extremely high probability of correct protection and reactivity control 1 l systems response to anticipated operational occurrences is maintained by j a thorough program of inservice testing and surveillance. Active com-

! ponents can be tested or removed from service for maintenance during reactor operation without compromising the protection or reactivity control j s, functions even in the event of a subsequent single failure. Components I

i 3.1-28 Am. No. 56, (3/81) .

ACNGS-PSAR s

g important to safety such as control rod drives, main steam isolation

) valves, residual heat removal pumps, etc., are tested during normal reactor

(

'- / operation. Functional testing and calibration schedules are developed using available failure rate data, reliability analyses, and operating experience. These schedules represent an optimization of protection and reactivity control system reliability by considering, on one hand, the failure probabilities of individual components and, on the other hand, the reliability ef fects during individual component testing on the portion of the system not undergoing test. The capability for inservice testing ensures the high functional reliability of protection and reactivity control systems should a reactor variable exceed the corrective action setpoint.

The capabilities of the protection and reactivity control systems to per-form their safety functions in the event of anticipated operational occur-rences are satisfied in agreement with the requirements of Criterion 29.

For further discussion, see the following sections:

l35(C)

a. Principal Design Criteria 1.2 1 56

b. Reactivity Control System 4.6 I

c. Main Steam Line Isolation Valves 5.5.5 l 35(C)

d. Residual lieat Removal System 5.5.7

e. Containment Systems 6.2

f. Emergency Core Cooling Systems 6.3

g. Reactor Protection System 7. 2

[ j h. Containment and Reactor Vessel Isolation Control System 7. 3

(/ i. Emergency Core Cooling Systems -

Instrumentation and Control 7.3

j. Neutron Monitoring System 7. 6

k. Process Radiation Monitoring 7. 6 1 teak Detection System 7.6

m. Accident Analysis 15.0 3.1.2.4 Group IV - Fluid Systems (Criteria 30 - 46) 3.1.2.4.1 Criterion 30 - Quality of Reactor Coolant Pressure Boundary Components which are part of the Reactor Coolant Pressure Boundary shall be designed, fabricated, erected, and tested to the highest quality stand-ards practicable. Means shall be provided for detecting and, to the extent practicable, identifying the location of the source of reactor coolant leakage.

3.1.2.4.1.1 Evaluation Against criterion 30 By utilizing conservative design practices and detailed quality control procedures, the pressure retaining components of the Reactor Coolant Pressure Boundary are designed and fabricated to retain their integrity during normal and postulated accident conditions. Accordingly, components which comprise the Reactor Coolant Pressure Boundary are designed, fabri-

[\ ,}/ cated, erected, and tested in accordance with recognized industry codes and standards listed in Chapter 5. Further, product and process quality (C)-Consistency 3.1-29 Am No. 56, (3/81)

ACNGS- PS A R 3.1.2.4.7 Criterion 36 - Inspection of Emergency Core Cooling System

\

(Q The Emergency Core Cooling System shall be designed to permit appropriate periodic inspection of important components, such as spray rings in the re-actor pressure ve ssel, wa ter inject ion nozzles , and piping, to assure the integrity and capability of the system.

3.1. 2.4 . 7.1 Evaluation Against Criterion 36 The Emergency Core Cooling Systems are as discussed in Criterion 35. The engineering and design ef fort for these systems include inservice inspec-tion considerations. The spray rings within the Reactor Vessel are acces-sible for inspection during each refueling outare. Removable plugs in the Reactor Vessel Shield and/or panels in the insulation provide access fo r examination of nozzles. Removable insulation is provided on the Emergency Core Cooling Systems piping out to and including the first isolation valve outside containment. Inspection of the Emergency Core Cooling System is in accordance with the intent of Section XI of the ASME Code. Sect ion 5.2 defines the Inservice Inspection Plan, access provisions, and areas of re-stricted access.

During plant operations, the pumps, valves, piping, inst rumentation, wir-ing, and other components outside the drywell can be visually inspected at any t ime . Components inside the drywell can be inspected when the drywell is open for acces s. When the Reactor Vessel is open, for refueling or other purposes, the spargers and other internals can be ins pe c t ed . Poc-tions of the ECCS which are part of the Reactor Coolant Pressure Boundary (O)

V are designed to specifications for inservice inspection to detect de fect s which r ight affect the cooling performance. Particular attention will be given to the reactor nozzles, core spray, and feedwater spargers. The de-sign of the Reactor Vessel and internals for inservice inspection, and the plant testing and inspection program ensures that the requirements of Cri-terion 36 will be met.

For further discussion, see the following sections:

a. Reactor Core Support Structures and 56 Internals Hechanical Design 3.9.4

b. Inservice Inspection Program 5.2.8

c. Reactor Vessel and Appurtenances 5.4

d. Emergency Core Cooling Systems 6.3 3.1.2.4.8 Criterion 37 - Testing of Emergency Core Cooling System l The Emergency Core Cooling System shall be designed to permit appropriate l periodic pressure and functional testing to assure (1) the structural and leaktight integrity of its components, (2) the operability and performance j of the active components of the system, and (3) the operability of the t

system as a whole and, under conditions as close to design as practical, the performance of the full operational sequence that brings the system j into operation, including operation of applicable portions of the protec-tion system, the trans fer between normal and emergency power sources, and

/] the operation of the associated cooling water system.

3.1-37 Am. No. 56, (3/81)

ACNGS-PSAR For further discussion, see the following sectioris of the SAR:

a. General Plant Description 1.2

b. Water Systems 9.2

c. Initial Tests and Operation 14.0

d. Technical Specifications 16.0 3.1.2.5 Group V - Reactor Containment (Criteria 50-5 7) 3.1.2.5.1 Criterion 50 - Containment Design Basis The reactor containment structure, including access openings, penetrations, and Containment lieat Removal System shall be designed so that the Contain-cent structure and its internal compartments can accommodate, without ex-ceeding the design leakage rate and, with sufficient margin, the calculated pressure and temperature conditions resulting from any loss of coolant ac-cident. This margin shall reflect consideration of (1) the effects of po- ,

tential energy sources which have not been included in the determination of the peak conditions, such as energy in steam generators and as required by 56 paragraph 50.44, energy f rom metal-water and other chemical reactions thac may result from degraded emergency core cooling functioning, (2) the limited experience and experimental data available for defining accident phenomena and Containment responses and (3) the conservatism of the calculational model and input parameters.

f 3.1.2.5.1.1 Evaluation Against Criterion 50 O

j ( / The Containment structure, including access openings and penetrations, is designed to accommodate, without exceeding the design leak rate, the tran-sient peak pressure and temperature associated with a LOCA up to and in-cluding a double ended rupture of the largest reactor coolant pipe.

The Containment structure and Engineered Safety Features Systems have been evaluated for various combinations of energy release. The analysis ac-counts for system thermal and chemical energy, as required by paragraph 56 50.44, and for nuclear decay heat.

The maximum temperature and pressure reached in the drywell and Containment i during the worst case accident are shown in Chapters 6 and 15 to be well 4

below the design temperature and pressure of this structure.

The cooling capacity of the Containment Heat Removal Systems are adequate to prevent overpressurization of the structure, and to return the Contain-i ment to near atmospheric pressure.

For further discussion, see the following sections of the SAR:

a. Classification of Structures, Components 3.2

~

and Systems

b. Steel Containment System 3.8.2

c. Containment Functional Design 6.2.1

d. Containment Heat Removal Systems 6.2.2 r

e. Accident Analyses 15.0 (C)-Consistency 3.1-46 Am. No. 56, (3/81) t

_ . .__..,m,_,. . _ . _ , , . . _ ,. - ,. - . _ . . . - _ _ , - - . _ . . . _ . . . . _ . , - - . . _ . _ , . . . .m. --

. _ . - . . -- - -- -. . _- = _ . _ . - . . _ _ . _ _ _ _ _ _ , . ~

i ACNGS-PAR j t) 1) Lines 3/a in. and smaller which are part of the Reactor Coolant l Presaure Eaundary shall be Safety Class 2/ Quality Group B.

2) All instrument lines which are connected to the Reactor Coolant 5 Pressure Boundary and are utilized to actuate or to monitor safety Q1-

, systems shall be Safety Class 2 from the outer isolatio- valve or 3.13 i the process shutoft valve (root valve) to the sensing instrumentation. .

3) All instrument lines which are connected to the Reactor Coolant 5 Pressure Baundary and are not utilized to actuate or to monitor Ql-i safety systems shall be "other" from the outer isolation valve or 3.13

, the process shutof f valve (root valve) to the sensing instrumentation.

I l

4) All other instrument lines:

(a) through the root valve shall be of the same classification as the system to which they are attached

l (b) beyond the root valve, if used to actuate a safety system,

shall be of the same classification as the system to which they are attached f[

j (c ) beyond the root valve, if not used to actuate a safety system, shall be Non Safety Class.

l

5) All sample lines from the outer isolation valve or the process root valve through the remainder of the sampling system shall be Non Safety Class.

.) The control rod drive insert and withdraw lines from the drive flange up j to and including the first valve on the hydraulic control unit shall be Safety Class 2. Section 4.6.2.3.2 gives the analysis which shows that loss of coolant from a postulated failure of a CRD Hydraulic System line l 56 would be less than 10 gpm which is much less than the leakage from a 3/4 inch line break. Based on 3/4 inch RCPB line break below the water level the leakage would be greater than 200 gpm.

i

, t) The Hydraulic Control Unit (HCU) is a General Electric factory-assembled engineered module of valves, tubing, piping, and stored water which controls a single control rod drive by the application of precisely I

timed sequences o; pressures and flows to accomplish slow insertion or withdrawal of the control rods for power. control, and rapid insertion for reactor scram.

Although the hydraulic control unit, as a unit, is field installed and connected to process piping, many of its internal parts differ markedly from process piping components because of the more complex functions i they must provide. Thus, although the codes and standards invoked by g the SC 1,2,3 and other pressure integrity quality levels clearly a v

3.2-31 Am. No. 56, (3/81) l l

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/"%

] 'N

\d Y \

ACNCS-PSAR i TABLE 3.2-10 ACTIVE VALVES IN SEISMIC CATECORY I SYSTEMS (Note: Environmental design criteria are discussed in Section 3.11.)

Valve Valve Valve Sire, Actuator Quantity Code Service Ident. Type In. Type

• per Unit Class NUCLEAR BOILER - FIC. 5.1-3a,b Main Steam Isolation F022A,B,C.D Clobe 26 A0 4 1 MSLIV Accumulator F024A,B,C,D Check N/A Rev. Flow 4 3 Main Steam Isolation F028A,B,C.D Clobe 26 A0 4 1 MSLV Accumulator F029A,B C.D Check N/A Rev. Flow 4 3 Sa fety/ Relief, ADS F0418,C; 438,C Relie f 6x10 A0 4 1 ADS Accumulator F036,39 Check N/A Rev. Flow 16 3 Non-ADS Accumulator F036 Check N/A Rev. Flow 13 3 Safety / Relief, ADS 10 F045A.D; 478,C Relief 6x10 A0 4 1 Sa fety/Relie f, Non-ADS F0425,C.C; 44A,B,D Re lie f 6x10 A0 6 1 Sa fet y/Relie f. Non- ADS F046A,B C; 488,C,D H Relief 6x10 A0 7 1 Feedwater Isolation F065A,8 Cate 20 M 2 1 Feedwater Supply F010A,5 Check 20 Rev. Flow 2 1 Fredwater Supply F032A,B Check 20 A0 Rev. Flow 2 1 Main Steam Drains F016 F019 Cate 3 MO 2 1 Main Steam Drains F067A.R.C.D Globe 1 1/2 MO 4 1 56 REACTOR RECIRCULATION - FIC. 5.5-2a,b Y Pump Seal F013A,B Check 3/4 Rev. Flow 2 2

$ Pump Seal F017A,8 Check 3/4 Rev. Flow 2 2 Sample F019 Diaph. Control 3/4 A0 1 2 Sample F020 Diaph. Control 3/4 A0 1 2 CONTROL ROD DRIVE -FIG. 4.6-5a,b,c 56 Scram Discharge Vent F010 Diaph. Control 1 AD 1 2 Scram Discharge Drain F0ll Diaph. Control 2 A0 1 2 CRD Supply F083 Cate 2 M0 1 2 l

CRD Supply F122 Check 2 Rev. Flow 1 2 .

STANDBY LIQUID CONTROL - FIC. 4.6-9 SLC Supply F033A Check 1 1/2 Rev. Flow 2 2 SLC Supply F004A,B Explosive 1 1/2 Electric 2 1

, y SLC Supply F006 Check 1 1/2 Rev. Flow 1 I 8

SLC Supply F007 Check 1 1/2 Rev. Flow 1 I

.E O

3

• A0 = Air Operated MO = Motor Operated Rev. Flow = Reverse Flow NA = Not available; t o be provided later.

,.m . __ .. . . . - - _ _ _ . . - . _ _ _ - _ _ . _ _ _ . _ _ _ _ _ _ . . . . _ _ . . _ _ . . . _ _ _ . . . _ . _ _ _ _ _ _ _ _ _ _ _ . . . _ _ - _ _ _ _ . . - - _ . _ _ _ .

O O O AsCS-PSAR TABLE 3.2-10 (Cont'd)

Valve Valve Valve Sire, Actuator Quantity Code Se rvice Ident. Type In. Type

• per Unit Class RESIDUAL HEAT REMOVAL - FIC. 5.5-Ila,b ILI Outlet F003A,8 Cate 18 MO 2 2 RHR Pump Suction F004A,B.C Cate 24 MO 3 2 Shutdown Suction Relief FG05 Relief I Syst. Press. I 2 Shutdown Suction F006A,B Cate 20 MO 2 2 Shutdown Suction F008 Cate 20 MO 1 1 Shutdown Suction F009 Cate 20 MO 1 1 Condensate to Pool F0llA,B Clobe 4 MO 2 2 Pump Suction Relief F011A,B Relief N/A Syst. Press. 3 2 10 Head Spray F019 Check 6 Rev. Flow 1 1 Q1 3,5 C Loop Test F021 Clobe 18 MO 1 2 l Head Sprav F023 Clobe 6 NO I I A, B Loop Test F024A,B Clobe 18 MO 2 2 Pump Discharge Relief F025A,B,C Relie f N/A Syst. Press. 3 2 RHR HX Consensate F026A,8 Cate 4 M0 2 2 Pump Discharge F031A,B,C Check 18 Rev. Flow 3 2 Condensate to Pool Relie f F036 Relief I Syst. Press. I 2 RHR HX to Radwaste F040 Globe 4 MO 1 2 LPCI to Reactor F041A,B,C Check 12 Rev. Flow 3 1 LPCI to Reactor F042A,B.C Cate 12 MO 3 1 Pump Min Flow Bypass F046A,B,C Check 3 Rev. Flow 3 2 u HX Inlet F047A,8 Cate 18 MO 2 2 y HX Bypass F048A,8 Clobe 18 4

MO 2 2 s- RHR HX to Radwaste F049 Cate MO 1 2 u

l t

I I >

! s

l. m 10 Ql-8.5 -

C

• A0 = Air Operated S Syst. Press a System Pressure O MO = Motor Operated Rev. Flow = Reverse Flow l NA = Not available; to be provided later.

l

ACNGS-PSAh l s of the shield wall will be approximately 28 feet with a total height above the RPV pedestal of approximately 52 feet. The shield wall will be pro-vided with continuous ring plates welded to the concentric cylinders at the top and the bottom of the wall. Vertical and horizontal stiffeners will be provided as required by design. Openings for the RPV nozzle penetra-tions will be r ovided with local reinforcement and removable shielding i 35 sections tc- allow nozzle inservice inspection. The removable shields at 8 the openings will be designed with restraining members to prevent them from becoming missiles under the pressure loading in the space between the RPV and the shield wall.

In addition to the shielding function, the wall will be used as a support for pipe whip restraints, and as a support for the RPV insulation. The biological shield wall outline is shown on Figure 3.8-6.

3.8. 3.1. 7 Pedestal for Reactor Pressure Vessel and Shield Wall The reactor vessel pedestal will consist of two concentric steel cylinders having diameters of approximately 20 and 32 feet, respectively. The annular space between the cylinders will be filled with concrete. A con- l56 tinuous steci plate ring will be provided at the top of the pedestal; the cylinders will be anchored to the concrete mat at the bottom. The free standing RPV will be anchored to the pedestal by welding or bolting the RPV support skirt to the top pedestal ring. The reactor shield wall will also be supported on the RPV pedestal. Vertical and horizontal stiffeners will be provided throughout the height of the pedestal as required. Local

\ stiffeners will be installed at the large rectangular openings, necessary for control rod drive mechanism operation, maintenance and removal. Where continuous horizontal diaphragm stiffeners are required by design, suffi-cient number of prout holes will be provided for concrete fill work. The bottom liner of the Containment will be seal welded to the inner and outer 35(C) faces of both cylinders.

The outline of the pedestal embedment details are shown on Figure 3.8-3.

An outline of the pedestal structure is shown on Figure 3.8-5.

3.8.3.1.8 Description of the Drywell Vacuum Relief System The sizing basis of the drywell vacuum breaker is to limit the maximum negative drywell pressure (containment to drywell) to prevent suppress-ion pool back flow into the drywell. The negative drywell pressure limit 26 is the hydrostatic water head corresponding to the minimum weir wall free board height. The vacuum relief capacity incorporates redundant flow paths 35(C) to accommodate the single failure criterion. The vacuum relief system lines are designed to ASME Section III, Subsection NE-7000 standards.

26 Each vacuum relief assembly will consist of a 10" check valve in series with an 18" automatic valve (see Figure 6.2-65). Actuation of the auto-matic valve will be controlled by differential pressure between the dry- 17,22, well and the containment. A transmitter will sense the differential pres- 23 sure and provide a signal to open the valve at the preset differential 1-9.22 pressure (-0.2 psid). A permissive signal of high drywell pressure (+2.0 ,;-9.21

[ psid) is required to open the automatic vacuum relief valve. Thus, small 26

( operational variations in the drywell to containment differential pressure (C)-Consistency ,

3.8-21 Am. No. 56, (3/81)

)

4 ACNGS-PSAR ,

i I

e i

i l

1 i

i 1

i.

i l l

l l l i

i 1

I i

l,

. l 1 i

' l t

l l

I

Figure 3.8-10 I i (No Figure Assigned) 56 i ,

1 I l

d 4

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Am. No. 56, (3/81)

l I

j ACNCS-PSAR l 3.9.4 REACTOR CORE SUPPORT STRUCTURES AND INTERNALS MECHANICAL DESIGN i

3.9.4.1 Design Bases 3.9.4.1.1 General Design Bases

3. 9 . 4 .1.1.1 Safety Design Basea The reactor core support structures and internals shall meet the following safety design bases:

a) Shall be arranged to provide a floodable volume in which the core can be adequately cooled in the event of a breach in the nuclear system process barrier external to the reactor vessel, b) Deformation shall be limited to assure that the control rods and Core Standby Cooling Systems can perform their safety functions.

c) Mechanical design of applicable structures shall assure that safety design bases (a) and (b), above, are satisfied so that the safe shutdown of the plant and removal of decay heat are not impaired.

3.9.4.1.1.2 Power Generation Design Bases l 56 The reactor core support structures and internals shall be designed to the O following power generation design bases:

a) They shall provide the proper coolant distribution during all anticipated normal operating conditions to allow power opera-tion of the core without fuel damage.

b) They shall be arranged to facilitate refueling operations.

c) They shall be designed to facilitate inspection.

1 3.9.4.1.2 Specific Design Characteristics 56 3.9.4.1.2.1 Design Loading Combinations The design loading combinations of the reactor vessel internals are covered in Section 3.9.4.3.1.1.

3.9.4.1.2.2 Stress, Deformation and Fatigue Limits for Reactor 56 Internals (Except Core Support Structure)

! The stress, deformation and fatigue criteria listed in Tables 3.9-9, 3.9-10 J

3.9-11, and 3.9-12 shall be used or the criteria shall be based on the cri-teria established in applicable codes and standards for similar equipment, l 56 by manufacturers' standards, or by empirical methods based on field experi-ence and testing. For the quantity SF . (minimum safety factor) appear-ing in those tables, the following valEeE listed shall be use,1.

3.9-9 Am. No. 56, (3/81)

ACNGS-PSAR Design SF .

Condition *9 Normal 2.25 Upset 2.25 Emergency 1.5 Fault 1.125 3.9.4.1.2.3 Stress, Deformation and Fatigue Limits for Core Support l56 Structures The stress, deformation and fatigue criteria presented in Tablen 3.9-13, l 56 3.9-14 and 3.9-15 shall be used. These criteria shall be supplemented, where applicable, by the criteria for the reactor internals in the previous section, but in no case shall the criteria presented in Tables 3.9-13 e I

56 3.9-14 and 3.9-15 be exceeded for core support structures.

3.9.4.1.2.4 Fuel Assembly Restraints l56 The fuel assembly structural design shall demonstrate sufficient dimen-sional stability and sufficient fuel rod support to maintain core geometry thus avoiding fuel damage for both planned operation and abnormal opera-tional transients.

3.9.4.1.2.5 Material Selection p The material used for fabricating most of the reactor vessel core support I and reactor internal structures are solution heat-treated, unstabilized Type 304 austenitic stainless steel conforming to ASTM and ASME specifica-tions. Weld procedures and welders are qualified in accordance with the intent of Section IX of the ASME Boiler and Pressure Vessel Code. Further controls for stainless steel welding are covered in Section 5.2.5.

All the materials of construction exposed to the reactor coolant are to be resistant to stress corrosion in the BWR coolant. Conservative corrosion allowances are to be provided for all exposed surf aces of carbon or low alloy steels.

Contaminants in the reactor coolant are controlled to very low limits by the reactor water quality specifications. No detrimental effects shall i

occur on any of the materials from allowable contaminant levels in the l high purity reactor coolant. Radiolytic products in a BWR shall have no adverse effects on the construction materials.

3.9.4.1.2.6 Radiation Effects l ~56 Where feasible, the design will be such that irradiation effects on the material properties will be minimized. Where irradiation effects cannot be minimized, the design of the reactor vessel internals will either have pro-visions for replaceable components, or the design will be shown to satisfy l

a set of stress and fatigue design limits that have been arrived at consid-ering the ef fect of irradiation damage on the fracture toughness, ductility and tensile properties of the materials.

3 M

3.9-10 Am. No. 56, (3/81)

i i ACNGS-PSAR i l 56 i 3.9.4.1.2.) Shock Loaos

The components shall be designed so as to accommodate the loadings dis- l l cussed in Section 3.9, " Mechanical bystems and components."

l56 I 3.9.4.2 bescription

{ The core support structures and reactor vessel internals include (exclusive

ot f uel, control rods, and in-core nuclear instrumentation) the following a components

a) Core bupport Structures

1) bhroud

)

2) Shroud support i

3) Core support and holddown bolts  ;

4) Top guide (including wedges, bolts and keepers) t 5) Fuel support pieces i

! 6) Control rod guide tubes 1 b) Reactor Internals i

1) Jet pump assemblies and instrumentation

2) Shroud heaa and steam separator assembly (including shroud head bolts)

3) 5 team dryers I

I 1

4) Feedwater spargers

5) Vessel head cooling spray nozzle i

6) Differential pressure and liquid control line

7) In-core flux monitor guide tubes and stabilizers  :

8) Initial startup neutron sources [

9) Surveillance sample holders

10) Core spray lines and spargers A general assembly drawing of the important reactor components is shown in 156 I

Fih ure 3.9-3.

The floodable inner volume of the reactor pressure vessel can be seen in ,

Figure.3.9-4. It is the volume inside the core shroud up to the level of l 56 the jet pump suction inlet.

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ACNGS-PSAR 3.9.4.2.1 Core Support Structure The core support structure consists of the shroud, shroud support, core support, fuel support pieces, control rod guide tubes, and top guide.

This structure is used to form partitions within the reactor vessel, to sustain pressure differentials across the partitions, to direct the flow of the coolant water, and to locate laterally and support the fuel anaem-blies, control rod guide tubes, and steam separators. Figure 3.9-4 shows l 56 the reactor veasel internal flow paths. .

3. 9. 4. 2.1.1 Core Shroud l 56 The core shroud is a stainless steel cylindrical assembly that provides a  !

partition to separate the upward flow of coolant through the core f rom the +

downward recirculation flow This partition separates the core region from the downcomer annulus, thus providing a floodable region following a recir-culation line break. The volume enclosed by the shroud is characterizeo by three regions. The urner shroud surrounds the core discharge plenum, which is bounded by the shrcad head on top and the top guide below. 1he central portion of the shroud surrounds the active fuel and forms the longest sec-tion of the shroud. This section is bounded at the bottom by the core sup-port. The lower shroud, surrounding part of the lower plenum, is welded to the reactor pressure vessel shroud support. (See Section 5.4, " Reactor Vessel and Appurtenances.")

3.9.4.2.1.2 Shroud Head and Steam Separator Assembly The shroud head and steam separator assembly is bolted to the top of the upper shroud to form the top of the core discharge plenum. This plenum provides a mixing chamber for the steam-water mixture before it enters the steam separators. Individual stainless steel axial flow steam separators, shown in Figure 3.9-3, are attached to the top of standpipes that are l 56 welded into the shroud head. The steam separators have no moving parts.

In each separator, the steam-water mixture rising through the standpipe ,

passes vanes that impart a spin to establish a vortex separating the water from the steam. The separated water flows from the lower portion of the steam separator into the downcomer annulus.

3.9.4.2.1.3 Core Support Plate The core support plate consists of a circular staintess steel plate with bored holes stiffened with a rim and beam structure. The plate provides lateral support and guidance for the control rod guide tubes, in-core flux monitor guide tubes, peripheral fuel supports, and startup neutron sources.

The last two items are also supported vertically by the core support plate.

The entire assembly is bolted to a support ledge between the central and lower portions of the core shroud. Alignment pins that engage slots and that bear against the shroud are used to correctly position the assembly before it is secured.

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ACNGS-PSAR 56 3.9.4.2.1.4 Top Guide The top guide is formed by a series of stainless steel beans joined at right angles to form square openings and are fastened to a peripheral rim.

Each opening provides lateral support and guidance for four fuel assemblies

or, in the case of peripheral fuel, one or two fuel assemblies. Notches are provided in the bottom of the beam intersections to anchor the incore flux cionitors and startup neutron sources.

The rim of the top guide is bolted to (and forms a section of) the shroud.

3.9.4.2.2 Fuel Support l56 ,

The f uel supports, shown in Figure 3.9-6, are of two basic types, namely, -

peripheral supports and four-lobed orificed fuel supports. The peripheral fuel support is located at the outer edge of the active core and is not adjacent to control rods. Each peripheral fuel supprt will support one fuel assembly and contains a single orifice assembly dedgned to assure proper coolant flow to the fuel peripheral assembly. Each 'our-lobed ori-ficed fuel support will support four fuel assemblies and is provided with I orifice plates to assure proper coolant flow distribution to each rod- l controlled fuel assembly. The four-lobed orificed fuel supports rest in the top of the control rod guide tubes which are supported laterally by the core support. The control rods pass through slots in the center of the four-lobed orificed fuel support. A control rod and the four adjacent fuel assemblies represent a core cell. (See Section 4.2.2, General Design Description." l56 3.9.4.2.3 Control Rod Guide Tubes U l56 The centrol rod guide tubes, located inside the vessel, extend from the top of the control rod drive housings up through holes in the core support plate. Each tube is designed as the guide for a control rod and as the vertical support for a four-lobed orificed fuel support piece and the four fuel assemblies surrounding the control rod. The bottom of the guide tube is supported by the control rod drive housing (See Section 5.4, " Reactor Vessel and Appurtenances"), which in turn transmits the weight of the guide tube, fuel support, and fuel assemblies to the reactor vessel bottom head.

A thermal sleeve is inserted into the control rod drive housing from below and is rotated to lock the control rod guide tube in place. A key is inserted into a locking Slot in the bottom of the control rod drive housing to hold the thermal sleeve in position.

3.9.4.2.4 Jet Pump Assemblies l56 The jet pump assemblies are located in two semicircular groups in the down-comer annulus between the core shroud and the reactor vessel wall. The design and performance of the jet pump is covered in detail in References 3.9-3 and 3.9-4. Each stainless steel jet pump consists of driving l56 nozzles, suction inlet, throat or mixing section, a diffuser (see Figure 3.9-7). The driving nozzle, suction inlet, and throat are joined together l56 as a removable unit, and the diffuser is permanently installed. High pressure water from the recirculation pumps (i.ee Section 3.5.1, " Reactor ,

Coolant Pumps") is supplied to each pair of jet pumps through a riser pipe 3.9-10c Am. No. 56, (3/81)

ACNGS-PSAR i

welded to the recirculation inlet nozzle thermal sleeve. A riser brace consists of cantilever beams extending from pada on the reactor vessel wall.

The nozzle entry section is connected to the riner by a metal-to metal, spherical-to conical seal joint. Firm contact is maintained by a holddown clamp. The tnroat nection is nupported laterally by a bracket attached to the riser. There is a slip-fit joint between the throat and diffuser. The dif fuser in a gradual conical section changing to a straight i

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cylindrical section at the lower end.

56

$ 3.9.4.2.5 Steam Dryers I

The steam dryers remove moisture f rom the wet steam leaving the steam sepa-retors. The extracted moisture flows down the dryer vanes to the collec-ting troughs, then flows through tubes into the downconer annulus (see Fig- 56 ure 3.9-8). A skirt extends from the bottom of the dryer vane housing to [

j the steam separator standpipe, below the water level. This skirt forms a j seal between the wet stesa plenum and the dry steam flowing from the top of the dryers to the steam outlet nozzles.

i' The steam dryer and shroud head are positioned in the vessel during in-stallation with the aid of vertical guide rods. The dryer assembly rests on steam dryer support brackets attached to the reactor vessel wall. Up-ward movement of the dryer assembly, which would occur only under accident conditions, is restricted by steam dryer hold-down brackets attached to the reactor vessel top head.

4 ,

56 i'

3.9.4.2.6 Feedwater Spargers .

e The feedwater spargers are perforated stainless steel headers located in the mixing plenum above the downconer annulus. A separate sparger is fit- l ted to each feedwater nozzle and is shaped to conform to the curve of the vessel wall. Sparger end brackets are attached to vessel brackets to support the spargers, and jack bolts position the spargers away from the vessel wall. Feedwater flow enters the center of the spargers and is dis-charged radially inward to mix the cooler feedwater with the downcomer flow from the steam separators before it contacts the vessel wall. The feed-water also serves to condense the steam in the region above the downconer annulus and to subcool the water flowing to the jet pumps and recirculation pumps.

3.9.4.2.7 Core Spray Lines l56 j

The core spray lines are the means for directing flow to the core spray nozzles which distribute coolant so that peak fuel clad temperatures of 2300 F are not exceeded during accident conditions.

I Two core spray lines enter the reactor vessel through the two core spray nozzles. (See Section 5.4, " Reactor Vessel and Appurtenances.") Each

! line divides immediately inside the reactor vessel. The two halves are

! routed to opposite sides of the reactor vessel and are supported by clamps l attached to the vessel wall. The lines are then routed downward into the downcommer annulus and pass through the upper shroud immediately below .

the flange. The flow divides again as it enters the center of the semi-circular sparger, which is routed halfway around the inside of the upper

shroud. The ends of the two spargers are supported by brackets designed

! to accommodate thermal expansion. The line routing and supports are de-signed to accommodate differential movement between the shroud and ves-l sel. The other core spray line is identical except that it enters the op-posite side of the vessel and the spargers are at a slightly different ele-vation inside the shroud. The correct spray distribution pattern is pro-vided by a combination of distribution nozzles pointed radially inward and l

l l 3.9-10e Am. No. 56, (3/81) l l

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ACNGS-PSAR i downward from the spargers. (See Section 6.3, " Emergency Core Cooling System.")

6 3.9.4.2.8 Vessel Head Cooling Spray Nozzle When reactor coolant is returned to the reactor vessel, part of the flow  :

can be diverted to a spray nozzle in the reactor head. This spray main-4 tains saturated conditions in the reactor vessel head volume by condensing j steam being generated by the hot reactor vessel walls and internals. The spray also decreases thermal stratification in the reactc- vessel coolant.

. This ensures that the water level in the Reactor Vessel can rise. The

! higher water level provides conduction cooling to more of the mass of metal

! of the reactor vessel and therefore limits thermal stress in the vessel a

during cooldown.

The vessel head cooling spray nozzle is mounted to a short length of pipe and a flange, which is bolted to a mating flange on the reactor vessel head nozzle. (See Section 5.5.I, " Residual Heat Removal System.")

1 3.9.4.2.9 Differential Pressure and Liquid Control Line l

The differential pressure and liquid control line serves a dual function within the reactor vessel; to provide a path for the injection of the 56 i liquid control solution into the coolant stream (discussed in Section 4.6.6, " Standby Liquid Control System") and to sense the differential pressure across the core support plate (described in Section 5.4, " Reactor Vessel and Appurtenances"). This line enters the reactor vessel at a point below the core shroud as two concentric pipes. In the lower plenum, the two pipes separate. The inner pipe terminates near the lower shroud with a perforated length below the core support plate. It is used to sense the pressure below the core support plate during normal operation and to inject liquid control solution if required. This location facilitates good mixing and dispersion. The inner pipe also reduces thermal shock to the vessel nozzle should the Standby Liquid Control System be actuated. The outer pipe terminates immediately above the core support plate and senses the pressure in the region outside the fuel assemblies. 56 3.9.4.2.10 In-Core Flux Monitor Guide Tubes In-core flux monitor guide tubes provide a means of positioning fixed detec-l tors in the core as well as a path for calibration monitors (TIP System).

The in-core flux monitor guide tubes extend from the top of the in-core j

flux monitor housing (see Section 5.4, " Reactor Vessel and Appurtenances")

! in the lower plenum to the tip of the core support plate. The power range detectors for tha power range monitoring units and the ury tubes for the Source Range Monitoring and Intermediate Range Monitoring (SRM/IRM) de-

' tectors are inserted through the guide tubes. A latticework of clamps,

tie bars, and spacers gives lateral support and rigidity, to prevent loos-ening during reactor operation.

56 3.9.4.2.11 Surveillance Sample Holders The surveillance sample holders are welded baskets containing impact and 3.9-10f Am. No. 56, (3/81)

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ACNGS-PSAR l tensile specimen capsules (see Section 5.4, " Reactor Vessel and Appurten-

• f

ances"). The baskets hang from the brackets that are attached to the in-  ;

side wall of the reactor vessel and extend to mid-height of the active core. The radial positions are chosen to expose the specimens to the same '

environment and maximum neutron fluxes experienced by the reactor vessel itself while avoiding jet pump removal interference or damage.

56' 3.9.4.2.12 Low Preasure Coolant Injection Lines l Three 33 1/3 percent capacity LPCI lines penetrate the core shroud through separate LPCI nozzles. Coolant is discharged inside the core shroud l immediately below the core spray spargers.

{ 3.9.4.3 Safety Evalcation l56 3.9.4.3.1 Evaluation Methods l56 To determine that the safety design bases are satisfied, responses of the  !

reactor vessel internals to loads imposed during normal, upset, emergency, and faulted conditions are examined. The effects on the ability to insert control rods, cool the core, and flood the inner volume of the reactor ves-sel are determined.

3.9.4.3.1.1 Input for Safety Evaluation l56 The operating conditions that provide the basis for the design of the re-actor internals to sustain normal, upset, emergency and faulted conditions

') as well as combinations of design loadings that are accounted for in design of the core support structure are covered in Table 3.9-16. l56 ,

In addition each combir ation of operating loads is categorized with respect to either normal, upset, emergency er faulted conditions as well as the as-sociated design stress intensity or deformation limits.

5 The bases for the proposed design stress and deformation criteria are also specified in Chapter 3.

3.9.4.3.1.2 Events To Be Evaluated y

Examination of the spectrum of conditions for which the safety design basis I

must be satisfied reveals three significant faulted events:

Recirculation Line Break (LOCA): the accident results in pressure a) differentials, within the reactor vessel, that may exceed normal loads. r Steam line break accident: a break in one main steam line between j b) the reactor vessel and the flow restrictor. The accident results in significant pressure diff erentials across some of the structures +

within the reactor.

1

-c) Earthquake: subjects the core support structures and reactor O

\

internals to significant forces.as a result of ground motion.

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l ACNCS-PSAR Analysis of other conditions existing during normal operation, abnormal operational transients, and accidents shows that the loads affecting the core support structures and reactor internals are less severe than these three postulated events.

3.9.4.3.1.3 Pressure Differential During Rapid Depressurization l 56 l 4 digital computer code (Reference 3.9-5) is aed to analyze the transient

conditions within the reactor vessel following the recirculation line break l 56 accident and the steam line break accident (see Reference 3.9-5). The analy-tical model of the vessel consists of nine nodes, which are connected to the necessary adjoining nodes by flow paths having the required resistance and inertial characteristics. The program solves the energy and mass con-servation equations for each node to give the depressurization rates and pressure in the various regions of the reactor. Figure 3.9-9 shows the 56 nine reactor nodes.

3.9.4.3.2 Recirculation Line and Steam Line Break 3.9.4.3.2.1 Accident Definition l 56 Both a recirculation line break (the largest liquid break) and an inside steam line break (the largest steam break) are considered in determining the design basis accident for the reactor internals. The recirculation line break is the same as the design basis loss-of-coolant accident de-scribed in Section 6.3, " Emergency Core Cooling Systems". A sudden, s complete circumferential break is assumed to occur in one recirculation loop.

The analysis of the steam line break assumes a sudden, complete circumfer- ,

ential break of one main steam line between the reactor vessel and the main steam line restrictor. This is not the same accident described in Chapter 15, " Accident Analysis," which has greater potential radiological effects.

A steam line break upstream of the flow restrictors produces a larger blow-down area and thus a faster depressurization rate than a break downstream of the restrictors. The larger blowdown area results in greater pressure differentials across the reactor assembly internal structures.

The steam line break accident produces significantly higher pressure differentials across the reactor assembly internal structures than does the recirculation line break. This results from the higher reactor depres-surization rate associated with the steam line break. The depressurization rate is proportional to the mass flow rate and the excess of fluid escape enthalpy above saturated water enthalpy,fh . Mass flow rate is inversely proportional to escape enthalpy, hg , and therefore the depressurization rate is approximately proportional to 1 - hf h,. Consequently, depres-surization rate decreases as h decreases, t at is, the depressurization rate is less for mixture flow Ihan for steam flow. Therefore, the steam line break is the design basis accident for internal pressure differ-entials.

3.9-10h Am No. 56, (3/81)

ACNGS-PSAR

, 3.9.4.3.2.2 Effects of Initial Reactor Power and Core Flow h6

/

'D] For purposes of illustration the maximum internal pressure loads can be considered to be composed of two parts: steady state and transient pres-sure differentials. For a given plant the core flow and power are the two major factors which influence the reactor internal pressure differentials.

The core flow essentially affects only the steady state part. For a fixed power, the greater the core flow, the larger will be the steady state pres-sure differentials. The core power affects both the steady state and the transient parts. As the power is decreased there is less voiding in the core and consequently the steady state core pressure differential is less.

Ilowever, less voiding in the core also means that less steam is generated in the reactor pressure vessel and thus the depressurization rate and the transient part of the maximum pressure load is increased.

Figure 3.9-10 is a power-flow map which defines the permissible operating h6 conditions of the reactor. From this range of operating conditions, it is necessary to determine the combination of core power and flow which will result in the maximum internal pressure loads. Consider the condition where the power is at 100 percent and the core flow is at 102 percent of rated conditions (i.e., the maximum point on the operating map). dince, as mentioned above, a decrease in power will result in higher transient pres-sure differentials, a more severe initial condition might be the condition ot 48 percent power, 108 percent flow. In going f rom 100 percent power, 34G) 102 percent flow to the 46 percent power, 108 percent flow condition, the steady state pressure differential has a net decrease. There is an in-

/ crease due to the slight increase in flow and a decrease due to the de-(7)

V crease in power (lower core pressure drop). This transient pressure dif-ferential increases due to the decrease in power. Ilowever, the maximum 34(G) pressure load (steady state plus transient) has a net increase for the flow power condition. If the power is decreased below 46 percent, the core flow must also be reduced. Analysis has shown that the decrease in flow and power reduces the steady state part of the maximum pressure load more than the corresponding increase in the transient part. Hence, the maximum pres-sure loads (steady state plus transient) are less if the core flow is re-duced from its maximum value. Therefore, the maximum internal pressure loads occur following an inside steam line break from an initial condition in which the reactor is at the minimum power associated with the maximum core flow (i.e., 46 percent power, 108 percent flow). 34(G)

Table 3.9-1; lists the maximum pressure loads occurring across the reactor internals during the accident for two cases. Case 1 is for an initial con- l 34(G) dition of 105 percent rated steam flow and 100 percent rated core flow.

Case 2 is for the maximum pressure loads and these occur at the initial condition of 46 percent power 108 percent flow. Comparison of Cases 1 and 34(G) 2 illustrates the generalized statements made above concerning the relation-ship between the maximum internal pressure loads and core power and flow.

It can be seen that the general comments, while basically applicable to all of the reactor components, are not equally true for all the components (e.h., the channel box is the same for Cases 1 and 2).

Realistically, if an inside steam line break were to occur, the maximum

/ T internal pressure loads would probably be closer to Case 1. This is be-(,/ cause the plant will most probably be operating at or near full power.

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% Also, the Case 2 condition, although possible, is rather abnormal in that rated core ilow is neither required nor desirable at such a reduced power condition.

3.9.4.3.2.3 Break Size Spectrum Analysis 56 It has been determined that the maximum internal pressure loads occur from an initial condition in which the reactor is at the minimum power associ-ated with the maximum core flow. It has also been concluded that these i maximums occur for an inside steam line break, the largest possible steam break. To verify this conclusion, the internal pressure differentials have been calculated for a spectrum of break areas. The initial reactor condi-tion for this break spectrina analysis is the worst case condition deter- 56 mined above (i.e., 48 percent power, 108 percent flow). Figure 3.9-11 is a plot of the maximum pressure loads versus steam line break area for the major reactor internal components. This figure shows that, without excep-tion, the maximum pressure loads decrease with decreasing break size. A typical forcing function acting on the reactor internal is shown in Figure 56 3.9-12. This is the transient loading acting on the reactor internal 1

structures following a break of the main steam line.

Conclusions 3.9.4.3 2.4 l56 It is concluded that the maximum pressure loads acting on the reactor in-ternal components result from an inside steam line break occurring while the reactor ir at the minimum power associated with the maximum core flow (Table 3.9-17, Case 2). This has been substantiated by the analytical com- l56 parison of liquiet versus steam breaks, by the investigation of the effects of core power and core flow, and by the break spectrtun analysis.

It has also been pointed out that, although possible, it is not probable that the reactor would be operating at the rather abnormal condition of minimum power and maximum core flow. More realistically, the reactor would be at or near a full power condition and thus the maximum pressure loads acting on the internal components would be as listed under Case 1 in Table l

3.9-17.

Response of Structures Within the Reactor Vessel to 56 3.9.4.3.2.5 Pressure Differences The maximum differential pressures are used, in combination with other structural loads, to determine the total loading on the various structures within the reactor. The structures are then evaluated to assess the extent

of deformation and buckling instability, if any. Of particular interest are

(1) the responses of the guide tubes and the metal channels around the fuel bundles, and (2) the potential leakage around the jet pump joints.

The guide tube is evaluated for buckling instability caused by externally

- applied pressure. Two primary modes of failure have been analyzed and are O

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ACNGS-P3AR e

described in Section 4.2.3.3.7. For a guide tube with minimum wall thick- l 56 nens and maximum allowed ovality, the pressure which causes yield strena in 93 pai compared to the design prensure of 37.5 pai. The design pressure in in all canen greater than any pressure ditferential the guide tube will experience including accident conditions. The stress the guide tube could experience would be approximately 5400 pai due to external prennure (37.5 pai), a 1.2g earthquake (including dead weight loading) and lateral loading due to coolant flow, while yield stress at 575 F in 17,500 pai. It in O

O 3.9-10k Am. No. 56, (3/81)

ACNGS-PSAR concluded that the guide tube will not fail under the assumed conditions.

r%

f l The f uel channel blowdown force analysis was completed for the worst case condition; that is, a steam line break accident at 105 percent rated steam 17 flow. This condition would cause the maximum P across the fuel channel to increase 2.8 psi. This results in a final P of 14 pai. Because this is O the worst case condition, the result is used for design purposes.

Thermal shock during depressurization and quenching following a postulated LOCA will not greatly alter the core geon.etry. The limits specified in the Interim Acceptance criteria and the new criteria of Appendix K (Reference l5 3.9-6) will assure that some ductility will remain in the zircaloy clad-ding as it goes through the quenching process. Therefore the core will remain essentially intact and in a condition amenable ta long term cooling.

In addition, data listed in Reference 3.9-7 provide a basis for cladding fragmentation. The data shows that zircaloy cladding will not fragment.

If less than 17 percent of the cladding oxidizes in the BWR/6 cores for the design basis LOCA, fragmentation from thermal shock will not occur.

The fuel assemb)y response to blowdown pressure loading is calculated using a general mechanical analysis computer code (such as Reference 3.9-8) for the evaluation of stress deflection within the assembly. The hg output stress and deflections will be compared with design limits as out-lined in the appropriate subsections of this chapter and the response to questions 4.39 and 5.7. Stress and deflection from these blowdown pres-sure loadings are expected to fall within upset design limits, but are required to fall within faulted limits. Calculated value will be available j after the detailed testing and analytical evaluations are completed.

V Jet pump joints have been analyzed to evaluate the potential leakage from within the floodable inner volume of the reactor vessel during the recircu-lation line break and subsequent LPCI reflooding. Because the jet pump diffuser is welded to the shroud support, the only remaining source of leakage from the lower plenum to the downcomer annulus is the jet pump throat-to-diffuser joint. These joints for all jet pumps leak no more than a total of 340 gpm. l 34 (G)

LPCI capacity is sized to accommodate 500 gpm l?akage at these locations.

It is concluded that the reactor vessel structures retain sufficient integ-rity during the recirculation line break accident to allow reflooding of the inner volume of the reactor vessel and in sufficient time to prevent significant increases in clad temperature.

3.9.4.3.3 Earthquake l 56 The seismic loads acting on the structures within the reactor vessel are based on a dynamic analysis of a model similar to that shown in Figure 3.9-13. Seismic analysis is performed by coupling this lumped mass model l56 of the reactor vessel and internals with the building model to determine the system natural frequencies and modes. The relative displacement, acceleration, and load response is then determined by either the time his-tory method or the response spectrum method.

, )

(G)-GESSAR 3.9-101 Am. No. 56, (3/81)

ACNCS-PSAR In the time history method , the dynamic response is determined for each mode of interest and added algebraically for each instant of time. Resul-g I ting response time histories are then examined, and the maximum value of displacement, acceleration, shears, and moments are used for design calcu-lations.

In the response spectrum method, the relative displacements, accelerations, shears, and moments are determined for each node of interest. The square root of the average of the squares of these individual responses are then used for design calculations.

The detailed descriptions of the earthquake analysis are given in Section 3.7.2.1.2.3. The time histories of the pressure differentials due to a steam line break are shown in Figure 3.9-12. The detailed description of the dynamic response analysis to these forcing functions are given in Sections 3.9.1.4 and 3.9.1.5. The peak dynamic loads from the SSE and the steam line break are combined in the square root of the sum of squares manner.

Since the length to width ratio of the fuel assembly is greater than 20, it is most appropriately modeled as a beam. In the beam model, both the flexural and shear deformations are accounted for. Since the sum of the fuel rod stiffnesses are less than ten percent of the channel box stiff- 17 ness, the flexural stiffness of the fuel assembly is dominated by the

• channel stif fness with the major portion of the mass being the fuel rod mass. Due to the stiffness of the channel box relative to the stiffnesses 5.4 of the fuel rod, it is quite clear that the channel box will resist the

_N major portion of the seismic loads. That is, the deformation of the fuel

) rods is controlled by the deformation of the channel box. During a seismic event, the fuel assembly, including the fuel rods and channel box, is set into vibratory motion with the channel providing most of the resistance to large amplitude oscillation.

The aforementioned model of the fuel assembly is used to determine its dynamic response. Then the motion of the channel box and the natural frequency of the channel box are used to compute the acceleration values of the fuel assembly as a function of the axial length. These acceleration values are then used to determine the stresses in the fuel rods and other internals of the fuel assembly. It should be pointed out that although a beam model is used to determine the dynamic response of the fuel assembly, the detailed stress calculations are performed with many other degrecs of freedom than the dynamic model. Thus the comprehensive dynamic analyses and the detail stress analyses jointly provide conservative stress and deformation values experienced by the fuel assembly during a seismic event.

The safety margins used in the design of the fuel assembly can be broadly catalogued into two areas; namely, conservatisms used in determining the seismic loads and conservatisms in the allowable load determination. The seismic analysis process is conservative in that safety margins are introduced at every step of the calculations. The determination of the ground acceleration, the determination of the response spectrum and the 7m generation of a synthetic time history which matches the response spectrum, the low damping values used for dynamic analysis of the soil structure

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3.9-10m Am. No. 56, (3/81)

ACNGS-PSAR system and the reactor dynamic analyses have built-in conservatisms. The introd6ction of the safety margins at each of the above ateps introduces k

j a large final safety factor since the safety factors are multiplicative.

Thus, the calculated seismic load is expected to be much higher than the actual earthquake loads which will be seen by the fuel assembly. Further-j more, the entire core is assumed to move in phase during an earthquake which is conservative since it is expected that due to nonlinear ef fects i

of the fluid the fuel channels would not move in unison thus significantly l

reducing the amplitude of vibration of the fuel channel. In the allowable j load computation, the unitradiated yield strength of the fuel channel 17 material at temperature is used. Since the irradiated yield strength are Q2-5.4 higher than the unitradiated yield str ength, the channel is expected to have a higher capability when an earthquake event actually occurs. The conservatisms in the calculation of the seismic load and the conservatisms used in calculation of the allowable loads together insure that the fuel assembly integrity during a seismic event is amply assured.

BWR/6 fuel assembly design is still in the process of being finalized.

The numerical results are not currently available but will be supplied at the FSAR stage.

Fluid sloshing will have very negligible effects on the fuel assembly.

Sloshing with significant amplitudes will occur only when a free surface with large areas exists. For the fuel assembly, the water entering the bottom of the fuel assembly is gradually changed into steam as it pro-gresses upwards. Thus, there exist no discernible free surf aces. Even if a free surface existed, the pressure change due to the sloshing within the small confines of the fuel channel is calculated to be about 0.002 psi.

This is evidently negligible compared to the 14 psi pressure dif ference '

across the channel.

56 3.9.4.3.4 Conclusions Response analyses of the reactor structures show that deformations are suf-ficiently limited to allow both adequate control rod insertion and proper operation of the core standby cooling systems. Sufficient integrity of the structures is retained during accident conditions to allow 3.9-10n Am. No 56, (3/81)

ACNGS-PSAR successful reflooding of the reactor vessel inner volume. The analyses considered various loading combinations, including loads imposed by exter-nst forces. Thus, safety design bases 3.9.4.1.1 are satisfied. 56 i 3.9.4.4 Inspection and Testing 56 Quality control methods are used during the fabrication and assembly of reactor vessel internals to assure that the design specifications are met.

The Reactor Coolant System, which includes the core support structures and reactor internals, in thoroughly cleaned and flushed before fuel is loaded initially.

During the preoperational test program, operational readiness tests are performed on various systems. In the course of these tests such reactor internals as the feedwater spargers, the core spray lines, the vessel aead cooling spray nozzle, and the Standby Liquid Control System line are func-tionally tested.

Steam separator-dryer performance tests are made during the startup test program of the first plant to use a specific design concept to determine carry-under and carry-over characteristics. Samples are taken from the inlet and outlet of the steam dryers and from the inlet to the main steam lines at various reactor power levels, water levels, and recirculation flow rates. Moisture carry-over is determined from sodium-24 activity in these samples and in reactor water samples. Carry-under is determined from meas-s ured flows, and temperatures are determined by heat balances.

Vibration analysis of reactor structures is included in the design to guard against potential vibration problems. When necessary, vibration is meas-ured during startup tests to determine the vibration characteristics of specific reactor structures. Vibratory responses are recorded at various recirculation flow rates and power levels using strain gages, accelerome-ters, and linear differential transducers as appropriate.

The vibration analyses and tests are designed to determine any potential, hydraulically induced equipment vibrations. The structures are analyzed for natural frequencies, mode shapes, and vibrational magnitudes that could lead to fatigue at these frequencies. With this analysis as a guide, the reactor structures are instrumented and tested to determine whether vibra-tion amplitudes are excessive. The cyclic loadings are evaluated using, as a guide, the cyclic stress criteria of the ASME Code,Section III. The se field tests are only performed on prototype reactor structures that re-present a significant departure from design configurations or operating conditions previously tested and found to be acceptable. The vibration programs established have be en designed to satisfy the requirements of AEC Regulatory Guide 1.20. " Vibration Measurements on Reactor Internals." Field test data are correlated with the analyses to ensure validity of the analytical techniques on a continuing basis.

GE satisifies all portions of RG 1.20 except paragraph C.1.b. Thigsub- 5 paragraph suggests the need to subject the reactor intervals to 10 O

3.9-10o Am. No. 56, (3/81)

ACNGS-PSAR cycles of vibration. The GE acceptance criterigare based upon the need to assurethatvjbrationwillbeacceptablefor10 cycles (40 years);

therefore, 10 cycles has no specific relevance in that context. During cold flow testing on prototype BWR's, the reactor intervals are exercised D

in excess of 10 cycles by reactor coolant flow. This is a sat .sfactory 5

apgroach to RG 1.20 since the ASME Code defines the endurance 1.. git as 10 cycles and since the degradation in fatigue strength from 10 to 7

10 cycles is very small.

The reactor vessel and the structures within the vessel are designed to assure adequate working space and access for inspection of selected com-ponents and locations. Criteria for selecting the components and locations to be inspected are based on the probability of a defect occurring or en-larging at a given location and include areas of known stress concentrations and locations where cyclic strain or thermal stress might occur.

The plant is designed to provide for in-service inspection as required by ASME Boiler and Pressure Vessel Code,Section XI.

a 1

l l

, Vb 3.9-10p Am. No. 56, (3/81)

- --- _. - .. - . - . , ,. .-_ ..- ~ __- - - . . ,

I l

ACNCS-PSAR 4

1 SECTION 3.9: REFERENCES l 3.9-1 Quinn, E. P., " Vibration of Fuel Pools in Parallel Flow," USAEC

, Report GEAP-4059, General Electric Cn. , Atomic Power Equipment  :

I Department, July 1962.

j 3.9-2 Quinn, E. P., "Vibratinn of SEFOR Fuel Rnds in Parallel Flow," USAEC j Report GEAP-4966, Atomic Power Equipment Department, September 1965.

>

j 3.9-3 " Design and Performance of G.E. BWR Jet Pumps," General Electric ,

Company, Atomic Power Equipment Department, APED-5460, July 1968. 56 i

3.9-4 Moen, R. H., " Testing of Improved Jet Pumps for the BWR/6 Nuclear ,

System," General Electric Co., Atomic Power Equipment Department.

NEDO-10602, June 1972.

i

, 3.9-5 Slifer, B. C., " Loss-of-Coolant Accident and Emergency Core Cooling Models for General Electric Boiling Water Reactors," NED0-10329,

April 1971.

s 3.9-6 Acceptance Criteria for Emergency Core Cooling Systems for Light

Water Conled Nuclear Power Reactors - Docket No. RM-50-1, Section j IIB, p. 33.

l 3.9-7 C. J. Scatena, " Full Cladding Embrittlement During a Loss of Coolant Accident ," NEDO-10674, October 1972.

3.9-8 Rashid, N. R., " Creep-Plast", GEAP-13262-1, March, 1973.

3.9-9 0'Donnel, W. J. and Langer, B. F., " Fatigue Design Basis for Zircaloy Components," Nuclear Science and Engineering, 1964.

a 3.9-10q Am No. 56, (3/81)

ACNCS-PSAR Q TABLE 3.9-9 56' DEFORMATION LIMIT (for reactor internal structures oniv)

Either One Of (Not Both) General Limit

a. Permissible def omation, DP g 0.9 Analyzed defomation SF,dn

~'

causing loss of function, DL t t 1

b. Pemissible deformation, DP g 1.0 Experiment deformation -SF min causing loss of function, DE where:

DP = permissible defomation under stated conditions of normal, upset, emergency or fault DL= analyzed function (p'formationwhichcouldcauseasystemlossof O DE = experimentally detemined defomation which could cause a system loss of function 1 Equation b will not be used unless supporting data are provided to the AEC by General Electric. The specific equation to be used on identified reactor internals will be presented in the FSAR.

2 " Loss of Function" can only be defined quite generally until attention is focused on the component of interest. In cases of interest, where deformation limits can af fect the function of equipment and components, they will be specifically delineated.

From a practical viewpoint, it is convenient to interchange some deformation condition at which function is assured with the loss of function condition if the required safety margins from the functioning conditions can be achieved. Therefore, it is often unnecessary to determine the actual loss of function condition because this interchange procedure produces conservative ,

and safe designs. Examples where deformation limits apply are:

control rod drive alignment and clearances for proper insertion, core support deformation causing fuel disarrangement or excess leakage of any component.

U 3.9-23 Am. No. 56, (3/81)

-= _ _ . . . _ _ . _- .. ~. . . . .. - .- __ __ _

ACNCS-PS AR

' 56 TABLE 3.9-10 .

PRIMARY STRESS LIMIT (for reactor internal structures only) t Any One Of (No More Than One Required) General Limit

a. Elastic evaluated primary stresses. PE 2.25 Permissib.e primary stresses, PN ,

- SF gg

b. Permissible load, LP , 1.5 Largest lower bound limit load, CL - SF ein i

c. ' Elastic evaluated 0 primary stress, PE I .75 SF min Conventional ultimate strength Jst temperature, US ,

d. Elastic-plastic evaluated 0.9 nominal primarv stress, EP I SFain Conventional ultimate strength ,

at temperature, US

- 1

e. Permissible load, LP 0.9 I SF s/m Plastic instability load, PL min 1

f. lPermissible load, LP g 0.9 Ultimate load from fracture - SF ein Analysis, UF ,

- 1

! g. Permissible load, LP g 1.0 Ultimate load or loss of function - SFmin load from test, LE where:

l PE = primary stresses evaluated on an elastic basis. "he effec-tive membrane stresses are to be averaged through the load carrying section of interest. The simplest average bending, I shear or torsion stress distribution which will support the external loading will be added to the membrane stresses at the section of interest.

PN = permissible primary stress levels under normal or upset con-i ditions under ASME Boiler and Pressure Vessel Code, Sec-tion III.

1 Equations e., f., and.m. will not be used unless supporting data are provided to theNRC by General Electric. The specific equa- l56

; tion to be used on identified reactor internals will be presented g / in the FSAR.

3.9-24 Am. No. 56, (3/81) i L_

ACNGS-PSAR i

i TABLE 3.9-10 (Cont'd) 56 l LP = permissible load under stated conditions of normal, upset, emergency or fault.

y CL = lower bound limit load with yield point equal to 1.5 S where o S e is the tabulated value of allowable stress at temperature of the i ASME III code or its equivalent. The " lower bound limit load" is here defined as that produced f rom the analysis of an ideally plastic (nonstrain hardening) material where deformations increase with no 4

further incresse in applied load. The lower bound load is one in which the material everywhere satifies equilibrium and nowhere ex-coeds the defined material yield strength using either a shear theory or a strain energy of distortion theory to related multiaxial yield ,

to the uniaxial case.

i US = conventional ultimate strength at temperature or loading which would cause a system malfunction, whichever is more limiting.

EP = elastic plastic evaluated nominal primary stress. Strain hardening of the material may be used for the actual monotonic stress strain curve at the temperature of loading or any approximation to the i

actual stress strain curve which everywhere has a lower stress for the same strain as the actual monotonic curve may be used. Either the shear or strain energy of distortion flow rule may be used.

PL = plastic instability load. The " plastic instability load" is defined here as the load at which any load bearing section begins to diminish

its cross-sectional area at a faster rate than the strain hardening can accommodate the loss in area. This type analysis requires a true i

stress-true strain curve or a close approximation based on monotonic loading at the temperature of loading.

UF = ultimate load from fracture analyses. For components which involve sharp discontinuities (local theoretical stress concentration 3)

' the use of a " fracture mechanics" analysis where applicable, utiliz-  !

ing measurements of plane strain fracture toughness may be applied to compute fracture loads. Correction for finite plastic zones and thickness effects as gross yielding may be necessary. The methods

! of linear elastic stress analysis may be used in the fracture l

analysis where its use is clearly conservative or supported by exper-inental evidence. Examples where "f racture mechanics" may be applied are for fillet velds or end of fatigue life crack propagation.

4 LE = ultimate load or loss cf function load as detarmined from experiment.

l - In using this method, account shall be taken of the dimensional j tolerances which asy exist between the actual part and the tested

' part or parts as well as differences which may exist in the ultimate tensile streng*h of the actual part and the tested parts. The guide i

to be used in each of these areas is that the experimentally deter-mined load shall use adjusted values to account for material property l

and dimension variations, each of which has no greater probability

'~'g than 0.1 of being exceeded in the actual part.

I 3.9-25 Am. No. 56, (3/81)

a i  :

} ACNCS-PSAR l

l TABLE 3.9-11 l 56 i

) BUCKLING STABILITY LIMIT i

(for reactor internal structures only)

I Any One of (No More han One Required) General Limit

~

l a. Permissible load LP g 2.25 l Code normal event permissible load, PN , SF,g, I

l

b. 'Persissible load, LP g 0.9 Stability analysis load, SL ,

- SF g, i c. Permissible load. LP 1.0 Ultimate buckling collapse load from test, SE; IQ l where:

{ LP = permissible load under stated conditions of normal, upset, j amargency or fault.

PN = applicable code normal event permissible load.

I SL = stability analysis load. The ideal buckling analysis is of ten sensitive to otherwise minor deviations from ideal

, geometry and boundary conditions. These effects shall be accounted for in the analysis of the buckling stability l

j loads. Examples of this are ovality in externally presa r-j ized shells or eccentricity on coluant members.

SE = ultimate buckling collapse load as determined from experi-j ment. In using this method, account shall be taken of the i dimensional tolerances which may exist bet'reen the actual

part and the tested part. De guide to be used in each of

! these areas is that the experiaantally determined load shall be adjusted to account for material property and dimension variations, each of which has no greater probability than 0.1

) of being exceeded in the actual part.

i l I i

1 Equatiert.c. will not be used unless supporting data are provided to the NRC by General Electric. The specific equation to be used l56 in identified reactor interaals will be presented in the FSAR.

t 3.9-26 Am. No 56, (3/81) i 4 . . . - . _ _ _ - - , - , . _ _ _ _ , . - - . _ - _ _ _ . _ _ _ _ _ _ . . . - . . . . _ _ _ _ . _ _ , , _ . _ . . _ . . . . _ . , . _ - _ _ . . _ _ _ - , _ , -

l ACNCS-PSAR

\

56

TABLE 3.9-12 FATIGUE LIMIT (f or reactor internr. structures only) t Sumnation of fatigue damage usage with design and operation loads following Miner hypotheses . . .(1)

Limit for Normal and Upset Design Any One Of (No More han One Required) Conditions

a. Mean f atigue(2'3) cycle usage from analyses 50.05

b. Mean fatigue (2.3) cycle usage from test 50.33

c. Design fatigue cycle usage from analysis 51.0 l56 using the methoo of Table 3.9 13.

I

(

l (1) Miner, M. A. , " Cumulative Damage in Fatigue," Journal of Applied Mechanics, Vol.12, ASME, Vol. 67, pp A159-A164, September 1945.

(2) Fatigue f ailure is defined here as a 25 percent area reduction for a l load carrying member which is required to function, or excess leakage,

! which is more limiting.

l (3) Equations a. and b. will not be used unless supporting data are pro-sided to the AEC by General Electric, he specific equation to be used on identified reactor internals will be presented in the FSAk.

! fg O

3.9-27 Am. No. 56, (3/81) l

1 56 TABLE 3.9-13 CORE SUPPORT STRUCTURES l STRESS CATECORIES AND LIMITS OF STRESS IlfrENSITY FOR NORMAL AND UPSET CONDITIONS l

PR'MA9Y STRE55F5 SECONDARY STRESSES PE AR STRESSES STRESS l wtmeeawt, P, twOtts a.7 a el et woewc, Pg twOf as a.7s a: et an. F twOtts 2 a at St$0ae .O e t a al l

I 1

n I

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, e, . ,, e, . e, . O e. e. O.*

l

9. i a I

, 5, 8. S S, 31, u

5  !

8 (LalTIC ttalteC E t alfiC E L a5f eC Os awat ulas 08 aw aL T185 04 awalr95 O' fMOf f 43 FaftCUt (NOTI 43 (MOT t 83 esso ? g g 3 g 9

]

T,a' .,t, ..,1, s t

- 3 I q UP5ET L lasti L test? PL a57BC j 08 aNaLY$35 04 am at v g g 04 amat v sel j (NO t t lot fMOt t IO) tMOT E St

q ....L. FOR CTCL t S t ($1

... .O.F ,

TMaw 9000. Ulf Prat 1

ft$t itst It alf tC.

] > IMOf f Ill fMOV E til tNOtt 13 Pt at tsC S

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i N ad v

4

ACNGS-PSAR O

q TABLE 3.9-13 (Cont'd) 56 NORMAL AND I!PSET CONDITIONS NOTE 1 -

This 41 citation applies to the range of stress intensity. When the secondary stress is due to a temperature excursion at the point at which the stresses are being analyzed, the value of S

shall be taken as the average of the S values tabulated m

i$ Tables I-1.1, I-1.2, and I-1.3 of ASE Boiler and Pressure Vessel Code,Section III, (ASME III) for the highest and the lowest temperature of the metal during the transient. When part of the secondary stress is due to mechanical load, the value of Sm shall be taken as the Sm value for the highest temperature of the metal during the transient.

NOTE 2 - The stresses in Category Q are those parts of the total stress which are produced by thermal gradients, structural discontin-ulties, etc., and do not include primary stresses which may also exist at the same point. It should be noted, however, that a detailed stress analysis frequently gives the combina-tion of primary and secondary stresses directly and, when ap-propriate, this calculated value represents the total of P,

+Pb + Q and not Q alone. Similarly, if the stress in Category F is produced by a stress concentration, the quantity F is the additional stress produced by the notch, over and above the g nominal stress. For example, if a plate has a nomir.a1 stress intensity, P,. s, pb = 0, Q = 0 and a notch with a stress concentration K is introduced, then F = Pm (K-1) and the peak stress intensity equals Pm + Pm (K-1) = KPm .

NOTE 3 Sa is obtained from the fatigue curves, Figures I-9.1 and I-9.2 of ASME III. The allowable stress intensity for the full range of fluctuation is 2 S,,

NOTE 4 -

The symbols P , pb, Q, and F do not represent single quanti-ties, but rat $ersetsofsixquantitiesrepresentingthesix stress components at, a y, er, *tl,*1r, 'rt.

NOTE 5 S,t denotes the structural action of shakedown load as defined in paragraph NB-3213.18 of ASME III calculated on a plastic basis as applied to a specific location on the structure.

NOTE 6 -

The triaxial stresses represent the algebraic sum of the three primary bination principal of stressstresses (al +Where components.

' 2 + '3) uniformfor the com-loading tension is present, triaxial stresses are limited to 4 S,,

NOTE 7 -

For configurations where com,'ressive stresses occur, the stress limits shall be revised to take into account critical buckling stresses (see paragraph NB-3211(c) of ASME III). For external

- pressure, the permissible " equivalent static" external pressure T shall be as specified by the rules of paragraph NB-3133 of ASME

(/ III. Where dynamic pressures are involved, the permissible ex-3.9-29 Am. No 56, (3/81)

-y- - w- y 3---+-- g ym- - =,y -

9- t----- "

g

l l

ACNGS-PSAR  ;

. 1

~s '

Q TABLE 3.9-13 (Cont 'd) 56 instability pressure.

NOTE 8 -

When loads are transiently applied, consideration should be given to the use of dynamic load amplification, and possible change in modulus of elasticity.

NOTE 9 -

In the f atigue data curves, where the number of operating cycles are less than 10, use the S, value for 10 cgeles; where the number of ogerating cycles are greater than 10 , use the S, value for 10 cycles.

NOTE 10 -

Lt is the lower bound limit load with yield point equal to 1.5 S is the tabulated value of allowable stress at tImp(where Serature"as contained in ASME III). The " lower bound limit load" is here defined as that produced from the analysis of an

, ideally plastic (non-strain hardening) material where deforma-tions increase with no further increase in applied load. The lower bound load is one in which the material everywhere satis-fies equilibrium and nowhere exceeds the defined material yield strength using either a shear theory or a strain energy of dis-j tortion theory to relate multiaxial yielding to the uniaxial case.

Note 11 -

For normal and upset conditions, the limits on primary membrane plus primary bending need not be satisfied in a component if it can be shown from the test of a prototype or model that the specified loads (dynamic or static equivalent) do not exceed 44 percent of L , where L is the ultimate load or the maxi-

, mum load or comSination uEed in the test. In using this method, i account shall be taken of the size effect and dimensional tol-

. erances which may exist between the actual part and the test part, or parts, as well as differences which may exist in the ultimate strength or other governing material properties of the actual part and tested part to assure that the loads obtained j from the test are a conservative representation of the load i carrying capability of the actual component under the postu-

lated loading for normal and upset conditions.

NOTE 12 - The allowable value for the maximum range of this stress inten-

, sity is 3S,except for cyclic events, which occur less than 1000 times during the design life of the plant. For this exception, in lieu of recting the 35 limit, an elastic-plastic fatigue analysis in accordance wi*th ASME III may be performed to demon-strate that the cumulative fatigue usage attributable to the combination of these low events, plus all other cyclic events, does not exceed a fatigue usage value of 1.0.

r m

3.9-30 Am. No. 56, (3/81) l l

,,e . ..-v --.,,,i ,-4, .c.--e , , . + , . . ,4 -- , , _ . - , ,--ea--,+-,-r -,e- y.-.,w...+w . . . , - -y-,w,. ,,ww em ww,e,--- -+c,,. y...,y--w.-.cw. --,.-g.m,v---

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56 TABIE 3.9-14 (Cont 'd) j CLRE SUPPORT STRUCTURFS ,

STRESS CATEGORIES AND LIMITS OF STRESS INTENSITY FDR EMERCENCY CONDITIONS 4 i PRfMARY STRESSE5 SECONDARY STRf 55E5 PE AK STRE55E5

ufu W ANf & Rf NfMNO P( AE,

' tit u*le ANE, P, (NO t t 18.7 4 seg fif MfMNr., Pg (NO T fl 1. 2 A 10, I

I P. P, P, 4

1 EL Alf tC E L ASitC

• 835 AN Aly ses 7255 AN AL vses (NOT f M (NQ][ 3)

+

OW OR p 4

• W Lfuti (tutt O i

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L g AN gwoggAL V4 $l,l Lg ANALT51$

(No t t di y

4

1. 04 OR

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3

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EMERC(NCT PLA41C PL A$ttC f ALU ATION EV ALU Af TON NOT g(Qute(D h

NOT REOUIRF O

'"OII 'I I I I. ANALT$t$ I II I, A N Ag, T $ $$

INOf f 4) (NO f (15&4 0e Oa i

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ACNCS-PSAR 4

TABLE 3.9-14 (Cont'd)

EMERGENCY CONDITIONS NOTE 1 -

The symbols P,, P , Q, and F do not represent single quan-tities, but rathek sets of six quantities representing the six stress components t a y , a ,ra ' *t l) *1r*'rt' NOTE 2 -

For configurations where compressive stresses occur, stress limits shall be revised to take into account critical buckling stresses. For external pressure, the permissable " equivalent static" external pressure shall be taken as 150% of that per-mitted by the rules of paragraph NB-3133 of ASME Boiler and Pressure Vessel Code,Section III ( ASME III) . Where dynamic pressures are involved, the permissible external pressure

shall satisfy the preceding requirements or be limited to

50* of the dynamic instability pressure.

NOTE 3 - The triaxial stresses represent the algebraic sum of the three primary principal stresses (ai+a2 + ' 3) f r the ceabina-tion of stress components. Were uniform tension loading is present, triaxial stresses should be limited to 6 S,.

NOTE 4 -

Lg is the lower bound limit load with yield point equal to g 1.5temperature at S. (where S Is is the tabulated contained in ASMEvalue III) of allowable

. The " lower stress bound limit load" is here defined as that produced from the analysis of an ideally plastic (nonstrain hardening) material where de-formations increase with no further increase in applied load.

The lower bound load is one in which the material everwhere sa-tisfies equilibrium and nowhere exceeds the defined material i yield strength using either a shear theory or a strain of dis-l tortion theory to relate multiaxial yielding to the uniaxial case.

NOTE 5 -

Su is the ultimate strength at temperature. Multiaxial ef-fects on ultimate strength shall be considered.

NOTE 6 -

This plastic analysis uses on elastic-plastic evaluated nomin-al primary stress. Strain hardening of the material may be used for the actual monotonic stress-strain curve at the temp-

erature of loading or any approximation to the actual stress-i strain curve which everywhere has a lower stress for the same strain as the actual monotonic wrve may be used. Either the shear or strain energy of distortion flow rule shall be used to account for multiaxial ef fects.

NOTE 7 -

For emergency conditions, the stress limits need not be satis-fled if it can be shown from the test of a prototype or model that the specified loads (dynamic or ' static equivalent) do not exceed 60% of L , where L is the ultimate load or ~ d2e max-Oi inum load or load combination used in the test. In using this method, account shall be taken of the size effect and dimen-3.9-32 Am. No. 56, (3/81)

. _.__. .- . . __- _ _. _ ._- . . _ _ .__ - .._ ~._. , _ .. - - . .

l A CNCS-PS AR TABLE 3.9-14 (Cont'd) sional tolerances which may exist between the actual part and the tested part or parts as well as differences which may exist in the ultimate strength or other governing material properties 1 of the actual part and the tested parts to assure that the t loads obtained from the test are a conservative representation i of the load carrying capability of the actual component under l postulated loading for emergency conditions.

NOTE 8 - Stress ratio is a method of plastic analysis which uses the

1. stress ratio combinations (combination of stresses that con-4 sider the ratio of the actual stress to the allowable plastic or elastic stress) to compute the maximum load a strain harden-ing material can carry. K is defined as the section factor; S,5 25 ,for. primary membrane loading.

NOTE 9 - Where deformation is of concern in a component, the deforma-i' tion shall be limited to two-thirds the value given for

!, Emergency Conditions in the Design Specification.

NOTE 10 - When loads are transiently applied, consideration should be given to the use of dynamic load amplification and possible change in modulus of elasticity.

\

l i

t 1

0 l

1 s

\

3.9 ,33 Am. No. 56, (3/81) 1

- --~,-n --e,--- , m.e-....w- - .-w-,.y ,~,.,,w - ...w,. w-- -ne- ,g,-,.y. , - . - -.--.-w,,--- - , - - , , - .--,m,.--r,.yswn4. -+n-,,,y ., - .--.n,,

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56 TABLE 3.9-15 J

CORE SUPPORT STRUCIURES STRESS CATEGORIES AND LIMITS OF STRESS INTENSITY FOR FAULT CONDITIONS 4

+

PRIMARY STRESSE5 SECONDARY STRE55E5 PE AK STRESSES g,

CATECORY PgAs, NtNeRANf,P,(NOffl3.2& M B f M DINO, Pg (NO T E S I.7 4 M Mf MBR ANE & BENDINO StCCNDART,O F e

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3. 3 3 L (NO f f M E

gNo g g as gNo v g g 3,gg, NOT R( W R(O NOT RENRf D OR OR PL A17tC TEST

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ACNCS-PSAR TABLE 3.9-15 (Cont'd)

NOTE 1 -

The symbols P,, Pb , Q, and F do not represent quantities but rather sets of six quantities representing the six stress components, a g, ay, ar' *t l' #

1r'

""d

• r t*

NOTE 2 -

When loads are transiently applied, consideration should be given to the use of dynamic load amplification and possible changes in modulus of elasticity.

NOTE 3 -

For configurations where compressive stresses occur, stress limits take into account critical buckling stresses. For external pressure, the permissible " equivalent static" external pressure shall be taken as 2.5 times that given by the rules of paragraph NB-3133 of ASME Boiler and Pressure Vessel CodesSection III (ASME III). Where dynamic pressures are in-volved, the permissible external pressure shall satisfy the preceding requirements or shall be limited to 75 percent of the dynamic instability pressure.

NOTE 4 - l g is the lower bound limit load with yield point equal to 1.5 S, (where S, is the tabulated value of allowable stress at temperature as contained in ASME III). The " lower bound limit load" is here defined as that produced from the analysis of an k

ideally plastic (non-strain hardening) material where deforma-tions increase with-no further increase in applied load. The lower bound load is one in which the material everywhere satis-fies equilibrium and nowhere exceeds the defined material yield strength using either a shear theory or a strain energy of dis-tortion theory to relate multiaxial yielding to the uniaxial Case.

NOTE 5 -

S, is the ultimate strength at temperature. Multiaxial effects on ultimate strength shall be considered.

NOTE 6 - This plastic analysis uses an elastic-plastic evaluated nominal primary stress. Strain hardening of the material may be used for the actual monotonic stress-strain curve at the temperature of loading, or any approximation to the actual stress-strain curve which everywhere has a lower stress for the same strain as the actual curve may be used: either the maximum shear stress or strain energy of distortion flow rule shall be used to account for multiaxial effects.

NOTE 7 - For fault conditions, the stress limits need not be satisfied if it can be shown from the test of a prototype or model that the specified loads (dynamic or static equi'21ent) do not exceed 80 percent of 17, where 1,7 is the ultimate lead or load combination used in the test. In using this method, account fm shall be taken of the size effect and dimensional tolerances ar well as differences which may exist in the ultimate strength or (V) other governing material properties of the actual part and the 3.9-35 Am. No. 56, (3/81)

ACNGS-PSAR b

k_,/ TABLE 3.9-15 (Cont'd) 56 s

tested parts to assure that the loads obtained from the test are a conservative representation of the load carrying capabil-ity of the actual component under postulcted loading for Fault Conditions.

NOTE 8 - Stress ratio is a method of plastic analysis which uses the ttress ratio combinations (combination of stresses that con-sider the ratio of the actual stress to the allowable plastic or elastic stress) to compute the maximum load a strain hard-ening material can carry. K is defined as the section factor; Sf is the lesser of 2.4 S ,or 0.75 S for primary menbrane Ioading.

NOTE 9 - Where deformation is of concern in a component, the deformation shall be limited to 80 percent of the value for Fault Conditions in the Design Specifications.

r l

l l

l O 3.9-36 Am. No. 56, (3/81)

. - - = = - . . == - .-. .-

4 ACNCS-PSAR O -

TABLE 3.9-16 56 DESIGN LOADING CONDITIONS AND COMBINATIONS

(

Operating Design Loading and StressCondition) Limits (I Conditions and Combinations Normal and Upset N and A r N and U D

Emergency N and R or other conditions which have a 40 year encounter probability from 10-1 to 10-3 Fault N and Am and R or other conditions which have a 40 year encounter probability from 10-3 to 10-6 where: N = Normal Loads U = Upset Loads excluding earthquake AD=SafeShutdownEarthquake/2[SSE)includingany associated transients. (2/

A, = Safe Shutdown Earthquake (SSE) including any associ-ated transients.

R = Automatic Blowdown or equivalent Auxiliary Pipe Rupure loading including any associated transients - Pipe Rup-ture loadings are not directly considered on piping it-self because this is handled by a failure mode analysis.

R = Primary Loadings which result from rupture of a main steam line or a recirculation line.

(1) The Design Stress, deformation and fatigue limits are:

For RPV and Appurtenances - ASME Section III.

For Core Support Structures - Tables 3.9-9 through 3.9-15 -

56 For Reactor Internal Structures - Tables 3.9-9 through 3.9-15. ,

O 3.9-37 A"* No. 56, (3/81)

i ACNGS-PSAR 56 TABLE 3.9-17

~-

PRESSURE DIFFERENTIALS ACROSS REACTOR VESSEL INTERNALS Pressure Differences a

Maximum Occurring Initial Steady During a Steam State Values Line Break psi Reactor Component Case 1 Case 2 Case 1 Case 2 56 Core Plate and 4

Guide Tube 22.5 21.5 27 28 Shroud Support Ring and Lower

, Shroud 26.7 24.4 36 43 34 Upper Shroud and (G)

Shroud Head 4.2 2.5 14 23 Channel Ec : 13.8 10.7 16 15 i Dryer

• 0.3 0.07 ** **

l l s_,/ Case 1 - Reactor Initially at 105% Rated Steam Flow.

34 Case 2 - Reactor Initially at 46% Rated Power, 108% Recirculation Flow.

g) 4 j

• Because the only consequence for a dryer failure would be (porsibly) interference with isolation valve closure and because isolation valve closure is not essential for a break inside the drywell, the dryer differential is evaluated for a steam line break outside the primary a containment.

• • An outside steam line break when the reactor is at hot standby would result in a dryer loading of 13 psi. The dryer design should be auch that this differential does not jeopardize steam line isolation 4

closures.

3 (G)-GESSAR 3.9-38 Am. No. 56, (3/81) '

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HOUSTON LIGHTING & POWER COMPANY Allens Creek Nuclear Generating Station Unit 1 TRANSIENT PRESSURE DIFFERENTI ALS FOLLOWING A STEAM LINE BREAK AT THE 105" RATED STEAM FLOW CONDITION FIGURE 3.9-12  ;

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i, ACNGS-PSAR j

! EFFECTIVE PAGES LISTING i Chapter 4  !

REACTOR i Page Amendment I

! l* 56 i 2* 56 3* 56 4* 56 5* 56 l

. 6* 56 l

7* 56 L

i 8* 56 l i 56  ;

j ii 56  ;

i iii 56 i iv 56  :

! v 56 .

l vi 56  ;

l- vii 56 .

viii 56 1x 56 l x 56 xi 56

! xit 56 xiii 56

!' . xiv 56 l xv 56 xvi 56 xvii 56 3 xviii 56 xix 56  :

xx 56  ;

xxi 56

• xxii 56 xxiii 56 xxiv 56 xxv 56 4.1-1 56 4.1-2 56 i 4.1-3 56 4.1-4 56 4.1-5 56  ;

4.1-6 56 4.1-7 56 4.1-8 56 i 4.1-9 56 4.1-10 56 i 4.1-11 56 f 4.1-12 56 4.1-13 56  ;

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• Effective Pages/ Figures Listings .

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ACNGS-PSAR EFFECTIVE PAGES LISTING Chapter 4 REACTOR PAGE AMENDMENT 4.1-15 56 4.1-16 56 4.1-17 56 4.2-1 56 4.2-2 56 4.2-3 56 4.2-4 56 4.2-5 56 4.2-6 56 4.2-7 56 4.2-8 56 4.2-9 56 4.2-10 56 4.2-11 56 4.2-12 56 4.2-13 56 4.2-14 56 4.2-15 56 4.2-16 56 4.2-17 56 4.2-18 56 4.2-19 56 4.2-20 56 4.2-21 56 4.2-22 56 4.2-23 56 4.2-24 56 4.2-25 56 4.2-26 56 4.2-27 56 4.2-28 56 4.2-29 56 4.2-30 56 4.2-31 56 4.2-32 56 4.2-33 56 4.2-34 56 4.2-35 56 4.2-36 56 4.2-37 56 4.2-38 56 4.2-39 56 4.2-40 56 4.2-41 56 4.2-42 56 4.2-43 56 4.2-44 56 4.2-45 56 4.2-46 56 4.2-47 56 2 Am. No. 56, (3/81)

I ACNGS-PSAR I

J EFFECTIVE PAGES LISTING Chapter 4 t

i REACTOR i PAGE AMENDf1ENT l i

i 4.2-48 56 4.2-49 56 4.2-50 56 4.2-51 56

] 4.2-52 56 .

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4.2-56 56 i 4.2-57 56  ;

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4.3-2 56 I 4.3-3 56 f 4.3-4 56 4.3-5 56 i 4.3-6 56 4.3-7 56 4.3-8 56 l 4.3-9 56

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j 4.3-12 56 l 4.3-13 56  ;

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4.3-20 56 4.3-21 56-4.3-22 56 4.4-1 56 i 4.4-2 56 l 4.4-3 56

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3 Am. No. 56, (3/81)

ACNGS-PSAR EFFECTIVE PAGES LISTING l Chapter 4 REACTOR PAGE AMENDMENT 4.4-14 56 4.4-15 56 4.4-16 56 4.4-17 56 4.4-18 56 4.4-19 56 4.4-20 56 4.4-21 56 4.4-22 56 4.4-23 56 4.4-24 56 4.4-25 56 4.4-26 56 4.4-27 56 4.4-28 56 4.4-29 56 4.4-30 56 4.4-31 56 4.4-32 56 4.4-33 56 4.4-34 56 4.4-35 56 4.4-36 56 4.4-37 56 4.4-38 56 4.5-1 56 4.5-2 56 4.5-3 56 4.5-4 56 4.5-5 56 4.5-6 56 4.5-7 56 4.5-8 56 4.5-9 56 4.6-1 56 4.6-2 56 4.6-3 56 4.6-4 56 4.6-5 56 4.6-6 56 4.6-7 56 4.6-8 56 4.6-9 56 4.6-10 56 4.6-11 56 4.6-12 56 4.6-13 56 4.6-14 56 56 4.6-15 4.6-16 56 4

Am. No. 56, (3/81)

ACNGS-PSAR EFFECTIVE PAGES LISTING Chapter 4 O REACTOR PAGE AMENDMENT 4.6-17 56 4.6-18 56 4.6-19 56 4.6-20 56 4.6-21 56 4.6-22 56 4.6-23 56 4.6-24 56 4.6-25 56 4.6-26 56 4.6-27 56 4.6-28 56 4.6-29 56 4.6-30 56 4.6-31 56 4.6-32 56 4.6-33 56 4.6-34 56 4.6-35 56 4.6-36 56 O 4.6-37 4.6-38 4.6-39 56 56 56 4.6-40 56 4.A-1 56 4.A-2 56 4.A-3 56 4.A-4 56 4.A-5 56 4.A-7 56 i

i O

5 Am. No. 56, (3/81)

.. - _ . - . - _ _ _ - . - - _ . - . - . - - - - . - . . _ . _ . . - . . - = . - . ... , - . - . - . _._..

ACNGS-PSAR EFFECTIVE FIGURES LISTING

• CHAPTER 4 REACTOR

• All figures, whether labeled " Unit 1" or " Units 1 and 2" are to be considered applicable to Unit No. 1.

Figure No. Amendment 4.2-1 56 4.2-2 56 4.2-3 56 4.2-4a 56 4.2-4b 56 4.2-5 56 4.2-6 56 4.2-7 56 4.2-8 56 4.2-9 56 4.3-1 56 4.3-2 56 4.3-3 56 4.3-4 56 4.3-5 56 1.3-6 56 4.3-7 56 4.3-8 56 4.3-9 56 4.3-10 56 4.3-11 56 4.3-12 56 4.3-13 56 4.3-14 56 4.3-15 56 4.3-16 56 4.3-17 56 4.3-18 56 4.3-19 56 4.3-20 56 4.3-21 56 4.3-22 56 4.3-23 56 4.3-24 56 4.3-25 56 4.3-26 56 4.3-27 56 4.5-28 56 4.3-29 56 4

6 Ae, No. 56,(3/81)

O

ACNGS-PSAR EFFECTIVE FIGURES LISTING CHAPTER 4 O Figure No.

REACTOR Amendment 4.4-1 56 4.4-2 56 4.4-3 56  ;

4.4-4 56 4.4-5 56

, 4.4-6 56 4.4-7a 56 j 4.4-7b 56 4.4-7c 56 i 4.4-8 56  !

4.6-1 36 4.6-2 56 4.6-3 56 4.6-4 56 4.6-Sa 56 4.6-5b 56 4.6-5c 56 4.6-6 56 4.6-7 56 4.6-8 56 4.6-9 56

4.6-10 56 i

4.6-11 56 4.6-12 56 4.A-la 56 4.A-lb 56 4.A-1c 56 4.A-Id 56 4.A-2a 50 4.A-2b 56

, 4. A-2c 56

4. A-2d 56 4.A-2e 56 4.A-3a 56 t 4.A-3b 56

)

4.A-3c 56 4.A-3d 56

4. A-3e 56 l 4. A-4a 56

4. A-4b 56

4. A-4c 56

4. A-4d 56

4. A-4e 56

4. A-Sa 56

4. A-5b 56 4.A-Sc 56

4. A-5d 56

4. A-Se 56

4. A-6a 56

4. A-6b 56 4.A-6c 56 4

7 Am.No. 56, (3/81)

ACNGS-PSAR EFFECTIVE FIGURES LISTING CHAPTER 4 REACTOR l l

Figure No. Amendment <

l 4.A-6d 56 i 4.A-6e 56 i 4.A-7a 56  !

4.A-7b 56 4.A-7c 56

4.A-7d 56 4.A-7e 56 4.A-8a 56 4.A-8b 56 4.A-8c 56 4.A-8d 56 4.A-8e 56 4.A-9a 56 4.A-9b 56 4.A-9c 56 ,

4.A-9d 56 4.A-9e 56 4.A-10 56 4.A-11 56 4.A-12 56 4.A-13 56 4.A-14 56 4.A-15 56 4.A-16 56 8 Am. No. 56,(3/81)

i l

i ACNGS - PSAR t

}

l TABLE OF CONTENTG

]

CHAPTER 4 ,

} l REACTOR l D*ction Title Paf;e

)

4 ,

l 4.1 SU!U ARY DESCRIPTION 4.1-1 (

i

{'

4.1.1 Reactor Vessel k.1-1  !

}

j h.1.2 Reactor Internsl Components h.1-1 l

h.1.2.1 Reactor Core 4.1-1 '

4.1.2.1.1 General 4.1-1 J

4.1.2.1.2 Core Configuration h.1-3 l

4.1.2.1 3 Fuel Assembly Description 4.1 4 i 4.1.2.1 3 1 Fuel Rod 4.1 4 4.1.2.1 3 2 Fuel Bundle 4.1-4 4.1 2.1.4 Assembly Support and Control Rod Location 4.1-4 f 4.1.2.2 Core Shroud 4.1-5 f l ,

j 4.1.2 3 Shroud Head and Steam Separators k.1-5 [

4.1.2.h Steam Dryer Assembly h.1-5 l 4.1 3 Reactivity Control Systems 4.1-5 4.1 3 1 Operation 4.1-5 4.1 3 2 Description of Control Rods 4.1-6 h l.3 3 Supplementary Reactivity Control 4.1-6 .

l h.1.4 Analysis Techniques 4.1-7 i 4.1.4.1 Reactor Internal Components 4.1-7 4.1.4.1.1 MASS (Mechanical Analysis of Space Structure) 4.1-7 d 4.1.4.1.1.1 Program Description 4.1-7 4.1.4.1.1.2 Program Version and Computer 4.1-7 4.1.h.1.1 3 History of Use h.1-7 4.1.4.1.1.4 Extent of Application 4.1-7 4.1.4.1.2 SNAP (MULTISHELL) 4.1-8 i Am. No.56, (3/81)

ItCNGS - PCAR TABLE OF CONTENTS (Continued)

S.etion Title Pace h.1.h.1.2.1 Pror. "m Description b.1-8 h.1.k.l.2.2 Prograu Vercion and Computer 4.1-8 4.1.h.1.2 3 History of Use h.1-8 h.1.h.l.2.N Extent of Application 4.1-8 4.1.h.1 3 GA3P h.1-8 4.1.h.1 3 1 Prodrau Description h.1-8 4.1.4.1 3 2 Program Version and Computer 1. 1-8 L.1.h.1 3 3  !!isti>ry of Use 4.1-9 h l.4.1 3.4 Extent of Application 4.1-9 4.1.4.1.4 N0llEAT 4.1-9 h.1.h.1.4.1 Program Description 4.1-9 L.1.4.1.h.2 Program Version and Computer 4.1-9 4.1.4.1.4.3 History of Use 4.1-9 4.1.4.1.h.4 Extent of Application 4.1-9 h.1.4.1 5 FINITE '4.1-9 4.1.4.1 5 1 Progran Description 4.1-10 h.l.4.1 5 2 Program Version and Computer 4.1-10 h.1.4.1 5 3 History of Use 4.1-10 h.1.4.1 5 4 Extent of Usage 4.1-10 h.l.h.1.6 DYSEA 4.1-10 4.1.4.1.6.1 Program Desc ription 4.1-10 l h.1.4.1.6.2 Program Vtraion and Computer 4.1-10 4.1.4.1.6.3 History of Use 4.1-11 4.1.4.1.6.h Extent of Application 4.1-11 4.1.k.l.7 SHELL 5 h.1-ll 4.1.4.1 7 1 Program Description 4.1-11 4.:..h.1 7 2 Program Version and Computer 4.1-11 ii Am No. 56,(3/81)

ACNGS - P3AR TABLE OF CONTENTS (Continued)

O Gection Title Pa6e 4.1.4.1.T.3 liiatory of Use 4.1-11 4.1.4.1 7 4 Extent of Application 4.1-11 4.1.4.1.8 HEATER 4.1-11 4.1.4.1.8.1 Program Description 4.1-12 4.1.4.1.8.2 Program Version and Computer 4.1-12 4.1.4.1.8.3 History of Use 4.1-12 4.1.4.1.8.4 Eotent of Application 4.1-12 4.1.4.1 9 FAP-71 (Fatigue Analysis Prograul) 4.1-12 4.1.4.1 9 1 Program Description 4.1-12 F.1.4.1 9 2 Program Version and Computer 4.1-12 4.1.4.1 9 3 History of Use 4.1-12 4.1.4.1 9 4 Extent of Use 4.1-13 ,

4.1.4.1.10 CREEP / PLAST 4.1-13 4.1.4.1.10.1 Program Description 4.1-13 4.1.4.1.10.2 Program Version and Computer 4.1-13 4.1.4.1.10 3 History of Use 4.1-13 4.1.4.1.10.4 Extent of Application 4.1-13 4.1.4.1.11 ANSYS 4.1-13

4.1.4.1.11.1 Program Description 4.1-13 4.1.4.1.11.2 Program Version and Computer 4.1-14 4.1.4.1.11 3 History of Use k.1-14

4.1.k.1.11.4 Extent of Application 4.1-14

(

4.1.4.1.12 CLAPS-02 4.1-14 4.1.4.1.12.1 Program Description 4.1-14 4.1.4.1.12.2 Program Version and Computer 4.1-14 4.1.4.1.12 3 History of Use 4.1-15 4.1.4.1.12.4 Extent of Application 4.1-15 iii

_ . _ _ . Am._ _No. 56, (3/81)

ACNGC - PSAR TABLE OF CONTENTS (Continued)

. etion Title Pace k.1.4.1.13 ASIOT 4.1-15 4.1.hol.13 1 Prog: us Description 4.1-15 h.l.h.1.13 2 Program Verulon and Computer 4.1-15 4.1.h.1.13 3 llistory of Use 4.1-15 4.1.4.1.13.h Extent of Application 4.1-15 4.1.4.2 Fuel Rod Thernal Analysis 4.1-15 4.1.4.3 Henator Systema Dynamics 4.1-16 4.1.4.4 Nuclear Analysin 4.1-16 h.l.4 5 Neutron Fluence Calculations h.1-16 h.1.4.6 Thermal llydraulic Calculations 4.1-16 4.1 5 References 4.1-17 4.2 FUEL SYSTEM DESIGN 4.2-1 4.2.1 General and Detailed Design Bases 4.2-1 4.2.1.1 General Design Bases 4.2-1 4.2.1.1.1 Fuel Assembly and Its Components 4.2-1 4.2.1.1.1.1 Safety Design Bases 4.2-1 4.2.1.1.1.2 Basis for Fuel Rod Safety Evaluation 4.2-2 4.2.1.1.1 3 Design Ratios 4.2-2 4.2.1.1.1.4 Maximum Allovable Stresses, Cycling and Fatigue Limits 4.2-4 4.2.1.1.2 Control Assembly and Its Components 4.2-5 l 4.2.1.2 Detailed Design Bases 4.2-5 i

4.2.1.2.1 Fuel Assembly and Its Components 4.2-5 4.2.1.2.1.1 Material Selection and Properties 4.2-6 h.2.1.2.1.2 Effects of Irradiation 4.2-6 4.2.1.2.1 3 Flow-Induced Vibration 4.2-7 4.2.1.2.1.4 Fuel Densification 4.2-8 i

l iv Am. No. 56, (3/81)

i i ACNGS - PSAR l

l TABLE OF CONTENTS (Continued) ,

Gection Title Page j

a

(

4.2.1.2.1 5 Fuct Rod Damage Mechanisms h.2-8 i l

4.2.1.2.1.6 Dimensional Stability 4.2-8 l l  !

f 4.2.1.2.1 7 Fuel Shipping and Handling 4.2-8 !

j 4.2.1.2.1.8 Capac Lty for F;ssion Gas Inventory 4.2-9 l i l 4.2.1.2.1 9 Deflection 4.2-9 I 4.J.l.2.1.10 Fretting Wear and Corrosion 4.2-10 4.2.1.2.1.11 Potential for Water-Logging Rupture k.2-10 h.2.1.2.1.12 Potential for Hydriding 4.2-10 ;

4.2.1.2.1.13 Stress-Accelerated Corrosion 4.2-11 ]

1 .

4.6 1 2.1.14 Fuel Reliability 4.2-11

]

h.2.1.2.1.15 Design Basis for Fuel Assembly Surveillance 4.2-12 4.2.1.2.2 Control Assembly and Its Components 4.2-12 h.2.1.2.2.1 Design Acceptability 4.2-12 I

i k.2.1.2.2.2 Control Rod Clearances 4.2-13 i

4.2.1.2.2 3 Mechanical Insertion Requirements 4.2-13 l

4.2.1.2.2.4 Material Selection 4.2-13 l 4.2.1.2.2 5 Radiation Effects 4.2-13 4.2.1.2.2.6 Positioning Requirements 4.2-13 4.2.2 General Design Description 4.2-14 h.2.2.1 Core Cell 4.2-14 4.2.2.2 Fuel Assembly 4.2-14 4.2.2.2.1 Fuel Assembly Orientation h.2-14 I

h.2.2 3 Fuel Bundle 4.2-15 h.2.2 3 1 Puel Rods 4.2-15 '

4.2.2 3 1.1 Puel Pellets 4.2-16 4.2.2 3 2 Water Rods 4.2-16 4.2.2 3 3 Fuel Spacer h.2-17 l

- _ _ _ . _ . __ __ __ _ ___V _ _

Am. No. 56, (3/81) I

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l AC: LGS - PSAR L l

TABLE OF CD:ITEiTS (Continued)

Cection Title Page 4.2.2 3.h Fuel Ch'innel 4.2-17 l l

l 4.2.2 3 5 Tieplaten 4.2-18

\

I h.2.2 3.6 Firiger Sprint;n 4.2-18 h.2.2.h Heac t,ivity Cont rol Accembly 4.2-19 4.2.2.4.1 Control Hods 4.2-19 u.2.2.4.2 Velocity Limiter 4.2-19 h.2 3 Design Evaluationa 4.2-20 h.2 3 1 Results of Fuel Rod Thermal-tiechanical Evaluations h.2 -20 4.2 3 1.1 Evaluation f4ethods 4.2-20 h.2 3 1.2 Fuel Damage Analysis 4.2-21 h.2 3 1 3 Steady-State Thermal-fiechanical Performance 4.2-21 <

4.2 3 2 Hesults from Fuel Design Evaluations 4.2-22 4.2 3 2.1 Flow-Induced Fuel Rod Vibrations 4.2-22 ,

4.2 3 2.2 Potential Damaging Temperature Effects During Transients 4.2-22 4.2 3 2 3 Fretting Wear and Corrosion 4.2-23 4.2 3 2.4 Fuel Rod Cycling and Fatigue Analysis 4.2-23 l 4.2 3 2 5 Fuel Rod Bowing 4.2-23 4.2 3 2.6 Fuel Assembly Dimensional Stability 4.2-23 4.2 3 2 7 Temperature Transients with a Waterlogged I

Fuel Element 4.2-2h 4.2 3 2,7 1 Energy Release for Rupture of Waterlogged Fuel Elements 4.2-2h 4.2 3 2.8 Fuel Densification Analyses 4.2-24 h.2 3 2.8.1 Power Spiking Analysis 4.2-25 4.2 3 2.8.2 Cladding Creep Collapse 4.2-26 4.2 3 2.8.3 Increased Linear Heat Generation Rate 4.2-26 4.2 3 2.6.4 Stored Energy Determination 4.2-26 vi Am No. 56, (3/81)

,. .- . . . ~ - - _ _ - . _ . - . _ _ . _ _ _ . _ . . _ _ . . _ . - . _ - -- . __ _ _ __. _ . - - _ _ - . _ _ -

i i I ACNGS - PSAR

! i i

I TABLE OF CONTENTS (Continued)

U Title Page Section 4.2 3 2 9 Fuel Cladding Temperatures 4.2-26 4.2 3 2.10 Incipient Fuel Center Melting 4.2-27 4.2 3 2.11 Energy Helease During Fuel El: ment Burnout 4.2-27 4.2 3 2.12 Fuel itod Behavior Effects fro:a Coolant Flow Blockage 4.2-29 f

4.2 3 2.13 Channel Evaluation 4.2-29 a

4.2 3 2.14 Fuel Shipping and Handling 4.2-29 4.2 3 2.15 Fuel Assembly - SSE and LOCA Loadings 4.2-29 4.2 3 3 Reactivity Control Assembly Evaluation (Control Rods) 4.2-30 I 4.2 3 3 1 Materials Adequacy Throughout Design

} Lifetime 4.2-30 4.2 3 3 2 Dimensional and Tolerance Analysis 4.2-30 l (,/

bi 4.2 3 3 3 Thermal Analysis of the Tendency to Warp 4.2-30 4.2 3 3.4 Forces for Expulsion 4.2-30 4.2 3 3 5 Effect of Fuel Rod Failure on Control Rod Channel Clearances 4.2-30

4.2 3 3.6 Effect of Blowdown Loads on Control Rod Channel Clearances 4.2-30 4.2 3 3 7 Mechanical Damage 4.2-31

4.2 3 3.T.1 First Mode of Failure 4.2-31 l

4.2 3 3.T.2 Second Mode of Failure 4.2-31 4.2 3 3.8 Analysis of Guide Tube Design 4.2-32 4.2 3 3 9 Evaluation of Control Rod Velocity Limiter 4.2-32 4.2.4 Testing and Inspection 4.2-32 4.2.4.1 Fuel, Hardware and Assembly 4.2-32 p 4.2.4.2 Testing and Inspection (Enrichment and I j Burnable Poison Concentrations) 4.2-33 D

4.2.4.2.1 Enrichment Control Program 4.2-33 4.2.4.2.2 Gadolinia Inspections 4.2-34 vi3 Am. No. 56, (3/81)

O' A 113 - PZAR l

TABLE OF CO:iTE*iTS (Continued) l l

Cection Jitle Page l

h. .4.2 3 Reactor Control hods L.2-3h l

4.2.4 3 Surveillance Inspection and Testing of Irradisted fuel Hoda 4.2-35

[ L.2 5 Operating and Developmental Experience k.2-36 l

l h.2 5 1 Fuel 0; crating Experience h.2-36 4.2 5 2 Fuel Development Experience 4.2-37 4.2 5 3 Fuel Rod Perforation Experience 4.2-33 L

k.2 5 4 Channel Operating Experience k.?-39 k.2.6 References L.2-39 k.3 !iUCLEAR DESIGIl (EQUILIBRIUM CORE) h.3-1 431 Design Bases 4.3-1 h.3 1.1 farety Design Bases h.3-1 4 3 1.1.1 Reactivity Basis 4.3-1 4.3 1.1.2 Overpower Bases 4 3-1

, k.3 1.2 Plant Performance Design Bases h.3-1 432 Description 4.3-2 4 3 2.1 Iluelear Design Description 4 3-3 4 3 2.1.1 Fuel !iuclear Properties 4 3-3 4.3 2.2 Power Distribution h.3 h

! h.3 2.2.1 Local Power Distribution 4.3 h h.3 2.2.2 Radial Power Distribution h.3-5 h.3 2.2 3 Axial Power Distribution 4 3-5 4.3 2.2.4 Power Distribution Calculations 4.3-6 h.3 2.2 5 Power Distribution Measurecents 436 4.3 2.2.6 Power Distribution Accuracy h.3-6 L.3 2.2 7 Power Distribution Anomalies 4 3-6 4.3 2 3 Reactivity Coefficients 4.3-7 Am. N . 56,(3/81)

-. . - _. __ ___ _. _._ .___,._.__ _ ____.. _ _,viii_________ _ _ __ ___ _ ._ __

f ACNGS - PSAR j TABLE OF CONTENTS (Continued) j Section Title Page 4.3 2 3 1 Void Heactivity Coefficients 4.3-7 4.3 2 3 2 Moderator Temperature Coefficient 4 3-8 43233 Doppler Reactivity Coefficient 4 3-8 4.3 2 3.4 Power Coefficient 4 3-9 4.3 2.4 Control Requirements 4 3-9 4 3 2.4.1 Shutdown Reactivity 4.3-9 4.3 2.4.2 Reactivity Variations 4 3-10 4.3 2 5 Control Rod Patterns and Reactivity Worths 4.3-11 l

[

4.3 2 5 1 Rod Control and Information System 4.3-11  !

4.3 2 5 2 Rod Pat'.ern Control System (RPCS) 4.3-11 4.3 2 5 3 Rod Worth Limiter (RWL) 4.3-12  ;

4.3 2 5 4 Control Rod Operation *.3-12 43255 Scram Reactivity 4 3-12 4.3 2.6 Criticality of Reactor During Refueling 4 3-12 4.3 2 7 Stability 4 3-13 4.3 2 7 1 Xenon Transients 4.3-13 4.3 2 7 2 Thermal Hydraulic Stability 4.3-13 4 3 2.8 Vessel Irradiations 4.3-13 4.3 3 Analytical Methods 4.3-14 434 Changes 4.3-14 4.3.4.1 Reactor Core 4.3-14 4.3.4.1.1 Active Core Volume Increase 4 3-14 4 3.4.1.2 Natural Uranium Utilized 4.3-15 4 3 4.1 3 Increase in Nonboiling Water Volume 4.3-15 4.3.4.1.4 Fuel Rod Diameter Reduction 4.3-15 4.3.4.1 5 Prepressurized Fuel Rods 4.3-15 4.3 5 References 4.3-16 ix- Am. No. 56, (3/81) l

AC?;GS - PSAR TABLE F CO:iTEtiTS (Cont inued)

Cection Title Page j

h.h TliEIciAL-l!YDRAULIC DESIGil 4.4-1 j l

4.4.1 Lesign irtsis 4.h-1 l

b.hol.1 Safety Design B:tses b.4-1 i 4.4.1.2 Puser Generation Design Bases 4.4-1 h.h l.3 liequirements for Steady-State Canditions 4.4-1

! 4.h.1.4 Requirements for Transient Conditions h.4-2 h.h l.5 Sum:ury of Design Bases h.4-2 l

l l

4.h.2 Description of Thermal-Hydraulic Design of the Reactor Core 4.4-2 l

4.h.2.1 Sun:stry Comparison 4.4-2 4.4.2.2 Critical Power Ratio 4.4-3 4.h.2.2.1 Boiling Correlations 4.h-3 h.h.2 3 Linear Heat Generation Rate (LHGR) 4.4-3 4.4.2 3 1 Design Power Distribution 4.h 4 4.4.2.3 2 Design Linear IIeat Generation Rates (LHGR) 4.4-4 h.h.2.4 Void Fraction Distribution 4.4 4 4.4.2 5 Core Coolant Flow Distribution and Orificing Pattern 4.4-5 h.h.2.6 Core Pressure Drop and Hydraulic Loads 4.4-5 4.4.2.6.1 Friction Pressure Drop 4.h-6 4.4.2.6.2 Local Pressure Drop 4.4-7 4.b.2.6.3 Elevation Pressure Drop 4.4-7 4.h.2.6.h Acceleration Pressure Drop 4.4-8 4.4.2 7 Correlation and Physical Data 4.4-8 4.4.2 7 1 Pressure Drop Correlations h.4-8 4.4.2 7 2 Void Fraction Correlation 4.4-9 1

h.h.2 7 3 Heat Transfer Correlation 4. 4-9 b.k.2.8 Therual Effects of Operational Transients 4.h-9 x Am. No. 56, (3/81)

1 l

l ACNGS - PSAR i

I TABLE OF CONTENTS (Continued)

Section Title Pas;e 4.4.2 9 Uncertainties in Estimates 4.4-9

! 4.4.2.10 Flux Tilt Considerations 4.4-9 i

i i

l h.4.3 Description of the Thermal and Hydraulic l Design of the Reactor Coolant System 4.4-9  ;

l

• 1 t

l 4.4 3 1 Plant Configuration Data 4.4-9  ;

I 4.4 3 1.1 Reactor Coolant System Configuration 4.4-10 l r

4.k.3 1.2 Reactor Coolant System Therntl Hydraulic Data 4.4-10 l,

j 4.4.3 1 3 Heactor Coolant System Geometric Data 4.4-10  !

i 4.4 3 2 Operating Restrictions on Pumps 4.4-10 l f

4 r i

4.4 3 3 Pover-Flov Operating Map 4.4-10  !

4.4.3 3 1 Limits for Normal Operation 4.4-10 4.4.3 3 1.1 Performance Characteristics 4.4-11 l 4.4.3 3 2 Regions of the Power Flow Map 4.4-11 l 4.4 3 3 3 Design Features for Pover-Flow Control 4.4-11 i

4.4.3 3 3 1 Flow Control 4.4-12 l

i 4.4.3.4 Temperature-Pover Operating Map (PWR) 4.4-13 4.4.3 5 Load-Following Characteristics 4.4-13 4.4.3.6 Thermal and Hydraulic Characteristics

! Summary Table 4.4-13 4.4.4 Evaluation 4.4-13 l

4.4.4.1 Critical Power 4.4-13 l

3 4.h.k.2 Core Hydraulics 4.4-14 1

l

4.4.4 3 Influence of Power Distributions 4.4-14

4.4.4.4 Core Thermal Response 4.4-14 l 4.4.4.5 Analytical Methods 4.4-14 l 4.4.4 5 1 Reactor Model 4.4-14  !

i 4.4.4 5 2 System Flow Balances 4.4-15 4

J

__ _ _ _ _ __ y i Am. No. 56' (3/81)

AC1:L - PGAR TABLE (W CON i'l'NTO (ConL : nued )

Joction Zitle htp: Oll

4. h .1. . ) . 3 'yate.a Heat ,, tlanct 3 4.h-16 l

4 . 4 . 4 . t> Thernal-Hyd raul i

• Stability Analy sis 4.h-17 l

(

f 4.h.4.6.1 Intro tac Lion 4.4-1(

l 6 .4.6.2 thf u e r i pt ion 4 . h - l '(

h h.h.6 3 Stabi Lity Criteria 4.4-18 f

h...h.6.h M tthematical Mo<lel 4.h-Id

[

l 4.h.4.6 5 Analytical Confirmation 4.4-19

\

4.h.h.6.0 Analysis Resulte L.h-19 4.h.4.6.6.1 Impact of Prepressurir.ed Fuel on Stability 4.4-21 l

4.4 5 Testind and Verification 4.h-21 4.h.6 Instrumentation Requirements h.4-21 L.h.T Re fe rences 4. h:-22 I

l h.S REACTOR MATERIALS 4 5-1 451 Control Rod System Structural Materials 4.5-1 4 5 1.1 Material Specifications 4.5-1 4 5 1.2 Austenitic Stainless Steel Components 4 5-2 h.5 1 3 Other Materials 4.5-4 4 5 1.4 Cleaning and Cleanliness Centrol 4.5-4 h.5 1.4.1 Protection of Materials During Fabrication, Shipping and Storage 4 5-4 4.5 2 Reactor Internal Materials 4 5-5 l

4 5 2.1 Material Specifications 4 5-5 l

f l h.5 2.2 Controls on Welding 4 5-6 h.5 2 3 Nondestructive Examinatton of Wrought Seamless Tubular Products 4.5 6 4 5 2.4 Fabrication and Processing of Austenitic Stainless Steel - Regulatory Guide Conformance 4 5-7 xii Am. No.'56, (3/81)

ACNGS-PSAR TABLE OF CONTENTS (Continued)

Section Title _ Page 4.5 2 5 Other Materials 4.5-18 4.5 3 Control Rod Drive llousing Supports 4.5-19 4.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTEMS 4.6-1 4.6.1 Information for Control Rod Drive System (CRDS) h.6-1 4.6.1.1 Control Rod Drive System Design 4.6-1 4.6.1.1.1 Design Bases 4.6-1 4.6.1.1.1.1 General Design Bases h.6-1 4.6.1.1.1.1.1 Safety Design Bases 4.6-1 4.6.1.1.1.1.2 Power Generation Design Basis 4.6-1 4.6.1.1.2 Description 4.6-1 4.6.1.1.2.1 Control Rod Drive tiechanisms 4.6-2 4.6.1.1.2.2 Drive Components 4.6-3 hI h.6.1.1.2.2.1 Drive Piston 4.6-3 4.6.1.1.2.2.2 Index Tube 4.6-3 4.6.1.1.2.2 3 Collet Assembly 4.6-3 4.6.1.1.2.2.4 Piston Tube 4.6-4 4.6.1.1.2.2 5 Stop Piston 4.6 4 4.6.1.1.2.2.6 Flange and Cylinder Assembly 4.6-5 4.6.1.1.2.2.7 Uncoupling Rod and Related Parts 4.6-5 4.6.1.1.2 3 Materials of Construction 4.6-6 4.6.1.1.2.3 1 Index Tube h.6-6 4.6.1.1.2.3 2 coupling Spud 4.6 6 4.6.1.1.2.3 3 Collet Fingers 4.6 6 4.6.1.1.2.3.4 Seals and Bushings 4.6-6 4.6.1.1.2 3 5 Summary 4.6 6

[ \

h.6.1.1.2.4 Control Rod Drive Hydraulic System 4.6-7 Am. No. 56, (3/81)

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xiii

I ACNGS-PSAR TABLE OF CONTE 7"3 (Continued)

Cection Title Page 4 . 6.1.1. 2. '4 .1 Hydraulic Requirements 4.6-7 h.6.1.1.2.4.2 System Description h.6-8 4.6.1.1.2.4.2.1 Supply Pump 4.6-8 4.6.1.1. 2. h . 2. 2 Accumulator Charging Pressure h.6-8 4.6.1.1.2.4.2.3 Drive Water Pressure 4.6-9 4.6.1.1.2.4.M.h Cooling Water Header k.6-9 h.6.1.1.2.4.2 5 Scram Discharge Volume 4.6-9 4.6.1.1.2.h.3 Hydraulic Control Units 4.6-10 h.6.1.1.2.4.3 1 Insert Drive Valve h.6-10 h.6.1.1.2.h.3.2 Insert Exhaust Valve h.6-10 4.6.1.1.2.4.3.3 Withdrav Drive Valve 4.6-11 4.6.1.1.2.4.3.4 Withdrav Exhaust Valve 4.6 11 4.6.1.1.2.h.3 5 Speed Control Units 4.6-11 4.6.1.1.2.4.3.6 Scram Pilot Valve Assembly 4.6-11 4.6.1.1.2.4.3 7 Scram Inlet valve 4.6-11 4.6.1.1.2.4.3.8 Scram Exhaust Valve 4.6-11 4.6.1.1.2.h.3 9 Scram Accumulator 4.6-12 4.6.1.1.2 5 Control Rod Drive System Operation 4.6-12 4.6.1.1.2 5 1 Rod Insertion 4.6-12 4.6.1.1.2 5 2 Rod Withdrawal 4.6-12 4.6.1.1.2 5 3 Scram 4.6-13 4.6.1.1.2.6 Instrumentation 4.6-14 4.6.1.2 Control Rod Drive Housing Supports 4.6-14 4.6.1.2.1 Safety Objective 4.6-14 h.6.1.2.2 Safety Design Bases 4.6-14 h.6.1.2.3 Description 4.6-lh h.6.2 Evaluations of the CRDS h.6-15 xiv Am. No. 56', (3/81)

ACNGS-PSAR TABLE OF C0!! TENTS (Continued)

Page a

) Section Title 4.6.2.1 Failure fiode and Effects Analysis h.6-15 h.6.2.2 Protection from Co: mon !1ide Failures 4.6-15

(

4.6.2.3 Safety Evaluation 4.6-16 4.6.2 3 1 Control Rods 4.6-16 4.6.2 3 1.1 f4aterials Adequacy Throughout Design Lifetime 4.6-16 4.6.2 3 1.2 Dimensional and Tolerance Analysis 4.6-16 4.6.2.3 1.3 Therraal Analysis of the Tendancy to Warp 4.6-16 4.6.2.3.1.4 Forces for Expulsion 4.6-16 4.6.2 3 1 5 Functional Failure of Critical Components 4.6-16 4.6.2 3 1.6 Precluding Excessive Rates of Reactivity Addition 4.6-16 4.6.2 3 1.7 Effect of Rod Failure on Control Rod m Channel Clearances 4.6-17 V 4.6.2 3.1.8 tiechanical Damage 4.6-17 4.6.2.3 1 9 Evaluation of Control Rod Velocity Limiter 4.6-17 4.6.2.3 2 Control Rod Drives 4.6-17 4.6.2.3 2.1 Evaluation of Scram Time 4.6-17 4.6.2.3.2.2 Analysis of !4alfunction Relating to Rod Withdrawal 4.6-17 4.6.2 3 2.2.1 Drive Housing Fails at Attachment Weld 4.6-17 4.6.2.3 2.2.2 Rupture of Hydraulic Line(s) to Drive Housing Flange 4.6-18 h.6.2 3 2.2.2.1 Pressure-under (Insert) Line Break 4.6-18 h.6.2.3 2.2.2.2 Pressure-over (Withdravr./ Line Break 4.6-19 4.6.2.3 2.2.2 3 Simultaneous Breakage of the Pressure-Over (Withdrawn) and Pressure-Under (Insert) Lines 4.6-19 O h.6.2 3 2.2.3 All Drive Flange Bolts Fail in Tension h.6-20 t

v/

xv Am. No. 56,(3/81)

ACNGS-PSAR TABLE OF CO:ITE:ITS (Continued)

Section Title Page k.6.2 3 2.2.+ Weld Joining Flange to Housing Fails in Tension h.6-20 h.6.2 3 2.2.5 Housing Wall Ruptures 4.6-21 L.6.2 3 2.2.6 Flange Plug Blows Out 4.6-22 4.6.2 3 2.2.7 Ball Check Valve Plug E tows Out 4.6-23 4.6.2.3.2.2.8 Drive / Cooling Water 'ressure Control Valve Closure (Reactor Pressure, o psig) 4.6-23 4.6.2.3 2.2 9 Ball Check Valve Fails to Close Passage to Vessel Ports h.6-23 4.6.2.3.2.2.10 Hvdraulic Control Unit Valve Failure s 4.6-24 4.6.2.3 2.2.11 Corlet Fingers Fail to Latch 4.6-24 4.6.2 3 2.2.12 Withdrawal Speed Control Valve Failure h.6-24 4.6.2 3 2.3 Scram Reliability h.6-2h 4.6.2 3 2.h Control Rod Support and Operation 4.6-25 4.6.2.3 3 control Rod Housing Supports 4 6-25 4.6.3 Testing and Verificatiors of the CRDs 4.6-26 4.6.3.1 Control Rod Drives 4.6-26 h.6.3.1.1 Testing and Inspection 4.6-26 l

l 4.6.3.1.1.1 Development Tests 4.6-26 l

l l h.6.3.1.1.2 Factory Quality Control Tests 4.6-26 l

4.6.3 1.1.3 operational Tests 4.6-27 4.6.3.1.1.4 Acceptance Tests 4.6-27 4.6.3 1.1 5 Surveillance Tests 4.6-28 4.6.3.1.1.6 Functional Tests 4.6-29 h.6.3 2 Control Rod Drive Housin6 Supports 4.6-30 h.6.3.2.1 Te ting and Inspection 4.6-30 h.6.4 Information for Combined Performance of Reactivity Control Systems 4.6-30 xvi Am. No. 56,(3/81)

I l J

a

! ACNGS-PSAR l TABLE OF CONTENTS (Continued)

I 1

Section Title Page j h.6.5 Evaluation of Combined Performance 4.6-30 l h.6.6 Standby Liquid Control System 1. 6-30 4.6.6.1 Design Bases 4.6-30 4.6.6.1.1 General Design Bases 4.6-30 i

4.6.6.1.1.1 Safety Design Bases 4.6-30

} i 4.6.6.2 Description h.6-31 l f 4.6.6.3 Safety Evaluation 4.6-3h i ,

i 4.6.6.4 Inspection and Testing 4.6-35 l

f l 4.6.6 5 Instrumentation h.6-36 I I 4.6.T References 4.6-36 l

< h.A.1 INTRODUCTION h.A-1 l

4.A.2 POWER DISTRIBUTION STRATEGY 4.A-1 l

4.A.2.1 Principle 4.A-1

4.A.2.2 Explanation of Principle 4.A-1  !

t I

h.A.2 3 Target Power and Exposure Shape k.A-2  :

i I l 4.A.2.4 Operational Implemer.'ation 4.A-2  ;

l i I

h.A.3 RESULTS OF CORE SIMULATION STUDIES 4.A-3 l 4.A.3 1 Description of Model 4.A-3 4.A.3 2 Cycle Analysis 4.A-3 I

[

4.A.3 3 Uncertainty Analysis 4.A-b h.A.h GLOSSARY OF TERM 3 4.A-15 f 4.A.5 REFERENCES h.A-17  !

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xvii Am. No. 56,(3/81)  ;

ACEGS-PSAR LIST OF TABLFS Table Title Page 4.2-1 Fuel Cladding Conditions of Design Resulting From -

In-Reactor Process Conditions Combined with Earthquake Loading 4.2-42 h.2-2 Ftel Cladding Stress Intensity Limits 4.2-43 h.2-3 Flel Cladding Estimated Number of Cycles per Each Cyclic Condition Used for Fatigue Analysis 4.2-kh 4.2-4 Fuel Data b.2-45 4.2-5 Material Properties 4.2 46 4.2-6 Post-Shipernt Fuel Inspection Plan 4.2 h7 4.2-7 Inspection Equipment 4.2 48 4.2-8 Summary of Experience in Production Zircaloy-Clad UO2 Fuel 4.2-k9 4.2-9 Summary of General Electric Operating Experience with Production Gadolinia-Bearing Fuel h.2-52 4.2-10 Ganeral Electric Developmental Irradiations Zircaloy-Clad 95% TD UO2 Pellet Fuel Rods 4.2-55 4.2-11 General Electric Developmental Irradiations Zircaloy-Clad 95% TD UO2 Pellet Capsules CE Test Reactor 4.2-56 4.2-12 Halden Irradiation Program Status 4.2-57 4.2-13 Fuel Rod Vibration Information b.2-58 4.2-14 Linea- Heat Generation Rate of Calculated 1%

Plastic Diameteral Strain for BWR/6 Fuel 4.2-59 4.3-1 Reactor Core Dimensions 4.3-17 h.3-2 Reactivity Data for the Cold, Xenon-Free State 4.3-18 4.3-3 Reactivity and Control Fraction for Various Reactor States 4.3-19 h.3-4 Summary of BWR/6 Design Revisions (GE Company Proprietary) 4.3-20 4.3-5 Calculated Neutron Fluxes (Used to Evaluate Vessel Irradiation) 4.3-21 0

xviii Am. No. 56,(3/81)

l l

ACNGS-PSAR l LIST OF TABLES (Continued) l Table Title Page 4.3-6 Calculated fleutron Flux at Core Equivalent f Boundary 4.3-22 ,

I I 4.4-1 Thermal and Hydraulic Design Characteristics of  ;

the Reactor Core 4.h-25

! 4.h-2 Axlal Power Distribution Used to Calculate

! MCPR Operating Limit 4.4-27 l 4.4-3 Void Distribution 4.h-28 4.4-4 Axial Power Distribution Used to Generate Void and Quality Distribut, ions 4.4-29 j

4.4-5 Flov Quality Distribution 4.4-30 f

4.4-6 Core Flow Distribution 4.h-31  ;

4.4-7 Calculated Vs Measured Core Plate Pressure Drops 4.4-32 i l

4.4-8 Typical Range of Test Data 4.h-33  !

l L h.4-9 Description of Uncertainties 4.h-3h 4.4-10 Bypass Flow Paths 4.4-36 4.4-11 Reactor Coolant System Geometric Data h.4-37 l f

i 4.4-12 Lengths of Safety Injection Lines 4.4-38 l 4.6-1 standby Liquid Control system operating ,

Pressure / Temperature Conditions Test Modes 4.6-37 [

l 4.6-2 Design Duty Cycle 4.6-39 f I

i 4.6-3 40-Year Design Life Test k.6 40 i

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F l

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I xix Am. No. 56, (3/81) j l i

t

ACNGS-PSAR LIST OF FIGURES _

Figure Title 4.2-1 Sc.ematic of Four-Bundle Cell Arrangment 4.2-2 Funl Assembly 4.2-3 Fuel Assembly Cross Section 4.2-ha Control Rod Assembly 4.2 hb Control Rod Information Diagram for BWR/6 Lattice 4.2-5 Control Rod Velocity Limiter 4.2-6 Fuel Cladding Average Tenperature at a Fuel Column Axial Gap 4.2-7 Cladding Temperature Versus Heat Flux, Beginning-of-Life 4.2-8 Cladding Temperature Versus Heat Flux, End-of-Life 4.2-9 Fuel Energy Release as a Function of Time 4.3-1 Equilibrium Core Loading Map h.3-2 BWR/6 Lattice Nominal Dimensions,120-mil Channel 4.3-3 Rod Type Distribution (CE Company Proprietary) Refer to Figure 4.3-4 4.3 4 Axial Fuel Rod Enrichment and Gadolinia Distribution (GE Company Proprietary) 4.3-5 Uncontrolled k-Infinity as a Function of Exposure at Various Void Fractions, Dominant Fuel Type 4.3-6 Weight Fraction - U238 as a Function of Exposure Dominant Fuel Type, 40% Voids 4.3-7 Weight Fraction as a Function of Exposure, Dominant Fuel Type, 40% Voids 4.3-8 Fission Fraction as a Function of Exposure Dominant Fuel Type, 40% Voids 4.3-9 Neutron Generation Time as a Function of Exposure at 40% Voids O

xx Am No. 56,(3/81)

ACNGS-PSAR LIST OF FIGURES (Continued)

Figure Title 4.3-10 Delayed Neutron Fraction as a Function of Exposure at 40% Voids 4.3-11 Variation of Maximum Local Power Peaking as a Function of Exposure Dominant Fuel Type, 40%

Voids, Uncontrolled 4.3-12 Uncontrolled Local Power Distribution as a Function of Exposure at 40% Voids, Dominant Fuel Type (GE Company Proprietary) 4.3-13 Uncontrolled Local Power Distribution as a Function of Void at 0.0 Lattice Exposure Dominant Fuel Type (GE Company Prorietary) h.3-14 Controlled Local Power Distribution at 40%

Voids, 0.0 Lattice Exposure Dominant Fuel Type (GE Company Proprietary) 4.3-15 Uncontrolled R-Factor Distribution at 40%

Voids, 0.0 Lattice Exposure Dominant Fuel Type (GE Company Proprietary)

% 4.3-16 Variation of the Maximum Uncontrolled Bundle-Integrated R-Factor as a Function of Bundle Average Exposure 4.3-17 Radial Power Factors for Beginning-of-Equilibrium-Cycle and Optimal End-of-Equilibrium Cycle Conditions h.3-18 Peginning-of-Equilibrium Cycle and Optimal End-of-Equilibrium-Cycle Ccre Average Axial Power 4.3-19 Equilibrium Cycle Void Coefficient for Stability Analysis as a Function of Percen?. Voids h.3-20 Dynamic Void Reactivity Coefficient as a Function of Percent Voids at End-of-Equilibrium-Cycle 4.3-21 Doppler Coefficient as a Function of Average Fuel Temperature at End-of-Equilibrium-Cycle 4.3-22 Cold Shutdown Reactivity as a Function of Cyle Exposure Strongest Rod Withdrawn, No Xenon 4.3-23 Banked Position Withdrawal Sequence, RPCS Groups 1-4, Sequence A

p. 4.3-24 Banked Position Withdrawal Sequence, RPCS

( Groups 5-10, Sequence A l

xxi Am. No. 56,(3/81)

L

ACNGS-PSAR LIST OF FIGURES (Continued)

Figu_re Title 4.3-25 Banked Position Withdrawal Sequence, RPCS Groups 1-4, Sequence B 4.3-26 Banked Pouition Withdrawal Sequence, RPCS Groups 5-10, Se luence B 4.3-27 Hot Operating, End-of-Equilibrium-Cycle Scram Re:tetivity (S), as a Function of Control Fraction 4.3-28 Model for one-Dimensional Transport Analysis of Vessel Fluence 4.3-29 Ra ital Power Distributions Used in the Vessel Fluence Calculation 4.4-1 Schematic of Reactor Assembly Showing the Bypass Flow Paths 4.4-2 Damping Coefficient versus Decay Ratio (Second Order Systems) 4.4-3 Hydrodynamic and Core Stability Model 4.4 4 Comparison of Test Results with Reactor Core Analysia 4.4-5 Power-Flov Operating Map 4.4-6 Total Core Stability 4.4-Ta 10 Psi Pressure Regulator Setpoint Step at

-515% Rated Power (Natural Circulation) 4.4-7b 10 Cent Rod Reactivity Step at -51 5% Rated Power (Natural Circulation) 4.4-7c 6-Inch Water Level Setpoint Step at -51%

Rated Power (Natural Circulation) 4.4-8 Relative Bundle Power Histogram for Power Distribution Used in Statistical Analysis (Basis in Reference 1) 4.6-1 Control Rod to Control Rod Drive Coupling

4. -2 Control Rod Drive Unit 4.6-3 Control Rod Drive Scheuatic 4.6-4 Control Rod Drive Unit (Cutaway) 4.6-Sa Control Rod Drive Hydraulic System P&ID xxii Am. No. 56,(3/81)

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

ACNGS-PSAR LIST OF FIGURES (Continued)

Figure Title 4.6-5 b Control Rod Drive Hydraulic System P&ID 4.6-Sc Control Rod Drive Hydraulic System P&ID 4.6-6 Control Rod Drive System Process Diagram and Data 4.6-7 Control Rod Drive Hydraulic Control Unit 4.6-8 Control Rod Drive Housing Support 4.6-9 Standby Liquid Control System Piping and Instrumnt Diagram 4.6-10 Sodiun Pentaborate (NapB 100 16 10H2 O) Solution Volume - Concentration Requirements 4.6-11 Saturation Temperature of Sodium Pentaborate Solution 4.6-12 Boron Requiremnts of Standby Liquid Control System 4.A-la Surrary of Haling Condition 4.A-lb Relative Axial Power and Exposure at (Haling) 6.69 GWd/st Cycle Exposure i

4.A-lc Integrated Power per Bundle (Haling) at 6.69 G GWd/st Cycle Exposure 4.A-ld Average Bundle Exposure (Haling) at 6.69 GWd/

st Cycle Exposure 4.A-2a Summry of 0.2 GWd/st Condition 4.A-2b Relative Axial Power at 0.2 GWd/st Cycle Exposure k.A-2c Relative Axial Exposure at 0.2 GWd/st Cycle Exposure 4.A-2d Integrated Power per Bundle at 0.2 GWD/st Cycle Exposure 4.A-2e Average Bundle Exposure at 0.2 GWd/st Cycle Exposure 4.A-3a Summary or 1.0 GWd/st Condition 4.A-3b Relative Axial Power at 1 GWd/st Cycle Exposure p 4.A-3c Relative Axial Exposure at 1 GWd/st Cycle Exposure xxiii Am. No. 56,(3/81)

r ACNGS-PSAR s LIST OF , FIGURES (Continued)

{

i Fi,ure Title

4. A- 1,3 In+.egrate ! Power per Bundle at 1.0 GWd/st j Cycle Exposure 4.A-3e Average Bundle Exposure at 1.0 gWd/st i' Cycle Exposure 4.A-ha Su= wry of 2.0 GWd/st Condition ll L.A-hb Relative Axial Power at 2 GW3/st l Cycle Exposure L.A-ke Relative Axial Exposure at 2 GWd/st Cycle Exposure 4.A-hd Integrate 3 Power per Bundle at 2.0 l GWd/st Cycle Exposure 4.A-4e Average Bundle Exposure at 2.0 GWd/st Cycle Exposure  :

I h.A-Sa Gumary of 3 0 GWd/st Condition h.A-5b Relative Axial Power at 3 GWdiat Cycle Exposure ,

b.A-Sc Relative Axial Exposure at 3 GWd/st Cycle  !

Exposure i i

h.A-5d Integrated Power per Bundle at 3.0 GWJ/st Cycle Exposure 4.A-Se Average Bundle Exposure at 3 0 GWd/st Cycle Exposure I h.A-6a Su:anry of k.0 GWd/st Condition  :

I  !

h.A-6b Relative Axial Power at b GWd/st Cycle Exposure 4.A-6c Relative Axial Exposure at h GWd/st Cycle Exposure

4. A-6d Integrated Power per Bundle at 4.0 GWd/st Cycle Exposure k.A-6e Average Bundle Exposure at h.0 GWd/st Cycle Exposure h.A-Ta Su:r.ary of 5 0 GWd/st Condition h.A-7b Relative Axial Power at 5 GWd/st Cycle Exposure L.A-Tc Relative Axial Exposure at 5 GWd/st Cycle Exposure l

I I

xxiv Am. No. 56,(3/81) l l

l l

l ...

~ . _ - . ~ . - - - .

_ _ - - . _ . - - . _ = . . --. _ - _ _ . _.- -.---.._.- _.- . =._-_-. - -

ACNGS-PSAR LIST OF FIGURES (Continued)

Figure Title r

k.A-7d Intedrated Power per Bundle at 5 0 GWd/st Cycle Exposure l

I 4.A-Te Averade Bundle Exposure at 5 0 GWd/st Cycle Exposure ,

5 h . A-8a Summary of 5 0 Gud/st Condition h.A-8b Relative Axial Power at 6 GWd/st Cycle Exposure 4.A-8c Relative Axial Exposure at 6 GWd/st Cycle Exposure 4.A Jd Integrated Power per Bundle at 6.0 GWd/st Cycle Exposure t

k . A-8 e Average Bundle Exposure at 6.0 GWd/st Cycle l Exposure i h.A-9a Sutmary of 6.6 GWd/st Condition [

l 4.A-9b Relative Axial Power at 6.6 GWd/st Cycle Exposure 4.A-9c Relative Axial Exposure at 6.6 GWd/st Cycle Exposure b.A-9d Integrated Power per Bundle at 6.6 GWd/st Cycle Exposure

) 4.A-9e Average Bundle Exposure at 6.6 GWD/st Cycle Exposure

< 4.A-10 Sequence A and B Designations h.A-11 Minimum Critical Power Ratio as a Function i of Cycle Exposure 9

4.A-12 Maximum Linear Heat Generation Rate as a j Function of Cycle Expcaure j

h.A-13 Achieved End-of-Equilibriu

n-Cycle Axial Exposure and Target Haling Distributions l

4.A-lh Under-Reactive Model 0.2 GWd/St Cycle Exposure k.A-15 Under-Reactive Model 3 GWd/st Cycle Exposure k.A-16 Over-Reactive Model 3 GWd/st Cycle Exposure

\

xxv Am. No. 56,(3/81)

ACNGS - PSAR

h. REACTOR 4.1

SUMMARY

DESCRIPTION The reactor assasly consists of the reactor vessel, its internal components o r che core (shroud, steam separator and dryer assemblies) and jet pumps. Also included in the reactor assembly are the control rods.

control rod drive (CRD) housings and the control rod dri.ves. Figure 5.1.-1 (Reactor Vessel Cutavay) shows the arrangement of reactor assembly components. A summary of the important design and performance characteristics is given in Subsection 1 3 1.1, " Nuclear System Design Characteristics". Loading conditionc for reactor assembly components are specified in Subsection 3 9 4.1.1 Resctor Vessel The reactor vessel design and description are covered in Section 5.h.

4.1.2 Reactor Internal Connonents The major r3 actor internal components are the core (fuel, channels, control blades and instrumentation), the core support structure (including the shroud, top guide and core plate), the shroud head and steam separator assembly, the steam dryer assembly, the feedwater spargers, the core spray p, spargers and the jet pumps. Except for the Zircaloy in the reactor core,

\ these reactor internals are stainless steel or other corrosion-resistant alloys. Of the preceding components; the fuel assemblies (including fuel rods and channel), control 1, lades, in-core instrumentation, shroud head and steam separator assembly, and steam dryers are removable when the reactor vessel is opened for refueling or maintenance.

4.1.2.1 Reactor Core k.1.2.1.1 General The design of the boiling water reactor (BWR) core, including fuel, is based on the proper combination of many design variables and operating experience.

r These factors contribute to the achievement of high reliability.

The nuclear core design described herein is based on the equilibrium reload cycle rather than the initial cycle. The equilibrium cycle is chosen for the basis of the licensing product for two important reasons:

(1) The equilibrium cycle is more typical of the expected operating state over the life of the reactor (2) The use of the equilibrium cycle generally results in a more conservative licensing basis than the initial cycle. At the FSAR stage, all safety analyses vill be done based on e.ctual fuel design parameters.

Os

I The equilibrium cycle is defined as that reload cycle in which all characteristics are identical to the previous cycles. That is, the reload 4.1-1 Am. No. 56,(3/81)

ACNGS - PSAR bundles, the reload batch size, the reload pattern, the cycle energy, etc.,

and the core behavior remain the same from cycle to cycle.

Since the equilibrium cycle does, in fact, bound the initial core from a licensing point of view, only a submittal of the description of the initial core vill be needed under the rules of 10CFR50 59 A number of important features of the BWR core design are summarized in the following paragraphs:

(1) The BWR core mechanical design is based on conservative application of stress limits, operating experience and experimental test results. The moderate pressure level characteristics of a direct cycle reactor (approximately 1000 psia) result in moderate cladding temperatures and stress levels.

(2) The low coolant saturation temperature, high heat transfer coefficients and neutral water chemistry of the BWR are significant, advantageous factors in minimizing Zircaloy temperature and associated temperature-dependent corrosion and hydride buildup.

The relatively uniform fuel cladding temperatures throughout the core minimize migration of the hydrides to cold cladding zones and reduce thermal stresses.

(3) The basic thermal and mechanical criteria applied in the design have been proven by irradiation of statistically significant quantities of fuel. The design heat transfer rates and linear heat generation rates are similar to values proven in fuel assembly irradiation.

(4) The design power distribution used in sizing the core represents an expected state of operation.

(5) The General Electric thermal analysis basis (GETAB) is applied to assure that tore than 99 9% of the fuel rods in the core are expected to avoid boiling transition for the most severe moderate frequency transient described in Chapter 15 The possibility of boiling transition occurring during normal reactor operation is insignificant.

l (6) Because of the large negative roderator density coefficient of reactivity, the BWR has a number of inherent advantages: (a) uses of coolant flow for load following; (b) inherent self-flattening of the radial power distribution; (c) ease of control; (d) spatial xenon stability; and (e) ability to override xenon, in. order to follow load.

Boiling water reactors do not have instability problems due to xenon. This has been derrnstrated by calculations and by special tests which have been conducted on operating BWRs in an attempt to force the reactor into xenon instability. No xenon instabilities have ever been observed in the test 4.1-2 An. No. 56,(3/81)

ACNGS - PSAR m

results. All of these indicators have proven that xenon transients are highly das ed in a BWR due to the large negative power coefficient of reactivity .

Important features of the reactor core arrangement are as follows:

(1) The bottom-entry cruciform control rods consist of EgC in stainless steel tubes sarrounded by a stainless cteel sheath.

Rods of this design vere first introduced in the Dresden-1 reactor in April,1961,and have accumulated thousands of hours of service.

(2) The fixed in-core fission chambers provide continuous power range neutron flux monitoring. A guide tube in each in-core assembly provides for a traversing ion chamber for calibration and axial detail. Source and intermediate range detectors are located in-core and are axially retractable. The in-core location of the startup and source range instruments provides coverage of the large reactor core and provides an acceptable signal-to-noise ratio and neutron-to-gamra ratio. All in-core instrument leads enter from the bottom and the instruments are in service during refueling. In-core instrumentation is discussed in Subsection 7.6.1.6.

_ (3) As shown by experience obtained at Dresden-1 and other plants, the

[s]

x ,/

operator, utilizing the in-core flux conitor system, can maintain the desired power distribution within a large core by proper control rod scheduling.

(4) The Zircaloy 4 reusable channels provide a fixed flow path for the boiling coolant, serve as a guiding surface for the control rods and protect the fuel during handling operations.

(5) The mechanical reactivity control permits criticality checks during refueling and provides maximun plant safety. The core is designed to be suberitical at any time in its operating history with any one control rod fully withdrawn.

(6) The selected control rod pitch represents a practical value of individual control rod reactivity worth, and allows adequate clearance below the pressure vessel between CRD mechanisms for ease of maintenance and removal.

4.1.2.1.2 Core Configuration The reactor core is arranged as an upright circular cylinder containing a large number of fuel cells and is located within the reactor vessel. The coolant flows upward throu6h the core. The core arrangement (plan view) and the lattice configuration are shown in Figures 4.3-1 and 4.3-2, respectively.

O' I

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f 4.1-3 Am. No. 56,(3/81)

ACNGS - PS AR L.1.2.1.3 Puel Assembly Description As can be seen fron the referenced figures, the BWR core is composed of essentially two components -- fuel assemblies and control rods. The fuel assembly and control rod nechanical configurations (Figurec 4.2-2 and 4.2 ka, respectively) are basically the same as used in Dresden-1 and in all subsequent General Electric boiling water reactors.

4.1.2.1.3.1 Fuel Rod A fuel rod consists of UO2 pellets and a Zircaloy-2 cladding tube. A fuel rod is made by atacking pellets into a Zircaloy-2 cladding tube which is evacuated and backfilled with heliun at 3 atmospheres pressure, and sealed by velding Zircaloy end plugs in each end of the tube. The ASME Boiler and Pressure Vessel Code,Section III, is used as a guide in the nechanical design and stress analysis of the fuel rod. The rod is designed to withstand applied loads, both external and internal. The fuel pellet is sized to provide sufficient clearance within the fuel tube to acconnodate axial and radial differential expansion between fuel and clad. Overall fuel rod design is conservative in its accommodation of the nechanisms affecting fuel in a BWR environnent. Fuel rod design bases are discussed in nore detail in Sabsection 4.2.1.

L.l.2.1.3.2 Fuel Bundle Each fuel bundle contains 62 fuel rods and two water rods which are spaced and support ed in a square (8x8) array by seven spacers and a lower and upper tieplate. The fuel bundle has two important design features:

(1) The bundle design places minimum external forces on a fuel rod; each fuel rod is free to expand in the axial direction.

(2) The unique structural design permits the removal and replacament, if required, of individual fuel rods.

The fuel assemblies, of which the core is comprised, are designed to meet all the criteria for core performance and to provide ease of handling.

Selected fuel rods in each assembly differ from the others in uranium enrichment. This arrangement produces more w11 form power production across the fuel assembly, and thus allows a significant reduction in the amount of heat transfer surface required to satisfy the design thermal linitations.

4.1.2.1.4 Assembly Support and Control Rod Location A few peripheral fuel assemblies are supported by the core plate.

Otherwise, individual fuel asse=blies in the core rest on fuel support pieces mou ited on top of the control rod guide tubes. Each guide tube, with its fuel support piece, bears the weight of four assemblies and is supported by a control rod drive penetration nozzle in the botton head of the reactor vessel. The core plate provides lateral support and guidance at the top of each control rod guide tube.

4.1-4 Am. No. 56,(3/81)

ACNGS - PSAR f%

The top guide, mounted on top of the shroud, provides lateral support and guidance for the top of each fuel assembly. The reactivity of the core is controlled by cruciform control cads, containing boron carbide, and their associated mechanical hydraulic drive system. The control rods occupy alternate spaces between fuel assemblies. Each independent drive enters the core from the bottom, and can accurately position its associated control rod during normal operation and yet exert approximately ten times the force of gravity to insert the control rod during the s; ram mode of operation.

1bttom entry allows optimum power shaping in the core, ease of refuelin6 and cervenient drive maintenance.

4.1.2.2 Core Shroud t'he information on the core shroud is contained in Subsection 3 9.4.2.1.1.

4.1.2 3 Shroud Head and Steam Separators The information on the shroud head and steam separators is contained in Subsection 3 9 4.2.1.2.

4.1.2.4 Steam Dryer Assembly The information on the steam dryer assembly is contained in Subsection 3 9 4.2 5

-s s 4.1 3 Reactivity Control Systems

\

\ 4.1 3 1 Operation The control rods perform dual functions of power distribution shaping and reactivity control. Power distribution in the core is controlled during operation of the reactor by manipulation of selected patterns of rods. The rods, which enter from the bottom of the near-cylindrical reactor core, are positioned to counterbalance steam voids in the top of the core and effect significant power flattening.

These groups of control elements, used for power flattening, experience a somewhat higher duty cycle and neutron exposure than the other rods in the control system.

The reactivity control function requires that all rods be available for either reactor " scram" (prompt shutdown) or reactivity regulation. Because of this, the control elements are mechanically designed to withatand the l dynamic forces resulting from a scram. They are connected to bottom-mounted, hydraulically actuated drive mechanisms which allow either axial positioning for reactivity regulation or rapid scram insertion. The design of the rod-to-drive connection permits each blade to be attached or detached from its drive without disturbing the remainder of the control system. The bottom-mounted drives permit the entire control system to be left intact and operable for tests with the reactor vessel open.

(D

, \v) l Am, No. 56, (3/81)

! 4.1-5

ACNGS - PSAR h.l.3.2 Description of Control Rods The cruciform-shaped control rods contain 72 stainless steel tubes (18 tubes in each wing of the cruciform) filled with vibration compacted baron-carbide powder. The tubes are seal velded with end plugs on either end. Stainless steel balls are used to separate the tubes into individual compartments.

The stainless steel balls are held in position by a slight crimp in the tube. The individual tubes provide containment of the helium gas released by the boron-neutron capture reaction.

The tubes are held in a cruciform array by a stainless steel sheath extending the full length of the tubes. A top handle (Figure 4.2 ha) aligns the tubes and provides structural rigidity at the top of the control rod.

Rollers, housed in the handle, provide guidance for control rod insertion and withdrawal. A bottom casting is also used to provide structural rigidity and contains positioning rollers and a parachute-shaped velocity limiter. The handle and lower casting are velded into a single structure by means of a snall cruciform post located in the center of the control rod.

The control rods can be positioned at 6-in. steps and have a nominal withdrawal and insertion speed of 3 in./sec.

The velocity limiter is a device which is an integral part of the control rod and protects against the low probability of a rod drop accident. It is designed to limit the free-fall velocity and reactivity insertion rate of a control rod so that minimum fuel damage vould occur. It is a one-way device, in that control rod scram time is not significantly affected.

Control rods are cooled by the core leakage (bypass) flov. The core leakage flow is made up of recirculation flow that leaks through the several leakage flow paths, the most important of which are:

(1) the area between the fuel channel and the fuel assembly lover tieplate; (2) holes in the lower tieplate; (3) the area between the fuel assembly lower tieplate and the fuel support piece; (4) the area between the fuel support piece and the control rod guide tube; (5) the area between the control rod guide tube and the core support plate; and (6) the area between the core rtpport plate and the shroud.

h.1 3 3 Supplementary Reactivity Control The core control requirements are met by use of the combined effects of the movable control rods, supplementary burnable poison, and variation of reactor coolant flov. The supplementary burnable poison is gadolinia 0 ) mixed with UO2 in selected fuel rods in each fuel bundle.

(Gd23 4.1-6 Am. No. 56,(3/81)

ACNGS - PSAR d h.l.h Analysis Techniques

4.1.h.1 Reactor Internal Components Computer codes used for the analysis of the internal components are listed as follows:

(1) MASS (2) SNAP (MULTISHELL)

(3) GASP (4) NCHEAT

, (5) FINITE (6) DYSEA (7) SHELL 5 (8) HEATER j (9) FAP-Tl (101 CREEP-PLAST (11) ANSYS (12) CLAPS-02 (13) ASIST Detail description of these programs are given in the following sections.

4.1.h.l.1 MASS (Mechanical Analysis of Space Structure) 4.1.4.1.1.1 Program Description s

The program, proprietary of the General Electric Company, is an outgrowth of the PAPA (Plate and Panel Analysis) program originally developed by L. Beitch in the early 1960s. The program is based on the principle of the finite element mthod. Governing matrix equations are formed in terms of joint' displacements using a " stiffness-influence-coefficient" concept originally proposed by L. Beitch 2. The program offers curved beam, plate and shell elements. It can handle mechanical and thermal loads in a static analysis and predict natural frequencies and mode shapes in a dynamic <

i analysis.

4.1.h.1.1.2 Program Version and Computer i The Nuclear Energy Business Group is using a past revision of MASS. This revision is identified as revision "0" in the computer production library.

The program operates on the Honeywell 6000 computer.

h.1.h.1.1 3 History of Use ,

Since its development in the early 60s, the program has been successfully applied to a vide variety of jet-engine structural problems, many of which involve extremely complex geometries. The use of the program in the Nuclear Energy Business Group also started shortly after its development.

p) h.l.4.1.1.h Extent of Application t

'v Besides the Jet Engine Division and the Nuclear Energy Business Group, the Missile and Space Division, the Appliance Division, and the Turbine Division 4.1-7 Am. No. 56, (3/81)

ACNGS - PSAP.

of General Electric have also applied the progran to a wide rarge of engineering problems. The Nuclear Energy Business Group (NEBG) uses it nainly for piping and reactor internals analyses.

k . l . 4.1. 2 SNAP (!fJLTISHELL) 4.1.4.1.2.1 Progran Description The SNAP Progran, which is also called !fJLTISHELL, is the General Electric Code which determines the loads, deformations and stresses of axisynnetric shells of revolution (cylinders, cones, dises, toroids and rings) for axisynnetric thernal boundary and surface load conditions. Thin shell theory is inherent in the solution of E. Peissner's differential equations for each shell's influence coefficients. Surface loading capability includes pressure , average temperature and linear throughwall gradients; the latter two may be linearly varied over the shell meridian. The theoretical limitations of this progran are the same as those of classical theory.

4.1.h.1.2.2 Prodran Version and Ccaputer The current version naintained by the General Electric Jet Engine Division at Evandale, Ohio is being used on the Honeywell 6000 computer in GE/NEBG.

h.1.4.1.2 3 History of Use The initial version of the Shell Analysis Progran was complated by the Jet Engine Division in 1961. Since then, a considerable amount of rodification and a6dition has been made to acconnodate its broadening area of application. Its application in the NEBG has a history longer than 10 years.

4.1.4.1.2.h Extent of Application The progran has been used to analyze jet engine, space vehicle and nuclear reactor components. Because of its efficiency and economy, in addition to reliability, it has been one of the main shell analysis prograns in General Electric's NEBG.

h.l.4.1.3 GASP 4.1.h l.3 1 Program Description GASP is a finite element program for the stress analysis of axisyenetric or plane two-dimensional geonetries. The element representations can be either quadrilateral or triangular. Axisy=netric or plane structural loads can be input at nodal points. Displacements, temperatures, pressure loads and axial inertia can be acconmodated. Effective plastic stress and strain distributions can be calculated using a bilinear stress-strain relationahip by means of an interactive convergence procedure.

h.}.4.1 3 2 Program Version and Computer The GE version, originally obtained from the developer, Professor E. L. Wilson, operates on the Honeywell 6000 computer.

4.1-8 An. No. 56, (3/81)

ACNGS - PSAR 4.1.4.1.3 3 History of Use J

The program was developed by E. L. Wilson in 1965 3 . The present version in GE/NEBG has been in operation since 1967

} k.l.k.l.3.4 Extent of Application The application of GASP in GE/NEBG is minly for elastic analysis of

, axisymnetric and plane structures under thermal and pressure loads. The GE version has been extensively tested and used by engineers in General Electric Ccnpany.

4.1.4.1.4 N0 HEAT 4.1.4.1.4.1 Program Description +

1 The N0 HEAT program is a two-dimensional and axisymmetric. transient,

nonlinear temperature analysis program. An unconditionally stable numrical integration schei2e is combined with an iteration procedure to compute temperature distribution within the body subjected to arbitrary time- and temperature-dependen' boundary conditions.

l

'lhis program utilizes the finite element method. Included in the analysis 6 i are the three basic forms of heat transfer, conduction, radiation, and convection, as well as internal heat generation. In addition, cooling pipe

boundary conditions are also treated. The output includes temperature of all the nodal points for the time instants specified by the user. The program can handle multitransient temperature input.

k.1.4.1.4.2 Program Version and Computer The current version of the program is an improvement of the program

' originally developed by I. Farboomand and Professor E. L. Wilson of 4 . The program operates on the '

Univer.sity of California at Berkeley

Honeywell 6000 computer. i 4.1.4.1.4.3 History of Use The progran vas' developed in 1971 and installed in General Electric .

Honeyvell computer by one of its original developers, I. Farboomand, in  !

1972. A number of heat transfer problems related to the reactor pedestal ,

have been satisfactorily solved using the program.

{

k.1.4.1.4.4 Extent of Application The progran using finite element formulation is compatible with the finite elemnt, stress-analysis computer program GASP. Such compatibility simplified the connection of the two analyses and minimizes human error. '

4.1.4.1 5 FINITE i

l 4.1-9 Am. No. 56, (3/81) 1

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

ACNGS - PS AR b.1.4.1 5 1 Program Deucription FINITE is a ;eneral-purpose, finite element computer program for elastic stress analysis of two-dimensional structural problems, including: (1) plane stress; (2) plane strain; and (3) axisymmetric structures. It has provision for thermal, mechanical and body force loads. The materials of the structure may be honodenous or nonhomogenous and isotropic or orthotropic.

The development of the FINITE program is based on the GASP prodram (Subsection h.1.4.1 3).

4.1.4.1 5 2 Program Version and Conputer Tne present version of the program at GE/NUBG was obtained from the developer J. E. !!cConnelee of GE/ Gas Turbine Department in 19695 The NEBG version is used on the Honeywell 6000 computer.

h.1.4.1 5 3 History of Use Since its completion in 1969, the program has been widely used in the Gas Turbine and the Jet Engine Departments of the General Electric Company for the analysis of turbine components.

4.1.4.1 5.4 Extent of Usage The prodram 10 used at GE/NEBG in the analyis of axisymmetric or nearly-axisynnetric B'4R internals.

h.l.h.l.6 DYSEA 4.1.4.1.6.1 Program Description The DYSEA (Dynamic and Seismic Analysis) program is a GE proprietary program developed specifically for seismic and dynamic analysis of RPV and internals / building system. It calculates the dynamic response of linear structural systems by either temp 3ral model superposition or response spectrum method. Fluid-structure interaction effect in the RPV is taken into account by way of hydrodynamic mass.

Prodram DYSEA was based on program SAPIV with added capability to handle the hydrodynamic mass ef fect. Structural stiffhess and mass matrices are formulated similar to SAPIV. Solution is obtained in time domain by calculatind the dynamic response mode by mode. Time integration is performed by usind Newmark's p-method. Response spectrum solution is also available as an option.

4.1.4.1.6.2 Program Version and Computer The DYSEA version now operating on the Roneywell 6000 computer of GE, Nuclear Energy Systems Division, was developed at GE by modifying the SAPIV program. Capability was added to handle the hydrodynamic mass effect due to fluid-structure interaction in the reactor. It can handle three-dimensional dynamic problems with beam, trusses, and springs. Both acceleration time histories and response spectra may be used as input.

4.1-10 Am. No. 56, (3/81)

ACNGS - PSAR p

h.1.h.1.6.3 History of Use The DYSEA program was developed in the Summer of 1976. It has been adopted as a standard production program since 1977 and has been used extensively in all dynamic and seismic analysis of the RPV and internals / building system.

h l.h.1.6.4 Extent of Application The current version of DYSEA has been used in all dynamic and seismic analysis since its development. Results from test problems were found to be in close agreement with those obtained from either verifie1 programs or analytic solutions.

4.1.4.1 7 SHELL 5 4.1.4.1 7 1 Program Description SHELL 5 is a finite shell element program used to analyze smoothly curved thin shell structures with any distribution of elastic material properties, boundary constraints, and mechanical thermal and displacement loading conditions. The basic elemnt is triangular whose membrane displacement fields are linear polynomial functions, and whose bending displacement field is a cubic polynomial function 6. Five degrees of freedom (three displacements and two bending rotations) are obtained at each nodal point.

(j) Output displacermnts and stresses are in a local (tangent) surface coordinate system.

Due to the approximation of element membrane displacements by linear functions, the in-plane rotation about the surface normal is neglected.

Therefore, the only rotations considered are due to bending of the shell cross-section, and application of the method is not recommended for shell intersection (or discontinuous surface) problems where in-plane rotation can be significant.

h.l.h.1 7 2 Program Version and Computer A copy of the source deck of SHELL 5 is maintained in GE/NEBG. SHELL 5 operates on the internal computers.

4.1.4.1 7 3 History of Use SHELL 5 is a program developed by Gulf General Atomic IncorporatedT in 1969 The program has been in production status at Gulf General Atomic, General Electric, and at other major computer operating systems since 1970.

4.1.4.1 7.h Extent of Application SHELL 5 has been used at General Electric to analyze reactor shroud support and torus. Satisfactory results were obtained.

h.1.4.1.8 (O

v)

HEATER 4.1-11 Am. No. 56, (3/81)

l l l

A=ias - PSA:t l L.l.4.1.8.1 Program Descri;,t ion HEATEH is a computer program used in the hydraulic design of feedwater spargers and their associated delivery header and piping. The program l utilices test data cbtained by CE using full-scale ucckups of feedwater l spirgera combined with a series of models which represent the complex mixing processes obtained in the upper plenum, downcomer, and lover plenum. Mass l and energy balances throughout the nuclear steam supply system (N5SS) are mdeled in detail d. l k.l.L.l.8.2  :'rogram Version and Computer  !

i Thit program was developed at GE/!!EBG in F01CR/Ji IV for the Honeyvell 6000 cofaputer. ,

I l L.l.k.l.8.3 listory of Use The program was developed by various individuals in GE/:!EBG beginning in 1970. '~he present version of the program has been in operation since January 1972.

4.1.k.l.8.h Extent of Application The program is used in the hydraulic design of the feedvater spargers for each BWR plant, in the evaluation of design modifications, and the i

evaluation of unusual operational conditions.

h l.hol.9 FAP-71 (Fatigue Analysis Progran)  ;

b.1.4.1 9 1 Program Description

'lhe FAP-71 computer code, or F atigue Analysis '.'n.3ra=, is a stress analysis tool used to aid in performing ASML-III Nuclea- Vessel Code structural design calculations. Specifically, FAP-71 is used in determining the primary plus secondary stress range and number of allowable fatigue cycles at points of interest. For structural locations at which the 3S ) AS!E Ccde limit is exceeded, the program can perform either (or both)e (P+Q of two elastic-plastic fatigue life evaluations: (1) the method reported in ASME Paper o8-PVP-3, or (2) the present method documented in Paragraph NB-3228.3 of the 1971 Edition of the ASME Section III Nuclear Vessel Code. The program can acco=odate up to 25 transient stress states of as many as 20 structural locations.

4.1.k.1 9 2 Program Version and Computer The present version of FAP-71 was completed by L. Young of GE/NEBG in 1971. 9 The Program currently is on the NEBG Honeywell 6000 computer.

L.l.b.l.9 3 History of Use Since its completion in 1971, the program has been applied to several design analyses of GE BWR vessels.

O-4.1-12 Am. No. 56,(3/81)

ACNGS - PSAR 4.1.h.l.9.h Extent of Use The program is used in conjunction with several shell analysis programs in determining the fatigue life of BWR mechanical components subject to thernal transients.

k.1.h.1.10 CREEP /PIAST 4.1.4.1.10.1 Program Description A finite element program is used for the analysis of two-dimensional (plane and axisymmetric) problems under conditions of creep and plasticity. The creep formulation is based on the memory theory of creep in which the constitutive. relations are cast in the form of hereditary integrals. The material creep properties are built into the program and they represent annealed 304 stainless steel. Any other creep properties can be included if required.

The plasticity treatment is based on kinematic hardening and von !!ises yield criterion. The hardening modulus can be constant or a function of strain.

h.1.4.1.10.2 Program Version and Computer The program can be used for elastic-plastic analysis with or without the presence of creep. It can also be used for creep analysis without the presence of instantaneous plasticity. A detailed description of theory is given in Reference 11. The program is operative on internal computers.

h.1.h.1.10 3 ' History of Use This program was developed by Y. R. Rashidll in 1971.- It underwent extensive program testing before it was put on production status.

4.1.4.1.10.h Extent of Application The program is used at GE/NEBG in the channel cross-section mechanical analysis.

4.1.h.1.11 ANSYS h.1.4.1.11.1 Program Description ANSYS is a general-purpose finite element computer program designed to solve a variety of problems in engineering analysis.

The ANSYS program features the following capabilities:

(1) ' Structural analysis, including static elastic, plastic and creep,

{

dynamic, seismic and _ dynamic plastic, and large deflection and i stability analysis.

.

! (2) One-dimensional fluid flow analysis. .

4.1-13 Am. No. 56,(3/81)

ACNGS - PSAR (3) Transient heat trans fer analysis includirg conduction, convection, an1 radiation with direct input to thermal-stress analyses.

(4) An extensive finite element library, including gaps, friction interfaces, springs, cables (tension only), direct interfaces (compression only), curved elbows , etc. Many of the elements contain complete plastic, creep, and swelling capabilities.

(5) Plotting - Geonetry plotting is available for all elementa in the ANSYS library, including isometric and perspective views of three-dinensional structures.

(6) Restar t Capability - The ANSYS program has restart capability for several analyses types. An option is also available for saving the stif fness matrix once it is calculated for the structure, and using it for other loadire conditions.

4.1.h.1.11.2 Program Version and Computer The program is maintained current by Svanson Analysis Systems, Inc. of Pittsburgh, Pennsylvania and is supplied to General Electric for use on the Honeywell 6000.

4.1.4.1.11.3 History of Use The ANSYS program has been used for productive analysis since early 1970.

Users now include the nuclear, pressure vessels and piping, mining, structures, bridge, chemical, and automotive industries, as well as nany consulting firms.

h.l.h l.ll.4 Extent of Application ANSYS is used extensively in GE/NEBG for elastic and elastic-plastic analysis of the reactor pressure vessel, core support structures, reactor internals and fuel.

h.1.h.1.12 CLAPS-02 4.1.4.1.12.1 Program Description CLAPS-02 is a general-purpose, two-dimensional finite element program used to perform linear and nonlinear structural mechanics analysis. The program solves plane stress, plane strain and axisy=cetric problems. It ma,y be used to analyze for instantaneous pressure, temperature and flux changes, rapid transients and steady-state, as well as convention.1 elastic and inelastic buckling analyses of structural components subjected to mechanical loading.

4.1.4.1.12.2 Program Version and Computer CLAPS-02 is operational on the Honeywell 6000 computer and has a capacity of 500 nodal points, 400 elements, and 100 time steps.

O 4.1-14 Am No, 56,(3/81)

l ACNGS - PSAR C

h.1.4.1.12.3 History of Use CLAPS-02 is an improved version of CLAPS-01, which was developed primarily for the creep collapse analyais of BWR fuel rods. CLAPS-02 uses a more sophisticated element, has more flexibility with regard to mechanical loading conditions and generally results in less running time than CLAPS-01.

4.1.4.1.12.4 Extent of Application CLAPS-02 has been videly used for stress analysis of fuel assembly ccuponents.

4.1.4.1.13 ASIST h.1.4.1.13 1 Program Description Y

The ASIST program is a General Electric code which can be used to obtain load distribution, deflections, critical frequencies and mode shapes in the "in-plane" or " normal-to-plane" mdes for planar structures of any

orientation that: (1) are statically indeterminate; (2) can be represented by. straight or curved beams; and (3) are under basically any loading, thermal gradient, or sinusoidal excitation. Deformations and resulting load distributions are computed considering all strain energies (i.e. , bending ,

torsion, shear and direct). ASIST also considers the effects of the deflected shape on loads and provides deflections calculated for the

\ structure. In addition to this beam column (large deflection) capability,

the buckling- instability of planar structures can also be calculated.

i k.1.h.1.13 2 Program Varsion and Computer The current program version ( ASIST-02) is used on the Honeywell-6000 computer in the General Electric Nuclear Energy Business Group.

h.l.k.l.13 3 History of Use.

The initial version of the ASIST program was developed 'oy the General Electric Jet Propulsion Division. The program and its predecessors have been in use in the General Electric. Aircraft Engine group for more than 10 years. Its application in GE/NEBG has a history longer than 6 years. >

h.1.4.1.13.4 Extent of Application The ASIST program has been used to determine spring constants, stresses,

! deflections, critical frequencies and associated mdes shapes for frames, shafts, rotors, and other jet engine components. It has been used j extensively as a design and analysis tool for various components of nuclear 3 fuel assemblies.

h.l.k.2. Fuel Rod Thermal Analysis r

Puel rod thermal design analyses are perf rmed utilizing the classical l

O* relationships for heat transfer in cylindrical coordinate geometry with internal heat generation. Conditions of 100% and 116% of rated power are

l. analyzed cooresponding to steady-state and short-term transient . operation.

4.1 Am. No. 56,(3/81) e . - . . - . . - - . , - . _ . - - , . - - - . - - - . , , _ _ - . - - . . . . , , - . . . - -

ACNGS - PSAR Abnormal operation transients are also evaluated to assure that the damage O

limit of 1.07. cladding plastic strain is not violated. The strength theory, terminolody, and strain-stress categories presented in the ASitE Boiler and Pressure Vecsel Code Section III are used as a guide in the mechanical design and stress analysis of the fuel rods.

4.1.4.3 Reactor Systems Dynamics The analysis techniques and computer codes used in reactor systems dynamics are described in Section h of Reference 10. Subsection 4.4.4.6 also provides a complete stability analysis for the reactor coolant system.

4.1.h.4 Nuclear Analysis The analysis techniques are described and referenced in Subsection h.3.3.

The codes used in the analysis are:

Computer Code Function Lattice Physics !!odel Calculates average fev-group cross sections, bundle reactivities, and relative fuel rod powers within the fuel bundle.

BWR Reactor Simulator Calculates three-dimensional nodal power distributions, exposures and thermal hydraulic characteristics as burnup progresses.

h.1.4.5 Neutron Fluence Calculations Neutron vessel fluence calculations were carried out using a one-dimensional, discrete ordinates, Sn transport code with general anisotropic scattering.

This code is a modification of a widely used discrete ordinates code which will solve a vide variety of radiation transport problems. The program vill solve both fixed source and multiplication problems. Slab, cylinder, and spherical geometry are allowed with various boundary conditions. The fluence calculations incorporate, as an initial starting point, neutron fission distributions prepared from core physics data as a distributed source. Anisotropic scattering was considered for all regions. The cross sections were prepared with 1/E flux veighted, P sub (L) matrices for anistropic scattering but did not include resonance self-shielding factors.

Fast neutron fluxes at locations other than the core mid-plane were calculated using a two-dimensional, discrete ordinate code. The two-dimensional code is an extension of the one-dimensional code.

h.1.4.6 Thermal-Hydraulic Calculations The digital computer program uses a parallel flow path model to perform the steady-state BWR reactor core thermal-hydraulic analysis. Program input includes the core geometry, operating power, pressure, coolant flow rate and inlet enthalpy, and power distribution within the core. Output from the 4.1-16 Am. No. 56,(3/81)

AONGS - PSAR progran includes core preucure drop, coolant flow distribution, critical

'" power ratio, and axial variations of quality, density, and enthal for each i

channel type.

4.1 5 References

1. R. L. Crowther, " Xenon Considerations in Design of Boiling Water Reactors," June 1968 ( APED-56ho).

2. L. Beitch, "Shell Structures Golved Nunerically by Using a Network of Partial Panels," AIAA Journal, Volume 5, No. 3, !! arch 1967 3 E. L. Wilson, "A Digital Computer Program For the Finite Element Analysis of Solids with Non-Linear !!aterial Properties," Aerojet General Technical !!eno No. 23, Aerojet General, July 1965

h. I. Farboonand and E. L. Wilson , "Non-Linear Heat Transfer Analysis of Axisynnetric Solids," SESli Report SESt!T1-6, University of California at Berkeley, Berkeley, Cali fbrnia,1971.

5 J. E. ficConnelee, " Finite-Users !!anual", General Electric TIS Report DF 69SL206, thrch 1969

6. R. W. Clough and C. P. Johnson, "A Finite Element Approximation for the Analysis of Thin Shells," International Journal Solid Structures, Vol,

'~"

j 4, 1968.

7 "A Computer Progran for the Structural Analysis of Arbitrary Three-Dimensional Thin Shells," Report No. GA-9952, Gulf General Atonic.

8. A. B. Burgess, " User Guide and Engineering Description of HEATER Computer Progran," !! arch 1974.

9 Young, L. J. , "FAP-71 (Fatigue Analysis Progran) Computer Code," GE/NED Design Analysis Unit R. A. Report No. 49, January 1972.

10. L. A. Carnichael and G. J. Scatena, " Stability and Dynamic Performance of the General Electric Boiling Water Reactor," January 1977 (NEDO-21506 ) .

11. Y. R. Rashid, " Theory Report for Crerp-Plast Computer Progran,"

GEAP-10546, AEC Research and Development Report, January,1972.

+

4.1-17 An. No. 56,(3/81)

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

ACNGS-PSAR 4.2 FUEL SYSTD1 DESIGN 4.2.1 General and Detailed Design Bases 4.2.1.1 General Design Bt ses

] The following paragraphs de.'ine the general nechanical design bases that 4 are considered in defining the design of the fuel assembly and its j components and the control rod assembly and its components. In addition, the fuel design shall meet the Technical Specification limits on Linear Heat Generation Rate (LIUR) which shall not be exceeded during steady state operation.

h.2.1.1.1 Fuel Assembly and Its Components k.2.1.1.1.1 Safety Design Bases

^

The fuel assembly is designed to ensure, in conjuction with the core nuclear characteristics (Section h.3), the core thermal and hydraulic characteristics (Section 4.4), the plant equipment characteristics and the instrurentation and protection system, that fuel damage vill not result in the release of radioactive materials in excess of the guideline values of 10CFR20, 50, and 100.

l The mechanical design process emphasizes that:

(1) the fuel assembly shall provide substantial fission product retention capability during all potential operational modes, and (2) the fuel assembly shall provide sufficient structural integrity to prevent operational impairment of any reactor safety equipment.

Assurance of the design bases considerations is provided by the following fuel assembly capabilities:

o Pressure and Temperature Capabilities i

4 The fuel assembly and its components are capable of withstanding the predicted thermal, pressure, and mechanical interaction ,

! loadings occurring' during startup testing, normal operation, and abnormal operation without impairment of operational capability.

l o Handling Capability ,,-

l The fuel assembly and each component thereof is capable of withstanding loading predicted to occur during normal handling without impaiment of operational capability.

I o Earthquake Loading Capability (1/2 SSE) l The fuel assembly and each component thereof is capable of sustaining in-core loading predicted to occur from an Operating

(

i s

4.2 Am. No. 56,(3/81)

ACNGS-PSAR Basis Earthquake (OBE), when occurring during normal operating conditions without impairment of operational capability.

o Earthquake Loading Capability (SSE)

The fuel assembly and each component thereof is capable of sustaining in-core loading predicted to occur from a Safe Shutdown Earthquake (SSE) when occurring during normal operation

.rithout :

(1) exceeding deflection linits which allow control rod insertion, or (2) fragnentation or severance of any bundle component.

o Accident Capability The capability of the fuel assembly to withstand the control rod drop accident, pipe breaks inside containmnt accidents, fuel-handling accident, recirculation pump seizure accident, and pipe breaks outside the containmnt accidents is determined by analysis of the specific event.

The ability of the fuel assembly and its components to provide the preceding capabilities is evaluated by one or mre of the following:

(1) Analysis developed and design ratios formulated to masure results against acceptance criteria (Subsection h.2.1.1.1.3).

(2) Analytical procedures based upon classical methods which do not change, and are patterned after the ASME Boller and Pressure Vessel Code Section III. This procedure allows analytical comparisons of new and old designs and raintains consistency of design characteristics (Subsection 4.2.1.1.1.h ).

(3) Experience and testing (Subsections h.2.1.2, h.2.4, and 4.2 5).

h . 2 .1.1.1. 2 BasN For Fuel Rod Safety Evaluation Fuel damage is defined as a perforation of the fuel rod cladding which would permit the release of fission products to the reactor coolant.

The mechanisms which could cause fuel damade in reactor operational transients and which are considered in fuel evaluations are: (1) rupture of the fuel rod cladding due to strain caused by relative expansion of the UO2 pellet, and (2) severe overheating of the fuel rod cladding caused by inadequate cooling (Subsection h.2.1.2.1 5).

h . 2.1.1.1 3 Design Ratios Design ratios are defined by the following relationship: D.R. = A/L, where D.R. is the design ratio, L is the limiting paramter value, and A is the actual parameter value. Design ratios of less than one shall be demnstrated for component paramters influenced by loading ccnditions 4.2-2 Am. No. 56,(3/81)

ACNGS-PSAR f- s which mgy affect the structural or dimensional integrity of the fuel

) assembly or any component thereof.

(

O' (1) Limiting Parameter Values

a. Normal and Upset Design Conditions Limiting paremter values for each component shall be determined in the following manner as defined by Table 4.2-1.

For stress resulting from mean value or steady state loading, the limiting value shall be determined by consideration of the material 0.2% offset yield strength or the equivalent strain, as established at operating t emperature.

For stress resulting from load cycling, limiting parameter values shall be determined from fatigue linits.

For stress resulting from loading of significant duration, the limiting paramter shall be determined from considerations of stress rupture as defined by the Larson-Miller pararrters. It mtal temperatures are below the level of applicability of stress rupture for the caterial, or if the yield strength is more limiting, t 'en gN the limiting value of stress shall be determined from consideration of the anterial 0.2% offset yield strength or

(

N- '

) the equivalent strain, as established at operating temperatures.

Where stress rupture and fatigue cyclind are both significant, the fbilowing limiting condition shall be applied:

n __

actual time at stress

{ allovable time at stress I=1 m

+ actual number of cycles at stress allovable cycles at s tress -- I=1 Critical instability loads shall be derived from test data, when available, or from analytical mthods when applicable test data are not available. Deflection limits shall be those values of component deformation which could cause an undesirable event such as impairmnt of control rod movemnt or an excessive bypass fg - flow rate. ! )

    ~-
       /

4.2-3 Am. No. 56,(3/81)

ACNGS-PSAR ba Emergency and Faulted Design Conditions Limiting parameter values shall be determined in the following manner, as defined by Table k.2-1: (1) Stress linits shall be determined from consideration of the ultinate tensile strength or equivalent strain of the caterial, as established at operating temperatures. (2 ) Critical instability loads shall be determined froa test data when available or from analytical nethods when applicable test data are not available. (3) Deflection limits shall be those values of deformation that, if occurring, could lead to a acre serious con. sequence such as prevention of control rod ir.se rtion. (2) Actual Parameter Values Actual parameter values shall be determined from the following considerations:

a. Effective stresses shall be determined at each point of interest using the theory of constant elastic strain energy of distortion:

2ae 2,(,X , ,Y )2 , ( ,Y , ,Z )2 ( ,Z _ ,X)2 2

                            + 6(r 2XY  +7 YZ 7 2)

ZX

b. Stress concentration may be applied only to the alternating stress co=ponent.

c. Design values of instability loads shall be scaled up to allow for uncertainty in ranaer of load application, variat .cn in modulus of elasticity, and difference between the actual case and theoretical one.

d. Calculated values of deflection for comparison with deflection limits may be based on the resulting permanent set after load removal; if load removal or curs before, dacage c47 result.

4.2.1.1.1.4 Maximun Allovable Stresses, Cycling and Fatigue Limits The strength theory, terminology and stress categories presented in the ASME Boiler and Pressure Vessel Code, Section III, are used as a guide in the mechanical design and stress analysis of the reactor fuel rods. The mechanical design is based on the maximun shear stress theory for conbined stresses. The equivalent stress intensities used are defined as the difference betvean the most positive and least positive principal stresses 4.2-4 Am lio. 56,(3/81)

ACNGS-PSAR in a triaxial field. Thus, stress intensities are directly comparable to [-s ) stren6th values found from tensile tests. Table 4.2-2 presents a summary \s_,/ of the basic stress intensity linits that are applied for Zircaloy-2 cladding. The fatigue analysis utilizes the linear cu=ulative damage rule (Miner's hypothesis, !!eference 1) and the Zircaloy fatigue design basis of Reference

2. This correlation includes a safety factor of 2 on stress or 20 on cycles (whichever is more conservative). The fatigue analysis is based on the cycles shown in Table 4.2-3 The expected time duration for each of the subject cyclic loadings is not specified and, for the startup and reduced power cycles, can vary accordin6 to the_ reactor status and power demand. The cyclic condition relating to overpower transients would result from an operator error or equipment malfunction and would, therefore, be expected to be of short duration (less than 8 hours). Additional information regarding the basis for fatigue analysis is presented in Section 6 of Reference 4.

4.2.1.1.2 Control Assembly and Its Components Safets Decign Bases The reactivity control mechanical design shall include control rods and gadolinia burnable poison in selected fuel rods within fuel assemblies and shall meet the following safety design bases: f% (1) The control rods shall have sufficient mechanical strength to ( ) Ns_ / prevent displacement of their reactivity control material. (2) The control rods shall have sufficient strength and be so designed as to prevent deformation that could inhibit their motion. (3) Each control rod shall have a device to limit its free-fall velocity sufficiently to avoid damage to the nuclear system process barrier by the rapid reactivity increase resulting from a free-fsll of the control rod from its fully inserted position to the position where the drive was withdrawn. The design basis of the initial core supplementary fuel / reactivity control rods (UGd) 02 is the same as UO2 fuel rods. Additional information on urania-gadolinia physical and irradiation characteristics and material properties is provided in Reference 29, 4.2.1.2 Detailed Design Bases 4.2.1.2.1 Puel Assembly and Its Components The following paragraphs present the detailed bases which are considered in defining the design of the fuel assembly and its components. (. m 4.2-5 Am. No. 56,(3/81)

1 ACNGS-PS AR h.2.1.2.1.1  !!aterial Selection and Properties The materials will be compatible with BWR conditions and retain their design capability during reactor operation. The mechanical, chemical, thermal and radiation properties utilized as design bases are presented in Section 3 of Reference 3. The basic materials used in fael assemblies are Zircaloy-2 and Zircaloy 4. Type-304 stainless steel, Alloy X-750 and ceramic uranium-dioxide and gadolinia. 4.2.1.2.1.2 Effects of Irradiation (1) Cladding Properties, Fuel Swelling Irradiation affects both fuel and claddind material properties. The effects include increased cladding strength and reduced cladding ductility. In addition, irradiation in a thermal reactor environment results in the buildap of both gaseous and solid fission products within the UOp fuel pellet which tend to increase the pellet diameter (i.e. , fuel irradiation swelling). The irradiation swelling model is based on data reported in References 5 and 6, as well as an evaluation of all applicable hidh exposure dataT. Pellet internal porosity and pellet-to-cladding gap are specified such that the thernal expansion and irradiation swelling are accommodated for the worst-case dimensional tolerances throughout li fe. Observations and calculations based on this refined model for relative UOg fuel / cladding expansion indicate that the as-fabricated UOg pellet porosity is adequate (without pellet dishing) to accommodate the fission product induced UO2 swelling out to, and beyond, the peak exposures expected. (2) Fuel Pellet-to-Cladding Gap and Gap Conductance The primary purpose of the gap between the UO2 fuel pellet and Zircaloy cladding is to accommdate differential diametral expansion of fuel pellet and cladding and, thus, preclude the occurrence of excessive gross diametral cladding strain. A short time af ter reactor startup, the fuel cracks radially and r edistributes out to the cladding. Experience has shown, however, the gap volume remains available in the form of radial cracks to accomrmdate gross diametral fuel expansion. The value of pellet-to-cladding thermal conductance used in BWR fuel design is 1000 Btu /hr-f t2 *F. This design value is empirically derived from postirradiation data on exposed fuel with an initial pellet-to-cladding gap which is significantly larger than that employed in the General Electric fuel design. The use of the constant value of 1000 Btu /hr-ft 2 *F for the pellet-cladding thermal conductance has been found to be a conservative assumption when applied in conj unction with the integral fuel design models. Specifically, the design fission gas release model employed in the determination of fuel rod

4. 2-6 Am. No. 56,(3/81)

ACNGS-PSAR plenum size and cladding vall thickness has been shown to ( overpredict available data on fission gas release when applied with a pellet-cladding thermal conductance model for relative fuel cladding expansion (pellet-to-cladding interaction). It also has been shown to be very conservative relative to available data when a value of 1000 Stu/hr-ft2 *F is used for

pellet-cladding thermal conductance 7. Additional discussion and evaluation of the pellet-to-clad gap conductance of GE BWR fuel prepressurized to 3 atmospheres is contained in References 3h through 37 (3) Axial Ratcheting Axial ratcheting of fuel cladding is not considered in BWR fuel rod design. Prototypical fuel rods have been operated in the No Halden test reactor with axial elongation trangducers.

significant axial ratcheting has been observed . (4) Fuel Melting Temperature Fission product buildup tends to cause a slight reduction in fuel melting temperature. The melting point of UO2 is considered to decrease with irradiation at the rate of 32*C/10,000 mwd /Te based on data from Reference 9 , (5) Fuel Thermal Conductivity s

        )                          In the temperature range of interest (500*C), the fuel thermal conductiviv is not considered to be significantly affected by i

irradiation 10, (6) Fission Gas Release A small fraction of the gaseous fission products is released from

the fuel pellets to produce an increase in fuel rod internal gas pressure. In general, such irradiation effects on fuel performance have been characterized by available data and are

' considered in determining design features and performance. Thus, the irradiation effects on fuel perfbrmance are inherently l considered when determining whether or not the stress intensity 4 limits and temperature limits are satisfied. h.2.1.2.1 3 Flow-Induced Vibration Flow-induced fuel rod vibrations depend on such parameters as flow - velocit;y, fuel rod geometry, fuel spacer pitch, fundamental rod frequency and the excitation fbrees due to fluctuating pressures. The stresses j resulting from flow-induced vibrations are considered in the mechanical l design and evaluations of the fuel rods. These stresses are compared to stress intensity limits as noted in Subsection h.2.1.1.1.h. The flow-induced vibration affecting other fuel assembly components is based upon operational experience which, to date, has shown no significant

          ,      adverse effects.
     /

l 4.2-7 Am No. 56,(3/81)

    .-,-..---. r    . , , ,     --    ,#         ,,,, ,      +     ,  , , - . - , , , - , , , , - - , , , - - , -
                                                                                                                         , ,- --,--~.   ,~ - - +     -s ,, --.-.=-- ..

ACNGS-PS AR 4.2.1.2.1.4 Fuel Densification The fuel densification design bases include the effects of: (1) power spikes due to axial gap fbrnttion; (2) increase in LH3R due to pellet length shortening; (3) creep collapse of the cladding due to axial gap forrution; and (4) changes in stored energy due to decreased pellet-cladding thermal conductance resulting fron increa ?d radial gap s ize. The General Electric fuel densification models are lescribed in References 11, 12, and 13. References 11 ani 12 contain a description of the most rent densification models, as nodified and approved by the NRC. Analyses of the ef fects of fuel densification on the design are contained in Subsection 4.2 3.2.8. h.2.1.2.1 5 Fuel Rod Damage Mechanisms As notel in Subsection 4.2.1.1.1.2, tha lectanisms which could cause fuel damage are (1) rupture of the fuel rol cladding due to strain caused by relative expansion of the UO 2 , pellet, and (2) severe overheating of the fuel rod claddire due to inadequate cooling. A value of 1% plastic strain of the Zircaloy cladding has been defined as the limit belov which fuel damage due to overstraining of the fuel cladding is not expected to occur. The 15 plastic strain value is based on General Electric data on the strain capability of irradiated Zircaloy cladding segments from fuel rods operated in several BWRs7 None of the data ebtained falls below the 1% plastic strain value; however, a statistical distribution fit to the available data indicates the 1% plastic strain value to be approximately the 95% point in the total population. This distribution implies, there fbre, a small (<5%) probability that some claiding sednents may have plastic elongation less than 1% at failure. The Fuel Cladding Integrity Safety Linit32 ensures that fuel damage due to severe overheating of the fuel rod cladding , caused by inadequate cooling, is avoided. h.2.1.2.1.6 Dimensional Stability The fuel assembly and fuel components are designed to assure dinensional stability in service. The fuel cladding and channel specifications include provisions to preclude dimensional changes due to residual stresses. In addition, the fuel assembly is designed to accommodate dinensional changes that occur in service due to thermal differential expansion and irradiation e f fe cts. For exanple, the fuel rods are free to expand axially independent of one another. h.2.1.2.1 7 Puel Shipping and Handling Daring shipnent, the fuel bundle is in a horizontal position with flexible packing separators installed between the fuel rods so that the weight of the fuel rods is supported by the shipping container rather than the spacers. Puel bundle shipping procedures are qualified by a test performed on each new design, and each individual bundle is inspected relative to important 4.2-8 Am, No. 56,(3/81)

ACNGS-PSAR p dimnsional characteristics following shipment to verify that no (vj dimensional deviations have occurred. The two major handling loads considered are (1) the loads due to m'txinum upward acceleration of the fuel assembly while grappled, and (2) the loads due to impact of the fuel ansembly into the fuel support while drappled. 4.2.1.2.1.8 Capacity for Fission Gas Inventory A plenum is providel at the top of each fuel rod to accommodate the fission gas release.1 from the fuel during operation. The design basis is to provide sufficient volume to linit the fuel rod internal pressure so that cladding stresses do not exceed the limits given in Table k.2-2 durind normal operation, and fbr short-term transients of 16% or less above the peak normal operating conditions. (1) Fuel Rod Internal Pressure Fuel rod internal preosure is due to the hellun which is backfilled durind rod fabrication, the volatile content of the UO2 , and the fraction of gaseous fission products which are released from the UO . The most limiting combination of 2 dimensional tolerances is assumed in defining the hot plenum volume used to coppute fuel rod internal gas pressure. The available fission gas retention volume is consertatively determined and the fuel rod internal pressure is calculated using (q) , the perfect gas law. \- (2) Fission Gas Generation and Release A quantity of 135 x 10-3 gram moles of fission gas is produced per ffWd of power production. In fuel rod pressure and stress calculations, h% of the fission gas produced is assumed to be released from any UO2y lume at a temperature less than 3000*F. The above basis has been demonstrated by experiment to be conservative over the complete range of design temperature and exposure conditions 7 (3) Plenum Creepdown and Creep Collapse Creepdown and creep collapse of the plenum are not considered because significant creep in the plenum region is not expected. The fuel rod is designed to be free-standing throughout its li fetime. The temperature and neutron flux in the plenum region are considerably lower than in the fueled region; thus, the nargin to creep collapse is substantially greater in the plenum. Direct measurements of irradiated fuel rods have given no indication of significant creepdown of the plenum. h.2.1.2.1 9 Deflection The operational fuel rod deflections considered are the deflections due to: ( ) (1) annufacturing tolerances; x - 4.2-9 Am. No. 56,(3/81)

AC?iGS-PSAR (2) flow-induced vibration; (3) tnemal ef fects; and (h) axi al losd. Two criteria linit the ragnitude of these dermetions: (1) the claddird stress limits r:ast be satisfied; and (2) the fuel rod-to-fuel rod and fuel rod-to-channel clearances rest be sufficient to allow free passage of cochtn t water to all heat transfer surfaces. The fuel rod-to-fuel rod spacing linit of 0.060 in, and fuel rod-to-channel spacing limit of 0.030 in, are based upon the range of clearances that hav e , in the past, been used in boiling transition testing. More recent testing to clesrances below these values would indicate that a lover linit is acceptable. 4.7..l.2.1.10 Fretting Wear and Corrosion Fretting wear and corrosion have been considered in establishird the fuel mechanical design basis. Individual rods in the fuel assembly are held in position by spacers located at intervals along the length of the fuel rod. Springs are provided in each epacer cell so that the fuel rod is restrained to avoid excessive vibration. h.2.1.2.1.11 Potential for Water-Logging Rupture For waterloggird to occur, the fuel cladding nust have a small pinhole. Pinholes are eliminated during production by 100. leak check of fuel as s enb lies. The 'eak Detector System consists of a high vacuun systen capable of attaining pressures less than 1 x 10-4 torr, and a nacs spectromter capable of detecting leaks as lov as 2 x 10-11 cc/sec. The fue 10-4,1torr. bundle is placed After in the vacuun the vacuun pressurechanber, and evacuated is attained, the rnas to less than 1 x spectromter tuned to the hellun nass range is switched into the systen. The output mter of the cass spectromter vill indicate the presence of any hellun gas in the chamber. This production procedure for the fuel is considered to preclude the potential for a waterlogging rupture. h.2.1.2.1.12 Potential for Hydriding The fuel design bases relative to the clad hydridirg nechanism are to assure, through a combination of ergineerird specifications and strict mnufacturing controls, that production fuel vill not contain excessive quantities of noisture or hydrogenous inpurities. An engineering specification linit on misture content in a loaded fuel rod is defined which is well below the threshold of fuel failure. Proceduzal controls are utiliced in rnnufacturird to prevent introduction of hydrogenous inpurities such as oils , plastics , etc. , into the fuel rod. Hot vacuun outgassird (drying) of each loaded fuel rod just prior to final end plug velding ia employed to assure that the level of misture is well below the specification limit. As a further assurance against chemical attack fron the inadvertent adnission of moisture or hydrogeneous impurities into a fuel rod d2rire mnufacture, a hydrogen getter ruterial is enployed in the 4.2-10 An, No. 56,(3/81)

ACNGS-PSAR f ~'\ upper plenum of all fuel rods. This getter material has been proven effective in both in-pile and out-of-pile tests. The getter material is a (V i zirconium alloy in the form of small chips. These getter chips are loosely packed in a utainless steel tube of which one end is capped, and the other end is covered by vire screening. Additional in formation regardiq; the getter is presented in Section 8 of Reference 4. h.2.1.2.1.13 Stress-Accelerated Corrosion Stress corrosion cracking, the phenamenon whereby a ductile material, such as Zircaloy-2, experiences nonductile fracture, has been identified as a factor in pellet-cladding interaction fuel failure. The aimaltaneous action of certain corrosive agents and local stresses for an extended period of time han been observed to embrittle Zircaloy-2 at tcmperatures typical of those achieved in light water reactors. Samples of Zircaloy-2 fractured in the presence of cadmium or iodine in out-of-pile tests, for example, show very little reduction in area and the fracture surface appears non-ductile. Pellet-cladding interaction type failures also exhibit very little reduction in area and a nonductile fracture surface appearance. Complete understanding of the influence of stress corrosion over certain time donnias, stress levels, environments, and cladding exposure on fuel cladding damage is limited by a developing technology. Consequently, stress-accelerated corrosion is not directly addressed in design analysis assessing the impact of the pellet-cladding interaction me chanism. However, qualitative observations of fuel failures and their relationship to steady-state linear power resulted in the fuel design

  '~'    thange to an 8x8 fuel rod natrix to reduce linear powers.
  '-    4.2.1.2.1.14     Puel Reliabiliqr The fuel component characteristics which can influence fuel reliability include:    (1) fuel pellet thermal and nechanical properties, dimensions, density, and U-235 enrichment; (2) Zircaloy cladding thermal and mechanical properties, dinensions, and defects; (3) fuel rod internal void volume and impurities; (k) fuel rod-to-fuel rod and fuel rod-to-channel spacing; and (5) spring constants of the fuel rod spacer springs which maintain contact between the spacer and the fuel rods. Important fuel pellet, cladding , and associated hardware characteristics, and dimensions are provided in Table 4.2 4 and Figure 4.2-3 The large iolume of irradiation experience to date with GE WR fuel indicates only a few mechanisms which have actually had a direct impact on fuel reliability; namely, cladding defects, excessive deposition of system corrosion products, cladding hydriding resulting from hydrogen impurity, and pellet-cladding interaction.

The cladding defects have been virtually aliminated through implementation of improved quality inspection equipment and more stringent quality control requirements during fuel fabrication. Excessive deposition of corrosion products has also been virtually eliminated through improved control of corrosion product impurities in the reactor feedwater, i O Cladding hydriding is the result of excessive amounts of hydrogenous impurities (moisture and/or hydrogenous materis1) inadvertently introduced (v) l l ! 4.2-11 Am. No. 56,(3/81) I

ACNGS-PSAR into the rod durind the fuel fabrication process. The fuel fabrication process currently includes the following steps to minimize possible fa ilures from this mchanism: (1) drying of components and pellets prior to rod loading; (2) hot vacuun outgassind of all loaded fuel rods prior to the final end-plug veld; and (3) strict control of hydrogenous mterials durind fabrication. In addition, as noted in Subsection h.2.1.2.1.12, every fuel rod contains supplemntary protection in the form of a hydrogenous impurity getter which is placed in the plenum. In early 1972, General Electric made design changes in the 7x7 fuel to reduce the incidence of pellet-claddina interaction in future production. The " improved 7x7" design incorporated a reduced pellet length-to-diamter ratio, chanfered pellet ends and the elimination of pellet dishing to reduce the nac,nitude of pellet distortions contributing to local cladding strains. This desid n also employed an increased claddind heat treatment temperature to reduce the statistical variability in cladding mechanical p rope rt ies . Additional information regarding this cladding material is provided in Gection b of Reference 4. These short-term design chandes have been coupled with the londer term design ef fort which culninated in the 8x8 design, which vss introduced into operatind reactors in the spring of 1974. '41th the 8x8 fuel, peak linear power is reduced by more than 25% relative to the 7x7 fuel design to address the strong dependence of PCI failures on bundle power. The favorable fuel performance of both early General Electric B'4R fuel designs with low linear heat rates, and current 8x8 reload fuel provides assurance of the improved reliability of the 8x8 fuel design. h .2 .1.2.1.15 Design Br sis for Fuel Assembly Surveillance General Electric maintains an active fuel assenbly surveillance program specifically intended to monitor performance in operating reactors to identify and characterize unexpected phenomna which can influence fuel integrity and performance. Outade-oriente 1 examinations are performed contingent on fuel availability as influenced by plant operation. Typically, peak duty fuel assemblies (with respect to exposure, linear heat generation rate, and the combination of both) are designated as lead assemblien for a particular design, and are selectively inspected. Numrous other assemblies are routinely inspected employing the nondestructive techniques discussed in Subsection h.2.4.3 Additional informtion redarding fuel surveillance is contained in Subsection h.2.4.3 h .2.1. 2. 2 Control Assemb:y and Its Components The followind paragraphs present the detailed bases which are considered in definind the design of the control assembly and its components. 4.2.1.2.2.1 Design Acceptability The acceptability of the control rod and control rod drive under scram loadirq; condition is deronstrated by functional testing instead of annlysis or adherence to formally defined stress limits. O 4.2-12 Am. No. 56,(3/81)

1 l l ACNGS-PSAR 4.2.1.2.2.2 Control Rod Clearances is that The basis of the nechanical design of the control rod clearances of the there shall be no interference which will restrict the passage control rod. tion of Layout stadies are performed to assure extrene detail part that, given the worst con In addition, which will restrict the passade of control rods. t to t. preoperational verification is nade on each control rod sys Mechanical Insertion Requirements 4.2.1.2.2 3 l t d to Mechanical insertion requirements durind norual operation i d to be able are se ec e provide adelatte operability and load fbliovind capabil ty, an f peak shutdown to control the reactivity addition resulting from burnout o xenon at 100% power. i Scram insertion requirements are chosen to providel sufficient transients.negat ve reactivity to rret all safety criteria for plant operationa 4.2.1.2.2.h Material Selection based upon The selection of naterials for use in the control rod design isThe irradiated their in-reactor properties. J ortion of the austenitic asseabhr, B4 stainless steel which comprises the na oforthe design pC Iowder, alloy X well known and are taken into account in establishing theThe basic crucifo control rod components. l Electric naterials have been operatind successfully in all Genera reactors. 4.2.1.2.2 5 Radiation Effects f The corrosion rate and the physical properties (e.g. l , density, modulus o teel, and elasticity, diraensional aspects, etc.) of austenitic stain ess s in Alloy X-750 are essentially unaffected by the dirradiation experien the BWR reactor core. l known and yield strength, ultimate tensile strength, percent elonga 4 are considered in nechanical design. The B4C claddind is include the release of Gaseous products and swelling. internal pressure buildup due designed to sustain the resultiGd lishel to products, and the lifetime of the control rod has been estab minimize the ef fects of swelling. 4.2.1.2.2.6 Positioning Requirenents increments (not lengths) are selected to provide adequate Rod positionind The combination of rod speed and notch length power-shapind capability.must also meet the liniting reactivity addition rate crite Am. No. 56,(3/81) 4.2-13

ACNGS-PSAR 4.2.2 General Design Description 4.2.2.1 Core Cell A core cell innediately consists surround it. of a control rod and the four f core cell is associated with a four-lobed fuel supportFigur piece. Around the Each e L.3-2 p outer edge of the core, certain fuel assemblies are not adjacent support to a control rod and are supportel byv indiinnediately pieces. id ual peripheral fuel The 4-bundle top duide cells. is a solid stainless steel plate nachi when seated, sprinds moun ted at the tops of the chaThe and, four fuel ass intobeans grid the (Figure corners of the cell such that the h.2-1). sides of thnnels force the ch e channel contact the 4.2.2.2 Fuel Assembly A fuel assembly it (Figure 4.2-2). consists of a fuel bundle and the approximate a right circular cylinder inside the coThe e to fuel assenblies a assembly is supported by a fuel support piece and re shroud. the t Each fuel op guide. The general confi d uration of the fuel assembly a d n the detailed evolutionary chande in custoner, performanceconfigurations u s of the of the design conception. serviceability requiremats , nanufacturing , and and the experience obtai in Table 4.2-k. A sunnary of fuel assembly mechanical data is prned since the initial esented r h.2.2.2.1 Fuel Assembly Orientation Proper orientation of fuel assemblies in the react verified byloading. visual obset tation and is assured or by core is readily verifi durind core cation procedures assembly orientation exist:Five separate visual indications of proper fuel (1) The channel fastener assemblies, including the s pring and guard used to mintain clearances between channels , are located at one corner rod. control of each fuel assembly adjacent to the centere of th (2) The identification the adjacent controlboss rod. on the fuel assembly handl e pd nts toward (3) The channel passage area. spacing buttons are adjacent to the co t n rol rod (4) The assembly identification numbers fuel assembly center handles, are all readable from the di, which are located on the of the cell. rection cf the (5) There is cell-to-cell replication. 4.2-14 Am. No. 56,(3/81)

ACNGS-PSAR Experience has demnstrated that these design features are clearly visible (qv) so that any misoriented fuel assembly would be readily distinguished durind care loading verification. 4.2.2 3 Fuel Bundle A fuel bundle contains 62 fuel rods and 2 vater rods which are spaced and supported in a square (8x8) array by 7 spacers and the lover and upper tieplates. The lower tieplate has a nosepiece which supports the fuel assembly in the reactor. The upper tieplate has a handle for transferring the fuel bundle from one location to another. The identifying assembly number is engrsvea on the top of the handle and a boss projects from one side of the handle to aid in assurind proper fuel assembly orientation. Both upper and lover tieplates are fabricated from Type-304 atainless steel castings. Finger sprinds, of the same design previously used with 7x7 and 8x8 initial core and reload fuel, are also employed with the BWR/6 fuel desi6n. The finger sprinds are located between the lover tieplate and t' e channel for the purpose of controlling the bypass flow through the flow-path (Subsection h.2.2 3.6). Zircaloy-4 fuel rod spacers equipped  ; with Alloy X-750 springs mintain fuel rod-to-fuel rod spacing. - i 4.2.2 3 1 Fue1 Rods Each fuel rod consists of high dena!.ty (95% TD) UO 2 fuel pellets stacked in  ; a Zircaloy-2 cladding tube which is evacuated, back-filled with helium at 3 atmospheres pressure, and sealed by Zircaloy end plugs velded in each end. The 150-in. active fuel column includes a 6-in. zone of naturally enriched pI v (0 711 vt% U-235) pellets at both the top and bottom. The fuel rod b cladding thickness is adequate to be essentially free-standing under the 1000 psia BWR enviror.nent. Adequate free volume is provided within each fuel rod in the form of pellet-to-cladding gap and a plenum region at the top of the fuel rod to accommodate thermal and irradiation expansion of the UO2 and the internal pressures resulting from the helium fill-gas, impurities and gaseous fission products liberated over the design life of the fuel. A plenum spring, or retainer, is provided in the plenum space to i prevent wement of the fuel column inside the fuel rod during fuel chipping and hndling (Figure 4.2-2). A hydrogen getter is also provided in the plenum spee as assurance against chemical attack from the inadvertent admission of roisture or hydrogenous impurities into a fuel rod during manufacture. . i Two types of fuel rods are utilized in a fuel bundle: tie rods and standard rods (Figure 4.2-3). The eight tie rods in each bundle have lower end plugs which thread into the lower tieplate casting and threaded upper end plugs which extend through the upper tie plate casting. A stainless steel hexagonal nut and lockind tab are installed on the upper end plug to  ; hold the fuel bundle together. These tie rods support the weight of the acaembly onl,y durind fuel handling operations when the assembly hangs by the handle; during operation, the fuel rods are supported by the lower tieplate. Fifty-four rods in the bundle are standard fuel rods. The end plugs of the standard rods have shanks which fit into bosses in the tieplates. l

   /  \
   \V) 4.2-15                        Am. No. 56,(3/81) p                                                                                              ,

L _

ACNGG-PSAH An Alloy X-750 exptnsion sprind in located over the upper end plud shank of each rod in the naaenbly to keep the rods sentel in the lower tiepinte while allowind independent axinl exptnaion by niiding within the holes of the upper tieplate. Additional information concerniry the fuel rc>d exptnsion oprind is provided in Section 7 of Reference 4. The fuel bundles incorporate the use of anall arounta of dadolintun na a barnable poison in selectel standard fuel rods. The irradiation producto of thin procena are other dadolintun isotopen hav ing low croan ocetions. The control nuda ntation ef fect dinappears on a predeternined schedule without cheuges in the chemical composition of the fuel or the phynical makeup of the core. Gone nonenblies contain Tre dadolinia than othera to inprove transverue power flattenind. Alno, ione usenblies contain axially distributed gadolinin to inprove axial power flattenind, Gd 033 in un i tb rnly diatributel in the UOg pellet and forno a no lid nolut ion. The d'ulo linia-uran in fuel roda are fabricated as try characteristic extended end pluga. These extenled eni plugn pernit a positive visual check on the location af each grulolinia-bearind rod after bundle nonenbly. h.".2 3.1.1 Pue l Pellets The fuel pelleto consist of hid h dennity ceranic uranian-dioxide ntnufactured by conptctind and nintering uranium-dioxide powder into right cylindrical pelleta witn flat ends and chanfered ed de3. Som" of the pelleto contain small nnounto of gadolinta na a burnable poinon. The average pellet inneraion denoity in approximately 95% of the theoretical denalty of U0g Ceramic uraniun-diox ide in chemically inert to the ela iding at operatiru tenperaturou and in reolatant to attack by water. Several U-235 enrichmnta are uaed in the fuel anaenblica to reduce the local power penkiry factor. Puel elerynt denign und nanu facturing procedures have been developed to prevent errorn in enrichmnt locat ions within a fuel nonenbly. h.2.2 2.2 Water Rods Twc roda in each fuel bundle are hollow wstcr tuben, one of which ( the spacar-positionind water rod) positions the aeven Zirenloy-h fuel rod apteera axially in the fuel bundle. The vnter rodo are nade fron Zirculoy-2 tubits of olightly larger diamter and thinner wall than the fue l roda. Goveral holen are punched around the circunference of each of the water roda near each end to allow coolant water to flow through the rod. Both water roda have square lower end plugn to prevent rotation. The spacer-positionity water rod in equipped with 114 taba which are welded to its exterior. The aptcer-ponitionind water rod and fuel apacern are nosenblel by niidirt the water rod through the appropriate spacer cell with the welded taba orientel in the direction of the corner of the npacer cell away from the spacer oprind. The rod in then rotated so that the taba are positioned above and below the apteer utructure. The opacer-poottioning rod tu prevented frou rotatirg and unlockind the spacers by engagerent of its aquare lower end plud with n 34u'tre hole in the lower tieplate. Dif ferential thernal exptnoton between the fuel roda und the unter roda can introduct- nxial loadirus into the water roda through the frictional forces between the fuel roda nnd the spacern. The tea t ing whi ch van pe r fo rned to 4.2-16 An. No. 56,(3/80

ACNGS-PSAR address this condition, and to verify the water rod / spacer conceptual O design, is discussed in Section 2 of Reference k and in Reference 14. h.2.2 3 3 Fuel Spacer The primary function of the fuel spacer is to provide lateral support and spacing of the fuel rods, with consideration of thermal-hydraulic performance, fretting wear, strength, neutron economy, and producibility. The spacer represents _ an optimization of these considerations. 11echanical design of the BWR/6 spacer is similar, in concept, to that of the current TxT and 8x8 spacers. The mechanical loadings on the spacer structure during normal operation and transients result from the rod-positioning spacer spring forces, from local loadings at the water rod-spacer positioning device, and a small pressure drop loading. During a seismic event, the spacer must transmit the lateral-acceleration loadings from the fuel rods into the channel, while maintaining the spatial relationship between the rods. As noted, the spacer represents an optimization of a number of cons iderations. Thermal-hydraulic development ef fort has gone into designing the particular configuration of the spacer parts. The resultant configurations give enhanced hydraulic perfbrmance. Extensive flow testing has been performed employing prototypical spacers to define single-phase and two-phase flow characteristics. During the blowdown portion of the postulated loss-of-coolant accident (LOCA), the hydraulic (pressure differential) forces on the spacer are of

 \s    about the came magnitude as those present during normal or transient operation of the fuel. There are no significant lateral hydraulic fbrces
     ,. on the spacer, because the fuel channel maintains the normal flow path daring the blowdown.

h.2.2 3.4 Puel Channel The fuel channel enclosing the fuel bundle is fabricated from Zircaloy 4 and perfbrms three functions: (1) the channel provides a barrier to separate two parallel flow paths - one for flow inside the fuel bundle and the other for flow in the bypass region between channels; (2) The channel guides the control rod and provides a bearing surface for it; and (3) the channel provides rigidity fbr the fuel bundle. The channel is open at the bottom and makes a sliding seal fit on the *.over tieplate surface. At the top of the channel, two diagonally opposite corners have velded tabs, which support the weight of the channel from raised posts on the upper tieplate. One of these raised posts has a threaded hole, and the channel is attached using the threaded channel fastener assembly, which also includes the fuel assembly positioning spring. Channel-to-channel spacing is provided for by l means of the fuel assembly positioning spring and the spacer buttons which are located on the upper portion of channel adjacent to the control rod l passage area. Axial differential expansion between the fuel bundle and its channel in accannodated at the lower tieplate. In addition to meeting design limits, assurance is provided that the ekannels maintain their dimensional integrity, strength, and spatial CJ) 4.2-17 Am. No. 56,(3/81)

l ACWGS-FSAR l l i position throughout their lifetine through specifications on the channel rnterials and manufacturind processes and by quality measurements and process qualifications to ensure compliance with these specifications. Under situations of adverse tolerance stackup, differential thermal expansion bet.reen the stainless steel tieplates and the Zircaloy channel can result in an interference fit; however, the resultant stress and strain levels in the channel do not exceed design limits. The 1e. ads and resultant stress imposed on the fuel channel in the event of control rod interference are also within design limits. 4.2.2 3 5 Tieplates The upper and lower tieplates serve the functions >f supportin6 the veid ht of the fuel and positioning the rod ends during all phases of operation and handling. The loading on the lover tieplate durind operation and transients is comprised of the fuel weight, the weight of the channel, and the forces from the expansion sprinds at the top of the fuel rods. The loading on the upper tie plate durin6 operation is due to the expansion spring force. The expansion sprinds permit differential expansion between the fuel rods without introdacing high axial forces into the rods. Most of the loadin ; on the lover tieplate is due to the veidht of the fuel rods and the channi 1, which are not cyclic loadings. Durin6 accidents, the tieplates are subJe :ted to the normal operational loads plus the blowdown and seismic loading ~. Durind handling, the tieplates are subjected to acceleration and impact loadings. 4.2.2 3.6 Finger Springs Finder springs are employed to control the bypass flow through the channel-to-lover tieplate flow path. They have bee 1 used in the initial core and reload fuel of one BWR/3 and all BWR/4 and later plants. They have also been employed on sorm reload fuel in some additional BWR/2 and BVR/3 plants to control bypass flow through the lover tieplate to channel flow path. Increases in channei vall permanent deflectan at the lower tieplate resultind from creep de forention at operat' ud conditions result in increased bypass flow through the channel to lower tieplate flow path. Changes in the flow through this path affect the total core bypass flov , which, in turn, affects the active coolant flow, void coefficient and operational transients. Finger sprind seals are employed to provide control over the flow through this path uer a wide range of channel vall deflections by maintaining a nearly constant flow area as the channel vall defo rms. The finger springs are located between the lower tieplate and the channel; a rure detailed rechanical description is contained in Section 9 of Reference 4. O

4.2-18 Am. No. 56,(3/81)

                                                                    ;

ACNGS-PSAR m 4.2.2.4 Reactivity Control Assembly I

 %_-   4.2.2.4.1   Control Rods The control rods perform the dual function of power shapind and reactivity control. A design drawing of the control blade is seen in Figure h.2 ha and b. Power distribution in the core is controlled during operation of the reactor by manipulating selected patterns of control rods. Control rod displacement tends to counterbalance steam void effects at the top of the core and results in significant power flattening.

The control rod consists of a sheathed cruci form arrgf of stainless steel tubes filled with boron-carbide powder. The control rods are 9 804 in. in total span and are separated uniformly throughout the core of a 12-in. pitch maxiumu. Each control rod is surrounded by four fuel assemblies. The main structural nenber of a control rod is made of Type-304 stainless steel and consists of a top handle, a botton casting with a velocity limiter and control rod drive coupling, a vertical cruciform center post, and four U-ahaped absorber tube sheaths. The top handle, bottom casting, and center post are velded into a single skeletal structure. The U-shaped sheaths are resistance velded to the center post, handle and castings to fonn a rigid housind to contain the boron-carbide-filled absorber rods. Rollers at the top and bottom of the control rod guide the control rod as it is inserted and withdrawn from the core. The control ("N rods are cooled by the core bypass flow. The U-shaped sheaths are (O) perforated to allow the coolant to circulate freely about the absorber tubes. Operatind experience has shown that control rods constructed as described above are not susceptible to dimensional distortions. The boron-carbide (B4 C) powder in the absorber tubes is compacted to about 70% of its theoretical density, The boron-carbide contains a minimum of 76.5% by weight natural boron. The boron-10 (B-10) minimum content of the boron is 18% by weight. Absorber tubes are made of Type-304 stainless - steel. Each absorber tube is 0.220 in. in outside diameter and has a 0.027 in. vall thickness. Absorber tubes are sealed by a plug velded into each , end. The boron-carbide is longitudinally separated into individual compartments by stainless steel balls at approximately 17-in. in tervals. The steel balls are held in place by a slight crimp of the tube. Should boron-carbide tend to compact further in service, the steel balls will distribute the resulting voids over the length or the absorber tube. 4.2.2.4.2 Velocity Limiter , The control rod velocity limiter (Figure 4.2-5) is an integral part of the , bottot assembly of each control rod. This engineered safeguard protects i against high reactivity insertion rate by limiting the control rod velocity in the event of a control rod drop accident. It is a one-way device in that the control rod scram velocity is not significantly affected, but the control rod dropout velocity is reduced to a permissible limit. f'~s The velocity limiter is in the form of two nearly mated, conical elements . ( ) that act as a large clearance piston inside the control rod guide tube.  ! Ns_/ l 4.2-19 Am. No. 56,(3/81)

            ,                      ACNGS-FSAR The lower conical elemnt Ls separated from the upper conical elemnt by four radial spacers 90 degrees apart and is at a 15-degree angle relative to the upper conical element, with the peripheral separation less than the central separation.

The hydraulic drad forces on a control rod are proportional to approxinately the square of the rod velocity and are negligible at normal rod withdrawal or rod insertion speeds. However, during the scram stroke, the rod reaches high velocit' , and the drad forces must be overcome by the drive mechanism. To limit control rod velocity durind dropout, but not during scram, the velocity limiter is provided with a streamlined profile in the scram (upward) direction. Thus, when the control rod is scrammed, water flows over the smooth surface of the upper conical element into the annulus between the guide tube and the limiter. In the dropout direction, however, water is trapped by the lover conical element and discharged through the annulus between the two conical sections. Because this water is jetted in a partially reversed direction into alter floving upward in the annulus, a severe turbulence is createl, thereby slowing the descent of the control rod assembly to less than 3 11 ft/sec. h.2 3 Design E cluations 4.2 3 1 Results of Fuel Rod Therral-Mechanical Cvaluaticns 4.2 3 1.1 Evalwition Methods Current methods for predicting fuel /claddind interaction in fuel design analysis have been discussed and compared to data in Reference 7 Important material properties used for analysis are provided in Table 4.2-5 Additional in formation regarding evaluation nethods is provided in Section 11 of Reference 4. The mechanical evaluations reported here were perfbrmed at a power letel equal to the license limit plus a power spike allowance, wherever apylicable, which assures with 95% confidence that <1 fuel rod in the core v:.11 exceed the maxinum LH3R for which the fuel has been designed. Additional details regardind this method are presented in Appendix B of Reference 15 Additional discussion of the analysis for fuel prepressurization to 3 atmospheres is contained in References 36 and 37 In the design of BWR Zircaloy-clad UO2 pellet fuel, continuous functional variations of mechanical properties with exposure are not employed, since the irradiation effects become saturated at very lov exposure. At be61nnind of life, the cladding mechanical properties employed are the unirradiated values. At subsequent times in life, the cladding nechanical properties employed are the saturated irradiated values. The only exception to this is that unirradiated nechanical properties are employed above the temperatures for which irradiation effects on cladding mechanical properties are assumed to be annealed out. The values of clad yield strendth and ultimate tensile strendth enployed represent the approximate lover bound of data on claddind fabricated by General Electric. 4.2-20 Am No. 56,(3/81)

J - ACNGS-PSAR

    < ~g In the design analysis, the calculated stress and the yield strength or
 ;
   'v}   ultimate strength, are combined into a dimensionless qu6ntity called the design ratio. This quantity is the ratio of calculated stress intensity to the design stress linit for particular stress category. The design stress limit fbr particular stre s category is defined as a fraction of either the yield strength or ultimate strength, whichever is lower. Thus, the design ratio is a mensare of the fYaction of the allowable stress represented by the calculated stress.

Analyses are performed to show that tPi stress intensity linits given in Table 4.2-2 are not exceeded during continuous operation with linear heat  ; generation rates up to the desh3n operating limit, or during tranaient operation above the design operating limit. Stresses due b) external coolant pressure, internal gas pressure, thermal ef fects, spacer contact, flow-induced vibration, and runufactaring tolerances are considered. Cladding nechanical properties used in streas analysis are based on test data of fuel rod cladding for the applicable temperature. 4.2 3 1.2 Puel Damage Analysis As noted in Subsection h.2.1.1.2, fuel damage is defined as a perforation of the fuel rod cladding which would permit the release of fission products to the reactor coolant. For fresh UO 2 fuel, the calculated linear heat generation rate (LHGR) corresponding to 1% diametral plastic strain of the cladding has been

     g calculated and the results are presented in Table 4.2-14. Due to depletion of fissionable material, the high exposure fuel has lesa nuclear capability
   '- ') and will operate at correspondingly lower powers, so that a significant margin is maintained throaghout life between the operating LHGR and the LIUR calculated to cause 1% cladding diametral strain.

The addition of small amounts of gadolinia to UO2 results in a reduction in the fuel thermal conductivity and melting temperature. The result is a reduction in the LHGRs calculated to cause 1% plastic diametral strain fbr gadolinia-urania fuel rods 35 Hos ever, to compensate for this the ' gadolinia-urania fuel rods are designed to provide margins similar to standard UO2 rods. 4.2 3 1 3 Steady-State Thermal-Mechanical Performance The fuel has been designed to operate at core rated power with sufficient design margin to accommodate reactor operations and satisfy the mechanical design bases discussed in detail in Subsection 4.2.1. In order to accomplish this objective, the fuel was designed to operate at a maximum steady-state LHGR of 13.4 kW/ft, plus an allowance, wherever applicable, for densification power spiking. Thermal and mechanical analysis have been performed which demonstrate that the mechanical design bases are mt for the maximum operating power and . exposure combination throughout fuel life. Design analysis have been perfbrmed for the fuel which show that the stress intensity limits given in l (% j Table h.2-2 are not exceeded during continuous operation with LHGRs up to l t the operating limit of 13.4 kW/ft, nor fbr short-term transient operation V i e 4.2-21 Am. No. 56,(3/81)

ACNGS-PSAR 1 1 up to itM above the peak operating linit or 13.4 kW/f t (i.e. ,15.6 kW/f t), ' plus a.. wance for dena t.fication power opiking. Streuceo due to external coolant pressure, internal gas presaure, thernal gradienta, spacer con tact , flow-induced vibration and manufacturing tolerances were cons ider ed. The tnxinum internal pressure is applied coincident with the niniman applicable coolant presaure. Additional information regardind thin type of analysis is provided in Referencea 4 (Section.11), 34, and 36. The calculatei rnximun fission gaa release fraction ili the highest design power density UOg rod is less than 25T.. This calculation is conservative becuuei it ausunea the rest linitin; peaking factors appliet to thin rod. The percentage of total fuel rod radioactivity released to the rod plenan la much less thaa 25% beca tse of radioactivity deeny durim; di! fasion from the UO 2. h.2.3.2 Results from Fuel Desida 9 valuations 4.2 3 2.1 Flow-Induced Fv l Rod Vibrationa Exper# men tal data on nultiple rod vibratious under two-phane flow condit iona have been used to develop a correlation between muinum rod displacement and fuel rod natural frequencyl5 These data indicate a decreasing nuinum displacemnt anplitude with increasind calculated fuol rod na tural frequency. The calculated rnximun vibrational amplit tde for the BWR fuel la 0.0007 in ch. The utreas leiela resultin c; from vibration are ney,ligibly lov and well below the endurance limit of all affected componenta. The de flection resu lti nd fr o vibration is combined with deflections from other loads and la used to demonstrat e adherence to the deflection criteria of Subsection 4.9.1.2.19 Flow-induc"4 fuel rc d vibration in not considered to be a viable li fe-limitind or fa . lure rechanism ia 2 BWR fuel desi nad baoed on ex tens ive fuel operatind experience. Fuel inapections, both visual inspections during normal refueling outagen, and rore detailel nondestructive examinationu as a part of GC surveillance programi, have not indicated any anomalous perforrance assactated with fuel rod vibration. Table 4.2-13 precents the lover ex treme of GE operatind experience compared to recen t fuel designs in terms of the calculated fuel rod natural freq ue n cy. The calculated rod natural frequency for the BWR/4, 5 and 6 fuel designa is clearly well within the experience bane of previouuly operate i GE BWH fue 1. 4.2.3.2.2 Potential Damagind Temperature Effects Durind Tranalenta There are no predicted algnificant temperature ef fecta during a power trani.ient resultind from a single operator error or uindle equipment malf un ction. For purposes of rnintaining ade< plate thernal enrgin durind normal steady-state operation, the !!inimum Critical Power Ratio (ItCPR) nust not be less than the required !!CPR olerating linit, and the flaxinum Linear Heat Generation Rate (f:LHGR) in naintained below the dealgn LHGR for the plan t. The core and fuel design banes for a teady-atate operation (i.e. , itCPR and LHGH limits) have been defined to provide mard in between the 4.2-22 Am. No. 56,(3/81)

ACNGS-PSAR

,m         steady-state operating condition and any fuel damage condition to accommodate uncertainties and to assure that no fuel damage results even (v}         during the vorst anticipated transient condition at any time in life.

Specifically, the !!CPR operating limit is specified such that at least 99 97. of the fuel rods in the core are expected not to experience boiling transition duriv the nost severe abnormal operational transient. The , calculated fuel rod cladding strain for this class of transients is I significantly below the calculate 1 damage limit. 4.2 3 2 3 Fretting Wear and Corrosion Tests of typical designs, representative of the BWR/6 fuel design, have [ been conducted both out-of-reactor as well as in-reactor prior to  ; application in a complete reactor core basis. All tests and post-irradiation exaninations have indicate 1 that fretting corrosion does , not occur. Post-irradiation examination of nany fuel rods has indicated , only minor fratting wear. Excessive wear at spacer contact points has never been obserted with the current spacer configuration. Additional informtion regarding these tests and inspection of operating furl is , presented in Section 10 of Reference 4 and Sections 4.T and h.8 of  ; Reference 14. 4.2.3.2.4 Fuel Rod Cycling and Fatigue Analysis Daring fuel life, less than 57, of the a'. lovable fatigue life is consumed.  ; Additional information regarding this type of analysis ia provided in , 3 Section 12 of Reference 4. \v ) h.2.3 2 5 Fuel Rod Boving i Fuel inspections, both visual inspections during normal refueling outages and r: ore detailed nondestructive examinations as a part of General Electric's active fuel surveillance program, have provided no indication of rod bowing as a viable failure or life-liniting mchanism. This successful operatir , experience has been supported by fuel mechanical analyses which predict an insignificant amount of fuel rod boving (< approximtely 20 i mils). These analyses consider the influence of initial bow, tubing eccentricity, fast neutron flux and thernal gradients on the potential for in-reactor creep boving. In addition, full scale thermal-hydraulic tests have been conducted '>y General Electric to assess the effects of gross fuel rod bowing. Based on results of these tests, it has been concluded that, even for severe rod bowing in the most limiting rods in the assenbly, there is a negligible effect on critical power performance. 4.2.3.2.6 Fuel Ass embly Dimnsional Stability  ! Piechanical analyses have been performed to assess the effects of the differential thernal expansion between the tieplates and spacer grids. The differential thermal expansion introducas a bending stress of less than h00 ' psi at the end of the fuel tube. Additional in formtion regarding the , model employed in this calculation is presented in Section 4.4 of Reference 3 N

%s 4.2-23                        Am. No. 56,(3/81)

ACNGS-PSAR 4.2 3.2 7 Temperature Transients with a Waterlogged Fuel Elemnt As indicated in Subsection 4.2.1.2.1.11, the potential for waterl  % ng i is considered in the fuel design. For waterlogging to occur, the fuel cbsdding must have a small pinhole. Pinhcles are clininnted durind

• production by a 1007. leak check of fuel assenblies. The Inak Detector System enployed is dencribed in Subsection L.2.1.2.1.11. Since waterlcgging is not expected and since it has not been observed in commercial power BWR fuel, no specific analj sis of the consequences is pe r fo rned.

In the unlikely event that a waterlogged fuel element does exist in a BWR core, it shaald not hwe a significant potential for claddind burst (due to internal pressure) daring a transient power increase unless the transient started from a cold or very low power condi. tion. Nornal reactor heatup rates are sufficiently slow ( <100*F/hr increase in coolant temperature) that water-vap;r formed ins ide a waterlogged fuel rod von 11 be expected to evacuate the rod through the sane passn6e it entered, allowing internal and external pressures to equilibrate as the coolant temperature and pressure rise to the ratel conditions. Once the internal and external pressu es ce at equilibriun, at rated coobtnt pressure and temperature, transient power increases should, in general, have the ef fect of only slightly reducing tLe laternal fuel rod plenum volum due to dif ferential therral expansion between fuel and cladding, thus ef fecting a snali chmt-term increase in in ternal fuel rol pressure. The potentia] abort-term increase in pressure due to this ef fect would, in general, be snall (e.g. , a power increase fron the cold condition to peak rated power would increase internal pressure less than 1% in the peak power fuel rod). For the range of anticipated transients, the cladding primry nenbrane stress resulting fron the temporary increase in internal pressure above the coolant pressure would not be expected to exceed the claddind stress design limits of Subsection h.2.1.1.1.4. 4.2 3 2 7 1 Energy Release for Rupture of Waterlogged Fuel Elenents Experimnts have been performed to show that waterlogged fuel elemnts can fail at a lower damade threshold than nonvaterlogged fuel durind rapid reactivity excursions from the cold condition 16,lT (i. e. , approxinttely 60 cal /gu as compared to >300 cal /gn). However, it has been shown28 that the resultan t mch ..i' cal energy release for waterlogged rods, even for sidnificant energy depositions (approximately 400 cal /gn), is of little consequence and is well below the erergy released for nonwaterlogged rods subjected to comparable energy depositions. 4.2 3.2.8 Fuel Densification Analyses The amoun t of densification employed in the following rrdels was determined through the use of rodels defined in References 11,12, and 13. 1 O 4.2-24 Am. No. 56,(3/81) l

ACNGS-PSAR 4.2 3 2.8.1 Power Spiking Analysis j \ ( ) The equation employed to calculatt maximun esp size is as described in

'% /   Reference 12:

AL AP where 7 = 7 + 0.0025 AL = maximum axial gap length; , L = fuel column length; AP = the average change in density as measured by thermal sinulation for 24 hours at 1700 C;  ; 2 = anisotropic factor appliet to denuification; and 0.0025 = allowance for irradiation induced cladding growth and  ;

                                                                                                   ~

axial str.in caused by fuel cladding mechanical inter-action. The resulting power spiking penalty at the top of the core is 2.2%. The power spiking penalty as a function of elevation from the bottom of the core can be conservatively expressed by:

                       'a p'           IaP'  X P                  L N                   -
                             "X           "L

(/ i where M

• power spiking penalty at elevation X from botton P of core;

              . - y

_ power spiking penalty at top of core;

              . P.

X = elevation from bottom of core; and L = fuel colurn length. The power increase described by the above equation as a function of axial position added to the license limit LHGR (13.? W/ft) has been considered in design and safety analysis, wherever applicatale. This ensures, with better than 95% confidence, that no raore than 7ne rod will exceed the power evaluated due to random occurrence of power spikes resulting from axial fuel colu:an gaps. The results of the power spiking analysis for normal operation have been utilized in the analysis of transients and accidents wherever applicable. The control rod drop accident is unique in the respect that it begins at w-4.2-25 ' Am, No. 56,(3/81

ACNGS-PSAR the cold condition and is not affected by normal operating power level. Further, the existence of fuel column gaps can result in power spiking in the cold condition during a control rod drop which should thus be considered in the evaluation of this accident. For this purpose, a separate power spiking analysis has been performed using the sane assumptions as indicated above, but ecploying a po.ier spike versus gap size calculatel to occur in the cold condition with zero voids. This analysis was performed with the maxinun gap size predicted at the top of the core in order to naximize the power spiking ef fect. This analysis yielded a 99% probability that any given fuel rod would have a power spike < 5%. h.2 3 2.8.2 Clad ling Creep Collapse A cladding collapse analysis has been performed employing the standard General Electric finite alecent mdell3 Figure b.2-6 presents the cladding midwall temperatu e versus time employed in the analysis. No credit is taken or intmaal gas pressure due to released fis . ion gas or volatiles. The aternal preasure due to helium backfill during fabrication is considere,1. Based on the analysis results, cladding collapse is not calculated to occur. h.2 3 2.8.3 Increased Linear Heat Generation Rate A fuel pellet expands 1.2% in going from the cold to hot condition at 13.h kW/ f t. While this increase in lengtl from the cold to hot condition is not taken credit for in either design calculations or in the process c c core performance analysis during reactor operation, the expansion more than offsets the decrease in pellet length due to densification. The following expression is employed to calculate the decrease in fuel column length due to densification in calculatior of an increase in linear heat generation rate: A L _ A_ P_ T- 2 where AP= the average change in density as neasur ed by t!.ernal simulation for 2h hours at 17000 C, and 2= anisotropic factor applied to densification. Using the above equation, the pellet decrease in length due to densification is less than the increase in length due to thermel expansion of the pellet in going from cold to hot condition. Therefore, r.o power increase is calculated dua to densification. h.2 3 2.8.h Stored Energy Deternination The ef fects on stored energy due to densification are accounted for in the LOCA evaluation. h.2.3 2 9 Fuel Cladding Tenperatures 4.2-26 Am. No. 56,(3/81)

ACNGS-PSAR "ael claddind temperatires as a function of heat flux are shown in Figure [~) 4.2-7 for bedinning-of-life (BOL) conditions and in Figure 4.2-8 for , late-in-li fe condi tions. The temperatures employed in mechanical design ("/ evaluations are calcuated using a conservative design allowance for the increase in resistance to surface heat transfer due to the accumulation of system corrosion products on the surface of the rod (crud) and cladding corrosion (zirconium oxide formation). 4.2 3 2.10 Incipient Fuel Center fielting l Incipient center meltind is expected to occur in fresh UO2 fuel rods at a LHGR of approximtely 20 5 kW/ft. This condition corresponds to the i ntegral: r T -

                          /nelt      kdT  =   93 W/cm 32 P The value of the above integral decreases slightly with burnup as a asult of the decrease in fuel melting ter erature with increasind exposure.

l 4.2 3 2.11 Energy Release During Fuel Element Burnout Boiling transition does not necessarily correspond to a fuel danade threshold. In-reactor experiments to assess the effect of operation of Zircaloy-clad UO p fuel rods after the onset of transition boiling have been conducted by a ndmber of different experimenters23-27 Post-irradiation i examinatiom conducted on the fuel tested verified thct no cladding failure

• Q #

and no appreciable cladding degradation occurred for fuel that experienced peak claddind temperatures less than approxirately 2000 F. The metal-water chemical reaction between zirconium and water is given by: Zr + 2H 2O =

                                        >Zr02 + 2H2 - H vhere -aH = 140 cal /g-mole. The reaction rate is conservatively given by the familiar Be'ter-Just rate equation:

2 6 ( 45y 00)  : W = 31 3 x 10 rexp g t where W= milligrams of zirconiun reacted per cm of surface aren; t = time (sec);  ; R = the gas constant (cal / mole OK); and T = is the temperature of zirconium ( K). i Q ,/ L i 4.2-27 Am, No. 56,(3/81)

ACNGS-PSAR Reference (18) shows this rate equation to be conservatively high by a factor 01 2. The above equ.stion can be dif ferentiated to give the rate at which the thickness of the claddind is oxidized. This yields:

                            ^1          2 th   =

g exp - Where th = rate at which the cladding thickness is oxidizing; AX = oxidized cladding thickness; A,A y 2

                  =   aPPr Priate constants; and T    =   reaction temperature.

The reac tion rate is inversely proportional to the oxide buildup; therefore, at a given cladding temperature the reaction rate is self-limiting as the oxide builds up. The total energy release from this chemical reaction over a time period is given by: T Q T JN g (-aH) C L P AX dt where O.

     ?l rods
                =   nunber of rods experiencing boiling transition (at temperature T);
         -aH    =   heat of reaction; C   =   cladding circunferences; L   =   axial length of rod experiencing boiling transitions and P =     density of zirec .iun.

This equation can be integrated and compared to the normal bundle energy release if the following conservative assumptions are nade: (1) At an axial plane, all the rods experieace boiling transition and are at the sane temperature. Thia is highly conservative since, if boiling transition occurs, it .111 normal.ly occur on the high power rod (s). (2) Boiling transition is assumed to occur uniformly around the circunference of a rod. This generally occurs only at one spot. O 4.2-28 Am. No. 56,(3/81)

ACNGS-PSAR (3) The rods are assumed to reach some temperature T instantaneously  ! and stay at this temperature for an indefinite amount of time. This integration has been performed per axial foot of bundle and the total energy release as a functioa of time has been compared to the total energy release of a high power buncle (approximately 6 MW) over an equal amount of time. The results are shown in Figure '4.2-9 For e w.mple, if the temperature of all the rods along a 1-ft 3ection of s ne bundle were instantly increased to 1500 F. the total amount of energy that has been released at 0.1 sec is 0.47,of the total energy that has been released by , the bundle (6 MW x 0.1 sec). Note that the fractional energy release decreases rapidly with tiue even though a cona:snt temperature is maintained. Tais is because the reaction is self-limiting as previously discussed. The amount of energy released is dependent on the temperature transient and  ! the surface area that has experienced heatup. This 'f course, is dependent on the initiating transient. For example. .f boiling transition were to occur during steady-state operating conditions, the cladding surface temperature would range from 1000 tc 1500 F, depending on the heat  ! fluxes and heat transfer coefficient. Even assuming all rods experience boiling transition instantaneously, the magnitude of the energy release is seen to be insignificant. Significant boiling transition is not possible at normal operating conditions or under conditions of abnormal operational transients because of the thermal margins at which the fuel is operated. It can, therefore, be concluded that the ener ;y release and potential for a chemical reaction is not an important consideration during normal operation O or abnormal transients. 4.2 3 2.12 Fuel Rod Behavior Effects from Coolant Flow Blockage The behavior of fuel rods in the event of coolant flow blockage is covered in Reference 19 l 4.2 3 2.13 Channel Evaluation l Channel analytical models and evaluation results are contained in Reference i 20. i 4.2 3 2.14 Fuel Shipping and Handling Analyses of the ma,jor handling loads have been performed and the resulting fuel assembly component stresses are within design limits. Additional information on fuel handling and shipping loads is presented in Reference

31. .

t 4.2 3 2.15 Fuel Assembly - SSE and LOCA Loadings , i An evaluation of combined Safe Shutdown Earthquake (SSE) and Loss-of-Coolant Accident (LOCA) loadings is contained in Reference 30. O 4.2-29 Am. No. 56,(3/81)

ACNGS-FSAR 4.2.3 3 Reactivity Control Assembly Evaluation (Control Rods) 4.2 3 3 1 thterials Adequacy Throughout Design Lifetime The adequacy of the control rod mterials throughout the design life vau evaluated in the design of the control rods. The primry mterials (B gC power and Type-30% austenitic stainless steel) have been found to perforn adequately for the lifetine of the control rod.

   ,2,3 3 2 Dinens tonal .nd Tolerance Analysis Lay'>ut studica are done to assure that, given the vorst conbination of ext ene detail part tolerance ranges at assenbly, no interference exists shich will restrict the passage of control rods.      In addition, pre-operational verification in made on each control rod systen to show that acceptable 1Nels of operational performnce are net.

4.2.3.3.3 Thermal Analysis of the Tendency to Warp The various psrts of the control rod assembly remin at approximtely the sane temperature during reactor operation, negating the problem of distorcion or varpage. Ibchanical design allows for what little differential thernal growth can exist. A ninimun gap is maintained between absorber rod tuben and the control rod frane assembly for the purpose. In addition, to further this end, dissimilar netals are avoided. 4.2 3 3.h Forces for Expulsion An analysis has been performed which evaluates the mxinun pressure forces which could tend to eject a control rod from the core. If the collet remains open, which is unlikely, calculations indicate that the steady-state control rod withdrawal velocity would be 2 ft/see for a pressure-under line break (the liniting case for rod withdrawal). 4.2.3.3 5 Effect of Fuel Rod Failure on Control Rod Channel Clearances The control rod drive mechanical design ensures a sufficiently rapid insertion of control rods to preclude the occurrence of fuel rod failures which could hinder reactor shutdown by causing significant distortions in channel clearances. 4.2.3 3.6 Effect of Blowdown Loads on Control Rod Channel Clearances The fuel channel load resulting from an internally applied pressure is evaluated utilizing a 1 xed bean analytical nndel under a uniforn load. Tests to verify the applicability of the analytical rodel indicate that the model is conservative. If the gap between channels is less than the thickness of the blade or the diameter ( e the roller, the roller and/or blade vill deflect the channel valls as it makes its way into the core. The friction force is a snall percentage of the total force available to the control rod drives for overconing such friction, and it is concluded that the min steanline break accident does not impede the insertability of the control rod. 4.2-30 Am. No. 56,(3/81) t

ACHGS-PSAR h.2.3 3 7 Mecitanical Damage Analysis has been performed for all areas of the control system showing

• that system mechanical damage does not affect the capability to continuously provide reactivity control.

The following dEscussion summarizes the analysis performed on the control - rod guide tube. l The guide tube can be subJteted to any or all of the following loads: , i (1) inward loss due to pressure differential; (2) lateral loads due to flow across the guide tube; (3) dead weight;  ; (h) seismic (vertical and horizontal); and I' ($) vibrat on. In all cases, analyses were performed considering both a recirculation line bceak and a steamline break, events which resu't in the lardest hydraulic loadings on a control rod duide tube. Two primary modes ot failure were considered in the guide tube analysis: 3 exceedind allowable stress and excessive clastic deformtion. It was found  ! O that the allowable stress limit will not be exceeded and that the elastic deformations of the guide tube never are great enough to cause the free movement of the control rod to be jeopardized.

                                                                                 ;

! 4.2.3 3.T.1 First Mode of Failure The first mode of failure is evaluated by the addition of all the stresses resulting from the mximum loads for the faulted condition. This results  ; in the :aaximum theoretical stress value for that condition. Making a linear supposition of all calculated stresses and comparing this value to i the allowable limit defined by the ASME Boiler and Pressure Vessel Code yields a factor of safety of approximately 3 For faulted conditions, the factor of safety is approximately 4.2. 4.2.3.3 7 2 Second Mode of Failure Evaluation of the second mode of failure is based on clearance reduction f between the guide tube and the control rod. The minimum allowable  : clearance is about 0.1 inch. This assumes maximum ovality and minimum l' diameter of the guide tube and the maximum control rod dimeasion. The analysis showed that, if the approximate 6000 psi for the faulted condition were entirely the result of differential pressure, the clearance between the control rod and the guide tube vould reduce by a value of approximately 0.01 inch. This gives a design margin of 10 between the theoretically  ; calculated maximum displacement and the minimum allowable c'earance. O . 4.2-31 Am. No. 56,(3/81)

ACNGS-PSAR l

                                                                                         )

4.2 3 3.8 Analysis of Guide Tube Design Two types of instability were considered in the analysis of guide tube d es ign . The first vau the classic instability associated with vertically loaded columns. The second was the diametral collapse when a circular tube experiences external to internal di fferential pressure. The liniting axially applied loa.1 is approximtely 77,500 lb resulting in a naterial compressive streus of 17,450 psi ( code allowable stress). Comparing the actual load to the yield strens level gives a design mrgin greater than ?Q to 1. From these values, it can be concluded that the guide tube is not an unstable column. When a circular tube experiences external to internal dif ferential pressure, two rede cf fsilure are possible depending on whether the tube is "long" or "short". In the analysis here, the guide tube is taken to be an infinitely lord tube with the mxinun allowable ovality and mininum wall thickness. The conditions will result in the lowest critical pressure calculation for the guide tube (i.e. , if the tube was "short", the critical pressure calculatica would give a higher nunber). The critical pressure is approximtely 1h0 psi. !Iovever, if the mxinun allowable stress is reachte at a pressure lower than the critical pressure, then that pressure is liniti ng. This is the case for a BWR guide tube. The allowable stress of 17,h50 psi vill be reached at approximately 93 psi. Comparing the naximun possible pressure diffferential for a steanline break to the limiting pressure of 93 psi gives a design rargin dreater than 3 to 1. There fo re , the guide tube is not unstable with respect to differential pressure. h.2 3.3 9 Evaluation of control Rod Velocity Limiter The control rod velocity ' niter limits the free-fall velocity of the control rod to a value that cannot result in nuclear systen process barrier danage. 4.2.h Testing and Inspection 4.2.k.1 Fuel, Hardware and Assenbly Rigid quality control requirenents are enforced at every stage of fuel mnufacturing to ensure that the design specifications are net. Written manufacturing procedures and quality control plans define the steps in the mnufacturing process. Fuel cladding is subjected to 1007, dimensional inspection and ultrasonic inspection to reveal defects in the cladding wall. Destructive tests are perforned on representative samples from each lot of tubing, including chemical analysis, tensile and burst tests. Integrity of endplug velds is assured by standarization of weld processes based on radiographic and metallographic inspection of welds. Fuel rod inspection includes metallographic and radiographic examination of fuel rods on a sample basis. Completed fuel bundles are helium leak tested to detect the escape of helium through the tubes and endplugs or welded reg ion s . Sanple tests are performed for qualification. Production samples are tested as a check on the process and process controls. UO2 Powder characteristics and p211et densities, composition, and surface finish are O 4.2-32 Am. No. 56,(3/81)

k

ACNGS-PSAR i

controlled by regular sampling inspections. UC2 weights are recorded at

,  O               every stage in nanufacturing.                                                                                                                      ;

I Each separate pellet group is characterized by 4 sin 6 l e stamp. Fuel rods  ; I are individually serialized prior to fuel loading to: (1) identify which. [ l pellet group (s) is to be loaded in each fuel rod: (2) identify which (

!                  position in the fuel assembly each fuel rod is to be loaded; and (3)                                                                               ;

facilitate total fuel material accountability for a given project. Each [ finished fuel rod is gamma scanned to detect an enrichment or rod pellet t

,                  loading deciations which exceed design specification.                                                                                              l

] l The fuel rod upper endplugs are designed to control placement of fuel rods j vithin a given assembly. Primary control is provided by alphanumeric  ! symbols on the end of the upper endplug which corresponds to a specific  ; enrichment and gadolinia content of the fuel rod. Secondary control is l provided by sizing of the upper endplugs such that a rod of high enrichmnt

;
;                  cannot be positioned in a significantly lower enrichnent location within                                                                           [

the fuel bitndle. Additionally, the gadolinia-bearing fuel rods have  ! l extended endplugs with characteristic markings on the shank to permit [ i visual identification of gadolinia rod location. Correct placemnt is j l verified by recording the fuel rod serial number on the lower endplug, the [ alphanumeric symbols on the upper endplug, visual inspection and placemnt j of the upfer tieplate, which has been machined to accept a specific pattern t . of endplus diamters. ,

                                                                                                                                                                      ;

Fuel assembly inspections consist of complete dimensional checks of [ channels and fuel bundles prior to shipmnt. Fuel bundles are given  ! another dimensional inspection of- significant dimensions at the reactor [ !,. \ site prior to use. The samplind rate, mthod and tools of the i post-shipment fuel inspection are outlined in Tables 4.2-6 and 4.2-7 I

                                                                                                                                                                      ?

4.2.h.2 Testing and Inspection (Enrichment and Burnable Poison f

!                             Concentrations)                                                                                                                         ;

The snutdown reactivity requirement is verified during initial fuel loading i and at any time that core loading is change . Nuclear limitations for  ; l control rod drives are verified by peric " cally testing the individual [ ! system. Test capabilities are described in the appropriate subsections.  ; l The following serves to identify the various tests and inspections employed  ! by General Electric in verifying the nuclear characteristics of the fuel l and reactivity control systems. f

                                                                                                                                                                      ?

h.2.k.2.1 Enrichment Control Program j i  !

                                                                                                                                                                      )

The incoming UF6 and the resultant UO2 p wder are verified by emission  : l spectroscopy for impurities. l [

The sintered pellet is also sampled for impurities by emission t I

spectroscopy. Chemical verification of impurities is also performed [

including vet chemistry for oxygen-uranium ratio determination.

{ i l  !

l

,  ; i  ! ! 4.2-33 Am. No. 56,(3/81) l

ACWGS-PSAR The enrichnent-blended material and the green pellet enrichnent are verified b/ the gamma scan enrichnent analyzer. Each rod is gamm scanned to screen out any possible, but unlikely, enrichment deviations. All ausenblies and rods of a given project are inspected to asure overall accoun tability of fuel quantity and placenent for the proj ect. 4.2.h.2.2 Gadolinia Inspec' ions The sane ridid quality control requirements observed for s tandard UO2 I"*1 are employed in nanufacturing gadolinia-urania fuel. Gadolinia bearing UO p fuel pellets of a given enrichnent and a gadolinia concentration are mintained in separate groups throudhout the manufacturind process. The percent enrichnent and ga lolinia concentration characterizing a pellet group is identifiel by a stanp on the pellet. Fuel rods are individually numbered prior to loading of fuel pellets into the fuel rods to: (1) identify which p< ilet group is to be loaded in each fuel rod; (2) identify which position in the fuel assenbly each fuel rod is to be loaded; and (3) facilitate total mterial accountability for a given proj e ct . Correct orientation of gadolinia-ber. ring rods within the fuel assembly is further assured by the longer upper endplug shanks for these rods. The following quality control inspections are nade: (1) gadolinia concentration in the gadclinia-urania powder blend is verified; (2) sintered pellet UO,,-Gd 0 solid solution homogeneity across a fuel pellet is verifiek $y examination of metallographic specimens; (3) gadolinia-urania pellet identification is verified; and (4) gadolinia-urania fuel rod identification is checked. 4.2.h.2 3 Reactor Control Rods Inspections and tests are conducted at various points during the mnufacture of control rod assenblies to assure that design requirements are being met. All boron carbide lots are analyzed and certified by the supplier. Anung the items tested are: (1) chemical composition; (2) boron wei 6ht percent; (3 3 boro *1 isotopic content; and (4) particle size distribution. O 4.2-34 Am. No. 56,(3/81)

      .   . ~ -       .--            .       . . -        .   .

t ALM-PSAR i Following receipt of the boron carbide and review of material certificates, i f Q / additional samples fron each lot are tested including those previously listed. Control is maintained on the BgC powder through the remaining j steps prior to loading into the abearber rod tubes. , t [ Certified test results are obtained on other control rod components. The [

absorber rod tubing is subjected to extensive testing by the tubing ,

supplier, including 100#. ultrasonic examination. !!etallographic f examinations are conducted on several tubes randomly selected fron each lot  ; to verify clean.iness and absence of conditions resulting frors improper j fabrication, cleaning, or heat treatrent. Other checks are nade on the i subassemblies and final control rod assembly, including veld joints j - inspected and B 4C loading.  ; l h.2.h.3 Surveillance Inspection and Testing of Irradiated Fuel Rods  ! f General Electric has an active program of surveillance of both production and developmental BWR fuel. The schedule of inspection is, of course, l contingent on the availability of the fuel as influenced by plant i j operation. The lead fuel rods (with respect to exposure, LHGR, and the [ combination of both) are selectively inspected. Inspection techniques used j include i (1). leak deter , ion tests, such as " sipping"; f (2) visusl inspection with various aids such as binoculars,  ! borescope, periscope, and/or underwater TV with a photographic

      \                     record of observations as appropriate;                                                                                                    ;

I' (3) nondestructive testing of selected fuel rods by ultrasonic test techniques; and - 1 f (h) dimensional measurements of selected fuel rods.  ; j Unexpected conditions or abnormalities which may arise, such as  ; j distortions, cladding perforation, or surface disturbances are analyzed. l { Resolution of specific technical questions indicated by site examinations  :

may require examination of selected fuel rods in Radioactive !!aterial {

! Laborntory facilities. 4  ; 1 The fuel channels are also under surveillance in continuing programs. f i These surveillance programs are designed not only for the evaluation of  : present day products, but are also providing data in the areas of alternate l i ruterials and design r.odeling.  ; i j The results of the program are used to evaluate the BWR fuel design methods i j and criteria used by General Electric and are generally reviewed with the , l Division of Reactor Licensing and documented in generic fuel experience r licensing topical reports.  ! f I j In addition to the fuel surveillance program, lead test assemblies ,

 !               essentially identical to the BWR/h and BWR/5 assemblies have been                                                                                  -l '

.p characterized in detail prior to irradiation and have been placed in

"/) J service in two operating reactors (Peach Bottom 2 and Vermont Yankee). 1  : i 4.2-35 Am. No. 56,(3/81)  ! 4  : ,i

ACNGS-PSAR These lead test bundles will provide approxinately two years of operating experience with this fuel design prior to full core loadind in a BWR/h or 5 reactor. A surveillance program similar to that described in the precedind paragraphs has been incorpeted in the lead test assembly program to provide extensive perfornan ;e ronitoring. The fuel rod design for the BWR/6 fuel assenblies is the sane as that employed in the lead test assemblies; therefore, the experience obtained with these assemblim is applicable to BWR/6 fuel. A prepressurized test assenbly enployind tha same fuel rod design was placed in operation in April 1977 For a description of thie fuel assembly, see Reference 36. L.2 5 Operating and Developmental Experience 4.2 5 1 Fuel Operating Experience A larde volume af experience with Zircaloy-clad UO2 pellet fuel has been obtained since 1960. The largest portion of this experience has been obtained in operating connercial power B'4Rs at LHGRs representative of, or higher than, carrent 8x8 fuel perforennce requiremnts. The large volune or production experience, startind with the first load of fuel in Dresder-1 Nuclear Power Statior in 1960, has provided feedback on the adequacy of the desid n for, and the effects cf, operation in a connercial power reactor environmnt. Table 4.2-8 presents a sunnary of B'4R experience with General Electric production Zircaloy-clad UO 2 pellet fuel. Overall, 73 production fuel types have been desidned, manufactured and operated in 32 BWRs. When all production fuel types are considered, a total of rmre than 1,250,000 Zircaloy-2 clad UOp fuel rods have been operated in General Electric designed BWRs. The most recent fuel experience documented in Reference 21 indicates that, although the available BWR fuel experience base has increased by over 50% in the last two years, no new fuel failure mechanisms have been observed. Peak LHGR, fron approximately 10 kW/ ft to approximately 18 5 kW/ft, have been experienced with the production fuel. Individual fuel assemblies have achieved averade exposures greater than 25,000 frid/t and have accumulated tore than nine years in-core residence. In comparison, the current 8x8 fuel designs have the followind proposed operating cheracteristics : (1) 13.4 kW/ft maxinum LHGR (operating limit); (2) 40,000 ff4d/ t maximun local exposure; (3) 30,000 ff4d/t maxinum assembly exposure; and (4) 4 to 6 years in-core residence tine. Fuel rod diamte- in the range of 0.425 in. to 0 593 in. outside diameter with claddin. thickness from 30 to ho mils and pellet-to-cladding gaps from 3 to 12 mils have been used in production fuel. Active fuel column lengths have varied form 59.8 to 146.0 in. vith rission gas plenum volume per unit of fuel volume from 0.013 to 0.11. O 4.2 36 Am. No. 56,(3/81)

ACNGS-PSAR Production fel rods employing gadolinia-urania fuel pellets have been in /3 use since 1965 During this time, a substantial number of gadolinia-urania ('"' ) rods have been successfuly irradiated to appreciable exposures. Table 4.2-9 sunnarizes this experience. Of these irradiated gadolinia-urania rods only a anall number have experienced failure, none of which could be attributed to the fact that they contained gadolinia bearing fuel pellets. h.2 5 2 Fuel Development Experience The production of Zircaloy-clad UO 2 pellet fuel experience described in Subsection h.2 5 1 is supplemented by a large amount of in-pile and out-of-pile developmntr.1 work. The developmntal work to date has been enployed to test a wide rande of design characteristics, to investigate various mechanisms affecting the performnce of th fuel rod and to extend irradiation experience to higher local combination of fuel rod power and exposure than covered by production fuel. The following presents a discussion of the pertinent developmental fuel experience which, in combination with the production fuel experience, provides the basis for the current BWR fuel design and operating limits. Tables 4.2-10 through 4.2-12 present a summry of design details and , performance conditions for Zircaloy-clad UO2 pellet fuel rods and capsules irradiated under General Electic or USAEC-General Electric development test programs. These data complement the BWR production fuel experience by providing additional data at higher local combinations of fuel rod power and exposure. Overall, more than 800 fuel pins with design characteristics similar to the current BWR fuel have been irradiated under General Electric \

    }
    /

or USAEC-General Electric programs. The irradiations have bee 4 perfo rmed vith BWR environment in both test reactors and in commercial power BWRs. Test reactors employed in General Electric developmental irradiations summarized in tables 4.2-10 through 4.2-12 are the Vallecitos Boiling Water Reactor (VBWR), the General Electric Test Reactor (GETR) and, more recently, the Halden Reactor. Developmental fuel irradiations hava also been performed in the Consumers Power Company Big Rock Point and Dresden Unit 1 commrcial power BWRs. The range of peak performance conditions covered by the various development irradiations goes beyond the design performance conditions for fuel in this class of reactor. The development performance conditins include: (1) 13 0 to 58.0 kW/ft maxinun LHGR, and (2) 1,500 to 100,000 mwd /Te maximum local exposure. The corresponding conditions for fuel to be operated in the BWR/6 class of reactor are: (1) 13.4 kW/ft maximum LER, and (2) approximtely 45,000 ffdd/Te maximum local exposure.

 -m     A capsule, as used herein, refers to a test fuel rod, or a group of rods combined, with all features similar to production fuel rods except for

[j\ \ havind reduced active fuel length (as lov as approximately 3 inches). 4.2-37 Am. No. 56,(3/81)

ACNGS-PSAR The rande of desidn characteristics and dinensions covered by the various developmntal irradiations also encompasses the characteristics and dimensions employed in the current BWR fuel design. The rande of design characteristics and dimensions covered by the various developmntal irradiations include the following: (1) fuel rod outside diameter (0.250 to 0 700 in. ); (2) cladding vall thickness (0.025 to 0.060 in. ); (3) pellet-claddind gap (0.0014 to 0.016 in. ); and (h) pellet lendth (0 3 to 0 95 in.). The correspondind fuel design characterirties for this c'nss of reactor are: (1) fuel rod outside diameter (0.483 in. ); (2) cladding wall thickness (0.032 in. ); (3) pellet-cladding dianetral gap (0.009 in. ); and (h) pellet kndth (0.41 in. ). It has been concluded that, for the complete rande of power levels and for peak fuel burnups, the calculated fuel perforrance has been adequately verified by experience. 4.2 5 3 F'tel Rod Perforation Experience The early General Electric BWP fuel experience has been extensively described in previous reports. In general, the Zircaloy-2 cladding performance in the very carly plants was good; however, some fuel failure nechanisms were exposed and corrected and are not significantly affecting current fuel performance. Details of this experience are provided in References 7, 21 and 22. Hydriding and pellet-cladding interaction are the failure mechanisms which have continued to affect fuel perforennce. Hydriding defects were ibntified by significant nunbers of rods perforating at relatively low power and exposure. The pellet-cladding interaction problen currently beind experienced in operatind fuel did not become appreciable until a statistically significant number of Zircaloy-clad fuel rods had experienced relatively higher power. The current fuel design incorporates improvements in desi;n and ranufacturing which provide confidence that a high degree of reliability can be expected. Operation with failed fuel rods has shown that the fission product release rate from defective fuel rods can be controlled by regulating power level. The rate of increase in released activity apparently associated with progressive deterioratior of failed rods has been deduced from chronological plots of the offgas activity measurements in operating plants. These data indicate that the activity release level can be lowered 4.2-38 Am. No. 56,(3/81)

ACHGS-PSAR by lovering the local power density in the vicinity of the fuel rod

p. failure. These reasured data also indicate that sudden or catastrophic

                                                                                               ;

failure of the fuel assembly does not occur with continued operation and that the presence of a failed rod in a fuel assembly does not result in propagation of failure to neighboring rods. Shutdown can be scheduled, as requir ed , for repairing or replacing fuel assemblies that have large defects. Evaluation of the fission product release rate for failed fuel rods shows a , wide variation in the activity release levels. Designers have attempted to relate the release rates to defect type, size and specific power level. , These data support the qualitative observations that fission product i release rates are functions of power density and that progressive deterioration is a function of time. h.2 5.4 Channel Operating Experience General Electric Company has more than 7000 Zircaloy channels in operating reactors, and surveillance of their perforcance is ongoing. The t preponderance of the experience has been with channels that are 5 278 in. inside width with 0.080 in. vall thickness. Channel sizes ranging from f 4.290 to 6 543 in. inside width and with 0.060 to 0.100 in. valls are included. The BWR/6 channel is 5 215 in. inside width, with 0.120 in.  ! I walls. The performance of the channels currently in operation has shown no tendency for gross in-cervice deformations, although long-term creep r I g deformacion has been identified as a potential lifeginiting phenomenon. y/ Separate reports on this subject have been provided , I h.2.6 References  ; i

1.  !!. A. ?!iner, " Cumulative Damage in Fatigue", Applied tiech. ,12, Tramsactions of the AS!!E, 67 (1945).

2. W. F. O'Donnell and B. F. Langer, " Fatigue Design Basis for Zircaloy Components", Nuclear Science and Engineering, Vol. 20,1964, pp 1-12.

3 BWR/6 Fuel Design, June 1976 (NED0-20948-P). i

4. " General Electric Boiling Water Reactor Generic Reload Application for i 8x8 Fuel", f4 arch 25,1976 (NEDO-20360-lP), Revision h.

5 "Effect of High Burnup on Zircaloy-Clad, Bulk UO Plate Fuel Elerent 2 Samples", September 1962 (WAPD-m-283).

6. " Irradiation Behaivor of Zircaloy Clad Fuel Rods Containing Dished End UO 2 Pellets", July 1967 (WAPD-m-629).

7 H. E. Williamson and D. C. Dittore, " Experience with BWR Fuel Through I September 1971", fiay 1972 (NEDO-10505).

8. D. C. Ditrore and R. B. Elkins, "Densification Considerations in BWR  !

   .         Fuel Design and Performance", December 1972 (NED?f-10735).

U 4.2-39 Am. No. 56,(3/81)

                                                      -~.

ACNGS-PSAR 9 J. A. Christenson, "f telting Point of Irradiated Uraniun Dioxide", Februsry 1965 (WACP-6065).

10. " Thermal Conductivity of Uranian Dioxide", Technical heport Series No.

59, IAC A, Vienna,1966.

11. Supplenent 1 to the Technical Report on Densification of General Electri: Reactor Fueln, December 1973

12. V. A. fbore, letter to I. S. t!itchell, "ifodified GE tiodel for Fuel Densification", Docket 50-321, !! arch 22,197h.

13. " Creep Collapse Analysis of BWR Fuel Uning Gare Collapse !!odel",

Audust 1974 (NEDE-20606) (Proprietary) . (NEDO-20605) (Nonproprietary).

14. "8x8 Fuel Bundle Developnent Support", February 1975 (NED0-20377).

15 " General Electric Boiling Water Reactor Generic Reload Application for 8x8 Fuel", !! arch 1976 (NEDO-20360), (Sapplement h).

16. L. A. Stephan , "The Responce of Waterlodged UO, '

Fuel Rods to Power Bursts", April 1969 (ID0-ITH-105). 17 L. A. Stephan , "The Ef fects of Clwiding fiaterial and Heat Treatanent on the Response of Waterlodded UO 2 Fuel Rods to Power Bursts", January 1979 (IN-ITR-111).

18. " Thermal Response and Cladding Performance of an Internally Pressurized, Zircaloy Clad, Simulated BWR Fuel Bundle Cooled by Spray Under Loss-of-Coolant Conditions", April 1971 (GEAP-13112).

19 " Consequences of a Postulated Flow Bloc 4ade Incident in a BWR", October 1977 (NEDO-10174) - Rev.1.

20. "BWR Fuel Channel flechsnical Design and Deflection", September 1976 (NEDE-21354-P) (GE Proprietary), (NEDO-21354) (Non-Proprietary).

21. R. B. Elkins , " Experience with BWR Fuel Through December 1976", July 1977 (NEDO-21660).

22. H. E. Williamson and D. C. Ditnore, " Current State of Kno .-ledge High Performnce BWR Zircaloy Clad UO Fuel", !!ay 1970 (NEDO-10173 ).

2 23 H. P. Olson, "In-Pile Burnout Protection Demonstrated at HBWR", Euro Nuclear, December 1964.

24. J. E. Boyden , S. Levy, !!. F. Lyons, T. Sorlie, " Experience with Operating BWR Fuel Rods Above the Critical Heat Flux", Nucleonics, April 1965, Volume 23, No. 4.

25 G. Kjaerheim, E. Rolstad, "BWR Burnout Experiments", Nuclear Engineering International, December 1968. O 4.2-40 Am. No. 56,(3/81)

ACNGS-PSAR

    -s   26. T. Sorlie, " Consequences of Operating Zircaloy-2 Clad Fuel Rods Above I    i       the Critical Heat Flux", October 1965 (APED-4986).

t 27 W. Redpath, "Winfrith SGHWR In-Reactor Dryout Tests", Journal of the British Nuclear Energy Society, 1974,13(1) p. 87-97 l i

28. L. B. Tnomposn, et. al. , " Light Water Reactor Fuel Behavior Pro 6 ram t Description; RIA Fuel Behaivor Experimnt Requiremnts", USAEC Report RS-S-76-170, September 1976.  ;

29 G. A. Potts, "Urania-Gadolinia Nuclear Fuel Physical and Irradiation Characteristics and thterial Properties", January 1977 (NEDE-20943) l Proprietary, (NEDO-20943) Nonproprietary.

30. "BWR/6 Fuel Assembly Evaluation Combined SSE and LOCA Loadings",

November 1976, (NEDE-21175-P).

31. Fuel Assembly Evaluation of Shippind and Handlind Loadings", March 1977, (NEDE-23542-P, NED0-235k2 ).

32. General 31ectric Thermal Analysis Bases (GETAB): Data, Correlation ,

and Design Application , General Electric Co. , November 1973, i (NED0-10958 ) . I i

33. BWR/4 and BWR/S Fuel Design Amendnent General Electric Co., January ,

1977, (NED0-209h4-1, Non-Broprietary).  ;

34. R. B. Elkins, " Fuel Rod Prepressurization Anendment 1", May 1973

,                (NED0-23786-1 ).

35 Letter, E. D. Fuller to 0. D. Parr, "NRC Request for Additional  ! Infbrnation on Fuel Rod Prepressurization", June 8,1978.

36. Letter, E. D. Fuller to 0. D. Parr, "NRC Requent for Additional  !

Information on Fuel Rod Prepressurization", August 14, 1978. I 37 R. B. Elkins, " Fuel Rod Prepressurization", March,1978 ( NEDE-23786-1-P). t

   %d 4.2-41                       Am. No. 56,(3/81)

l ' ! ACNCS-PSAR l Table 14.2-1 l FUEL CLADDING CONDITIONS OF DESIGN RESULTING FR0t! IN-REACTOR j PROCESS CONDITIONS C0f!BINED WITH EARTHQUAKE LCADING 1 i i l CONDITIONS OF DESIGN i 1 1 Reactor Initial ibrcent of Safe Shutdown Farthiluake Imposed l Conditions _ 0% 50% 100% j Startup testing Upset -- -- l i Normal Normal Upset Faulted ! Abnormal Upset -- -- 3 l i d I l i r i I i O i } i i i 1 I l l i l l l l p l l O 4.2-42 Am. No. 56,(3/81)

ACNGS-PSAR Table 4.2-2 S FUEL CLADDING STRESS INTENSITY LIMITS 1 l Yield Strength r  ! Sy Ultimate StrengthTe$aile g l

                                                                                                                                         . - -                               ;

1 Pri ary Membrana Stress 2/3 1/2 Primary Membrane Plus Bending Strev.i Intensity 1 1/2 to 3/4 Prienry Plus See andary . Stress Intensity 2 1.0 to 1 5 I i

9 i

1 1 , t 1 l i i i i j l 9  ! l 4.2-43 Am, No. 56,(3/81)  ! I

i i l l

;

l !! ACNCS-PSAR  ! Table 4.2-3  ! 1  ; f FUEL CLADDING ESTIMATED NUMBER OF CYCLES FOR EACH CYCLIC l CONDITION USED FOR FATIGUE ANALYSIS j

I

. Cyclic Condition Estimated Cycles i i f l Room Temperat tre to 100% Power approximately h/yr  ; Hot Standby t> 100% Power approxiuntely 12/yr j 50% Power to 100% Power approximately 60/yr i 75% Power to 100% Power approximately 250/yr { i 1005 Power to 116% Power atproximately 1/2 yr i l l 5 h

                                                                                                                                                                                     ;

l 9 l 1 l 4 o G r i >9 t i f l 4.2-44 Am. No. 56.(3/81) } l

ACNCS-PSAR r~'s Table 4.2-4 (d) FUEL DATA Core Number of Fuel Assemblies 748 Fuel Cell Spacing (Control Rod Pitch) (in.) 12.0 Total Number 4 Weled Rods

• 46376 Core Power Density (Rated Power) (kW/1.) 54.1 Total Core Heat Transfer Area (ft2) 73303 Fuel Assembly Data Nominal Active Fuel Length (in. )*** 150 Fuel Rod Pitch (in. ) 0.636 Fuel Rod Spacing (in.) 0.153 Fuel Bundle Heat Transfer Area (ft2) 98 Fuel Channel Wall Thickness (in.) 0.120 Channel Width (Inside) (in.) 5 215 Fuel Rod Data Outside Diameter (in.) 0.483 Cladding Inside Diameter (in.) 0.419

 /      ,        Cladding Thickness (in.)                                         0.032

( ) Fission Gas Plenum Length (in.) 9 48

   ' ~'           Pellet Imnersion Density (".TD)                                95 Pellet Outside Diameter (in. )                                  0.k10 Pellet Length (in. )                                            0.h10 Water Rod Data Outside Diameter (in. )                                          0 591 Inside Diameter (in.)                                           0 531
             *Does not include two water rods in each assembly
            ** Includes 6 in of Natural U at the top and bottom of the fuel column.
                                                         ~
 )        i i
   \,_/

1 l 4.2-45 Am. No. 56. (3/81)

i l l

                                                                                                                                                                                                                                                              )

i ACNGS-PSAR l 4 Table 4.2-5  ! ! l l MATERIAL PROPERTIES

• l

!                                                                                                                                                                                                                                                             I i                                                                                                                                                                                                                                                              l l                                                         Zircaloy-2 Cladding Ther:nal Conductivity for T = (600 to 800*F),' k = 9-10 (Btu /hr-ft *F) 1 Coafficient of Linear Thermal Expansion l

approximately 3 x 10 6 (.p-1)

                                                                                                                              ~

l Total Elongatio'n (Irradiated) > 1% _

) Pelleta UO2 l } ihermal

                                                                                " ""~ *

6) 6$+T = + 6.02366 x 10-lE (T + 460)3 (Btu /hr-ft *F) tivity g t

• Melting Temperature = 5080 - 63 5 x 104 E (*F)  !

i ) i (where E = Exposure mwd /t) g l

i i I

• Additional information on material properties is presented in Section 3 of l Reference 4 and Section h of Reference 3

! l I i  ! .;

                                                                                                                                                                                                                                                              \

i r i i i i I l 4 I l i l  ! .i 1 I

                                                                                                                                                                                                                                                              ;

P 1

I 1 f j 4.2-46 Am. No. 56,(3/81) i

'                                                                                                                                                                                                                                                             t t

t __ , _ ~ . _ _ _ _ _ _ . . _ . . - - _ . _ _ . . _ _ _ _ . _ . , _ - - . . . . , _ _ . . _ . , . . . . - . _ , _ _ _ , . _ _ , _ _ . . , _ . . . , _ , _ .__,,_,.....-_.._.__..t

1 l l ACNGS-PSAR Table h.2 6 l 1 POST-SHIPMENT FUF4 INSPEC" ION PLAN I i j Characteristic Method Frequency l 1 t Container Damage Visual 100%  ! l  ! ! and Leak ! Bundle Damage Visual 100% ) i Shipping Sepa.a- Visual 100% j tors Remove: eleanliness lia a l 100". l Rod Integrity Vicus1, Gauge 100f, i when required  ! Lock Tab Washers Visual 100% j Channel Integrity Visual 190% i

Channel Cleanliness Visual 100% i j Guard Integrity Visual 6.. 100% l

{ and Int.tallation 'larque Wrench i i Spacer Damage Visual 100% for first 5 bundles and I i every 20th thereafter, other-f wise the middle 3 spacers i i Rod to Rod Feeler Gauge 100% of first i bundles and i f every 20th the 9after, l 8 otherwise twa metions, all , spacers, alternate the { G Rod-to-Simulated Simulated sections 100% of first 5 bundles l 1 Channel Channel and and every 20th thereafter, Feeler Gauge otherwise 2 sections, h sides per section, alternate sec- , tions excluding end sections. l Expansion Spring Visual 100% for all bundles j j Length l Gauge 100% for first 5 bundles and { i every fourth thereafter, j otherwise visual inspection. l l

\

Finger Spring Visual 100% for all bundles. Seated in Pocket Gauge 100% for first 5 bundles and every fourth thereafter, otherwise visual inspection. Note: Deviations required 10C/, inspection of the next 5 bundles for that characteristic. Two deviations for 1 characteristic within 6 consecutive bundles require revision of the AQL (acceptable quality level) with the General Electric, Wilmington, North Carolina, U.S. A. facility. When a reduced ir.epection was performed, all inspection steps shall be [ designated S OK (stamped OK). 4.2-47 Am. No. 56,(3/81)

ACNGS-PSAR i l l Table 4.2-7 INSPECTION EQUIPMENT l (1) Sling, four-legged to lift containers and remove lids from chipping boxes. l (2) Bundle hold-down bars (drawing 107Ch713) for clamping on to (, containers before pickup to the vertical position (2 each) ' I l (3) Special spreader bar sling arrangement consisting of: 1 ' I < i f a. spreader bar (drawing 10704707); i

b. bridle sling, similar to Bethlehem No. 206-C (5/16 in. diameter by 2-foot length, two legs, thimble on load ends); and l c. two single-leg Bethlehem No. 110-C safety swivel hooks (4-foot length by 3/B-in. diaceter wire) and 4 (7/16-in.) screw pin anchor shackles.

(4) Special holding bar and cable sling (QCF-0015A) to support channel for channeling operation. (5) Stop plate for tilting container at bundle unloading station. i (6) Safety straps, with safety swivel shaps on each end, and takeup G j buckle to secure container in the vertical position. l (7) Feeler gauges, nylon: 0.105- and 0.100-in. j (8) Shim gauge: 0.015-in. l (9) Spring length gauge. f l I l (10) Simulated channel gauge. s (11) Male thread gauge, 5/16-18 UNC, go-no-go type. O 4.2-48 gm, 30, 56,(3/81)

I \ I O TWS k.*A O O I l l f TrftAW OF M?511?:17 17 PM177D7 *.IY4L%CLAS MS ML ' I i ( WMne't 11, 10'6) l Number nnposare 1 Sesign Aetive Seseent(S) l 3 Nel or Rods Maposure Average Time in Peak Nel 'lo43 Cla111y Pellet-to. l Assembly Core LOS '!!avier "'h t c kne s s Cla141rtg Cap Length Still t Class of Tw.1 10 of 'eak 'ellet la Core Df Ft /t ) (*f41/* ) fyears) kW / ft (in.) (mils) (it,91nni mils) (in.) Total ha:t or %ac t or M yntles 91,mq 8,3% %9 -14.k  % 567 il T.0 106.5 77.184(S) 0 Dres en i 1 536 27,000 5 w.19,500 5 %5 3 1%k 9.555 35 T.5 109 912 1,080 4 Itin m, iv nk n,5 _ ? , ,02 5 e.5 3 n.5 ,. m 5 35 n we.n ..,44 s 2,556' 4

                                                                                         ?8,000 5                       19, % 05                             6.53               15.5                   %%25                              35                10         109.?$     3,816 i                                                                     V          1%

4 9,090 +1 %q00 0.%9 ll 5 5 % 75 7,200(S) 0

             -     L'%WIk                                                       101                                                                             -                 -

M ,600 14,300 11.9 17.3 0.5 % U $ 195.T 16,848 0 1 lanciano 4 W1 66 19,600 8.? I k .f.  % 593 37 11 107 4,224 3,200 r4 ?9.500 S3 72  ? k .9% 15,500 6.1 14.6 0.593 P 11 ICT 4,608 3.200 k.0 1k.6 0.511 37 11 107 3,968 3,968 j SC 6? 17.300 17.600 k6 T,300 4,300 1.? 13.4 0.593 37 11 107 2.944 2,944 i r.D ' 76 ~ 3,400 - -  %%T 30 5 56.75 5,472(5) 0

             -     Jm                                                                        -

l D 35,k00 2 3,k00 50 15.0 0.4k9 34 8 70.0 3,630 0

1 Consovra ( m') B k7 11 70.0 3,321 0

! 9 41 16,000 9,k00 3.? 17.7 0.5625 m 33 23,700 17.3 % F.1 1T.7 0. % 25 k9 11 70.0 2.673 0 l 40 11.5 7%0 6.885 2,916 * '  !' R5 70,3 % 14,%q 4.1 17.7 0. % 25 ' ! hmbolit 11 169 23,000 14,600 9.0 17.1 %4% 33 D 79.0 6.084 0 O l I 111 176 'k,000 12,900 k.k 16.8  % 563 3? 11 79 0 6.336 4,212 h 1 eta A 371 M ,500 15, % 0 7.7 15.9  %%?5 35 17 139 13.356 504 C/3 3 40 71,500 12,500 k.0 15.9 0.563 39 11 130 648 288 8 l E9 Tarapar 1

• lot ?9,6M 17,000 T.0 15.9 0 M?5 35 1%5 Ikk 10,836 2,628 @

l 1 23,600 13,600 k.T 15.8 0.53 6 39 10.5 1h2.25 2,412 1,906

• j A 9eloa1 T4 6T
• 11,000 0.k 15.9 %563 32 1%5 142.25 2,664 2,160 :XS

' *9 TN 1% 400 g 17,500 7.9 15.8  % 56?5 15 10.5 1kk 11.088 5,184 e 1 Tarapur ? T D9  ? % 000 1k2.?5 T4 17,000 11,'00 k.0 15,8 n.%3 3' 1%5 1,332 540 A Relaa1 37 1k2.25

      @                                                               T9           14     13,900                           R,300g                             2.k                15.R                   %%3                               3?               1%5                       504            432 4 D ,9%g                         19,700                               7,6                17.'                    % 570                             35.5             11         144      27.440           9,018 9    fryst er Creek                                    JC         560 JC4        156      '0,k00                         12,600                               5,1                17,?                    %%)                               3?               12         144        7,644          7,448' 1elost 77,7%                          19,500                               7.3                17.5                    0.570                             35.5             11          1kb     26,068           8,862
               ?    1tne ' tile Point 1                               tt         $1?

17 Ikk 2,744 1elos t itA  % 25.100 16,200  %? 17.5 %563 3? 2.744 OSA 40 29,400 1h T00 4,5 17.5  %%3 3' I? Ikk 3,960 1,666

                                                                      !W         109      5 ,1%                        16,000                                3.5               17.5                    %%3                               37               I?          14%       5.292          5,292 TD           96     19,700                         P,3%                                 0.5                 13.4                   0.k33                             %                  9         Ikk       6,0 8          6,048

9,900 5,900 13.4 %k93  % 9 1k4 12,600 12,600 1,Y MO t.0 4

               ?                                                      Jt         3ik      ?6,600                         19,TM                                 T.3               17.5                    % 570                             % .5             l'          ikk      15,386             539 TsuruCa                                                                                                                                                                                                                                             Ikk 9elons                                         JAA          49      D ,600                        P1,500                               6.1                 17.5                   0.%3                              37               12                    2,352             882 J U)         %      2 3 .'M                        15,600                                k.1                17.5                   0.563                             3?               12          144       4,165           t,911 JA7          76     M ,$M                          n ,600                                3.5                17.5                   %%)                               37               I?          Ikk       3,724           3,381 J 40        53      13,M7                            T.700                              2.4                17.5                   %%3                               37               1?          Ikk       2,597          2,597 S?        9,6M                           5,4%                               1.3                17.5                   %%1                               37               12          Ikk       4,018          4.018 JAE f4y         36       k .1%                          ?,600                              0.3                13.4                   0.493                             %                  9         1kk       2,268          2,268 3,3%                                                  17.5                   %%3                               3'                17          144     35,476                0 3    1ra . Fn '                                        Tl        7'%         6.6%                                                              0.3 ni,a1                                         %            'O       ',l'O                          1,%q                               %5                 17.5                   %%3                               37               P           Ikk       I,044                0
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                                                                                          -.            - - _ _ . - . .        . _ .          . . - ~ . _ . . . . _ . ~ . - _ .                - _ - . _ . . . . . - . - _ -                . - ~ _ . - - . _ - - ~ - - - . .-                    -----

O O O

                                                                                                                                     *4'4LF k.'      % ("ont inuei)

Trfit1Y OF EY"e91F1"E Ils P'10W"171 IIMALOY-CLAS 1707 F"IEL ( DE"E't3E9 31, 1976) Numbe r

                                                                                                                        *'aposure       I                                Sasign                                                                   Active                        Se gment e( S )

r aposure Average Time in Peak Fual 901 I 71s111ng I Pellet-to- Fuel or Rods Assent:1y Core LU9 Gia.vter Thickness Cladfir.g Cap Length Still

                                           *16ss of                          Fael     *lo. of      Peak Pellet In Core
                                           % etnr             7eactor        gye,    Nniles         (*N s /t )            ("W11tl           (yea rs )                    kW / ft      fin.)             (mils!               ('hminal alls)         (in.)                  Total 13,000               7,600                     ?.6                    19.5     0.563                37                                         la                18,424                   0 k      Verent Yar.Kee        V7           376 1,9w              1,960 41       n,#0                 0,6%                       3.1                   19 . '.   . /,3               3?                       l'               164
                                                          'talaa i           "E7 28,602            20,538 3?9 /1,.#

7,0% '.0/0.3 13.4 0.493  % 1 1kk LF l,300

                                                                                                                           ?,M0                      0.3                    13.4     0.493                 p

• 150 126 126 i Lf L*4  ?  ?,900 37,436 3'/37 1? Ikk 37,436

                                                                             ?Y      168/596 7         11,570              6,700                     3.59                   19.5     0.563                                                                                                                  ,

4 3rowns Ferry 2 f 2.41 19.5 0.563 3'/ r' I? Ikk 37,436 37,436 t 3rowns Ferry 2  % 169/5 %? 5 .9% 3.400 0 Ik6 48,132 48,132 4 k Browns Ferry 3 BF 764 0.3 13.4 0.493 34  ! 3.3 19.5 0.563 3'/37 12 144 37,436 28,224 i !

• Peac% Botton ? M 169 /SM M ,100 13,700 150 252 L"" L 5,400 3,053 0.5 13.4 0.493 34 9 252 l 5,'00 3,400 0.5 13.4 0.k93 34 9 1k4 11,592 11.592 LI) 19k j

19,000 10.700  ?.4 19.5 0.563 P 1? 146 37,436 37,436 k Peach Bottom 3 PS 764 26,411 ) l 4 Fukashina ? I'*J 554 16,300 9,T90 3.9 19.5 0.561 $2 12 1kk 27,126 ' 14k 4

                                                                                                                            ?,k00                     1.9                    19.5    0.563                37                       12                                        44g              44g y

! F"! A 0 3,500 cooper C l'9/k?q2 17,900 17,400 2.9 19.5 0.563 32/37 12 1h4/146 26,852 20,972 O l 3 1,7 1?O - - 0.1 13.4 0.403 1h 9 lbk/146 7,560 7,560 $ l 4 Duane Arnold A'3 369 19,hoo 11,100 2.9 19.5 0.563 37 12 14b 15.032 13,720 ya 9aloas GE7 k 7,600 5,500 0.9 19.5 0.563 37 12 Ikk 196 196 8 { 9 5,292 N LT 94 6,600 k,900 0.9 13.h 0.k93 34 144 5.292 I 9,130  ?.3 19.5 0.563 37 19 144 27,440 27,440

                                    ,O          k       Hatch                FfX           560 7

I?,900 Ikk 27,440 27,440 k F D zPat ric k EA 132/4?9 I?,600 7,?00  ?.1 19.5 0.563 3?/37 12 N 1.6 0.563 12 Ikk 27,440 27,440 8 k Brunswtek ? B1 560 7,600 k,500 19.5 37 1,900 1,000 0.6 19.5 0 563 37 12 Ikk 196 196

                                    $                      Sclnat            GED               b Assembly average exposare .br those assenblics renaining in the core or assembly average 11scharge esposure v'Len no assemblies remain in the core.                                                                                                          ,

j hpproximately 1/4 (vre wss tra old 7x7 fuel design an1 the remaining 43/4 core the " improved" Tx7 fuel design. ! 3" Improved" 7.7 fuel ro1 has a ' .563 inch diameter and a 37-mil wall thickness. The 919 rod has a 0.493 inch diavter and a 3k-m11 vall thickness. l d fInformationasof'%rch29,lh Inf ormation as of September 31, 1974 l l I ) i l e i

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$ i i 9 9 TABLE h. STi'4ARY OF GENERAL ELECTRIC OPERATING EXDE9IENCE i WIT 9 PRODUCTION GADOLINIA-BEA9I M 7UEL (JvME 30, 1976) Number of Number of Exposure l Design Active Gadolinia Gadolinia Average Time in Peak Fuel ' lod Cladding Pellet-to- Fuel Class of Fu sl Bearing Bearing Assembly Core LO9 Diameter Thickness Cladding Gap Length Reactor Reactor  ?;pe Bundles Rods (?Nd/t) (yearsi (kW/ft) (in.) (mils) (nominal mils) (in.) 1 Dresden 1 111F 104 1042 - 23,000 9.5 15 5 0.5625 35 10 153.25 V 106 324 -1S,000 6.5 15 5 0 5625 35 a 108.25 1 Big Rock Point m 38 1802 12,000 k.5 17.7 0 5625 ko 11 To i F 85 340 8,000 2.8 l' - 7 0.5625 ko 11.5 70 1 Humboldt 111 1 T6 352 12,800 4.4 16.8 0.563 32 11 79 1 Tarapur 1 TA 6T 134 10,400 4.1 15 8 0 563 3? 12 1h2.25 j TB T4 148 9,100 3.0 15.8 0.563 32 12 142.5 1 Tara,ur 2 TA 37 74 8,900 35 15.8 0 563 32 12 142.25 3 ' TB 14 28 T,500 2.0 15.8 0 563 32 12 1h2.25 I 2 Oyster Creek JCA lj6 62h 12,600 4.5 17 5 0.563 32 12 14k g a j 2 Nine Mile Point 1 N'1A 56 22h 1k,000 k.6 17.5 0 563 32 12 144 Y' l GEA 40 120 17,500 4.0 17 5 0.563 32 12 144 3  ! j j N'4C 108 432 Ik,700 3.0 17 5 0 563 37 12 1h4 ]; l

• N'O 96 384 7,900 2.0 13.4 0.493 34 9 144 j

{ ! .? IJO 108 k3? 3,000 0.5 13.4 0.k91 3h 9 14k i LJ2 92 368 3,000 05 13.4 0.493 34 9 144 l $3 i 2 Tsuruga JAA 48 192 14,100 5.6 17 5 0.563 32 12 Ikk l JAB 85 340 10,600 3.6 17 5 0 563 32 12 144 JAC T6 304 9,000 30 17 5 0.563 37 12 1kk JAD 47 188 6,000 2.0 17 5 0 563 37 12 144 j JAE T2 288 h,000 07 17 5 0.563 37 12 144 hk2 12,800 50 17.5 0 563 32 12 144  ; l 3 Dresden 2 CY 215 DN 509 1018 T,700 4.0 17.5 0.563 32 12 144 i GEB 32 96 4,100 1.0 17.5 0 563 37 12 144 i LJO 108 h32 2,900 1.0 13.4 0.k93 34 9 144 l 3 GBR 8 32 1,000 0.3 13.4 0.493 34 9 144 i

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  ,5    3     Dresden 3         GEB        52         156           9,400      3.0       17.5   0 563         37            12        Ikh       l DDB        4h         11 6          8,100      2.0       13.4   0.k93         34             9        Ikk       i l

IJO 108 432 4,200 0.7 13.h 0.k93 34 9 Ikk j i 7 ikk 1 O GBH 8 32 3,000 0.7 13.h 0.493 3h 9  ; ! g LJ2 24 96 4,100 0.7 13.4 0.493 34 9 Ikk , ~ Fukushima 1 TXA 60 240 6,300 h.5 17.5 0 563 3? 12 144 l 3 i j l TXB 111 444 5,100 2.8 17 5 0 563 37 12 144 , TXC 40 160 1,500 05 17.5 0 563 37 12 144 I I 1 l t

I 9 TABLE k.2-9 (C 9nued) 9 I SU!iMARY OF Gr'.NERAL ELECTRIC OPERATING EXPE9IENCE WITH PRODUCTION GADOLINIA-BEA9I47 WEL (JUNE 30, 1976) Number of Number of Exposure l Design Active , Gadolinia Gadolinia Average Time in Peak Fuel Rod Cladding Pellet-to- Fuel  ! Class of Fuel Bearing Bearing Assembly Core LM7R Diameter Thickness Length Reactor Reactor Type Bundles Rods (mwd /t) (years) (kW/ft) (in.) (mils) Cladding (nominal Gap mils ) (in.) 3 Millstone GEA 82 246 13,700 3.3 17.5 0.563 32 12 1h4

GEB 30 90 14,500 3.3 17 5 0 563 37 12 1kk j tGB 125 500 9,200 1.6 17.5 0.563 37 1? 1hk

! MO 12G 480 h,900 0.6 13.h 0.493 3h 9 1% M2 23 92 k,800 0.6 13.1, 0.h93 34 9 1h4 3 Monticello GEB 20 60 15,000 3.0 17 5 0 563 37 12 In tfI'B 116 h64 9,000 2.0 13.h 0.h93 34 9 1% GFB 80 320 6,800 1.4 13.h 0.h93 34 9 in U2 267 801 3,600 0.6 13.4 0.k93 3h 9 IW LJ3 1 3 3,200 0.6 13.4 0.k93 34 9 lu 3 Nucienor GEA 28 84 15,500 37 17.5 0.563 32 12 lu ' NUB 68 272 13,500 3.0 17.5 0 563 37 12 1% NUC 96 3% 9,500 2.0 13.k 0.493 34 9 in MO 60 2ho 6,500 1.0 13.h 0.h93 34 9 IW 3i 9 3 Pilgrim BEA 20 80 k,600 2.0 13.h 0.493 3h 9 lu 8* l o l 3 Quad Cities 1 CX T2h IT60 13,300 h.T 17 5 0.563 32 I? IW s" l . GEB 28 84 8,600 2.0 17 5 0.563 37 12 lu i 7 GEH 36 lu 9,1h0 2.0 13.4 0.k93 34 9 1% l t", MO 12 k8 1,hoo 0.3 13.k 0.493 3h 9 lu M2 lu 576 1,600 0.3 13.h 0.493 34 9 1kk l 3 Quad Cities 2 CY T24 1760 13,200 k.2 17 5 0 563 32 12 lu CX N 31 92 12,hoo 07 17 5 0.563 3? 12 1% MO 148 592 5,800 1.2 13.k 0.493 3h 9 1kh 4 Brovns Ferry 1 TY 596 2717 5,TTk 3.0 18.5 0 563 37 la ik. h Browns Ferry 2 TZ 596 2717 2,156 2.0 18.5 0.563 37 12 lu k Brunswick BR hho 1392 h,500 1.0 18.5 0 563 37 12 IW s h Cooper C h20 1932 9,600 2.5 18.5 0 563 37 12 1h6 h Duane Arnold AR 288 816 T,300 2.3 18.5 0.563 37 12 lu y GED 4 12 1,500 0.2 18.5 0 563 37 12 10 k FitzPatrick EA 428 1970 h,300 1.6 13.5 0.563 37 12 lu O 4 Fukushiraa 2 m k36 2000 7,koo 3.3 18 5 0 563 37 12 in FUA 9 36 1.300 13 18.5 0.563 37 12 in 4 Hatch 1 HXO 560 2008 6,300 1.8 18.5 0 563 37 12 lu

TABLE h.2-9 (Co ed) l i GU' NARY OF GENERAL ELETRIC OPERATING EXPE9IENCE j WITH PRODUCION GAD 0LIMIA-BEA' LIM PJEL ' (JUNE 30, 1976) Nunber of Exposure l Design Active Number of Gadolinin Average Time in Peak Fuel Rod Cladding Pellet-to- Fuel Gadolinia L91R Dianeter Thickness Lenrth Fuel Bearing Bearing Assembly Core Cladding Class of Reactor Reactor Type Bundles Rods (?Nd /t) (yea r_s) (kW/ft) (in.) (mils) (nominal Gap mils ) (in.) { i 12 144 k KKM GED 12 36 10.800 25 18.5 0.563 37 I ' 108 432 11,400 1.7 13.h 0.493 34 9 lk4 A!u 0.8 13 4 0.493 3h 9 144 LI2 40 200 6.100 3050 10.200 2.8 18.5 0 563 3T I? Ikk l k Peach Bottom 2 PH 596 3 LIL h 20 <100 0.1 13.4 0.493 32 9 1pc ,

                                                                                                                                                         <100                                0.1   13.h                     0.h93          34                              9                 144 LI3                    18h         920 l

h Peach Bottom 3 PB 596 2717 7,300 2.0 18.5 0 563 37 12 146 . 40 120 8,100 2.6 18 5 0.563 37 12 Ikh 4 vermont Yankee GED 9,600 15 13.4 0.493 34 9 144 LIO 328 964 l i > i 1 Assembly sverage exposure for those assemblies reciaining in the core. Q

• 2 Includes 98 standard assemblies with 1 Gd230 -Alumina rod; 2 special assemblies with 1 Gd ?o3-Alumina segmented rod; and 4 special assemblies S V 1 d23 0 -Urania segmented rod. o

                 .'?                                                                                                                                                                                                                                                                                 W g            3     Inclades 35  standard assemblies with h Gadolinia-UO2 rods, 4 special E assemblies with 4 Gadolinia-tn2 rods, a ad 3 special EG assemblies                                                                                                                      M g with 8 Gadolinia-UO2 rods.

Quad Cities ? fuel assembly type CX are a portion of the CX assemblies originally loaded into Quad Cities 1. a I

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                                                                                                                               "'able 4.2-10 e

OENERAL ELTTRIC DEVELOP *fENT IRRADIATIONS ZI9CALOY-CLAD 95% TD UO2 PELLT FUP.L 90DS l Number Fuel Rod Clad Vall Pellet-to- Peak Heat Peak Peak I of Diameter "'hicknes s Cladding Flux LM09 Exposure Name Reactor Rods (in.) (in.) Gap (mils) Btu /h-ft2 (kW/ft) (?f4d /"'e) Status j l l l Dresden Prototype VB'4R 9 0.565 0.030 3.0-16.0 460,000 10 94 12,000 Completed l Yuel Cycle (R&D)* VBW9 144 0.h2h 0.022 2.0 9.0 h09,000 16.6 13,800 Completed Dresden Prototypes V' 5? 0 565 0.029 5 0 9.0 h07,000 17.64 10,001 Completed I h l High Performance 12 0 565 0.030 h.0 i.0 630,000 27 0 1,50? Completed UO b 1,1.>4,qoq hq,9 2 High Pcrformance in 2 0.565 0.030 h.0-11.0 1,135,000 58.0 14,000 Completed

• UO b 2

SA-lc Dresden 1 98 0.k2k O.022 h.0 8.0 400,000 13.0 40,000 Completed i l D-1,2,3 d consumers 363 0.424 0.030 7.0 43h,000 14.2 30,000 completed 2 9 D-50 f Consumers 36 0 570 0.035 12.0 507.000 22.0 15,400 g,i 8

 . D-52,53                consumers                                    58                        0 700                       0.040          13.0      525,000   27 0      4,600         i             h l 7 l 5 "USAEC Contract AT(Oh-3) - 189, Project Agreement 11 I

{ bUSAEC Contract AT(04.3) - 189, Project Agreement 17 USAEC Contract AT(Oh-3) - 189, Project Agreement 41 dVSAEC Contract AT(04-3) - 351

    ' Hollow Pellet                                                                                                                                                                                      .

I USAEC Contract AT(Ok-3) - 199, Project Agreement 50

 . 8Eight fuel rods tested during second operating cycle due to abnornal crud and scale deposition z h
 . One rod failure at 49 KV/ft                                                                                                                                                                        i i

7" Fuel assemblies presently out of reactor pending approval for reiratertion C O i l

I'l 3  !' I f! i l,iI 1' !l, , k; O _ Ns,iu N e e e e e e e _ s t t t e t e t e t e t e _ u e e l l l t l l l l t a p m n p p s: p p m m o n no p p _ S o o o oC o _ C C C C C C _ e) ) re 0 0 0 0 0 ' 0 k u? 0 0 0 0 0 0 0 a s/ eod 0, 0, 0, 0, 5, 0 0, P pJ f 9 4 5 0 6 0 9 x 9 3 6 5 3 0 2 _ E (' 1 q 0 tJ

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f 9 T L m 'J l l 2 Lo M as 3 E 'A 0 V s _ I C e) . 2 + 6 6 5 0 C gn 0 3 '3 3 0 3 0 1 0 3 0 D nkn - 0 0 _ "E ici k 0 0 0 O L t i( 2 0 0 0 I L th 0 . a' P l" 0 CE L C _ S L d r _ _ A oe) _ t l

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5 e n 2 8 S 5 5 5 9 l - wi k. 5 5 2 '4 S ea( k. k. G ui 0 0 0 0 0 0 0 FD _ r e s bfd 3 1 6 5 5 5 3 noo a R N _ e l u F s A B C D E p a C O  :  ; ? # cDO l' i jr fal! i!. i!l:I  ! ll- j lll ,1I!

_ _ _ _ _ _ _ ..__ _ __ _ _ _ _ . _ _ _ _ = . . _ . . . _ . . . _ . _ _ _ _ . . . . . _ _ . . .__. _ . _ . . . . . _ _ _ . . _ . e i. a i G O I Table k.2-12 4 j HALDEN IRRADIATIO*! PROC %'t STATUS I i Peak LH1R Peak Exposure l Number Fuel Rod Cladding Wall Pellet-to- (kW/ft) I '* /T ) d of Diameter Thickness Cladding Gap (ac of tas of Paraneter o. Assembly Rods (in.) (in.) (nils 9/27/74) (9/27/Th Interest i l IFA-131 6 0.563 0.032 8--14 16.8 37,300 Pellet . geonetry IFA-213 7 0 563 0.032 10 16.2 2T,600 cladiing heat I treatment IFA-214 7 0 563 0.032 - 0.060 8--10 8.6 18,500 Hollow pellets , i mixed oxide i IFA-237 1 0 563 0.032 -- 18.0 28,000 vipae powder 4 i IFA-238 7 0 563 0.032 - 0.040 8 - 10 19.3 21.400 Cladding thickness and heat treatment i ! IFA-236 6 0.h93 0.034 8 12.2 16.100 Fuel dem ity and i hollow pellets N IFA kO8 11 0 563 0.03T 1 - 10 16.8 12,900 vipae, mixed oxide, g l * " i hellow pellets

   ?                                                                                                                                                                                                           5 12                          0.493         0.034                                   9        15 4                              11,300                                Densification and    M
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Draft 2 1 1/13/81 i ACNGS-PSAR i Table 4.2-13 i FUEL ROD VIBRATION INFORMATION Rod OD/Cled Calculated Rod Fuel Design Lattice Thickness In. Natural Frequency Rove - Yankee

• 16 x 16 0 362/0.021 25 9 Nuclenor* 7x7 0 570/0.0355 29 8 l

Humboldt Reload 7x7 0.486/0.033 29 9 i I Big Rock Point 11 x 11 0.449/0.034 31 3 I l Reload BWR/2-5 8x8 0.493/0.034 32.C - 35 0 BWR/2-6 8x8 0.483/0.032 31.6 - 33 7

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Draft 2 ACNGS-PSAR 1/13/81 Table h.2-14 LINEAR HEAT GENERATION RATE OF CALCULATED 1% PLASTIC DIAMETRAL STRAIN FOR BWR/6 FUEL ,

                                                                                                                ;

1 i Exposure LHGR at Calculated 1% l (mwd /t) Plastic Strain (kv/ft)* , UO2 Gd** i 0 >24.7 >21 9 20,000 >23 0 >20.h h0,000 })96 })73 "The values reported have been reduced by an amount equal to the calculated power spiking penalty (%)

      **Results for gadolinia are applicable for maximum concentration used in BWR/6.                                                                                               l
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ACNG3-PSAR 100 O G E ASSUMPTIONS: (1) ALL RODS IN THE BUNDLE ARE AT THE SAME TEMPERATURE 2 (2) RODS REACH INDICATED y TEMPERATURE INSTANTANEOUSLY v

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          ~                                                                                                    l 1000'F 0.01            ' ' I'!!              !      !    ! ! ! III                I               I I ! ! II 0.1                 1.0                                   10                                     100 TIME AT INDICATED TEMPERATURE (sec)

Am. No. 56,(3/81) HOUSTON LIGHTING & POWER COMPANY Allens Creek Nuclear Generating Station Unit 1 F;.21 Energy Release as a Function of Time Figure 4.2-9

ACNGS-PSAR 4.3 HUCLEAR DESIGH I m

   \

\ / The nuclear core design presented herein is based on the equilibrium cycle rather than the initial cycle. The justification for this feature is i presented in Section 4.1. The design bases and licensing requirements are independent of whether an initial or equilibrium cycle is used. Only the i description and the core characteristics are different. 4.3.1 Design Bases The nuclear design bases are conveniently divided into two specific categories. The safety design bases are those that are required for the  ; plant to operate fron safety considerations. The second category is the < plant perfornance design bases that are required in order to meet the objective of producing power in an efficient manner.  ! 4.3.1.1 Safety Design Bases The safety design bases are requirements which protect the nuclear fuel from a violation of the design integrity liuits. In general, the safety bases fall into two categories: (1) the reactivity basis which prevents an uncontrolled positive reactivity excursion, and (2) the overpower bases, which prevent the core from operating beyond the fuel integrity limits. 4.3 1.1.1 Reactivity Basis The nuclear design shall meet the following basis: The core shall be 4 i capable of being rendered subcritical at any time or at any core conditions , ( ,/ with the highest worth control rod fully withdrawn. 4.3 1.1.2 Overpower Bases The nuclear design shall meet the following bases: (1) The void coefficient shall be negative over the entire operating range. (2) The Technical Specification limits on Linear Heat Generation Rate (LHGR), Minimum Critical Power Ratio (MCPR), and the Maximum Average Planar Linear Heat Generation Rate (MAPLHGR) shall not be exceeded during steady-state operation. (3) The nuclear characteristics of the design shall exhibit no tendency toward divergent operation. 4.3 1.2 Plant Performance Design Bases The nuclear design shall meet the following bases: (1) The design shall provide adequate excess reactivity to attain the

                                                                                             +

desired cycle length. (2) The design shall be capable of operating at rated conditions  ! (m) without exceeding Technical Specification limits. 1 4.3-1 Am No. 56,(3/81)

ACHGS-PSAR (3) The nuclear design and the reactivity control system shall allow continuous, stable regulation of reactivity. (4) The nuclear design shall have adequate reactivity feedback to facilitate normal operation. (5) Separate bases on the void coefficient are defined wit'. respect to core stability analysis and plant transient analyses.

a. The uagnitude of the uaximum void coefficient at any point in the cycle shall not exceed a specified design limit used to evaluate core reactivity stability.

b. The dynamic void reactivity coefficient, which is used to measure the core transient response, shall not exceed a design limit established to ensure optimum plant performance. The dynamic void reactivity coefficient is defined as the nuclear void reactivity coefficient, expressed in cents of reactivity, multiplied by the core average void fraction, expressed in percent.

(6) The scram reactivity response shall not violate a design limit established to ensure optimum plant performance. This design limit must be met during the first 60% of scram insertion, since the transient response is insensitive to the remainder of the scram curve. The scram reack.ity is evaluated at end-of-cycle for the 105% steam flow reactivity condition. (7) The Doppler reactivity coefficient shall not violate a design limit established to ensure optimum plant performancer. This design limit is established for the end-of-cycle condition at 105% steam flow. 4.3 2 Description The BWR core design utilizes a light-water moderated reactor, fueled with slightly enriched uranium-dioxide. The use of water as a moderator produces a neutron energy spectrum in which fissions are caused principally by thermal neutrons. At normal operating conditions, the moderator boils, producing a spatially variable distribution of steam voids in the core. The BWR design provides a system for which reactivity changes are invursely proportional to the steam void content in the moderator. This void feedback effect is one of the inherent safety features of the BWR system. Any system input which increases reactor power, either in a . local or gross sense, produces additional steam voids which reduce reactivity and thereby reduce the power. The fuel for the BWR is uranium-dioxide enriched to approximately 3 vt% in U-235 Early in the fuel life, the fissioning of the U-235 produces the majority of the energy. The presence of U-238 in the uranium-dioxide fuel leads to the production of appreciable quantities of plutonium during core l operation. This plutonium contributes to both reactivity and reactor power production (i.e., approximately 50% at end-of-life). In addition, direct fissioning of U-238 by fast neutrons yields approximately 7 to 10% of the 4.3-2 Am. No. 56,(3/81)

ACNGS-PSAR total power and contributes to an increase of delayed neutrons in the core.

 /G    Since the U-238 has a strong negative Doppler reactivity coefficient, the i     peak power during an excursion is limited.

The reactor core is arranged roughly as a right circular cylinder containing a large number of fuel assemblies and control rods. The dimensions of the reactor core are shown in Table 4.3-1. 4.3.2.1 Nuclear Design Description At each refueling period, approximately 25% of the fuel bundles are dischar6ed from the core and replaced with an equivalent number of fresh fuel assemblies. The fuel bundles having the highest exposure (i.e., the lowest reactivity) are discharged starting with the highest exposure and , moving toward less exposure. The remaining bundles and new fuel assemblies are then shuffled in order to minimize radial power peaking and maximize the end-of-cycle reactivity. This is accomplished by loading the lowest reactivity fuel on the periphery, loading the relatively high reactivity fuel in a region next to the periphery toward the core center, and loading the medium reactivity fuel in the central region of the core. Within each of these zones, the fuel bundles are arranged in a nearly homogeneous  ; uanner in order to minimize reactivity mismatch. A diagram of the equilibrium core bundle loading pattern is shown in Figure 4.3-1. Each bundle contains 62 fuel rods and 2 vater rods. The water rods have a slightly larger diameter than the fuel rods. The layout and dimensions of g the fuel bundle are presented in Figure 4.3-2. Gadolinium, in the form of t 1 Gd230 , is mixed with the UO2 and placed in selected fuel rods to provide

 ,b    reactivity control. The fuel rod distribution for the reload bundle is shown in Figure 4.3-3. The axial distribution of enrichment and gadolinia for the rods in the fuel bundle is shown in Figure 4.3 4. Each fuel and burnable poison rod has natural uranium e.t the top and botton.

4.3 2.1.1 Fuel Nuclear Properties The bundle reactivity is a complex function of several important physical properties. The impcrtant properties consist of the average bundle enrichuent, the gadclinia rod location and gadolinia concentration, the void fraction and the accumulated exposure. The variation in reactivity of the infinite lattice data (k-infinity as a function of void fraction and exposure) for the dominant segment of the bundle is presented in Figure 4.3-5 At low exposu 2, the reactivity effect due to void formation is readily apparent; however, at higher exposure, due to the effect of void history, the curves cross. The primary reason for this behavior is the greater rate of plutonium formation at the higher void fraction. The variation of the isotopic concentrations as a function of exposure for the douinant fuel type is presented in Figures 4.3-6 and 4.3-7 for the important heavy element isotopes. Early in the fuel bundle life, approximately 93% of the power is produced i by fissions in U-235 with the remainder coming from fast fissions in U-238. . l At high bundle exposures, typical of discharge, the power production due to O plutonium exceeds that of the U-235 The fraction of total fissions in the ! f ) important isotopes is shown in Figure 4.3-8 for the dominant fuel type. v  ; 4.3-3 Am. No. 56,(3/81)

I ACHGS-PSAR Other bundle paraueters such as neutron generation time and delayed neutron fraction as a function of exposure at approximately core average voids are shown in Figures 4.3-9 and 4.3-10, respectively, for the doninant segment of the bundle. The variation of the core-wide nuclear characteristics is a function of the characteristics of each bundle in the core. With the various reload situations, any description of the gross core characteristics can only be expressed in terms of the overall core performance. 4.3.2.2 Power Distribution The core is designed such that the resultant operating pawer distributions neet the plant Technical Specifications. The three criteria for thermal limits are the Maximum Linear Heat Generation Rate (MLl!GR), the Minimum Critical Power Ratio (MCPR), aad a Maximum Average Planar Linear Heat Generation Rate (MAPLHGR)l. Each of these thermal limits is a function of both the gross three-dimensional power distribution and the local rod-to-rod power distribution. Sufficient design calculations are performed to insure the core meets these criteria. For design convenience, separate target peaking factors are used for the local and the gross power distributions. The local peaking factor is defined as the ratio of the power density in the highest power rod in the lattice, at a cross section throudh the bundle, to the average power density in the lattice at that location. In addition, the local effects on the Critical Power Ratio (CPR) are characterized by a quantity designated as R-Fector2 . For the BWR/6 nuclear fuel design, the target local peaking factor is 1.13 and the target R-factor is 1.05 The gross power peaking is defined as the ratio of the maximum power density in any axial segment of any bundle in the core to the average power density in the core. The target gross pover peaking limit for this BWH/6 fuel design is 2.00. Variations from these target peaking factors are considered acceptable providing the Technical Specifications are not exceeded anywhere in the core. That is, the real criteria are the LHGR, the MAPLHGR, and MCPR; not the power peaking factors. The peaking factors are used merely as a design convenience. Appropriate design allowance's are included at the design stage to provide a high degree of assurance that the Technical Specification limits will be met during plant operation. During operation of the plant, the power distributions are measured by the in-core instrumentation system, and thermal margins are calculated by the process computer. 4.3.2.2.1 Local Power Distribution The local rod-to-rod power distribution and the associated R-Factor distribution are functions of the lattice fuel rod enrichment distribution. Near the outside of the lattice, where the thermal flux peaks due to interbundle water gaps, lower enrichment fuel rods are utilized to minimize power peaking. Closer to the center of the bundle, higher enrichment fuel rods are used to increase the power generation and flatten the power distribution across the bundle. In addition, two water tubes containing unvoided water are at the center of.the lattice. The combination of these factors results in the relatively flat local power distribution. The fuel 4.3-4 Am. No. 56,(3/81)

ACNGS-PGAR rods which contain gadolinia produce relatively little power early in bundle life; however, as the gadolinia is depleted, the power in these rods g increases to approximately the lattice average. The variation of the maximum local peaking factor as a function of exposure at 40% voids is shown in Figure 4.3-11. The high power rods deplete at a greater rate and the local power peaking factor decreases with exposure. The local rod-to-rod power distribution for 40% voids at beginning-of-cycle, at an exposure typical of end-of-equilibrium-cycle, and end-of-bundle life is shown in Figure 4.3-12. The variation in local power distribution for various lattice void fractions is shown in Figure 4.3-13. In general, the local power distribution tends to flatten with increasing void fraction. It can be seen fron these data that the target local peaking factor of 1.13 is met over the range of interest. Although the

 ;

1.13 target value is exceeded at very lov exposures, the reactivity of the

 ~

bundle is sufficiently lov that the bundle is never limiting. At exposures where this bundle produces significant power, the local peaking factor is well within the target. The presence of a control blade adjacent to the 3 bundle significantly skews the local power distribution within the bundle. I The local power distribution for the controlled lattice is shown in Figure 4.3-14 Although the local peaking factor is quite large in this case, the gross power in a controlled bundle is sufficiently low such that a controlled lattice is never limiting. Figure 4.3-1S presents the uncontrolled R-Factor for each fuel rod at a planar elevation through the dominant fuel type at beginning-of-cycle at (p) V 40% voids. The variation of the maximan bundle-integrated R-Factor with exposure is presented in Figure 4.3-16. The data shown demonstrate that the 1.05 target limit is not exceeded for the exposure range of interest. 4.3.2.2.2 Radial Power Distribution The integrated bundle power, commonly referred to as the radial power, is the primary factor for determining MCPR. At rated conditione, the !!CPR is directly proportional to the radial power peaking. The radial power distribution is a complex function of the control rod pattern, the fuel bundle type, the loading pattern and the void condition for that bundle. The BWR simulator 3 is used to calculate the three-dimensional power distribution in the core and the power is axially integrated to determine average bundle power. The bundle radial power distributions for typical beginning and end-of-cycle conditions are presented in Figure 4.3-17 t The radial power distribution is influenced by both the radial reactivity i zones and the control rods. The control rods are selectively inserted, or withdrawn, to flatten the radial power distribution consistent with the reactivity control needed. Near the end-of-cycle, the region of high reactivity adjacent to the periphery provides the necessary radial power flattening without recourse to control rods. l 4.3.2.2.3 Axial Power Distribution For the reload situation, the axial exposure shape existing in the bundles , sj vhich remain rrom 1.revious cycles, in combination with the selective use of l 4.3-5 Am. No. 56,(3/81) l

ACNGS-PSAR shallow control rods, provide the necessary axial power flattening. Current reload bundle designs contain axially uniform gadolinia. The effect of voids is to skew the power toward the bottom of the core; the effect M the bottom entry control rods is to reduce the power in the bottom of the core; and the effect of the exposure distribution is to flatten tne power. Detailed three-dimensional calculations are performed to deteruine the axial power distributioa. The control cod patterns help to achieve an end-of-cycle exposure distribution which approximates the Haling shapelh. This, in turn, provides a relatively flat axial power shape with all control rods withdrawn. A typical beginning-of-equilibrium-cycle axial power shape is shown in Figure h.3-18 along with the optimal end-of-equilibrium-cycle power shape. h.3.2.2.h Power Distribution Calculations A full range of calculated power distributions along with the resultant exposure shapes and the corresponding control rod patterns are shown in Appendix hA for a typical BWR/6 equilibrium cycle. h.3 2.2 5 Power Distribution Measurements tlc techniques for a measurement of the power d: stribution within the reactor core, together with instrumentation cori elations and operation limits, is discussed in Reference 4. h.3 2.2.6 Power Distribution Accuracy Tre accuracy of the calculated local rod-to-rod power distribution is discussed in Reference 5 The accuracy of the radial, axial and the gross three-dimensioral power 'Astribution calculations is discussed in Reference 6. h.3 2.2.7 Power Distribution Anomalies Stringent inspection procedures are utilized to ensure the correct rearrangement of the core following refueling. Although a misplacement of a bundle in the core would be a very improbable event, calculations have been performed in order to determine the effects of such accidents on linear heat generation rate (LHGR) and critical power ratio (CPR). These results are presented in Chapter 15 The inherent design characteristics of the BWR are well suited to limit gross power tilting. The stabilizing nature of the large moderator void j coefficient effectively reduces perturbations in the power distribution. In addition, the in-core instrumentation system, together with the on-line computer, provides the operator with prompt information on power distribution so that he can readily use control rods or other means to limit the undersirable effects of power tilting. Because of these design characteristics, it is ne necessary to allocate a specific margin in the peaking factor to accour Cor power tilt. If, for some reason, the power distribution could not ue unintained within normal limits using control 4.3-6 Am. No. 56,(3/81)

ACNGS-PSAR rods, then the operating power limits would have to be reduced as prescribed in Chapter 16 (Technical Specifications). ('~'} V h.3 2 3 Reactivity Coefficients Reactivity coefficients, the differential changes in core reactivity produced by differential changes in core conditions, are useful in calculating the response of the core to external disturbances. The base initial condition of the system and the postulated initiating event determine which of the several defined coefficients are significant in evaluating the response of the reactor. There are three primary reactivity coefficients which characterize the dynamic behavior of BWRs over all operating states: (1) Doppler reactivity coefficient; (2) moderator temperature reactivity coeffic'ent; and (3) moderator void reactivity coefficient. Also associated fich the BWR is a power reactivity coefficient; however, this coefficient is merely a coubination of the Doppler and void reactivity coefficients in tne power oyerating range. In order to assure optimum plant performance determined by the predicted plant transient response, design limits have been established on those nuclear paraueters which have a significant effect on the plant transients. Specifically, design limits have been established for the void coefficient, the Doppler coefficient, and the scram reactivity response. These limits were established such that the propar balance was made between the nuclear design and the plant desigr.. The vesign linits presented herein represent 7s

    )

the extreme values expected to occur during normal operation of the core \s_ / over the lifetime of the plant. These reactor dynamics calculations are indicative of core responses to individual phenomena. The code used to calculate transient response utilizes all of the specific inputs to predict the core response to a particular transient. 4.3 2 3 1 Void Reactivity Coefficients The most important of the reactivity coefficients is the void reactivity coefficient. The void coefficient must be large enough to prevent power oscillation due to spatial xenon changes yet small enough that pressurization transients do not unduly limit plant operation. In addition, the void coefficient in a BWR has the ability to flatten the radial power distribution and to provide ease of reactor control due to the void feedback mechanism. The overall void coefficient is always negative over the couplete operating range since the BWR design is undermoderated. The reactivity change die to the formation of voids results from the reduction in neutron sloving down due to the decrease in the water-to-fuel ratio. A detailed discussion of the methods used to calculate void reactivity coefficients, their accuracy and their application to plant transient  ; analysis is presented in Reference 7 A comparison of a detailed model using spatial variation of the important parameters and the point model is also shown.

,0 hv) 4.3-7                         Am. No. 56,(3/81)

ACNGS-PSAR The uoderator void reactivity coefficient as a function of percent voids is presented in Figure 4.3-19 for the exposure at which the void coefficient reaches its r:nximum. Also shown in this figure is the design limit used for core stability analysis. 't is clear that the magnitude of the void coefficient is less than the design limit. The most limiting transient response occurs at end-of-cycle. The variation of the calculated dynamic void reactivity coefficient, as a function of core average void percent is shown in Figure 4.3-20 for the end-of-cycle conditior.. Also shown is the couparison of the design limit and the calculated nominal value at 105% steam flow conditions multiplied by 1.25, the safety conservatism factor. These data demonstrate that the design limit value is not exceeded. 4.3.2.3 2 Moderator Temperature coefficient The moderator temperature coefficient is the least important of the reactivity coefficients, since itr effect is limited to a very small portion of the reactor operating >snge. Once the reactor reaches the power producing range, boiling begins at the moderator teuperature remains

ssentially constant. As with tht void coefficient, the moderator temperature coefficient is associa ed with a change in the moderating power of the water. The temperature coe ficient is negative for most of the operating cycle; novever, near the end-of-cycle the overall moderator temperature coefficient beccmes slightly positive, due to the fact that the uncontrolled BWR lattice is slightly overmoderated in the unvoided state at high exposures. This, combined with the fact that more control rods must be withdrawn fror the reactor core near the end-of-cycle to establish criticality, results in the slightly positive overall moderator temperature coefficient.

The range of values of moderator temperature coefficients encountered in current BWR lattices does not include anv that are significant from the safety point of view. Typically, the teuperature coefficient may range from +4 x 10-5 Ak/kOF to -14 x 10-5 Ak/keF, depending on base temperature and core exposure. The small magnitude of this coeffici-;t, relative to that associated with transfer of heat from the fuel to the coolant, makes the reactivity contribution of moderator temperature change insignificant during rapid transients. For the reasons stated previously, current core design criteria do not impose limits on the value of the temperature coefficient, and effects of minor design changes on the coefficient in members of the same class of core usually are not calculated. A measure of design control over the temperature coefficient is exercised, however, by applying a design limit to the void coefficient. This constraint implies control over the water-to-fuel ratio of the lantice; this, in turn, controls the temperature i coefficient. Thus, imposing a quantitative limit on the void coefficient effectively limits the temperature coefficient. 4.3.2.3 3 appler Reactivity Coefficient The Dopplei reactivity coefficient is the change in reactivity due to a change in tua temperature of the ' fuel. This reactivity change is due to 4.3-8 Am No. 56,(3/81)

L

                                               -ACNGS-PSAR                                                           {

l t the broadening of the resonance cross sections as the fuel temperature increases.- At beginning-of-life, the Doppler contribution is pri'marily due  ; k to U-238; however, the buildup of Pu-240 with exposure adds to the Doppler  ! coefficient. A detailed discussion of the methods used to calculate the  ! Doppler coefficient, their accuracy, and application to plant transient i

analyses is presented in Reference 7 The application of the Doppler [

coefficient to the analysis of the Rod Drop Accident is discussed in i Reference 8. {, l' J The variation in the core average Doppler reactivity coefficient as a  ; function of average lattice fuel temperature at end-of-equilibrium cycle { and with the multiplier of 1.25, the safety conservatism factor, are shown l in Figure h.3-21. Also shovn is the design limit for the Doppler l coefficient. These data demonstrate that the design limit on the Doppler l coefficient is not violated. f j L 3.2 3.4 Power Coefficient 4 1 The power coefficient is determined from the composite of all the j significant individual sources of reactivity change associated with a l differential change in reactor thermal pcVer assuming xenqn reactivity  ! reuains constant. At end-of-equilibrium-cycle, the power coefficient at 105% steam flow conditions is approximately -0.05 Ak/k + AP/P. This value is well within the range required for adequately damping power and  ; spatial-xenon disturbances. f h.3 2.4 Contro] Requirements

                                                                                                                     ]

The nuclear design in conjunction with the reactivity control system l provide an inherently stable system for BWRs.  ; i The control rod system is designed to provide adequate control of the i maximum excess reactivity anticipated during the equilibrium cycle  ! operation. Since fuel reactivity is a maximum and control is a minimum at l ambient temperature, the shutdown capability is evaluated assuming a cold, l xenon-free core. The safety design basis requires that the core, in its  ! maximuu reactivity condition, be suberitical with the control rod of the t highest worth fully withdrawn and all others fully inserted. f 4.3 2.h.1 Shutdown Reactivity To assure that the safety design basis for shutdown is satisfied, an i { additional design margin is adopted: k-effective is calculated to be less  ; l that or equal to 0 99 with the control rod of highest worth fully i

withdrawn. The cold shutdown reactivity as a function of cycle exposure is ( ! shown in Figure 4.3-22. i  ! l The shutdown reactivity curve shows the calculated values of k-effectiv  ; l for the condition 2000, highest worth rod withdrawn, no xenon. The minimum l

shutdown margin occurs at beginning-of-equilibrium-cycle (BOEC) when [ bundles which have been in the core one cycle are at their peak reactivity. l f Haating the reactor to hot cond tions will increase the shutdown margin by , [ 0.02 Ak to 0.03 Ak. For this reason, shutdown margin calculations are not , l necessary for hot conditions. [ w y I i 4.3-9 Am. No. 56,(3/81) l

                                                                                                                      ;

_ . _ - _ , ,m._- . , - , - . - . . , _ -_- - . - - , . _

I ACNGS-PSAR 1 The accuracy with which shutdown reactivity is calculated is discussed in Reference 6. Basically, the accuracy is characterized as a bias and an uncertainty. The bias is a reactivity correction applied directly to the calculated results. For example: kerr (Expected) = ke rr (Calculated) + ak (Bias) This bias has been incorporated into the shutdown curve shown in Figure 4.3-22. The 1% design margin is satisfied after the bias correction is applied. Reduction of control rod effectiveness during the operating cycle is not a najor concern with the BWR. Using normal control rod sequencing, the control rod worth remains essentially constant over the BWR operating cycle. 4.3.2.4.2 Reactivity Variations The excess reactivity designed into the core is controlled by the control rei system supplemented by gadolinia-urania fuel rods. The reload fuel enrichment for the cycle is chosen to provide excess reactivity in the fuel assemblies sufficient to overcome the neutron losses caused by core neutron leakage, moderator heating and boiling, fuel temperature rise, equilibrium xenon and samariun poisoning, plus an allowance for fuel depletion. Control rods are used during the cycle partly to compensate for burnup and partly to flatten the power distribution. Reactivity balances are not used in describing BWR behavior because of the strong interdependence of the individual constituents of reactivity. Therefore, the design process does not produce components of a reactivity balance at the conditions of interest. Instead, it gives the k e rr representing all effects combined. Further, any listing of components of a reactivity balance is quite ambiguous unless the sequence of the changes is clearly defined. Consider, for example, the reactivity effect of control rods and burnable poison. The combined worth of these two absorbers vould be considerably different than the sum of their individual worths. Even this combined worth would be of questionable significance unless the path and conditions of other parameters (i.e. , temperature, void, xenon, etc. ) were completely specified. Many other illustrations could be presented showing that the reactivity balance approach, which may be appropriate in some types of reactors, is coupletely inappropriate in a BWR. This is related to the large potential excess reactivity in a BWR combined with the dependence of interaction (shadowing) factors on reactor state. In order to understand the various reactivity effects in a BWR design, certain reactivity states can be defined which provide information about BWH behavior. Typical data are presented in Table 4.3-2 and show the predicted reactivtity, kerf, ror various cold, xenon-free conditions. The r sctivity and control fraction values for a variety of operating conditions are shown in Table 4.3-3. The worth of various reactivity 4.3-10 Am No. 56,(3/81)

                                             -                                       s ACHGS-PSAR effects can be estimated by taking the differences between reactivity states with all but one variable consi..a.      Estimates of the temperature

(_}

 \,,,/   defect, the power defect, the xenon defect and the excess reactivity can be inferred.

4.3.2 5 Control Rod Patterns and Reactivity Worths For BWR plants, control rod patterns are not uniquel; .. lied in advance; . rather, durind norual operation, the control rod patt .ns are selected based on the measured core power distributions, within the constraints imposed by the systems indicated in the following sections. All rod i patterns vill be such that the limits specified in the Technical ' Specifications (Chaptcr 16) vill be met throughout the cycle. Typical control rod patterns are calculated during the design phase to insure that . all safety and performances criteria are satisfied. Control rod patterns and the associated power distributions for a typical BWR are presented in Appendix 4A. These control rod patterns are calculated with the BWR Core Sinulator3 The ability of this model to predict control rod worth can be inferred from the detailed reactivity data presented in Reference 6. The comparisons of calculated and measured reactivity for the cold condition in both an in-sequence critical, where roughly 25% of the control rods are withdrawn, and the stuck rod measurement, where only one or two rods are withdrawn, show the ability of the model to predict rod worth. The data presented in Table 7 of Reference 6 shows that no significant bias exists between these r~ss two configurations; therefore, it is concluded that the worth of the rods

       ) is accurately predicted for the cold condition. Figure kh of Reference 6
 \__/    shows the calculated critical reactivity for a variety of BWR cores and over a wide range of exposures. Since this represents a large variation of i

the number of control rods inserted and, since no significant bias is observed, it is concluded that rod worths for the hot operating condition - i are accurately predicted. 4.3 2 5 1 Rod Control and Information System Control rod patterns and associated control rod reactivity worths are regulated by the Rod Control and Information System (RCIS). This system utilizes redundant inputs to provide rod pattern control over the complete range of reactor operations. The control rod worths are limited to such an extent that the rod drop accident (RDA) and the power range rod withdrawal error (RWE) become unimportant. The RCIS provides for stable control of core reactivity in both the single rod or ganged rod mode of operation. The Rod Pattern Control Systea (RPCS) of the RCIS provides protection from a RDA frou startup to the low power setpoint (LPSP), about 20% of rated , power. The Rod Withdrawal Limiter (RWL) function provides protection from , tne RWE for all conditions above the LPSP. Each of these functions is  ! described in the following sections. h.3 2 5 2 Rod Pattern Control System (RPCS) l The RPCS restricts control rod patterns to prescribed withdrawal sequences from the all-rods-inserted startup condition until about 20% of rated . (\~~/

    'N)   power. The withdrawal mode, c>lled bank position withdrawal sequence,                    f F

2 ! 4.3-11 Am. No. 56,(3/81) w - - . -T'-f M r

ACMGS-PSAR minimizes control rod worths to the extent that they are not an importat.t concern in the operation of a BWR. The consequences of a RDA or a RWE in this range are significantly less severe than that required to violate fuel integrity limits. This system ia described in detail in Reference 9 Above 20% of rated power, control rod worths are very small due to the formation of voids in the moderator. Therefore, restrictions on control rod patterns are not required to minimize control rod worths. 4.3.2 5 3 Rod Withdrawal Limiter (RWL) Above the LPSP, the RCIS relies on the RWL function to provide regulation of control rod withdrawals in order to prevent the occurrence of a rod withdrawal error. This function limits the withdrawal of a single control rod or a gang of control rods to a predeternined increment, depending on whether the power level is above or below the high power setpoint (HPSP), typically 70% reactor power. The system senses the location of the rod or gang of rods and automatically blocks withdrawal when the preset increment is reached. The preset licit is determined by generic analysis such that the AMCPR and aLHGR for the RWE are less than the limiting transient. Above the HPSP, the rod will block typically at 12 in. withdrawal. Below the HPSP, the increment is typically 2h in. withdrawal. 4.3.2 5.4 Control Rod Cperatic The control rods can be operated either individually or in a gang composed of up to four rods. The purpose of the ganged rods is ta reduce the time required for plant startup or recovery from a scram. The RCIS provides regulation of control rod operation regardless of whether rode are being moved in single or ganged mode. The assignment of contcol rods to RCIS groups is shown in Figures 4.3-23 and 4.3-24 for sequence A and Figures 4.3-25 and h.3-26 for sequence B. Also shown in these figures is the division of the groups into gangs of 1 to 4 rods which can be moved simultaneously. 4.3.2 5 5 Scram Reactivity The Reactor Protection System (RPS) responds to certain abnormal operational transients by initiating a scran of all control rods. The RPS and the Control Rod Drive (Cn0) System act quickly enough to prevent the initiating disturbance from driving the fuel beyond transient limits. The scram reactivity curve at the end-of-equilibr 1m-cycle is shown in Figure h.3-27 Also shown is the design limit scram curve. At the hot-operating condition, tne control rod, power, delayed neutron and void distributions must all be properly accounted for as a function of time in the transient analysis. Therefore, the scram reactivity is calculated with an integrated one-dimensiot.< 1 reactor core computer model which is coupled to recirculation and major system control models10, h.3.2.6 Criticality of Reactor During Refueling The maximum allowable value ofe k rp is <l.000 at any time. For each reload cycle, the maximum core reactivity is calculated with the highest worth rod O 4.3-12 Am. No. 56,(3/81)

ACNGS-PSAM withdrawn to show at least 1.0% Ak margin. A control rod system interlock prevents the withdrawal of more than one rod while in the REFUEL code. ( ]] \U Another refueling interlock prevents the movement of the refueling platform over the core when all control rods are not fully inserted. Proposed Technical Specifications allow the bypassing of these two interlocks under procedural control: more than one control blade can be withdrawn in the refuel mode only if all four fuel assemblies in the cell surrounding every withdrawn blade are removed. These special controls result in a core k pte less than that of the fully controlled core without the procedural bypasses. The kerr of a fully controlled core is typically ~0 95 and this large margin a much uore than required for the small reactivity perturbations occurring during fuel shuffling. h.3.2 7 Stability 4.3.2 7 1 Xenon Transients Boiling water reactors do not han .nstability problems due to xenon. This has been demonstrated by opere' r BWRs for which xenon instabilities have never been observed (such instaollities would readily be detected by the LPRM's) by special tests which have been conducted on operating BWRs in an attempt to force the reactor into xenon instability, and by calculations. All of these indicators have proven that xenon transients are highly damped in a BWR due to the large negative power coefficient. Analysis and experiments conducted in this area are reported in Reference i 11. (b 4.3 d.7 2 Thermal Hydraulic Stability This subject is covered in Subsection h.h.h.6. 4.3 2.8 Vessel Irradiations The neutron fluxes at the vessel have been calculated using the one-dimensional discrete ordinates transport code described in Subsection 4.1.4.5 The discrete ordinates code was used in a distributed source mode with cylindrical geometry. The geometry described six regions from the center of the core to a point beyond the vessel. The core region was modeled as a single homogenized cylindrical region. The coolant water region between the fuel chsnnel and the shroud was described containing saturated water at 5500F and 1050 psi. The material compositions for the stainless steel in the shroud and the carbon steel in the vessel contain the mixtures by weight as specified in the ASME material specifications for ASME SA 2h0, 30hI, and ASME SA 533 grade B. In the region between the shroud and the vessel, the presence of the jet pumps was ignored. A simple diagram showing the regions, dimensions, and weight fractions are shown in Figure 4.3-28. The distributed source used for this analysis was obtained from the gross radial power description. The distributed source at any point in the core is the product of the power from the power description and the neutron (o v yield from fission. By using the neutron energy spectrum, the distributed 4.3-13 Am. No. 56,(3/81)

ACNGS-PSAR source is obtained for position and energy. The integral over position and energy is normalized to the total number of neutrons in the core region. The core region is defined as a 1 centimeter thick disc with no transverse leakage. The power in this core region is set equal to the maximum power in the axial direction. The radial and optinum axial power distributions are shown in Figures 4.3-29 and h.3-18, respectively. The neutron fluence is determined from the calculated flux by assuming that the plant is operated 90", of the time at 90% power level for 40 years or equivalent to 1 x 109 full power seconds. The calculated fluxes and fluence are shown in Table 4.3-5 The calculated neutron flux leaving the cylindrical core is shown in Table 4.3-6. h.3.3 Analytical Methods The analytien1 methods and nuclear data used to determine the nuclear characteristics are similar to those in use for design and analysis of water moderated reactors. The Lattice Physics Modell2 is used to calculate lattice reactivity characteristics, few group flux averaged cross sections and local rod-to-rod power and exposure distributions. These data are genera *,ed for various teuperature, vold, exposure and control conditions as requjred to represent the reactor core behavior. The BWR Core Simulator 3 is a large, three-dimensional code which provides for spatially varying voids, control rods, burnable poisons, xenon and exposure. This code is used to calculate three-dimensional power and exposure distributions, control rod patterns and thermal-hydraulic characteristics throughout core life. These methods have been compared extensively to experiments and plant operating data. The results are presented in Reference 5 and 6. 4.3.h Changes 4.3.h.1 Reactor Core Relative to the previous core design (documented in Reference 13) this reactor core incorporates the design changes described in the following sections. Although these features are new to the BWR/6 standard plant, they have already been incorporated into reload fuel and core designs and have been successfully demonstrated in numerous operating plants. 4.3.h.1.1 Active Core Volume Increase The active core height is increased by 2 inches. This is made possible by the lower fission gas release in the 8x8 design, which allows a reduction in the length of the fission gas plenum. From this increase in active core height, the core average power density is reduced which results in increased margins in fuel duty, thermal hydraulics and power peaking performance. Table 4.3 4 summarizes the changes in core design. O 4.3-14 Am. No. 56,(3/81)

ACNGS-PGAR i 4.3.4.1.2 Natural Uranium Utilized 4

 /N

('# ) Natural uranium is incorporated into the top and bottom 6 inches of the t lengthened active fuel region. This change increases cold shutdown margin and reduces the core average enrichment for a fixed energy production. h.3.4.1.3 Increase in Nonboiling Water Volume The number of water tubes was increased from one to two and their outside diameter has enlarged from 0.493 to 0.591 in. (Table 4.3-4) plus the size of the channel surrounding the fuel has been reduced. The larger water rods tend to reduce the maximum local power factor and decrease the amount of fissile inventory required to achieve a fixed energy production. The increase in the ratio of nonboiling to boiling water due to the channel and water rol changes results in an additional reduction in the magnitude of the void coefficient. The changes to the void coefficient, due to the two water tubes, flatten the axial power distribution which more than offsets the effect of the natural uranium on the axial power peaking. Hydraulic stability margins at natural circulation and the effects of pressurization transients are also improved by the reduction in the void coefficient. Cold shutdown mar1 6 n is improved since the large water rods increase the relative burnuo in the top of the core and decrease the hot-to-cold reactivity swing. Part of this reduction in hot-to-cold swing can be used - to increase hot excess reactivity to provide additional margin against uncertainties in nuclear predictions and to provide more power-shaping [N_,,}- flexibility early in the cycle. 4.3.4.1.4 Fuel Rod Dianeter Reduction To maintain standardization of design and fabrication, the fuel rod diameter is reduced forn 0.493 to 0.483 inches. The 10 mil reduction in fuel rod diameter was accomplished by reducing the pellet diameter by 6 mils and decreasing the cladding thickness by 2 mils. The revisions to the fuel rod design are shown in Table 4.3-4. The MLHGR is preserved at 13.4 kW/ft as a design basis; therefore, the maximum fuel centerline temperature at full power remains very nearly the saue. Although the fuel time-constant is slightly decreased as a consequence of the reduction in fuel rod diameter, analyses of core transient response have indicated the MCPR, peak fuel temperature and systeu pressure will be maintained below limits. A sumnary of fuel bundle design changes can be found in Table 4.3-4. The new fuel bundle design is also illustrated in Figure 4.3-2. 4.3.h.1 5 Prepressurized Fuel Rods Although the fuel has been propressurized to 3 atm, no effect on the nuclear response occurs. l s I sus l l l 4.3-15 Am. No. 56,(3/81)

ACNGS-PSAR h.3 5 References

1. " General Electric Company Model for Loss-of-Coolant Accident Analysis In Accordance with 10CFR50" Appendix K, Janaury 1976 (NEDO-20566).

2. E. C. Eckert, et al. , " General Electric Thermal Analysis Basis (GETAB): Data, Correlation, and Design Application", January 1977 (NEDO-10958A).

3. J. A. Woolley, "3D BWR Core Simulator", May 1976 (NEDO-20953).

h. J. F. Carew, " Process Computer Performance Evaluation Accuracy", June 1974 (NEDO-20340). l l

5 C. L. Martin, " Lattice Physics Me' '.ods Verification", August 1975 ) (NEDO-20939).  ; l

6. G. R. Parkos, "BWR Simulator Methods Verification", January 1977 (NEDO-209h6A).

7 R. C. Stirn, " Generation of Void and Doppler Reactivity Feedback For Application to BWR Plant Transient Analysis:, August 1975 l (NEDO-20964). '

8. R. C. Stirn, et al., " Rod Drop Accident Analysis For Large Boiling Water Reactors", General Electric Co. , Atomic Power Equipment .

Department, March 1972 (NEDO-10527). (Also Supplement 1, July 1972 I and Supplement 2, January 1973. ) 9 C. J. Paone, " Banked Position Withdrawal Sequence", September 1976 s (HEDo-21231).

10. K. W. Cook, " Qualification of the One-dimnnsional Core Trancient Model -

for Boiling Water Reactors", October 1973 (NEDO-2h154, Volume 1).

11. R. L. Crowther, " Xenon Considerations in Design of Boiling Water Reactors", June 1968 ( APED-56h0) . i I

12. C. L. Martin, " Lattice Physics Methods", February 1977 (NEDO-20913A). [

13 General Electric Standard Safety Analysis Report (GESSAR). f lb. R. K. Halind, Operating Strategy for Maintaining an Optimum Power f Distribution Throughout Life, Paper Presented at ANS Topical Meeting, j San Francisco, September 1963 (TID-7672). > e i 4.3-16 Am. No. 56,(3/81) ; I

ACHOS-PSAR Table 4.3-1 REACTOR CORE DIMENSIONS Active Core Height, ft (cm) 12.5 (381) 8x8 BWR/6-Lattice Control Rod Pitch, in. (cm) 12.0 (30.48) Control Rod Thickness, in. (cm) 0.328 (0.8331) Fuel Assembly Cross Section, in.2 (cm2 ) 5,h55x5,455 (13.86x13.86) . Fuel Assembly Pitch, in. (cm) 6.0 (15 24)  ! Channel I.D. , in. (cu) 5.215 (13.25) Channel Thickness, in. (cu) 0.120 (0 3048) Fuel Rod 0.D., in. (cm) 0.483 (1.227) Fuel Rod Clad Thickness, in. (cm) 0.032 (0.0813) Fuel Pellet 0.D., in. (cm) 0.410 (1.041) Humber of Fuel Rods 62 . Water Rod 0.D., in. (cm) 0 591 (1.501) Water Rod Clad Thickness, in. (cm) 0.030 (0.0762) Humber of Water Rods 2 Active Fuel Length, in. (cm) 150 (381) 2.540 cm = 1 in. l l' l 4.3-17 Am. No. 56,(3/81)

I ACHOS-PSAR l

Table h.3-2 REACTIVITY DATA FOR THE COLD, XENON-FREE STATE i f BPWS Rod G roups Condit. ion Withdrawn  % Cor. trolled BOC MOC EOC i l

        -                100           0 927      c.917    0 903 1&2                  75           0 994      0 982    0 968 a

1, 2, 3 & 4 50 1.032 1.019 1.600 1 All Rods Out 0 1.112 1.097 1.073 Highest Worth Rod Withdrawn 0 973 0 961 0.953 l h ! Highest Worth Rod Core Co-l Ord. (18,51) (18,51) (22,55) l BOC = Beginning of Cycle l I40C = liiddle of Cycle BOC = End of Cycle ! BPWS = Bank Position Withdrawal Sequence

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4.3-18 Am. No. 56,(3/81) i I

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Table 4.3-3 [ REACTIVITY AND CONTROL FRACTION FOR VARIOUS REACTOR STATES  ; 4 Beginning- Middle- End of-Cycle of-Cycle of-Cycle n Condition keff CF keff CF keff CF fI Cold, no xenon, 0 994 0.75 1.00h 0.61 1.000 0 50  ; critical *, zero i power }e

                                                                                                                                                           }

Hot, no xenon, 1.006 0 50 -1.000 0 50 1.010 0.44 , critical *, zero , power ,

                                                                                                                                                         } ,

f Hot, no xenon, 1.000 0.24 0 996 0.25 0 995 0.11  : I critical *, rated  ; power i Hot, with xenon, 0 998 0.16 0 999 0.13 0 999 0.0 . critical *, rated power- l i Cold, no xenon, 1.032 0 50 1.019 0 50 1.031 0.44 l

!                     zero power                                                                                                                           ,

1.079 0.24 1.078 0.25 1.062 0.11 i Hot, no xenon, zero power j l Hot, no xenon, 1.027 0.16 1.027 0.13 1.026 0.0 I rated power i l Hot, with 1.036 0.0 1.032 0.0 0.999 0.0  ! xenon l i

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 ,                     ' Control rod patterns adjusted approximately to critical.                                                                          j The deviations from k                          1.000 were allowed to minimize the                                                 i analyticaleffort.Th$$k=betweenconditionswiththesamecontrol                                                                        !

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I [ 5 ,sl l t i b i i i  ! 4.3-19 Am. No. 56,(3/81)  ! I

ACNGS-PSAR Table 4.3 4 [ SUfEARY OF BWR/6 DESIGil REVISIONS (GE C0!!PANY PROPRIETARY) BWR/6 Current Percent Original Design Change Core Dimensions i. Active Core Height 148 in. 150 in. +1.4 , Lattice and Puel Bundle Number of Water Rods 1 2 +100.0 Water Rod o.d. (in.) 0.493 0 591 +19 9 Vater Rod Clad Thickness (in.) 0.034 0.030 -11.8 Humber of Puel Roda 63 62 -1.6 Uraniun Content (approx) kg 188.6 182.7 -3.1 r Puel Rod Design Active Fuel Length (in.) 148 150 +1.4 Fission Gas Plenan Length (in. ) 9 48 12 -21.0 , Puel Rod o.d. (in.) 0.493 0.483 -2.0 Cladding Thickness (in. ) 0.03h 0.032 -5 9 Puel Pellet Diameter (in.) 0.416 0.410 -1.h  ; Use of Natural Uraniun in Ends

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ACNGS-PSAR Table 4.3-5 (~N CALCULATED NEU" DON FLUXES (USED TO EVALUATE VESSEL IRRADIATION) v) Flux at the Average Plux Flux at the Inside Surface Neutron Energy In the Core Core Boundary Vessel (MeV) (n/cm2 -sec) (n/cm2 -sec) (n/cm2_see)

                >3 0             1.4 x 1013       4.2 x 1012       1.1 x 109 1.0 - 3.0             3.6 x 1013       9 5 x 10 12      9 5 x 109 0.1 - 1.0             6.1 x 10 13      1 5 x 1012       1.6 x 109 10 "       (2)

Maximum Fluence 1.0 MeV = h.3 x 10 2 em Hotes:

1. The calculated flux is a maximum in the axial direction but average over the aninuthal angle.

2. The uaximum fluence is calculated using the flux and a capacity factor of 80% or 1 x 109 full power seconds. The fluence includes

([)N an azianthal peaking factor and a factor to cover analytical

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uncertainties. The azimuthal peaking factor is derived from the results of a two-dimensional analysis models the reactor bundle pattern in an r, 6 deometry. Fluxes are calculated at the cylindrical core shroud surrounding the core. The peaking factor used was 1.4. In addition to the angular peaking factor, a safety factor of h was applied to ensure that the predicted values are conservative. Y (o

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Am. No. 56,(3/81) 4.3-21

ACNGS-PSAR Table 4.3-6 CALCULATED HEUTRON FLUX AT CORE EQUIVALENT BOUNDARY Lover Energy Flux Group Bound (eV) (n/cm2 -sec) 1 10.0 x 106 a.6 x 1010 2 6.065 x 106 5 3 x 1011 3 3.679 x 106 2.0 x 1012 h 2.231 x 106 3 9 x 1012 5 1.353 x 105 4.6 x 1012 6 8.208 x 105 h.1 x 1012  ; 7 4.979 x 105 h.0 x 1012 8 3.020 x 105 2.8 x 1012 9 1.832 x 105 2.4 x 1012 10 6.738 x 104 3.4 x 1012 11 2.479 x 104 2.3 x 1012 12 9 119 x 103 2.3 x 1012 13 3.355 x 103 2.1 x 1012 i 14 1.23h x 203 2.1 x 1012 , 15 h.540 x 102 2.0 x 1012 16 1.670 x 102 2.1 x 1012 17 6.144 x 101 1 9 x 1012 18 2.260 x 101 1 9 x 1012 19 1.371 x 101 9 2 x 1012 , 20 8.315 9 2 x 1012 21 5 0h3 8.4 x 1012 l 22 3.059 8.7 x 1011 23 1.255 8.6 x 1011 24 1.125 8.5 x 1011 0.616 9 1 x 1011 25 26 0.000 3 2 x 1013 , 4.3-22 Am No. 56,(3/81) f

ACNGS-PSAR t I t J f S NOTE: LOADING PATTERN IS SHOWm FOR 1 QUARTER CORE ONLY. REFLECTIVE ~ SY MMETRY APPLIES TO REMAINDER _- 3 F F 1 F 1 2 3 1 3 1 F 2 F 2 F 2 1 3 i '

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l l 1 1 I I I i l I i l I l l 1- 1 3 5 7 9 11 13 15 17 19 21 23 25 27 29 j X, Y = PLANT COORDIN ATES 1, J = MODEL COORDINATES HOUSTON LIGHTING & POWER COMPANY Allens Creek Nuclear Generating Station g Unit 1 Banked Position Withdrawal Sequence, Am. No. 56,.(3/81) Figure 4.3-26 RPCS (Groups 5-10, Sequence B) l

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ACNGS-PSAR j o I I i 2 0 1 RE ACTOR CORE WATER 4 WATER

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3 SHROUD 5 VESSEL l t MATERIAL MATERIAL MATERIAL DENSITY RADIUS , INCHES 3 s NAME WATER 0.318 g/cm NO. , 3 00 2 2.334 gkm 1 REACTOR CORE 92.58 ZlRCONIUM 0378 okm 3 2 WATER 99 3 WATER 0.74 okm l 1 3 SHROUD 101S 304L STAINLESS STE EL FROM ASME SA 240 4 WATER 119D WATER O.74 gkm3 ' 5 VESSE L 125.0 CARBON STEEL FROM ASME SA 240 6 AIR AIR 1.3 x 10 g/cc i

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i Am. No. 56,(3/81) HOUSTON LIGHTING & POWER COMPANY ' Allens Creek Nuclear Generating Station l Unit 1 Model for One-Dimensional Transport Figure 4.3-28 Analysis of Vessal Fluence l

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                                                                                                                                                                          .                                 i uamod1viavu34uvlsu                                                   HOUSTON LIGHTING & POWER COMPANY Allens Creek Huclear Generating Station Unit 1 Radial Power Distributions                                              ;

Used in the vessel J Am. No. 56,(3/81) Figure 4.3-29 Fluence Calculation j t

I l l ACNGS-PSAR r l 4.4 THERMAL-IIYDRAULIC DESIGN j 4.h.1 Design basis ! l ! 4.4.1.1 Safety Desi;n t Bases i Thernal-hydraulic design of the core shall establish: (1) Actuation limits for the devices of the nuclear safety systems such that no fuel damage occurs as a result of moderate frequency trans,9nt events. Specifically, the !!inirma Critical Power Ratio (MCPR) operatind limit is specified such that at least 99 9% of I the fuel rods in the core are not expected to experience boiling transition during the uvst severe noderate frequency transient ' events. (2) The thertal-hydraulic safety limits for use in evaluating the safety margin relating the consequences of fuel barrier failure to public safety. l (3) That the nuclear system exhibits no inherent tendency toward , divergent or limit cycle oscillations which would compromise the 1 integrity of the fuel or nuclear system process barrier. 4.4.1.2 Power Generation Design Bases  !

O The thernal-hydraulic design of the core shall provide the following operational characteristics-(1) he ability to achieve rated core power output through-out the  !

design life of the fuel without sustaining premature fuel l failure. l (2) Flexibility to adjust core output over the range of plant load and load mneuvering requirements in a stable, predictable nanner without sustaining fuel damage. l 4.4.1 3 Requirements for Steady State conditions ! Steady-State Limits For purposes of mintaining adequate therml margin during norml i steady-state operation, the MCPR must not be less than the required MCPR l operating limit, and the MLilGR must be maintained below the design LliGR for i { the plant. This does not specify the operating power nor does it specify j peaking factors. These paraneters are determined subject to a number of constraints including the therml limits given previously. The core a. ( fuel design basis for steady-state operation (i.e., MCPR and LHGR limits ht.ve been defined to provide uargin between the steady-state operating conditions and any fuel damage condition to acccr:nnodate uncertainties and to assure that no fuel danage results even during the worst anticipated transient condition at any time in life. The design steady-state !!CPR operating limit and the peak LHGR is given in Table k.h-1. 4.4-1 Am. 56, 2/81

l 1 l , } j ACHGS-PSAR ) I k.h.l.k Requirements for Transient Conditiono Transient Linits The transient thernal limits are established such that no fuel damage is l expected to occur during the nost severe noderate frequency transient j event. Fuel danage is defined as perforation of the cladding that peruits j release of fission products. Mechanisms that cause fuel damage in reactor ! transients are: a (1) severe overheating of fuel cladding caused by inadequate cooling,

and i

l (2) fracture of the fuel cladding caused by relative expansion of the

;                                                  uranium-dioxide pellet insid2 the fuel cladding.

I l For desidn purposes, the transient linit requirenent ir net if at least j 99 9". of the fuel rods in the core do not experience boiling transition durind any t r >lerate frequency transient event. flo fuel damge would be l ! expected to occur even if a fuel rod actually experienced a boiling transition. l A value of 1", plastic strain of Zircaloy cladding in conservatively defined as the linit belov which fuel damce frou overstrainind the fuel cladding is not expected to occur. The LHGH required to cause this amount of ) c,laddind strain is approximtely 25 kW/ft in unirradiated UO2 fuel, but , decreaues with burnup to approximtely 20 LW/ft for UCp at a local exposure i of 40,000 !Nd/t. i h.h.l.S Sutmary of Design Bases In sunmry, the steady-state operating limits have been establinhed to l

assure that the design basis is satisfied for the nost severe noderate

! frequency transient event. There is no steady-state design overpower basis. An overpower which occurs during an incident of a noderate frequency transient event must neet the plant transient MCPR linit. Denonstration that the transient limits are not exceeded is nufficient to l conclude that the design basis is satisfied. The MCPR and peak LHGR limits

 !                    are sufficiently general so that no other linits need to be stated.                              For exanple, cladding surface temperatures will always be mintained within 10 to 150F of the coolant temperature as long as the boiling process is in the                                             l nucleate redine. The cladding and fuel bundle integrity criterion is                                                    )

assured as lond au MCPR and LHGR limits are net. There are no additional i decign criteria on coolant void fraction, core coolant flow-velocities, or flow distribution, nor are they needed. l The coolant flow velocities and void fraction becane constraints upon the nechanical and physics design of reactor components and are partially constrained by stability and control requirements. k.4.2 Description of Therml-Hydraulic Design of the Reactor Core b.4.2.1 Suunary Comparison 4.4-2 Am No. 56,(3/81)

ACNGS-PSAR An evaluation of plant performance from a thermal and hydraulic standpoint m)/ I is provided in Subsection 4.h.3. A tabulation of thernal and hydraulic parameters of the core' is given Table  ; 4.4-1, which give comparison of this reactor n th others of similar design. h.4.2.2 Critical Power Hatic There are three different types of boiling heat transfer to vater in a forced convection systen: nucleate boiling, transition boiling, and filn boiling. 13ucleate boiling, at lower heat transfer rates, is an extremely , efficient node of heat transfer, allowing large quantities of heat to be transferred with a very small temperature rise at the heated vall. As heat transfer rate is increased, the boiling heat transfer surface alternates between film und nucleate boiling, leading to fluctuations in heated wall teuperatures. The point of departure frou the nucleate boiling region into the transition boiling region is called te boiling transition. Transition boiling begins at the critical power and is characterized by fluctuations in cladding surfr.ce temperature. Film boiling occurs at the highest heat transfer rates; it begins as transition boiling comes to an end. Film boiling heat transfer is characterized by stable vall temperatures which are higher than those experienced during nucleate boiling. , k.4.2.2.1 Boiling Correlations The occurrence of boiling transition is a function of the local steam ( quality, boilind length, nass flow rate, pressure, flow geometry, and local y ) peaking pattern. Ganeral Electric has conducted extenc've experimental. U investigations of these parameters. These parametric studies encompass the entire design range of these variables. In the experimental investigations, a boiling transition event was associated with a 25oF rise in rod surfact tenperature. The (critical) quality at which boiling transition occurs as a function of the distance from the equlibrium boiling boundary is predicted by the GEXL (General Electric Critical Quality, X sub , (c) -Boiling Length) Correlation. This correlation is based on accurate test data of full-scale prototypic simulations of reactor fuel assemblies operating under conditions duplicating those in actual reactor designs. t The GEXL correlation is a "best fit" to the data and is used together with a statistical analysis to assure adequate reactor thermal margins l. The figure of mer it used for reactor design and operation is the Critical Power Ratio (CP: i. This is defined as the ratio of the bundle power which would produce equilibrium quality equal to, but not exceeding, the correlation value (critical quality), to the bundle power at the reactor L condition of interest (i.e., the ratio of critical bundle power to operating bundle power). In this definition, the critical power is deteruined at the same mass flux, inlet temperature, and pressure which exist at the specified reactor condition. h.h.2.3 Linear Heat Generation Rate (LHGR) , L The limiting constraints in the design of the reactor core are stated in terus of the LHGR limit and the MCPR operating limit for the plant. The g design philosophy used to assure that these limits are net involves the , l 4.4-3 Am. No. 56,(3/01)

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

1 i l i ACNGS-PSAR i I l l selection of one or more distributions which are nore limiting than l expecte1 operating conditions and subsequent verification that under these more stringent conditions, the design limits are net. Therefore, the

j " design power distribution" in an extreme condition of power. It is a fair ' and stringent test of the operability of the reactor as designed to comply with the forn;oing linits. Expected operating conditions are less severe than those represented by a design power distribution which gives the ntximun allowable idIGd :tud the MCPR operating limits for the plant.

                   !!oweve r , it nust be established that operation with a less nevere power distribution is not a necessary condition for the safety of the reactor.

Because there are an infinite number of operating reactor states which can exist (with variations in ro1 patterns, time in cycle, power level, distribution, flow, etc.) which are withir the design constraints, it is not possible to determine them all. However, constant monitoring of operating conditions using the available plant measurements can ensure coupliance wit h design objectives. The core average and fiLHGR are given Table h.h-l. h.h.2.3.1 Deuign Power Distribution Thernal desi t;n of the reactor - includint; the selection of the core size and effective heat transfer area, the design steam quality, the total recirculttion flow, the inlet subcooling and the specification of internal flow distribution - is based on the concept and application of a design power diutribution. The design power distribution is an appropriately connervative repreaentation of the most limiting thurnal operating state at rated conditionn and includes design allowances for the combined effects (on the fuel rod, and the fuel assembly heat flux and temperature) of the ' gross and local steady-state power density distributions and adjustments of the control rods. The deaign power distribution is used on conjunction with flow and pressure drop distribution couputations to determine the thermal conditions of the fuel and the enthalpy conditions of the coolant throughout the core. The design axial power distribution used in the calculation of the MCPR operating limit is given in Table 4.4-2. This distribution is consistent with that discussed in Reference 1. The deuign power distribution is based on detailed calculations of the neutron flux distribution as discussed in Appendix A of Gection b.a. 4.4.2.3.2 Design Linear Heat Generation Rates (LHGR) The maximun and core average linear heat generation rates are shown in l Table 4.4-1. The naximum linear heat generation at any locatlon is the product of the average linear heat generation rate at that location and the total peaking j factor. h.h.2.h Void Fraction Distribution The core average and maximuu exit void fractions in the core at rated condition are given in Table 4.4-1. The axial distribution of core void 4.4-4 Am, No. 56,(3/81)

ACNGS-PSAR fractions for the average radial channel and the maximum radial channel 9 (end of node value) for the core are given in Table h.h-3. The core average and naximum exit value is also provided. steam quality are provided in Table 4.h-5 The Gimilar distributiens for core average axial power distribution used to produce these tables is given in Table h.4-4. h.h.2 5 Core Coolant Flow Distribution and Orificing Pattern l t Correct distribution of core coolant flow anong the fuel assemblies is l j accomplished by the use of an accurntely calibrated fixed orifice at the inlet of each fuel assembly. The orifices are located in the fuel support piece. They .:erve to control the flow distribution and, hence, the coolant conditons within prescribed bounds throughout the design range of core operation. The sizing and design of the orifices ensure utable flow in each fuel asumbly during all phases of operation at normal operating 7 conditions. , t l I The core is .livided into two orificed flow zones. The outer zone is a ! narrow, reduced-power region around the periphery of the core. The inner l zone consists of the core center region. No other control of flow and . steau distribution, other than that incidentally supplied by adjusting the l

power distribution with the con' trol rods, is used or needed. The orifices can be changed during refueling, if necessary.

Design core flow distribution calculations are made using the design power i distribution which consists of a hot and average powered assembly in each f 9 of the two orifice zones. The design bundle power and resulting relative flow distribution are given in Table 4.h-6. The flow distribution to the fuel assemblies is calculated on the l asuuuption that the pressure drop across all fuel assemblies is the same. This assumption has been confirmed by measuring the flow distribution in a modern boiling water reactor as reported in Refer ence 2. There is reasonable assurance, therefore, that the calculated flow distribution throughout the core is in close agreement with the actual flow distribution of an operating reactor. The use of the design power distribution discussed previously ensures the

orificing chosen covers the range of normal operation. The expected shifts in power production during core life are less severe and are bounded by the design power distribution.

h.h.2.6 Core Pressure Drop and Hydraulic Loads The pressure drop across various core components under the steady-state design conditions is included in Table h.h-l. Analyses for the most ! limiting conditions, the recirculation line break and the steam line break, I are reported in Chapter 3 The components of bundle pressure drop considered are friction, local 9 elevation and acceleration. Core plate prascare drop measurements have been taken on several operating BWR/3 and 4 plants containing 7x7, 8x8, and 4.4-5 Am. No. 56,(3/81)

ACNGS-PbAR nixtureu of 7x7 and 8x8 fue' . Table 4.h-7 compares measured and calculated core plate pressure dropn. Phe measured and calculated valuen are in Good agreement. The data in predicted with an avernde error of 0.0h pai. The one signa error is 0.86 pai. The therml-hydraulic loads on the fuel rods during the steady-state operation, transient and accident conditionn are negligible, primarily because of the channel confinement, thereby ranulting in sntll cross flow betw2en roda (i.e., essentially constant preu :ure at any given elevation in the fuel bundle). rhe londo (i.e., horizontal) across the control bladen are niinimal or ned l egible primarily due to the flat interchannel velocity profile an given in Hererence 3. 4.4.2.6.1 Friction Pre.nur Drop

Friction preocure . trop in eniculated using the model relation: 3p = W'- fL 2 3 2P d D. A2 d ch TPF where l AP p = friction precoure drop (pui); i W = maus flow rate g = neccleration of gravity; P = Water density; D = channel hydraulic diameter; g ! A = channel flow area; ch L = lendth; l f = friction factor; and V TPF = two-phane friction multiplier This basic nodel is sinilar to that used throughout the nuclear power in du st ry. The foruulation for the two-phase multiplier is baned on data which coupare closely to that found in the open literature b. General Electric Company has taken significant amounts of friction pressure drop data in multirod geometries representative of modern BWR plant fuel bundles and correlated both the friction factor and two-phase multipliers on a best-fit basis using the above pressure drop Iormulation. Checks against more recent data are being made on a continuind l 'tG iB to ensure the beat models are used over the full range of interest to boiling water reactors. I 4.4-6 A m. No. 56, (:? /81)

I l l ACNGS-PSAR I k.4.2.6.2 Local Pressure Drop l , The local pressure drop is defined as the irreversible pressure loss associated with an area change such as the orifice, lower tieplatee, and  ! spacers of a fuel assemblf. l l The general local pressure drap nodel is similar to the friction pressure l drop .tnd is , ! dPg = E_ d2

2gn A 2 TPL l l l' \ } vhere l t s SPg = local pres: tre drop (psi); l K = local pres are drop loss coefficient;  ; I i A = reference area for local loss coefficient; and j 1 , p2TPL = two-phase local multiplier and v, g and p are defined i the same as for friction. This basic model is similar to that used l j throudhout the nuclear power industry. Le formulation for the two-phase J multiplier is si.milar to that reported in the open leteratureS with the l addition of enpirical constants to adjust the results to fit data taken by 1 j General Electric Company for the specific designs of the BWR fuel assembly. l Tests are performed in single-phase water to calibrate the orifice in the l } lover tieplate, and in both single- and two-phase flow to arrive at l best-fit design values for spacer and upper tieplate pressure drop. The range of test variables is specified to include the range of interest to boiling vater reactors. New data are taken whenever there is a significant design change, to ensure the uost applicable methods are in use at all times. h.4.2.6.3 Elevation Pressure Drop The elevation Iressure drop is based on the well known relation: Py

                                   =

gggL l i

                                   =

p(18)+QA p EdL l vhere AP E = elevation pressure drop (psi); ob = incremental length; fT= average water density; A= averade void fraction over the length L; P re Pg = saturated water and vapor density, respectively; and g = acceleracion of gravity. 4.4-7 Am. No. 56,(3/81)

                                            - .. ,                                                 _             . - - .___ 0

l 4 1 l ACNGS-PSAR I I 4.4.a.6.4 Acceleration Pressure Drop A reversible pressure change occurs when an area change is encounterea, and an irreversibl 2 loss occurs when the fluid is accelerated through the boilind process. The basic foruulation for the reversible pressure change resultint; fron a f'>w area change is d i ven by: o (1 - P) A,, 3p = -

                                                                                 " 2,4=
                                                                                      '       3.

ACC ,, 2 A

                                                                            % Pa,, ,3          1 where
                          'AP     g a       acceleration pressure drop; Aj      =        final flos area; an1 A,      =        initial flow area and other teri:a are as previously defined. The basic fornulation for the acceleration pressure change due to density change is:

3 L*W, P . , L, u w' 1 1 P GA$h II. 3UT "M. I U where 1 = X

                                                  +        (1-X)       -.- . , .

p M P a ( 1_ ,) t~ p = uomentuu density ; and g i i l x = stean quality and other terus are as previously defined. The total acceleration pressure l drop in BWHs in on the order of a few percent of the total pressure drop. I h.h.2 7 Correlation and Physical Data General Electric has obtained substantial amounts of physical data in support of the pressure drop and thernal-hydraulic loads discussed in Subsection 4.4.2.6. Correlations have been developed to fit these data to the foruulations discussed. 4.4.2.7 1 Pressure Drop Correlations General Electric Company has taken significant amounts of friction preocure drop data in multirod geonetries representative of modern BWR plant fuel , bundles and correlated both the friction factor and two-phase multipliers ) on a best-fit basis using the pressure drop formulations reported in l Gubsections h.4.2.6.1 and h.h.2.6.2. Checks ai;ainst more recent data are being made on a continuous basis to ensure the best nodels are used over the full range of interest to BWRs. 4.4-8 Am. No. 56,(2/81)

ACNGS-PSAR est" are perforued in ningle-phase water to calibrate the orifice and the G low r tieplate, and in both single- and two-phase flow to arrive at best-fit design values for saacer and upper tieplate pressure drop. The ! rant;c of test variables is pecified to include the range of interest to boiling vater reactors. New data are taken whenever there is a significant design change to ensure the most applicable methods are in use at all times. Applicability to the sind e-phase and two-phase hydraulic models discussed l l in Subsections 4.4.d.6.1 and 4.h.2.6.2 is confirmed by prototype (64-rod i bundle) flow tests. The typical range of the test data is cummarized in  ! ! Table 4.4-8. [ 4.4.2 7 2 Void Fraction Correlation The void fraction correlation used is a version of the Zuber-Findlay nodelll where the concentration paraneter and void drift coefficient are i based on coupurison with a large quantity of world-wide data13-24, j l i l 4.k.2 7 3 Heat Transfer Correlation l The Jens-Lotteu6 wall :uperheat equation is uned in fuel desin to determine  ! the cladding-to-coolant hea'- transfer coefficients for nucleate boiling. l i I h.h.2.8 Thernal Effects of Operational Transients l The evaluation of the core's capability to withstand the thermal effects resultind frou anticipated operational trannients is covered in Chapter 15 l9 (Accident Analyses). i l h.4.2 9 Uncertainties in Estimates l Uncertainties in thermal-hydraulic parameters are considered in the l stat 4 t'.:al analysis which is performed to establish the fuel cladding l ir ty safety limit such that at least 99 9% of the fuel rods in the  ! L r expected not to experience boiling transition during any moderate j fre gency transient event. The statistical nodel and analytical proceduce i are described in detail in Reference 1. The conservative power { distribution used for the statistical analysis is shown in Figure 4.k-8 in l terms of relative bundle power histogram. The uncertainties considered and l their input values for the analysis are shown in Table 4.h-9 l , 1 4.4.2.10 Flux Tilt Considerations l l For flux ;ilt considerations, refer to Subsection 4.3 2.2.T. h.4.3 Description of the Thermal and Hydraulic Design of he f a Re actor Coolant System i a The thernal and hydraulic desi6n of ;he reactor coolant system is described l , in this section. j 4.4.33 Plant Configuration Data 1 l l l 4.4-9 Am. No. 56,(3/81)  ; I

l ACHGS-PSAR h.4.3.1.1 Heactor Coolant Systen Configuration The reactor coolant systen is described in Section S.S and shown in ison tric perspective in Figure 5 5-1. The piping uizes and valves are listed in Table 5.S-1. 4.h.3.1.2 Heuctor Cool tnt Systen Therrxtl H: 3raulic Data l The stea ly-st, te distribution of temperature, preuuure and flow rate for j each flow pat't in the rt actor coolant systen is shown in Figure 5 1-1. I 4.h.3.1.3 Heuctor Coolant Systen Geonetric Data l Fluid Voluues of regions and components within the reactor vessel are shown in Vi dure 5.1 . Tab.e 4.4-11 govides the flow path length, height, liquid level, nininun elevations, nn 1 nininun flow areas for each rnjor flow path volute within the reactor vessel and recirculation loops of the reactor coolant systens. Table 4.4-12 provides the lengths and sises of all safety injection lines to the reactor coohnt systen. h.4.3.2 Operating Heutrictions on Puups Expected recirculation punp perfor:nnce curven are shown in Figure 5 5-17 These curves are va11.- for all conditions with a nornal operating range varyind from apprc xisttely 20% to 115% of rated puup flow. The pump characteristics, includind considerations of HP3H requirements, are the cane for the conditions of two-punp and one-punp operation as describei in Subsection S.S.1.3 Subsection h.h.3.3 gives the operating limits iuposed on the recirculation pumps by cavitation, punp loads, bearind de ign flow starvation, and punp speed. h.h.3.3 Power-Flow Operating Map I 4.h.3.3.1 Linits for Hornal Operation A BWH nust operate with certain restrictions because of pump net positive suction head (HP311), overall plant control characteristics, core thernal power linits, etc. The power-flow map for the power range of operation is shown in Figure 4.h-S. The nuclear systen equipment, nuclear instrunentation, and the reactor protection systen, in conjunction with operatind procedures, maintain operations within the area of this nap i , nornal operating conditions. The boundaries on this nap are as follows: Natural Circulation Line, A: The operating state of the reactor noves alond this line for the normal control rod withdrawal sequence in the absence of recirculation pump operation. 10S$ Steam Flow Hod Line or Hated Power (Whichever in Less): The 105% stean flow rod line passts through 104.2% power at 100% flow. The operatind state for the reactor follows this rod line (or similar l 4.4-10 Am. No. 56,(3/81) 1

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

ACliCS-PSAR l onen) during recirculation flcv changes with a fixed control rod patt.ern; however, rated power may not be xceeded. 105% steam flov j G rod line is based on a constant xenon concentrction at 104.2% power and rated flow. l Cavitation Protection L!ne: This line reaults from the recirculation f puup, flow control valv.: and jet pump IlPSH requirements.

l l 4.4.3 3 1.1 Performance Characteristics l I l 5 Other performance characteristics shown on the power-flow operating map I

j are: ' l' Constant Rod Lines: These lines show the change in power associated  : with flow changes, while mintaining constant control rod position. j Constant Position Lines for Flow Control Valve, B, C, D, and F: These i i l lines show the change in flow associated with power changes while ' maintaining flow-control valves at a constant position. i a j 4.4.3.3.P Regions of the Power Flow Map l l Nedion I - This region defines the system operational capability j with the recirculation pumps and mators being driven by I the low frequency motor-ganerator set at 25% speed. l Flow is controlled by the flow control valve and power { changes, during normal startup and shutdown, will be in this region. The normal operating procedure is to l start up along curve C - FCV vide open at 25% speed. 4 Redi on 11 - This region shows the area where 25% pump speed and j ] 100$ punp speed operating regimes overlap. The j

 ,                                                                                  switching sequence from the low frequency m-g set to I                                                                                     100% speed will be done in this region.                                 ;

2 Region III - This is the low power area of the operating map where i

 ;                                                                                  cavitation can be expected in the recirculation punps,

< jet pumps, or flow control valves. Operation within

this region is precluded by system interlocks which i trip the main motor from the 100% speed power source to I the 25% speed p% er source.

[ Hegien IV - This represents the normal operating zone of the nap j vhere power changes can be mde, by either control rod movement or by core flow changes, through use of the l flow control valves. ! 4.4.3 3.3 Design Features for Power-Flow Control l l The following limits and design features are employed to maintain

power-flow conditions to the required values shown in Figure 4.4-5 (1) Minimum Power Limits at Intermediate and High Core Flows: 9 To prevent cavitation in the recirculation pumps, jet pumps, and 4.4-11 Am. No. 56,(3/81) l L_-_____________ _ _ _ _ _ _ _ . _ _ , . _ , __ _. _ . _ _._ ___

!                                                       ACllGS-PSAR

! flow control valves, the recirculation systen is provided with an l interlock to trip off the 100% speed power source and close the ! 25% speed power source if the difference be* ween steanline ! ten c erature and recirculation pump inlet teuperature is less than I a preset va'ae (9.80f). This differential temperature is neacared using high accuracy RTDs with a sensing error of less l that. 0.20F at the two standard devi.ttion (2 ) confidence level. l This action is initiated electronically throu6h a 15-sec tine I delaj. The interlock is active while in both the autouatic and I rnnuitl operation modes. f (2) llininun Power Linit at Low Core Flow: During low power, low loop flow operation, the tenperature dif ferential interlock nay not provide sufficient cavitation protution to the flow control l

valves. Tnerefore, the systen is provided with an interlock to trip off the 100% speed power source and , ' ace the 25% speed

 ;                    power source if the feedwater flow falls below a preset level                                         r (22f of rate 1) and the flow control valves are below a preset position (20% open). The feedwater flow rate and recirculation I

flow control valve position are neasured by existing process control instrunents. The speed chan6e action is electronically initiated. This nterlock is active during both automatic and nanual nodes of operation. l (3) Puup Bearing Liuit: For punps as 1 trge as the recirculation puups, practical 'i-it s of pump bearing design require that uininun pump flow be linited to 20% of rated. To assure this nininuu flow, the systen is designed so that the minimun flow control valve position will allow this rate of flov. ! (4) Valve Position: To prevent structur2 or cavitation damage to the recirculation puup due to pump suction flow starvation, the ' l systen is provided with an interlock to prevent starting the j pumps, or to trip the punps if the suction or discharge block ! valves are at less than 90% open position. This circuit is activated by a position limit switch and is active before the puup is started, during nanual operation node, and during autouatic operation node. 4.h.3.3.3.1 Flow Control The principal modes of normal operation with valve flow control-low frequency notor Generator (LFt1G) set are sunrarized as follows: the recirculation punps are started on the 100% speed power source in order to unseat the puup bearings. Suction and discharge block valves are full open and the flow control valve is in the minimum position. When the punp is near full speed, the nain p>wer source is tripped and the pump allowed to coast down to approximately 25% speed, where the LFMG set will power the punp and notor. The flow control valve is then opened to the maxiaun position, at which point reactor heatup t.nd pressurization can connence. l When operatind pressure has been established, reactor power can be l increased. This power-flow increase will follow a line within Region I of j the flow control nap shown in Figure k.h-5 l O;1 ' 4.4-12 An. No. 56,(3/81) l

ACNGS-PSAR i t Uhen reactor power is (;reater than approximtely 20-28% of rated, the low G feedwater flow interlock is cleared and the min recirculation pumps can be  ; switched to the 100% speed power source. The flow control valve is closed  ; f to the ninimuu position before the speed change to prevent large increases l l in core power and potential flux scran. This operation occura within ! He6 i on II of the operating map. The system is then brought to the desired l power-flow level within the norna. ,oerating area of the mp (!h gion IV) by l opening the flow contr71 valves and u,, withdrawing control coda. i I Control rod withdrawal with constant flow control valve position will f result in power / flow thang s along lines of constant c sub (v) (cnnstant j j position). Flow contcol s .ive movement with conutant control rod position will renult in power, flow changes along, or nearly parallel to, the rated l clow contrci line. i 4.4.3.h Temperature-Power Operating Map (PWR) E i Not applicable.

f 4.h.3.5 Load-Following Characteriutics i Larde nedative operating reactivity co fficients inherent in the INR I i provide the followind important advancages: ) j (1) good load-following with well-damped behavior and little  ! ] undershoot in the heat transfer response; (2) load-following with recirculation flow control; and

(3) strong damping of spatial power disturbances. l

-                                                                                                                                                                                  1

Design of the BWR includes the ability to follow load demands over a i

;

reasonable range without requiring operator action. Henctor power can be j i controlled automatically by flow control over approximately a 255 power j rangc at, for example, approximtely 1% per second for a 10% step-load  ;

!             change.                                                                                                                                                               l i

j 4.4.3.6 Therual and 1(ydraulic Characteristics Summary Table l' The therml-hydraulic characteristics are provided in Table 4.4-1 for the core and tables of Section 5 5 and other portions of the reactor coolant j j system. l 4.4.4 Evaluation i The design basis employed for the therual and hydraulic characteristics j incorporated in the core design, in conjunction with the plant equipment ! characteristics, nuclear instrumentation, and the reactor protection [ system, is to require that no fuel damage occur during normal operation or du/ing abnormal operational transients. DemonsLration that the applicable therml-hydraulic limits are not exceeded is given by analyses. 4.h.4.1 Critical Power 9 4.4-13 Am. no. 56,(3/81)

_ . _ _ - - . - . . . _ . ~ . . . 1 1 l ACI!GS-PSAR 4 j The GEXL critical power correlation is utilized in thernal-hydraulic I evaluations. This correlation is discussed in more detail in Subsection j h.h.2.2.1. 4.4.4.2 Core Hydrau'ics Core hydraulic nodels and correlations are dincussed in Subsections

b.4.2.6, 4.4.2 7, and h.4.h.S.

l I 4.h.4.3 Influence of Power Distributions The influence of power distributions on the thernal-hydraulic design is liscussed in Hererence 1, Aapendix V. 4.h.4.h Cor- Thermal Response Jhe thert.ul response of the core for accidents and expected transient conditions is discunced in Chapter 15 ( Accident Analyses). l 4.4.4.5 Analytical t4ethods Jhe analytical methods, theruodynanic data, and hydrodyr.anic data used in deteruinind the ther..nl and hydraulic characteristics of the core are sinilar to those uurd throughout the nuclear power indut try. Core thert al-hydrwilic analyses are perforned with the aid of a digital couputer prograu. This progran nodels the reactor core throudh a hydraulic description of orifices, lower t,ieplates, fuel rods, fuel rod spacers, upper tie plates, fuel channel, and the core bypass flow paths. 4.4.h.S.1 Reactor Model The orifice, lower tieplate, fuel rod spacers, and upper tieplate ar" hydraulically represented as being separate, distinct local losses of zero thickneau. The fuel channel cross section is represented by a square section with enclosed area equal to the unrodded cross-sectional area of the acti ' lel channel. The fuel channel assembly consists of three ba31c axial regions. The first and nost inportant is the active fuel ed i on which consists of the 62 fuel rods, 2 nonfueled rods, and T fuel rod spacera. The second is the nonfueled region consisting of 64 nonfueled rods and the upper tieplate. The third region represents the unrodded portion of the fuel channel above the upper tieplate. "'he active fuel redi on is considered in 2h independent axial seguents cr nodes over which fuel thermal properties are as aned constant and coolant properties are assuned to vary linearly. The code can handle 12 fuel channel types and 10 types of bypass flow paths. In normal analyses, the fuel assemblies are modeled by b channel ' types - a " hot" central orifice region channel type, an average central orifice region channel type, a " hut" peripheral orifice renion type and an average peripheral orifice region type. Usually, there ia one fuel assenbly representind each of the " hot" types. The average types then nahe l up the balance of the core. O l 4,4_14 Am. No. 56,(3/81) l l l

ACMGS-PSAR 4 ! The coupater prodran iterates on flow through each flaw path (fuel l usseublies and bypass paths) until the total differential pressure (plenun ! to plenun) across each path is equal, and the sum of the flows throudh each { path equals the total core flow.

;

Orificing in selected to optimize the core flav distributior:u between j orifice ced ons i as discusse.1 in Subsection h.h.2 5 The core design j pressure is deteruined fron the required turbine throttle pressure, the

steamline pressure drop, staan dryer pressure drop, and the steam separator i pressure drop. 'Che core inlet enthalpy is determined frou the reactor and l ' carbine heat balances. The required core flow is then determined by l

upplying the procedures of this section and specifications nuch that the thernal linits of Reference i are satisfied and the noninal expected bypass j flow fraction is approximately 10%. The results of applying these nethods and spec trications are: (1) flow for each bundle type; i l (2) flow for each bypass path; i 1 ! (3) core presuare urop; I l (4) fluid property axial distribution for each bundle type; and { ' (S) CPR calculations for each bundle type. 4.4.k.5.2 Systen Flow Balances l O The lasic assumption used by the code in performing the hydraulic analysis is that the flow entering the core vill divide itself between the fuel l bundles and the bypass flow paths such that each assembly and bypass flow path experience the same pressure drop. The bypass flow paths considered are described in Table 4.4-10 and shown in Figure h.4-1. Due to the large flow area, the presst.re drop in t'.e bypass region cbove the core plate is essentially all elevation head. Thus, the sum of the core plate differential pressure and the bypass region elevation head is equal to the core differential pressure.  ! l l I j The total core flow, less the control rod cooling flow, enters the lower l 4 plenum throudh the jet pumps. A fraction of this passes through the j

various bypass paths. The remainder passes through the orifice in the fuel J t support (experiencind a pressure loss) where more flov is lost through the fit-up between the fuel support and the lower tieplate and also through the lower tieplate holes into the bypass region. The majority of the flow continues through the lower tieplate (experiencing a pressure loss) where sone flow is lost through the flow path defined by the fuel channel and lower tieplate, and restricted by the finger springs, into the bypass region. The flow through the bypass flow paths is expressed by the form: W-C 1 APl/2 + C 2 AP C +C 3 O 4 4-15 Am. No. 56,(3/81)

ACllGS-PGAR Full-acale tests have been perforned to establish the flow coefficients for the mjor flow paths 12 These tests simulate actual plant configurations which have neveral par.tllel flow paths and, therefore, the flow coefficients for the individual paths could nat be separated. !Iow eve r , analytical models of the individual flow path: were developed an an independent check of the tests. The modela were derivei for actual BWR j design dinonsions :tnd cons Hered the effects of dinensional variations. l Theu" aode lo predic ted the test results when the an-built dinennions were l applied. When using thene todeln fo r hy d rau '. i o design calculations, noui :al drawind dimensions are used. This ic done to yield the most accurate prediction of the expected bypass flow. With the large number of couponents in a typic:tl BWR core, deviatiens from the nominal dinensionn wili tend to statintically cancel, resulting in a total bypass flow best represented by that calculated usind nominal line nc ion n. The balance of the flow entera the fuel bundle fron the lowe r ' ieplate and pauues throt the fuel rod channel spaces. A nrn11 portion of the in-channel flow enters the nonfueled rods through three orifice holen in i i cach rod oust abovt th2 lower tieplate. Thir flow, normally referred to as  ; the water rod JLow. renixen with the active coolant channel flow below the upper tieplate. The uncertaintien in calculationa and the resultant uncertai nty in reactor coolant systen flow ute are provided in Table 4.h-9 i h.4.h.$.3 Synten Heat Balances Within the fuel annenbly, heat balances on the active coolant are perforued n od. tlly . Fluid propertien, ex} *e: aed as th, hun:lle average at the g rticular node of intereut, a 2 baned on Reference 7 In evaluatind fluid i properties, a constant preunure nodel is used.

 'Jhe core power is divided into two parts                                                               an active coolant power an'                                               ;

bypass flow poser. The bypass flow is heated by neutron-slowing down gauna heating in the water, and by heat trannfer through the channel k uls. lleat in also transferrred to the bypass flow fron structures and control elenentn which are, thennelves, heated by gunua absorption and by (7,a) reaction in the control material. The fraction of total reactor power deposited in the bypass redian is very nearly 2%. A uinilar phenonenon l occurs, with the fuel bundle, to the active coolant and the water rod flows. The net effect in that 96% of the core power ir conducted throudh the fuel claddind and appears as heat flux. I In design analyseu, the power is allocated to the individual fuel bundles usind a relative power factor. The power distribution along the length of the fuel bundle in specified with axial power factors which distribute the i bundle's power among the 2h axial nodes. A nodal l a t peaking factor is I used to establish the peak heat flux at cauh nodal location. l 1 I The relative ( radial) and axial power distributions when use.4 with .he i buadle flow deteruine the axial coolant property distribution resulting in ( cufficient information to calculate the pressure drop components within ) each fuel ascenbly type. Once the equal preucure drop criterion has been  ; natisfied, the critical bundle power (the power which would result in i critical quality existind at some point in the bundle using the correlation l i i 4.4 16 Am. No. 56,(3/81)  ! t

ACNGS-PSAR l expressed in Reference 9) is determined by an iterative process for each fuel type.

                       }
           \         j    In applying the above methods to core design, the number of bundles (for a specified core thermal power) and bundle geometry (8x8, rod diameter, etc. )

are selected based on power density and LHGR limits. 4.4.4.6 Thermal-Hydraulic Stability Analysis l l 4.4.4.6.1 Introduction i There are many definitions of stability, but for feedback processes and control systems it can be defined as follows: a system is stable if, following a disturbance, the transient settles to a steady, noneyclic state. A system uay also be acceptably safe even if oscillatory, provided that any limit cycle of the oscillations is less than a prescribed magnitude. j Instability, then, is either a continual departure from a final ! steady-state value or a greater-than-prescribed limit cycle about the final uteady-state value.

l The nechanism for instability can be explained in terms of frequency response. Consider a sinus aidal input to a feedback control system which, for the moment, has the feedback disconnected. If there were no time lags ! or delays between input and output, the output would be in phase with the f input. Connecting the output so as to subtract from the input (negative I

                        ) feedback or 180 degrees out-af-phase connection) would result in stable

j closed loop operation. Hove rer, natural laws can cause phase shift between output and input and, should the phase shift reach 180 degrees, the e feedback signal vould be reinforcing the input signal rather than i subtracting from it. If the feedback signal were equal to or larger than j the input signal (loop gain equal to one or greater), the input signal

could be disconnected and the system would continue to oscillate. If the l feedback signal were less than the in';ut signal (loop gains less than one),

the oscillations would die out. l It is possible for an unstab1:: process to be stabilized b, ,. ding a control

system. In general, novever, it is preferable that a process with inherent f feedback be designed to be stable by itself before it is combined with i other processes and control systems.

The design of the BWR is based on this premise, that individual system f components are stable. 4.4.4.6.2 Description Three types of stability considered in the design of boiling water reactor  ! are: (1) reactor core (reactivity) stability; (2) channel h"drodynamic  ! stability; and (3) total system stability. Reactivity feed %ck instability i of the reactor core could drive the reactor into power osefilations. l Hydrodynamic channel instability could impede heat transfr r to the  ; moderator and drive the reactor into power oscillations. The total system } j stability considers control system dynamics combined viti basic process J l l

                                                                                                                    ;

i i 4.4-17 Am. No. 56,(3/81) i

ACMGG-PSAR dynamics. A stab 12 systen is analytically demonstrated if nc inherent limit cycle or diverdent oscillation develops within the systen as a result of calculated step disturbances of any critical variable, such as steam flow, pressure, neutron flux, and recirculation flow. The criteria to be considered are ntated in terms of two compatible parameters. First is the decr,y ration x2/xo, designated as the ratio of the magnitude of the second overshoot to the first overshoot resulting from a step perturbation. This etaracteristic provides a graphic representation of the physical responsivenes of the system which is readily evaluated in a time-domain analysis. Secor.d is the damping coefficient &n, the definition of which correspoads to the pole pair closest to the J axis in the s-plane for the systeu closed loop transfer function. This parameter also applies to the frequency-d > main interpretation. The damping coefficient is related to the dacay ratio as shown in Figire 4.4-2. 4.4.4.6.3 Stability Criteria The assurance that the total plant is stable and, therefore, has significant safety margin shall be demonstrated analytically when the decay ratio, x2 xo,

         / is less than 1.0 or, equivalently, when the dampind coefficient, (n, is greater than zero for each type of stability discussed.

Special attention is given to differentiate between inherent system limit cycles and small, acceptable limit cycles that are always present, even in the nost stable reactors. The latter are caused by physical nonlinearities (deadband, striction, etc.) in real control system and are not representative of inherent hydrodynamic or reactivity instabilities in the reactor. The ultimate performance limit criteria for the three types of dynanic performance are summarized below in terms of decay ratio and damping coefficient: Channel hydrodynamic stability x2 /*0 #1' I n >0 Reactor core (reactivity) stability 0 x2I*0 '

                                                          #n Total system stability x2 /*0   '

n These criteria shall be satisfied for all attainable conditions of the reactor that may be encountered in the course of plant operation. For stability purposes, the most severe condition to which these criteria vill be applied correspond to the highest attainable rod line intersection with natural circulation flow. 4.4.4.6.4 Mathematical Model The mathematical model representirg the core examines the linearized reactivity response of a reactor sy stem with density-dependent reactivity feedback caused by boiling. In addition, the hydro-dynamics of various hydraulically coupled reacotr channels, or regions, are examined separately on an axially uultinoded basis by drouping various channels that are thermodynamically and hydraulically similar. This interchannel hydrodynamic interaction, or coupling, exists through pressure variations in the inlet plenum, such as can be caused by disturbances in the flow distribution between regions or channels. This approach provides a 4.4-18 Am. No. 56,(3/81)

ACNGS-PSAR reasonable accurate, three-dimensional representation of the reactor's i hydrodynamics. \ ) The core model 2 5-30, shown in block diagram foru in Figure 4.h-3, solves the dynamic equations that represent the reactor core in the frequency douain. From the solution of these dynamic equations, the reactivity and individual channel hydrodynamic stability of the boiling water reactor is deteruined for a given reactor flow rate, power distribution, and total power. This gives the most basic understanding of the inherent core behavior (and hence the system behavior) and is the principal consideration in evaluating the stable performance of the reactor. As new experimental or reactor operating data are obtained, the model is refined to improve its capability and accuracy. The plant model considers the ca. tire reactor system, neutronics, heat transfer, hydraulics, and the basic processes, as well as associated control sy.'tems such as the flow controller, pressure regulator, feedvater controller. etc. Although the control systems may be stable when analyzed individunly, final control system settings must be made in conjunction with the operating reactor so that the entire system is stable. The plant uodel yields results that are essentially equivalent to those achieved with the core model and allows the addition of the controllers, which have , adjustable features permitting the attainment of the desired performance. The plant model solves the dynamic equations that present the BWR system in the time domain. The variables, such as steam flor and pressure, are epresented as a function of time. The extensiveness of this model is

   } shown in Reference 10. The model is periodically refined, as new g/   experimental or reactor operating data are obtained, to improve its                    i capability and accuracy.

i 4.h.4.6.5 Analytical Confirmation Figure 4.h-4 demonstrates the competence and inherent conservatism of the core stability model. The relationship of the calculated damping coefficient from the reactor core dynamic analytical code is related to measured results from 14 rod oscillator tests performed at large operating BWH plants by the General Electric Company. The correlated, Most Probable Values, (MPV), based on a least squares analysis, and the line representing a 97 57. (two sigma) confidence level, below which the actual values will fall, are presented in Figure 4.h-4. The results show the analytical methods to be an effective and useful design tool, with significant conservatism in its application to boiling water reactor core evaluation. Heal and Zivi7 further confirm the effective application of essentially the same model to channel and core l analysis, as does Reference 6.  ! h.h.h.6.6 Analysis Results The analysis is rerformed using a bounding value void coefficient (Figure h.3-23) which is expected to cover several cycles of operation. For all operating conditions in which the actual void (oefficient is less than the (p) bounding value, the analysis remains valid. She most sensitive reactor 4.4-19 Am. No. 56,(3/81)

    --_.- _                             __              _=_.         .          .- - _ .                     --             . . . . _ ~ . ~ . . - -                    - ..                                      __ . . - _ - _ - -                                _ _ - - - - - _

ACNGS-PSAR operating condition in that correnluniding to the highent attninable rod line internection with natural circulation flow. A generic power / flow operat ing condition ( i.e. , '>l.'>$ nuclear boiler rated power) iu unalyzed to bound the typical power / flow operating region an ohown in Figure 1.1 ';. 6

                                     'lypical valueu of renetor core atability are an followu:

Natural Circulation llenc t o r 51.*>% Power Core utabil_ity, (10'>5 Hod Pattern) Decny ratio, (xg/x0) 0*96 llenannot frequency, (liz ) 0.13 4 The enlculated valuen ohow the reactor to be in compliance with the ultimate perfor mance criterin to the mont realx>nnive attninnble node nn cited for the reactor core ntability evaluntion. Figure 1.1-66 6 uhown the eniculated variation of the deeny ratio over the norrral power-ficw range for the bounding core conditions. The channel hydrodynnuic perforrnnce in evaluated at the nont liuiting condition that occurn at the bounding core condition, penked to the bottom of the core. The calculations yield decay ration an prenented below: Channel Natural Circulation Hyd rody nanic Performance ';1. ';% Power Decay-rnt10 (xp/x0) 0.96 Henonant frequency (11 ) 0 . 16 3 At thiu raout reuponnive attainable mode, the mont renponnive channel confortan with the ultinnte performance criterin of 1.0 decay ratio.

                                    '"he channel performance over the entire range of attninable operation in well below the threnhold of inntability.

Conforatnce with the ultinate perforrnnee criterion in further tested by nununing that the reactor in initially operating at the mont nennitive condition. The nuclear nyntcu in then nubjected to step dinturbancen from prenuure regulator netpoint, control rodu und 1avel controller netpoint. These time reuponuen are e'own in Figuren 1.!.-(n, b, and c. It in clear 4 l that the doeny ratio in l e t. ' han 1.0 and in conformance with the ultimate perfortance cri terion. The analynin reprenentu a typical HWH renponne with control nyutema nimilar to that denigned for thin plur.t. The control nyntem nettingn uued are typical valuen chonen within the rnage of equipment capability. The final control nyuten nettings will be entablinhed and optimized during pinut utartup and, hence, the finni nyutem reuponne will be nomewhat different than that repreuented in Figuren 1.1-7n, 6 6 b, and c. The renulta in thene figuren deuonstrate that a highly utable node of operation in attninnble 9 4 4-20 ^m. no. So,o/so .w-- ay y - - - - .eg--- -">-y--- W v'y va- t w Fw w r 7T vvew - - - - r--aw ---7-- --N--*w 'e-v'w-w----we-- a = - - - - - - - - - - - - - - - - - - - -- * - - - - - - - -

l l 4 ACNGS-PSAR I u :l within the range of Pernitted control systen settings permitted by the { l duqign. , h.4.4.6.6.1 Inpact of Prepressurized Fuel on Stability i The impact on stability parameters of GE BWR fuel prepressurized up to 3 atnospheres has been evaluated generally and docunented in References 31 I through 34. liased on these evaluations, it has been concluded that the i effects of prepressurization up to 3 atmospheres is bounded by the current j therml-hydraulic stability analysis and addi.tional analyses are not

l required.

} l 4.4 5 Testing and Verification 1 ] The testing and verification techniques to be used to assure that the ! planned therml and hydraulic design characteristics of the core have been provided, and will remin within required li . tits throughout core lifetine, are discussed in Chapter 14, (Initial Test l'cogran). A sunnary is as  ! follows: . (1) Preoperational Testing ' i

Tests are performed during the preoperational test progran to l confirm that construction is couplete and that all process and j l safety equipuent is operational. Baseline data are taken to i l assist in the evaluation of subsequent tests. Heat balance

instrunentation and jet pump flow and core temperature instrunentation are calibrated and set points verified. l l j (2) Initial Startup { Hot functional tests are conducted with the reactor between 5 and l 107, power. Core performance is monitored continuously to assure l that the reactor is operating within allowable limits (e.g., ( peaking factors, LHGR, etc.) and is evaluated periodically to j verify the core expected and actual performance margins. 1 i i 4.4.6 Instrunentation Requirements [ he reactor ves el instrumentation monitors the key reactor vessel j operating parameters during planned operations. This ensures sufficient l control of the parameters. Th2 following reactor vessel sensors are discussed in Subsect' . 7.6.1.2 and 7.6.1.6: (1) Reactor Vessel Teuperature (2) Reactor Vessel Water Level i (3) Reactor Vessel Coolant Flow Rates and Differential Pressures } (4) Reactor Vessel Internal Pressure (5) neutron Monitoring Systen l I i 4.4-21 Am. No. 56, (3/81)

1 ACllGS-PSAR

i 4.4.7 heferences i

1. " General Electric Therutt Analysis basis (CETAB): Data, Correlation,

and Design A; plication",s eneral Electric Company, January 19'(7

{ (ID:DO-10956 A ) . l l 2. " Core Flow D ,tribution in a Modern Boiling Water Reactor as lleasured l in Monticello", Audust 1976 (NEDO-10722A). 3 " Peach Botton Atonic Power Station Unitr 2 and 3, Safety Analysis Report for Plant Modi 'ications to Eliminate Significant In-Core i Vibration," Septenber 1975 (!iEDO-2099 41 ) . l 4 R. C. Martinelli and D. E. Nelson, "Preli.ction of Pressure Drops l l Durind Forced Cotrzection Boiling of Water", AGME Trans. , 70, pp , 695-702, 19LB. l f 5 C. J. Baroczy, "A Systematic Correlation for Two-Phase Pressure Drop", i Heat Transfer Conference (Los Angeles), AICLE, Preprint No. 37, 1966. i

 ;

6 W. 11. Jens and P. A. Lottes , Analysis .>f Heat Transfer, Burnout , Pressure Drop, and Density Data for High Pressure Water, USAEC Report

!        4627, 1972.

i 7. L. G. lleal and S. M. Livi, "The Stability of Boiling Water Reactors { and Loopu", Nuclear Science and Engineering, 30 p. 25,1967.

8. " Stability und Dynanic Performance of the General Electric Boiling Water Reactor", General Electric Company, January 1977 (NED0-21506).

9 3. Levy, et. al., " Experience with BWR Fuel Rods Operating Above Critical Flux", Nucleonics, April 1965

10. "Analytien1 tiethods of Plant Transient Evaluations for General Electric Boilind Water Reactor", General Electric Company, BWR Systens Department, February 197? (NEDO-10802).

11. H. Zuber and J. A. Findlay, " Average Volumetric Concentration in Two-Phase Flow Systens", Trans. ASME, Journal of Heat Transfer, Novenber 1965

12. " Supplemental Information for Plant Modification to Eliminate Significant In-Core Vibration", NEDE-21156 (Class III), January 1976.

13 H. S. Isbin, H. A. Rodriquez, H. C. Larson and B. D. Pattie, " Void Fractions in Two-Phase Flow", A.I. Ch.E. Journal, Volune 5, Ho. 4, pp. 427-h32, Decenber 1959

14. H. S. Isbin, U. C. Sher, K. C. Eddy, " Void Fractions in Tw> Phase Stean-Water Flow", A.I. Che.E. Journal, Volume 3, No.1, pp. 136-142, flarch 1957 i

15 J. F. Marchaterre, "The Effect of Pressure on Boiling Density in 1 Mu: tiple hectandular Channels", A4L-5522, February 1956. 4 4-22 Am. No. 56, (3/81)

l I ACHOS-PSAR l l

                                                                                                                     ;

16. E. Janseen and J. A. Lervinen, "Two-Phase Pressure Drop in Straight l Pipes and Channels; Water-Steam Mixtures at 600 to 1400 psia", May 1 f

9 17 1964, (GWAP-4616). W. If. Cook, " Boiling Density in Vertical Rectangular Multichannel l ! Section with Hatural Circulation", AHL-5621, Hovember 1956.

10. G. W. Mauer, "A Method of Predicting Steady-Utate Boiling Vapor Fractions in Henctor Coolant Channels", WAPD-B"'-19, June 1960.

, 19 S. Z. Rouhani, " Void Measurements in the Region of Subcooled and Low ! Quality Boiling", Symposium on Two-Phase Flav, University of Exeter, Devon, England, June 1965 4 MO. A. Firstenberg and L. G. Heal, " Kinetic Studies of Heterogeneous Water

Resctora", STL 372-38, April 15, 1966. , I

21. J. s. Ferrel, "A Study of Convection Boiling Inside Channels", North l Carolina State University Raleigh, H.C. , September 30, 1964 i

l

                      ?2. G. Z. Houhani, " Void Measurements in the Region of Subcooled and Low Quality Boiling", Part II, AE-RTL-788, Aktiebolaget, Atomenergi,                          j

! Studsvik, Sweden, April 1966. e l

3. H. Christensen, " Power-to-Void '.'ransfer Functions", AHL- 6385, July 1961. l

l 24. R. A. Ugsn, U. A. Dingee, J. W. Chastain, " Vapor Formation and Behavior in Boiling Heat Transfer", BMI-ll63, February 1957. 4 t 25 KAPL-2170 Hydrodynanic Stability of a Boiling Channel, by A. B. Jones, j i 2 October 1961. l

26. KAPL-2208 Hydrodynarde Stability of a Boiling Channel Part 2, by A. B.

Jones, 20 April 1962. . 27 KAPL-2290 Hydrodynamic Stability of a Boiling Channel Part 3, by A. B. Jones and D. G. Dight, 28 June 1963

28. KAPL-3070 Hydrodynamic Stability of a Boiling Channel Part 4, by A. B. l Jones, 18 August 196h. l 1

j 29 KAPL-3072 Reactivity Stability of a Boiling Reactor Part 1, by A. B. ) ! Jones and W. M. Yarbrough, lh September 1964.  !

30. KAPL-3093 Reactivity Stability of a Boiling Reactor Part 2 by A. B.

Jones, 1 March 1965

31. R. B. Elkins, " Fuel Rod Prepressurization Amendment 1", May 1978 (NEDO-23786-1).

32. Letter, E. D. Fuller to 0. D. Parr, "HRC Request for Additional Information on Fuel Rod Prepressurization", June 8, 1978.

4.4-23 Am. No. 56, (3/81)

\ l ACllGG-PSAR I i$. Letter, E. D. Fuller to 0. D. Parr, "llHC Hequent for Additional Inforuation on Fuel Rod Prepressurization", August lb, 1974.

34. R. B. Elkina, " Fuel Rod Prepressurization",riarch 1978 (IlEDE-23786-1-P) .

                                                                                                                                                           ,l I

i O l

l i O 4.4-24 Am. No. 56, (3/81)

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

1. i i

 <                                                                                                                                                             i I                                                                               ACNGS-PSAR                                                                      [

4 l Table k.k-1 1 THERMAL AND HYDRAULIC DESIGli CilARACTERISTICS j

0F THE REACTOR CORE [

f

                                                                                                                                                                 ;

i General Operating Conditions (238-748)

                                                                               .                                                                               ?

j Reference design thernal 3579

,             output (ffwt) i              Power level for end neered             i                                                        3730                                            !

safety features (Mwt) .[ Uteam flow rate, at 420 F final 15.h00 [ feedwater temperature (millions lb/hr) i Core coolant flow rate 10h.0 (millions lb/hr) Feedwater flow rate (millions 15 367 i i lb/hr) f I Uysteu pressure, nominal in 1040 i steam doue (psia) j [- i Uystem pressure, nominal core 1055 design (psia) j Coolant saturation temperature 551 ( l; at core design pressure ( F) {- 4 Average power density Sh.1 (kW/ liter) f 4 r l Maximum Linear lleat Generation 13.4 I

l. Rate (kW/ft) f Average Linear lleat Generation 5.9 Rate (kW/ft) 73,303 ,

} Corg) ( ft total heat transfer area  ! l l Maximum heat flux (Btu /hr-ft ) 361,600 Average heat flux (Btu /hr-ft2) 159,500 i Design operating uinimum 1.20 , critical power ratio (MCPR) l t i Core inlet enthalpy at 420 F 527 7 l FFWT (Btu /lb) l' l ! I i ) i  ; 4.4-25 Am. No. 56, (3/81) l

   -      ...        ~            - . . . - .              .. -. - , .- .                                 .        . _                                    .w

) ACNGS-PSAR L Table k.4-1 (continued)  ; i T!!ERMAL AIJD HYDRAULIC DESIGN CHARACTERISTICS OF THE REACTOR CORE  ; General Operating Conditions (238-748) [ i ! Core inlet teuperature, at 533  ! h20 F FFWT ( F) f i < Core uaximum exit, voids within 79 0 assemblies (b) Core averade void fraction, 0.k14  ; active coolant  ; i Maximum fuel temperature ( F) 3435 [ I  !

15 16k j

Active coolang) assembly (in. flow area per Core average inlet velocity 6.98 ) (ft/sec) l Maximuu inlet velocity (ft/sec) 8 54 2 ~ Total core pressure drop (psi) 26.4 f Core support plate pressure 22.0 drop (psi) Average orifice pressure drop 5 71 i Central region (psi) ) Average orifice pressure drop 18.68 Peripheral region (psi) Maximum channel pressure 15.40 i loading (psi)  ! - Average-power assembly channel 1h.1 f

pressure loading (bottom) (psi)  !

j t ) Shroud support ring and lower 25 7  ; shroud pressure loading i Upper shroud pressure loading 3.7  !

(psi) l i

4 1 i i { t L 4.4-26 Am. No. 56, (3/81) - r

 .    -            .                          =  -     .     ._ .-     .            . ,

l ACNGS-PSAR Table 4.4-2 AXIAL POWER DISTRIBUTION USED TO CALCULATE MCPR OPERATING LIMIT H< de Axial Power Factor Botton of core 1 0.47 2 0 55 3 0.64 4 0.74 5 0.85 6 0 97 7 1.10 , 8 1.21 9 1.29 10 1.34 11 1.38 12 1.40 13 1.39 1h 1.36 15 1.30 l 16 1.23 17 1.15 18 1.08 19 1.01 20 0 93 21 0.8h 22 0.74 23 0.60 Top of Core 2h 0.43

                                                                                        ;

i l I l t t l 4.4-27 Am. No. 56, (3/81)

ACNGS-PSAR 1 Table h.h-3 l VOID DISTRIBUTIO:1 Core Average Value - 0.h14 Maximum 1:xit Value - 0 790 Actuve Fuel Length - 150 inches Core Average  !!aximum Channel Node (Average Node Value) (End of Hode Value) Bottom of Core 1 0 0 2 0 0.008 3 0.008 0.084 4 0.042 0.20h 5 0.104 0.314 6 0.178 0.402 7 0.253 0.475 8 0 323 0.532 9 0.381 0.578 10 0.429 0.614 11 0.467 0.644 12 0.498 0.668 13 0 524 0.687 1h 0.545 0.703 15 0 563 0 718 0i 16 17 16 0.579 0 593 0.606 0.730 0 742 0.753 19 0.619 0 763 20 0.631 0.773 21 0.6ho 0 780 22 0.648 0.785 23 0.654 0.789 Top Core 24 0.656 0 790 0 . 4.4-28 Am. No. 56, (3/81)

I ACNGS-PSAR l Table k.4-h AXIAL POWEll DISTRIBUTION USED TO GENERATE f VOID A.lD QUALITY DISTRIBUTIONS l Node Axial Power Factor ' Bottom of Core 1 0.38 2 0.69 l i 3 0 93 l h 1.10 i j' 5 1.21 l 6 1.30 7 1.h7 8 1 51 9 1.49 10 1.h4 11 1.36 12 1.28 13 1.16 i 14 1.06 15 1.01 16 0 97 17 0 9h 18 0 97 19 0 96 20 0.91 21 0.77 22 0 59 23 0.38 ! Top of Core 24 0.12 i l l l 4 l

i l G 4.4-29 Am. No. 56,(3/81) '

ACNGS-PSAR l I Table 4.4-5 [ FLOW QUALITY DISTRI1UTIOM l Core Average Value ').079  ; f4aximum 1:xit Value - J.332  ! Active Fuel Length - 150 inahes  ; I

                                                                                                                                                   ;

Core Average- 14axinam Channel , Node (Average Node Value) (End of Node Value) , l Botton of Co e 1 0 0 2 0 0 l 3 0 0.004 i 4 0.001 0.014 i 5 0.004 0.031 I 6 0.011 0.050 i 7 0.020 0.072 8 0.031 0.096 l 9 0.043 0.118 l 10 0.055 0.140 l 11 0.066 0.161 i 12 0.077 0.181 i 13 0.087 0.199 . 14 0.097 0.215 l 15 0.106 0.231  ! 16 0.114 0.245 l 17 0.122 0.260  ! 18 0.130 0.275  ! 19 0.138 0.289  ! 20 0.146 0.303  ! 21 0.153 0 315  ! 22 0.159 0.324 l 23 0.163 0.330 i Top Core 24 0.165 0.332 l i t I f

                                                                                                                                                   ;
                                                                                                                                                   )

i i t i i l f v i 4.4-30 Am. No. 56,(3/81) l l  ! l

ACNGS-ITAR l ! Table 4.k-6 CORE FLOW DISTRIBUTION i i ! Orifice '.one Central Central Peripheral Peripheral l _ Description Hot Average Hot Average Relative Assei..bly Power 1.400 1.083 1.000 0 350 llelative Assenbly Flow 0 926 1.051 0 563 0.641 1 1 i i 4 I i i l 4 i I

l 1

l 1 e i l i i i f .i l 1 i I 4.4-31 Am. No. 56,(3/81) 'l l I_.___,_-_.._-__.-__...._.._._

ACHGS-PSAR Table 4.4-7

     )

j CALCULATED VS MEASUPED CORE PLAT!' PR1:3SURE DROP.i Test Condit ton Core Plate AP Plant size Powei ,($ rated) Flow (% rated) 11eas (psid), Calc (psid) 183-368 83 9 100.2 25 10 24.82 251-764 95 3 94.9 18.14 17.91 99 3 96.9 18.69 18.47 224-580 70 3 60.8 5 04 5.05 99 3 99 3 14.74 14.77  ; 218-540 86.7 100.6 17 17 19 30 90 3 96.0 16.13 17.85 218-560 66.4 59 9 7.47 6.73 79 2 94.4 18.24 17 38 251-764 46.9 68.0 6.13 7 55 51.3 103 3 18.50 18.00 , 46.6 48.0 3 99 3 52 64.9 70.3 9 42 8 90 O 75.9 57.8 70.1 101.0 46.4 T1.0 18.51 3 79 9 57 10.50 3.46 9.32 96.4 98.9 19 50 19 25 i l

 \s_ /                                                                                           i l

4.4-32 Am. No. 56,(5/81) l i

i ACNGS-PSAR i Yable 4.h-8 TYPICAL RANGE OF TEST DATA 14easured Parameter Test Condition i 1 Adiabatic Tests: l l 5 5 Spacer single pha..e loss flge. = 0 5 x 10 to 3 5 x 10 l coefficient 1 i Lower tieplate + ori rice T = 100 to 500 F l I single phase loss j coefficient f l Upper tieplate single-phase t friction factor i Upacer two-phase loss P = 800 to 1400 psia  ! coefficient  ! l 6 C Two-phase friction G = 0 5 x lg to 1.5 x 10 uultiplier lb/h-ft - X = 0 to 40% f I Diabatic Tests i k Heated bundle pressure drop P=800toIg00 psia 6 f G = 0 5 x lg to 1 5 x 10 l lb/h=ft 1 S. i i I e s t t l I i i

• Reynolds llunber l 4.4-33 Am. No. 56,(3/81) f t

ACliGS-PSAR l Table 4.4-9 DESCRIPT1011 0F UllCERTAINTIES Standard Deviation Quantity (% of Point) Connent Feedvater Flov 1 76 This is th> 1argest component nf i total reactor power uncertainty. l l Feedvater Temperature 0 76 These are the other significant ! Reactor Pressure 05 parane.ters in core power i determination. Core Inlet Tenperature 0.2 Affect quality and boiling langth. Core Total Flov 25 Flow is not measured directly, but is calculated from Jet punp AP. The listed uncertainty in flow corresponds to 11.2% ctan% rd deviation in each individual pump difference. i ! Channel Flow Area 3.0 This accounts for manufacturing and l service induced variations in the

free flow area within the channel.

Friction Factor 10.0 Accounts for uncertainty in the

         !!ultiplier                                                correlation representing two-phase pressure losses.

Channel Friction 50 Represents variation in the pressure Factor !!ultiplier loss characteristics of individual f channels. Flow area and pressure

! loss variations affect the core i flow distribution, influencing the I quality and boiling length in ! individual channels. 1 TIP Readings 6.3 These sets of data are the base from which gross power distribution l j is determined. The assigned I uncertainties include all electric and 6eometrical components plus a contribution from the analytical l { extrapolation from the chamber location to adjacent fuel assembly segment. Also included are uncer-tainties contributed by the LPR!! systen. LPRM readings are used to correct the power distribution O calculations for changes which have 4.4-34 Aw No. 56,(3/81)

i I I ACNGS-PSAR  ; I i Table 4.4-9 (Continued) j ( DESCI:PTION OF UilCERTAINTIlli l C',andard j Deviation j Quantity (f,of Point) Comnent , l occured since the last TIP survey. l The assigned uncertainty e.frects power distribution in the sans manner as the base TIP reading , uncertainty.  : 1 l H-Factor 1.5 This is the last of the three power I distribution related uncertainties.  ! It is a function of the uncertainty l in local fuel rod paver. j Critical Power 3.6 Uncertainty in the GEXL correlation expressed in terms of critical power. i l l l i l l. l t i l l b 4.4-35 Am. No. 56,(3/81) i l l

ACNGS-PSAR 7-_s Table k.4-10 ( l \s_s/ BYPASS FLOW PATHS Flow Path _ Description Driving Pressures Number of Paths la. Between Fuel Support and Core Plate One/ Control Rod the control Rod Guide Differential Tube (Upper Path) l b. Between Fuel SJupport Core Plate One/ Control Rod and the Control Rol Differential Guide Tube (Lower Path)

2. Between Core Plate and Core Plate One/ Control Rod Control Rod Guide Tube Differential

3. Between Core Gupport and Core Plate One/ Instrument the In-Core Gupport Differential Instruuent Guide Tube

4. Between Core Plate and Core Plate One Shroud 5 Between Control Rod Core Plate One/ Control Rod f' '3 Guide Tube and Control Rod Drive Housing Differential

( )

%/

6. Between Fuel Support Channel Wall One/ Channel and Lower Tieplate Differential Plus Lov Tieplate Differential 7 Control Rod Drive Independent of One/ Control Rod Coolant Core

8. Between Fuel Channel Channel Wall One/ Channel and Lover Tieplate Differential 9 Holes in Lover Tieplate Lower Tieplate/ Two/ Assembly Bypass Region Differential

>f1 N) 4.4-36 Am. No. 56,(3/81)

 . . _ . . _ _ . _ . _ _ . . _ _ . _ _ _ _ _ _ _ _ _ _ - - -                         . . _._.__ __. _ ______ .                    -                                .__._m           _ _ - __--._ _ _ _ _           -

p. t I l t

ACNGS-PSAR 1 ~ i

;                                                                                                Table 4.4-11 HEACTOH COOLANT SYSTEM GEOMETRIC DATA                                                                                                                                  ,

1 Height Elevation { 4 Flow and of Bottom Minimum i l Path Liquid of Each Flow

Length Level Volume

• Aregs j (in.) (in.) (in.) ( ft )

I A. Lower Plenan 213.5 213 5 -170 5 84.0 4 213 5 i  ! [ U Core 164 5 164 5 43 0 146.5 j 16h.5 includes l bypass i Upper Plenum and 207 5 57 5 i 4 C. 179 0 179 0 Separators 179 0  ; v D. Dome (Above Ilornal 289 5 289 5 386.0 309 0 i Water Level) 0  !

,                                                                                                                                                                                                                                         f E.          Downcomer Area                                       311 5                                             311 5    -27 5                 66.0 i                                                                                                                                                   311.5                                                                                  j a

F. Hecirculation Loops 114.0 ft 398.0 -392.0 132 5 in j

;                                           and Jet Pumps                                    (one Loop)                                            398.0                                                                                  l
!                                                                                                                                                                                                                                         l i
!                                                                                                                                                                                                                                         E E

I i i l i , t' >  ? I i

                                                                                                                                                                                                                                          ?

4 I 3 i i i

                                                                                                                                                                                                                                          )

• Reference Point is recirculation nozzle outlet centerline.  !

l 4.4-37 Am. No. 56, (3/81) [ t

      . . . . - . . . ,                             ,.       --,--c.---    - - . - . - . . - - , _ . , . - . . , - , _ . - _ . , _ . - - - - - - .                        . . . .               . . . - - . - - - . - , - - - , , - , _ .

l 1 l ACllGS-PSAR i f Table L.4-12 LE iG'JHS OF GAFETY IllJt'CTIt'l LI:lES Ilouinal Dianeter Pipe Leng+,h Loop Line (in) Jehedule (ft) l i 'iPCd HPCS 3 16 100 57 12 100 h8

l. HP;S 4 12 60 154 i

LPCS LPCG 2 14 40 92 i 12 k0 lh LPCU 3 12 80 140 1

LPC1"A" RhR 7 18 40 23 i RHR 12 18 40 4 l RER 9 18 40 116 li 14 40 45 t RHR 10

.2 80 66 I i LPCI"B" RifR 13 13 40 23 i l RHR 18 _3 h0 4 I l RIIE 15 18 40 104 l i ik 40 112 .

l Rhn 16 12 80 61 LPCI"C" RHR 21 18 40 96  ! , RiiR 22 14 80 216 l 12 80 58 l

                                                                                                                                                                                                                                                           ;
               !!O'.'E Lendths given are to the nearest foot, and are measured                                                                                                                                                                             l j                        frou the rgpropriate pump outlet no'zle to the RPV nozzle.

i  !

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                                                                                                                                                                                                                                                           ;

l 1 i l I l l t I t i ( 4.4-38 Am. No. 56, (3/81) t

                                                                                                                                                                                                                                                           ;

i

t I ACNGS-PSAR l NOTE: PERIPHE R AL FUE L SUPPOPITS ARE WELDED INTO THE CORE N SUPPORT PLATE. FOR THESE I BUNDLES, PATH NUMSERS 1 LOWER TIE PLATE 2,5 AND 7 DO NOT EXIST l

                               ,/'                                                                                       :

CHANNEL f s a- ---- a f I

                                                                                                                         ?

I 9 [ CORE SUPPORT f PLATE FUE L SUPPORT l / 7. [ t

                                               ~

[ IN-COR E GUIDE TUBE [ 6

                                              \
                                                                                        ,                                i y SHROUD CONTROL ROD                                                      ,

GUIDE TUBE  ! I b 4----  ! kI -- i h f P r l { g h

                       ~\                                            1.

2. CONTROL ROD GUIDE TUBE FUEL SUPPORT CONTROL ROD GUIDE TUBE CORE SUPPORT PLATE

3. CORE SUPPORT PLATE IN40RE GUIDE TUBE

4. CORE SUPPORT PLATE SHROUD  ;

, S CONTROL ROD 5. CONTROL ROD GUIDE TUBE DRIVE HOUSING j DRIVE HOUSING 6. FUEL SUPPORT LOWER TIE PLATE ,

7. CONTROL ROD DRIVE COOLING WATER

                  -M-                                               8. CHANNFL LOWER TIE PLATE                         I

9. ALTERN ATE FLOW PATH HOLES ,

HOUSTON LIGHTlHG & POWER COMPANY I Allens Creek Nuclear Generating Station f C\ i Unit 1 Schematic of Reactor Assembly Showing the Bypass Flow Paths l Am. No. 56, (3/81) Figure 4.4-1. , t

i ACaGS-PSAR  ! I

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                                          .,-,_e                                                                - _ _ . _ _ _ , _ _ . _ _ . , _ .

ACNGS-PSAll e NEUTHON s [ \ / j FLUX RE SPONSE

  %s A  L RE ACTIVITY PE RTUR BA TION KINE TICS
                          ' ()

4 s TOT AL RE ACTOR RE ACTIVITY FEEDBACK TOT AL INDIVIDUAL CH ANNE L TYPE REACTIVITY FEEDBACK

                               *          *                   +                FIE ACTIVITY TO POWEFI m     [           .m                  m               TR ANSFER FUNCTION      m
                          " (j'                       (j'                       AT CONSTANT INLET
                                   ;  L J  ,*

F ROM OTHE R [N f f CHANNEL TYPES \j HE ACTIVITY TO F LOW TRANSFER FUNCTION AT CONSTANT POWEFI 4 L F LOW TO POWER TR ANSF ER 1 FUNCTION 1 P TO OTHER CHANNEL 5 Am. No. 56, (3/81) HOUSTON LIGHTING & POWER COMPANY f)/ s V Allens Creek Nuclear Generating Station Unit 1

                                                                                  !!ydrodynamic and Core Stability Model Figure 4.4-3.

t AC?iGS-PSAR 12 i O KR5 DATA Q SENN DATA 1.6 -

                                                                                                            ;

1.4 - 1.2 -  ; I  ! 9 W 9 I 1.0 - 0 S , 0 o 0 O C  : OE - 2 o CONFIDENCE LINE 0 8 i . U 0.6 - O 4 O MOST PROBABLE CORRELATION LINE w a A Bx 0.4 - O O ' O O O ' O.2 - O O I O  ; l 0 0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 . DAMPING COEFFICIENT fn, FROM TEST DATA

                                                                                                             )

HOUSTON LIGHTING & POWER COMPANY l O Allens Creek Nuclear Gene ating Station Unit 1 Comparison of Test Results With Reactor Am. No. 56, (3/81) Fi e 4.4-4.

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9  ! 7 a: 8 , 0 0.6 - i 0.4 - f ti i 0.2 t l i I I a 100 0 20 40 60 80 , l PERCENT POWER HOUSTON LIGHTING & POWER COMPANY ' Allens Creek Nuclear Generating Station s Unit I Total Core Stability j I Am. No. 56, (3/81) Figure 4.4-6.

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I HOUSTON LIGHTING & POWER COMPANY y d Allens Creek Nuclear Generating Station I g g d

             -         -                                                                            Unit 1 "8"                                                     10 PSI Prassure Regulator                     ;

Setpoint Step at 51.5% Rated g Power (Natural Circulation)

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I - i n  : t - ( g g g d* HOUSTON LIGHTING & POWER COMPANY C - Allens Creek Nuclear Generating Station (,- M usi A D E ndad) Unit 1 i 6-In. Water Level Setpoint ,  ; Step at 51.5% Rated Power ( Am. No. 56,(3/81) Figure 4.4-7c.

ACNGS-PSAR ,

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4 I l s _ 3 i J 3 l I 2 I I 2 5 [ L1 - j L s 35m 1 i l- 25E I 1 1 0

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g 1 a _ ; I I I I I I I I I I I o. g g e e , n o e e , n oo S370Nn0 dO W39WAN HOUSTON LIGHTING & POWER COMPANY Allens Creek Huclear Generating Station Unit 1 Relative Bundle Power Ilistogram for Power Distribution Used in Am. No. 56,(3/81) Figure 4.4-S. Statistical Analysis (Basis in Reference 1)

l A;NGS - PSAR I w h.5 REACTOR fiATERIALS V) h.5.1 Control Rod Systen Structural !!aterials b.5 1.1  !!aterial Specifications

a. flaterial List The following material listing applies to the control rod drive mechanism supplied for this application. The position indicator and minor nanstructural itens are omitted.

(1) Cylinder, Tube and Flange Assembly Flange AS!!E SA182 Grade F304 Plugs AS!!E SA182 Grade F304 Cylinder AST!! A269 Grade TP 304 Outer Tube ASR1 A269 Grade TP 304 Tube ASt!E SA351 Grade CF-3 Spacer ASfE SA351 Grade CF-3 (2) Piston Tube Assembly Piston Tube AS!!E SAh79 or SA2h9 Grade Xft-19 Nose ASfiE SAh79 Grade Xft-19 Bare ASt1E SA479 Grade Xft-19 Ind. Tube ASfE SA312 Type 316 {'] Cap AS!!E SA182 Grade F316 (3) Drive Line Assembly  : Coupling Spud Inconel X-750 Conpression Cylinder AS!!E SA479 or SA249 Grade Xft-19 Index Tube AStiE SAh79 or SA2h9 Grade Xff-19 Piston Head arf!CO 17 4 PH or its equivalent Piston Coupling ASD1 A312 Grade TP 30h or ASD1 A269 Grade TP 304

                   !!agnet Housing           ASD1 A312 Grade TP 30h or AS31 A269, Grade, TP 304 or ASTf! A312, A249, or A213 TP 316L (4) Collet Assembly Collet Piston             AST! A269 TP 30h or ASM1 A312 TP 304 Finger                     Inconel X-750 Retainer                  ASti A269 TP 304 Guide Cap                 AST!! A269 TP 304 f')

4.5-l Am. No. 56,(3/81)

i ACfiGS - PGAR (S) !!incellaneoun Parta Stop Pinton AletCO 17 4 PH or its equivalent 0-Hing Spacer ASTf t A240 Type 304 Nut ASitE SA479 Grade Xft-19 Collet Spring Inconel X-750 Hire Flange AS!!E SA182 Grade F304 Buf fer Shaft A101CO l'f 4 Pil or its equivalent Bu ffer Platon ARI!CO 17-4 PH or its equivalent Buf fer Spring Inconel X-750 Nut (hex ) Inconel X-7SO The nuntenitic material 300 aeries stainleau atecin that are listed are all in the annenled candition (with the exception of the outer tube in the cylinder, tube ani flange ancembly), and their propertica are readily available. The outer tube is approximately 1/8 hard, and han a tennile of 90,000/125,000 pai, yield of 50,000/85,000 pai and mininun elongation of 2S%. The coupling upud , collet fin,r,ern , bu ffer spring , nut (hex ), and collet upring are fabriented from Inconel X '(50 in the annenled or equalized condition, and twed 20 houru at 1300 F to produce a tenaile of 165,000 pai mininum, yield of 105,000 pai minir.am, and elongation of 20% mininum. The piston head, stop picton, buffs ahnf t, and buffer piston are arf!CO 17 h PH (or its equivalent) in aondition H-1100 (aged 4 houra at 1100 F), with n tensile of 140,000 poi minimum, yield of 115,000 psi minimum, and elongation of 15% ctininum. These are videly used materinla, whose properties are well known. The parta are readily accensible for inspection and replaceable if necesanry. All materials, ex cept SA479 or SA249 Grade Xf t-19, have been unecenafully used for the pnat 10 to 15 years in sinili.r drive nochanisns. Extensive laboratory tests have demonntrated that aSf1E SAkT9 or SA249 Grade Xf t-19 are cuitable materialc tnd that they are renistant to stress corronion in a BWR env i ronnen t.

b. Special tinterinls No cold-worked nuatenitic stainlens steels with a yield strength greater than 90,000 psi are empl?/ed in the Control Rod Drive (CRD) system.

tktrtensitic precipitated hardened stainlens ateel, ARftCO 17-4 PH, is used for the pis ton hend, stop pis ton , buf fer shaf t, and buf fer pis ton. This material in iged to the H-1100 condition to produce resistance to stress corrosion crackine, in the BWR environments. ARtCO 17-4 PH (or ito equivalent) (H-1100) hna been succeanfully used for the past 10 to 15 years in BWR drive mechanions. h.5 1.2 Austenitic Stainless Steel Componenta

n. Processes, Inspections and Tentn Solution annenled 300 atninless steel material used in fabrienting control rod drive parts is verified as being correctly colution annenled by testing Am. No. 56,(3/81)

I ACNGS - PSAR i f s per .idTff-A262, " Detecting Susceptibility to Intergranular Attack in i I

     \    Stainless Steels."

l Two special Irocesses are employed which subject selected 300 Series  : stainless steel components to temperatures in the sensitization range: (1) The cylinder and spacer (cylinder, tube and flange assembly) and the retainer (collet assembly) are hard surfaced with Colmonoy 6 1 (or its equivalent). (2) The collet piston and guide cap (collet assembly) are nitrided to provide a wear-resistant surface. Nitriding of the above parts is accor:.;11shed usind a proprietary process called New !!alcomizing. Components are exposed to a temperature of about , 10300 F for about 20 hours during the nitriding cycle. l r Colmonoy hard-surfaced components have performed successfully for the past i 10 to 15 years in drive mechanisms. Nitrided components have been used in CRDs since 1967 It is normal practice to remove some CRDs at each  ; refueling outade. At this tine, both the Colnonoy hard-surfaced parts and nitrided surfaces are accessible for visual examination. In addition, dye  ; penetrant exaninations have been performed on nitrided surfaces of the longest service drives. This inspection program is adequate to detect any incipient defects before they could become serious enough to cause operating problems. i Welding of austenitic stainless steel parts is performed in accordance with Sections IX (Welding and Brazing Qualification) and Section II Part C (Welding Rod Electrode and Filler fletals of the ASt!P Boiler and assemblies l require solution annealing to minimize the possibility of the sensitizing. However, velded assemblies are dispesed fron this requirement when there is documentation that velds are not subject to sustained loads and assemblies l have been free of service failure. Other reasons, in line with the j reguls tory guide, for dispensing with the solution annealing are that , assemblies are exposed to reactor coolant during normal operation service  ! which is below 200 F temperature or assembliec are of material of low carbon  ; content, less th:m 0.025 percent. These controls are employed in order to  ; comply with the intent of the Regulatory Guide 1.44. { Regulatory Guide 1.h4 1 General Compliance or Alternate Approach Assessment: For commitment and  ; Revision Number, see Appendix C. l

b. Control of Delta Ferrite Content )

Control rod drive parts were fabricated after the issuance of Revision 2 of I Regulatory Guide 1 31. I All type-308 veld metal was purchased to a specification which required a minimum of 57. delta ferrite. Ferrite measurements were made with a h calibrated, magnetic instrument on undiluted weld pads for each lot and heat Q of weld filler metal. For the submerged arc welding process, measurements 4.5-3 Am No.56,(3/81)

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ACNGS - PSAR vere made for each vire-flux combination. These proceduz as conply with the requirenents of Revision 2 to Retpalatory Guide 1.31. Hegulatory Guide 1.31 General Conplience or Altet aate Approach Asnessment: For Connitnent and Revision Number, See Appendix C. h.5 1.3 other flaterials These are discuase.1 in Subsectin 4.5.1.1.b. 4.5.1.4 Cleanind and cleanli ness control L.5 1.4.1 Protecti m of !!aterials During Fabrication, Shipping and Storade All the CRD parts listed above (Subsection b.5 1.1) are fabricated under a

    ..rocess specification which limits contaninants in cutting, grinding ar:4 tapping coolants n,d lubric ints. It also restricts all other processind ruterials (narking inks, tape etc. ) to those which are completely renovable by the applied cleanind process. All contaninants are then required ta be renoved by the appropriate cleaning process prior to any of the following:

(1) Any processind which increases part tenperature above 200 F. (2) Assembly which results in decrease of accessibility for cleaning. (3) Release of parts for shipnent.

    "'he specification for packaging and shippind the Control Rod Drive provides the following:

The drive is rinsed in hot deionized water and dried in preparation for shipment. The ends of the drive re then covered witvh a vapor tight barrier with dessicant. Packaging is designed to protect the drive and prevent danate to the vapor barrier. The planned storage period considered in the desigt of the container and packaging is two years. This packaging has been qua ified and in use for a number of years. Periodic audits have indicated satinfactory protection. The degree of surface cleanliness obtained by these procedures neets the equirements of Regulatory Guide 1.37 Site or vareho,se storade specific.ations r:anire inside heated storage comparable to level B of ANSI N45 2.2. Semi $nnual exanination of the hunidity indicators of ten percaent of the units during inside heated warehouse storage is required to verify that the units are dry and in satisfactory condition. Position indicator probes are not subject to this ins pection. Regulatory Guide 1.37 l General Conpliance or Alterante Approach Assessment: For Connitnent and j Revision Number, see Appendix C. 4.5-4 Am. No. 56,(3/81) l

ACNGS - PSAR

   'O   h.5.2       Reactor Internal !!aterial 4.5 2.1     !!aterial Specifications
        !!aterials used for the Core Support Structure:

Shroud Gupport - Nickel-Chrome-Iron-Alloy, ASfE SB166 or SB168. Shroud, core plate, and top guide - ASt!E SA240, SA182, SA479, SA312, SA249, or SA213 (all Type 304L). Peripheral fuel supports - ASD1 A312 Grade TP-304, 30hL, or 316L. f Core plate and top guide studs and nuts, and core plate wedges - ASt4E SA479, SA193 Grade B8A, SA194 Grade 8A (all Type-304) Top guide pins - ASfE SA479 (Type Xff-19). i Control rod drive housing - ASitE SA312 TP-316L, SA182 Tp-316L, SA213 TP-316L, SB167 Type Alloy 600. Control rod guide tube - AStiE SA358 Grade 304, SA312 Grade TP-304, SA351 Grade CF3 or CF8, SA249 Grade TP-304 Orificed fuel support - AS31 A240 TP-30h or A276 TP-304, A2h0 TP-316L, ASfE SA479 Tp-316L, ASt4E SA351, Grade CF8.

        !4aterials Employed in Other Reactor Internal Structures.

(1) Steam Separator and Steam Dryer All materials are TP-30h, 30hL or 316L stainless steel. Plate, Sheet and Strip AS21 A240, TP-304, 30hL or 316L Forgings AS31 A182 Grade F304 or 30hL  ; Bars AST!! A276 TP-304 or 316L Pipe ASHf A312 Grade TP-30h Tube AS'Ifi A269 Grade TP-304 r Castings AS'Ili A351 Grade CF8 (2) Jet Pump Assemblies , The components in a Jet Pump Assembly are the Riser, Inlet flixer, , Diffuser, and Rise Brace. Platerials used for these components are  ; to the following specifications: s Castings ASD1 A351 Grade CF8 and AS31 SA351 Grade CF3 l l

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4.5-5 3 ,, go, 56, (3/81)

ACNGS - PSAR Bara ASTf1 A276 TP-304, ASU! A479 TP-316L ASU1 A637 Grade 688 Bo lts ASt1 A193 Grade B8 or B8t1 and ARIE SAh79 TP-316L Sheet and Plate ASUt A240 TP-30h, and AQ1E SA240 TP-30ht, 316L Tubi ng ASTtt A269 Grade TP-304 P'. pe AST!! A358 TP-304, 316L and AStiE SA312 Grade TP-304, 316L Forged or Rolled Parts ASt1E SA182, Grade F304, F316L, ASU{ B166, and ASTM A637 Grade 688. !!aterials in the Jet Pump Ansenblies which are not austenitic stainless steel are listed below:

a. The Inlet flixer Adaptor casting, the wedge casting, bracket casting adjusting screw casting, and the Diffuser collar casting are hard surfaced with Stellite 6 (or its equivalent) for slip fit joints.

b. The Diffuser is a binetallic compontent made by velding an austenitic stainless steel ring to a forged Alloy 600 ring, made to Specification ASTf1 B166.

c. The Inlet-!!ixer contains a pin, insert, and beam made of Alloy X-750 to satisfy Specification ASU1 A637 Grade 688 properties.

All core support structures are fabricated from AStiE specified materials, and designed in accordance with requirements of ASME Code, Section III, Subsection NG. The other reactor internals are noncoded, and t'.2y are fabricated from ASTil or AS!!E specification materials. flaterial requirements in the ASTf1 specifications are identical to requirements in corresponding AG1E material specifications. 4.5 2.2 Controls on Welding Core support structures are fabricated in accordance with requirenents of AStIE Code Section III, Subsection NG. Other internals are not required to meet A31E Code requirements. Requirements of AS!!E Section IX B and DFV Code, are followed in fabrication of core support structares and other internals. 4.5 2.3 Nondestructive Examination of Wrought Seamless Tubular Products Wrought seanless tubular products for CRD housings and peripheral fuel supports, were supplied in accordance with ASiE Section III, Class CS, which require examination of the tubular products by radiographic and/or ultranonic methods according to paragraph NG-2550, 4.5-6 Am. No. 56, (3/81)

ACNGS - PSAR p (' Wrought seamless tubular products 03r other internals were supplied in  ! accordance with the applicable ASMt or ASME naterial specifications. These specifications require a hydrostatic test on each length of tubing. i 1 4.5 2.h Fabrication and Processing of Austenitic Stainless Pteel Regulatory Guide Conformance Regulatory Guide 1 31: Control of Stainless Steel Welding Cold-worked stainless steels are not used in the reactor internals, except fbr the steam dryer unit vanes. All austenitic stainless steel veld filler materials were supplied with a minimum of 5% delta ferrite. This amount of ferrite is considered adequate to prevent micro-fissuring in austenitic stainless steel velds. Reactor internals were fabricated prior to the issuance of Rev. 2 to Regulation Guide 131. Ferrite measurements are nade in accordance with the requirements of the  ! ASME Code in ef fect at the time. This code requires the use of the chemical I composition in conj unction with the Schaeffler din?ran to verify that veld filler netal contains a mininun of 5% delta ferrite. i An extensive test program performed by General Electric Company, with the  !

    h   concurrence of the Regulatory Staff, demonstrated that the use of the Schaef fler diugram to control veld filler metal ferrite at 5% minimun was y     /

adequate to produce satisfactory production velds. The 357 production velds evaluated in this program were fabricated with filler metal controlled in accordance with the Schaeffler diagram to contain a minimum of 5% ferrite, All these production velds net the requirements of the Interin Regulatory Position to Reg. Guide 131 which was in effect at that time. + Regulatory Guide 1 3h: Control of Electroslag Weld Properties  ! Electroslag velding is not employed for any reactor internals. Regulatory Guide 1.36: Nonmetallic Therrnl Insulation for Austenitic Stainless Steel Inside the containment, nonmetallic thermal insulation vill not be used on austenitic stainless steel. Outside of the containment, where nonmetallic thermal insulation is used fbr austenitic stainless steel, the requirements of this regulatory guide vill be met. . t Regulatory Guide 1.bb: Control of the Use of Sensitized Stainless Steel All vrought austenitic stainless steel was purchased in the solution heat , treated condition. Heating above 8000 F was prohibited (except for velding) unless the stainless steel was subsequently solution annealed. For TP-304

    ,_s    steel with carbon content in excess of 0.035% carbon, purchase
  /      j specifications restricted the maxinum veld heat input to 110,000 Joules per
  \s_,/    inch, and the veld interpass temperature to 350 F maximum. Welding was 4.5-7                         Am. No. 56, /.3/81)

ACNGS - FSAR performed in accordance with Section IX of the ASt!E Boiler and Pressure Vessel Code. These controls were employed to avoid severe sensitization and comply with the intent of Regulatory Guide 1.kh. Regulatory Guide 1.71: Welder Qualification for Areas of Limited Accessibility There are few reutrictive velds involved in the fabrication of items described in this section. Mockup velding was performed on the velds with most difficult access. !!ockups were examine I with radiodraphy or by sectionind. Regulatory Guide 1 37: Quality Assurance Requirenents for Cleaning of Fluid Systems and Associated Components of Water-Cooled Nuclear Power Plants Exposure to containinant was avoided by carefully controlling all cleaning and processing materials which contact stainless steel during construction. Any inadvertent surface contamination was renoved to avoid potential detrinental effects. Special care vau exercised to insure removal of surface contaminants prior to any heating operation. Water quality for rinsing, flushing, and testing vns controlled and monitored. The degree of cleanliness obtained by these procedures meeta the requirements of Regulatory Guide 137 Regulatory Guides 1. 31, 1 3h , 1. 36, 1. h h , 1 71 an d 1. 37 General Compliance or Alternate Approach Assessment: For Commitment and Revision Nunber, see Appendix C. 4525 Other !!aterials Hardenable martensitic stainless steel and precipitation hardening stainless ateels are not used in the reactor internals.

 ?!aterials, other than Type-300 stainless steel, enployed in vessel internals are:

(1) SAh79 Type X!1-19 stainless atcel; (2) SB166,167, and 168, Nickel-Chrome-Iron ( Alloy 600); and (3) SA637 Grade 688 Inconel X-750. Alloy 600 tubing plate, and sheet are used in the annealed condition. Bar may be in the annealed or cold-drawn condition. Alloy X-750 components are fabricated in the annealed or equalized condition and aged 20 hours at 1300 F. e l l 4.5-8 Am. No. 56, (3/81)

l ACNGS - PSAR i Stellite 6 (or its erluivalent) hard surfacind is applied to austenitic stainless steel castings using the gas tungsten are velding or plasra arc i j i l O surfacind Processes. All materials, except SA 479 Grade Xft-19, have been successfully used for l l the past 10 to 15 years in BWR applications. Extensive laboratory tests j t have demonstrated that Xff-19 is a suitable material and that it is resistant 6 to stress corrosion in a BWR environment. { h.5 3 Control Rod Drive Housing Supports l I All CRD housing support subasser'blies are fabricated of AST!!-A-36 structural l steel, except for the following items:

                                                                                                        !!aterial i                                         Grid                                           AST1 Ahkl                                                                ,

t I j Disc springs Schnorr Pfpe BS-125-71-6 (or , j its equivalent) l l Hex bolts and nuts ASD1 A307 1 \ l 6 in. x 1 in. x 3/8 in. ASD1 A500 Grade B l j tubes > 1  ! ! For further CRD housing support information, see Subsection 4.6.1.2. ~O l i l

1 I l l . i I ! l l l l l l l l l I I i i O , 4.5-9 Am. No. 56, (3/81)

1 ACNGS-PSAR l l I h.6 FUNCTIONAL DESIGN OF REACTIVITY CONTROL SYSTFT.S The reactivity control cystems consist of control rods and control rod l drives, supplementary reactivity control in the form of a burnable poison l (Section 4 3), and the Standby Liquid Control Systen. (described in Subsection 4.6.6). h.6.1 Mformation for Control Rod Drive System (CRDS) h.6.1.1 Control Rod Drive System Design 4.6.1.1 1 Design Bases i f 4.6.1.1.1.1 General Design Bases 4.6.1.1.1.1.1 Safety Design Bases The CRD mechanical system shall meet the following safety design bases: (1) The design shall provide for a sufficiently rapid control rod insertion such that no fuel damage results from any moderately frequent event (cee Chapter 15). (2) The design shall include positioning devices, each of which l individually supports and positions a control rod. (3) Each positioning device shall:

9 a. prevent its control rod from initiating withdrawal as a result of a single malfunction; 4

b. be individually operated so that a failure in one positioning device does not affect the operation of any other positioning device; and

c. be individually energized when rapid control rod insertion (scram) is signaled so that failure of power sources external to the positioning device does not prevent other positioning i devices' control rods from being inserted. '

4.6.1.1.1.1.2 Power Generation Design Basis The control rod system drive design shall provide for positioning the control rods to control power generation in the core. l 4.6.1.1.2 Description l l' The Control Rod Drive System (CRDS) controls gross changes in core reactivity by incrementally positioning neutron absorbing control rods within the reactor core in response to manual control signals. It is also required to quickly shut down the reactor (scram) in emergency situations by rapidly inserting all control rods into the core in response to a cr.nual e or automatic signal from the Reactor Prctection Trip System. The CRDS l consists of locking piston CRD mechanisms, and the CRD hydraulic system , I I' 4,6-1 Am. No. 56, (3/81) I a

ACHCS-PSAR (includinr power supply and regulation, hydraulic control units, interconnecting piping, instntmentation and electrical controls ). 4.6.1.1.2.1 Control Rod Drive Mechanists The CRD mechanism used for positioning the control rod in the reactor core is a double-acting, mechanically latched, hydraulic cylinder usinC vater as its operating fluid (Figures 4.6-1, 4.6-2, 4.6-3, and 4.6-4). The individual drives are nounted on the bottom head of the reactor pressure , vessel. The drives do not interfere with refueling and are operative even when the head is renoved from the reactor vessel. The drives are also readily accessible for inspection and servicing. The bottom location takes maximum utilization of the water in the reactor as a neutron shield and gives the least possible neutron exposure to the drive components. Using water from the condensate treatment system, and/or condensate storage tanks as the operating fluid eliminates the need for special hydraulic flu id. Drives are able to utilize simple piston seals j vhose leakage does not contaminate the reactor water but provides cooling for the drive mechanisrm and their seals. The drives are capable of inserting or withdrawing a control rod at a slow, controlled rate, as well as providing rapid insertion when required. A mechanism on the drive locks the control rod at 6-in. increments of stroke over the length of the core. A coupling spud at the top end of the drive index tube (piston rod) engages and locks into a mating socket at the base of the control rod. The weight of the control rod is sufficient to engage and lock this coupling. Once locked, the drive and rod form an integral unit that must be manually unlocked by specific procedures before the components can be separsted. The drive holds its control rod in distinct latch positions until the hydraulic system actuates movement to a new position. Withdrawal of each  ! rod is limited by the seating of the rod in its guide tube. Withd rawal , beyond this position to the overtravel limit can be accomplished only if , the rod and drive are uncoupled. Withdrawal to the over-travel limit is annunciated by an alarm.  ; The individual rod indicators, grourc- one control panel display, correspond to relative rod locati t core. The Display Module is divi >' into two sections. The r. . piny section and a Rod Select l sectic which are physically sup rimposta but independent of each other. For display purposes, the control rods are consicered in groups of four adjacent rods centered arc'n i a common core volume. Each group is monitored by four LPRM strings (Sibsection 7 6.1.6, Neutron Monitoring System). Rod groups at the periphery of the core may have less than four rods. A white light indicates which of the four rods is tn.= one selected for movement. l 4.6-2 Am. No. 56, (3/81)  ; t

ACNGS-PSAR n\ h.6.1.1.2.2 3 rive Components

 /
'\_,e)       Figure 4.6-2 illustrates the operating principle of a drive. Figures 4.6-3 and 4.6 4 illustrate the drive in more detail. The main components of the drive and their functions are described below.

h.6.1.1.2.2.1 Drive Piston The drive piston is rounted at the lower end of the index tube. The function of the index tube is similar to that of a piston rod in a conventional hydraulic cylinder. The drive piston and index tube make up the main moving assembly in the drive. The drive piston operates between positive end stops, with a hydraulic cushion provided at the upper end only. The piston has both inside and outside seal rings and operates in an annular space between an inner cylinder (fixed piston tube) and an outer cylinder (drive cylinder). Because the type of inner seal used is effective in only one direction, the lower sets of seal rings are mounted with one set sealing in each direction. A pair of nonmetallic bushings prevents metal-to-metal contact between the piston assembly and the inner cylinder surface. The cuter piston rings are segmented, step-cut seals with expander springs holding the segments against the cylinder vall. A pair of split bushings on the outside of the piston prevents piston contact with the cylinder vall. The effective piston area for downtravel, or withdrawal, is approximately 1.2 in.2 versus 4.1 in.2 for uptravel, or insertion. This difference in driving area tends f e"%g to balance the control rod weight and assures a highe: force for insertion i f than for withdrawal. V h.6.1 1 2.2.2 Index Tube The index tube is a long hollow shaft made of nitrided stainless steel. Circumferential locking grooves, spaced every 6 in, along the outer surface, transmit the weight of the control rod to the collet assembly. The upper end of the index tube is threaded to receive a coupling spud. The coupling (Figure 4.6-1) accommodates a small amount of angular misalignment between the drive and the control rod. Six spring fingers allow the coupling spud to enter the mating socket on the control rod. A plug then enters the spud and prevents uncoupling. h.6.1.1.2.2 3 Collet Assembly The collet assembly serves as the index tube locking mechanism. It is located in the upper part of the drive unit. This assembly prevents the index tube from accidentally moving downward. The assembly consists of the

           , collet fingers, a return spring, a guide cap, a collet housing (part of the
  -          cylinder. tube, and flange) and the collet piston.

Locking is accomplished by fingers mounted on the collet piston at the top of the drive cylinder. In the locked or latched position the fingers g engage a locking groove in the index tube. ('~'/

   \s_-

4.6-3 Am. No. 56, (3/81)

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ACNGS-PS AR The collet piston is normally held in the latched position by a force of approximately 150 lb supplied by a spring. Metal piston rings are used to seal the collet piston f rom reactor vessel pressure. The collet assembly will not unlatch until the collet fingers are unloaded by a short. automatically sequenced, drive-in signal. A pressure, approximately 1b0 pai above reactor vessel pressure, must then be applied to the collet piston to overcome spring force, slide the collet up against the conical surface in the guide cap, and spread the fingers out so they do not engage a locking groove. A guide cap is fixed in the upper end of the drive assembly. This member provides the unlocking cam surface for the collet fingers and serves as the upper bushing for the index tube. If reactor water is used during a scram to supplement accumulator pressure, it is drawn through a filter on the guide cap. h.6.1.1.2.2.h Piston Tube The piston tube is an inner cylinder, or colurn, extending upward inside the drive piston and index tube. The piston tube is fixed to the bottom flange of the drive and remains stationary. Water is brought to the upper side of the drive piston through this tube. A buffer shaft, at the upper end of the piston tube, supports the stop piston and buffer components. h.6.1.1.2.2 5 Stop Piston A stationary pisten, called the stop piston, is mounted on the upper end of the piston tube. This piston provides the seal betwern reactor vessel pressure and the space . cove the drive piston. It also functions as a positive end stop at the upper limit of control rod travel. Piston rings and bushings, similar to those on the drive piston, are mounted on the upper portion of the stop piston. The lower portion of the stop piston form a thinwalled cylinder containing the buffer piston, its metal seal ring, and the buffer piston return spring. As the drive piston reaches the upper end of the scram stroke, it strikes the buffer piston. A series of orifices in the buffer shaft provides a progressive water shutoff to cushion the buffer piston as it is driven to its limit of travel. The high pressures generuted in the buffer are confined to the cylinder portion of the stop piston and are not applied to the stop piston and drive piston seals. The center tube of the drive mechanism forms a well to contain the position indicator probe. The probe is an aluminum extrusion attached to a cast aluminum housing. Mounted on the e> trusion are hermetically scaled, magnetically operated, reed switches. The entire probe assembly is protected by a tL n-valled stainless steel tube. The switches are actuated by a ring magnet located at the bottom of the drive piston. The drive piston, piston tube and indicator tube are all of nonmagnetic stainlers steel, allowing the individual switches to be operated by the magnet as the piston passes. Two switches are located at each position corresponding to an index tube groove, thus allowing redundant indication at each latching point. Two additional switches are located at each 4.6-4 Am. No. 56, (3/81)

ACNGS-PSAR , mid' nt between latching points to indicate the intermediate positions (T

   \_,)

dur ; drive motion. Thus, indication is provided for each 3 in. of travel. Duplicate switches are provided for the full-in and full-out s positions. Redundant overtravel switches are located at a position below the normal full-out position. Because the limit of downtravel is normally provided by the control rod itself as it reaches the backseat position, the i i drive can pass this position and actuate the overtravel switches only if it is uncoupled from its control rod. A convenient means is thus provided to verify that the drive and control rod are coupled after installation of a  : drive or at any time during plant operation. h.6.1.1.2.2.6 Flange and Cylinder Assembly I A flange and cylinder assembly is made up of a heavy flange welded to the drive cylinder. A sealing surface on the upper face of this flange forms the seal to the drive housing flange. The seals contain reactor pressure and the two hydraulic control pressures. Teflon-coated, stainless steel

• rings are used for these seals. The drive flange contains the integral ball, or two-way, check (ball-shuttle) valve. This valve directs either the reactor vessel pressure or the driving pressure, whichever is higher,

 ,      to the underside of the drive piston. Reactor vessel pressure is admitted to this valve from the annular space between the drive and drive housing through passages in the flange.

I Water used to operate the collet piston passes between the outer tube and the cylinder tube. The inside of the cylinder tube is honed to provide the surface required for the drive piston seals. ( Both the cylinder tube and outer tube are velded to the drive flange. The upper ends of these tubes have a sliding fit to allow for differential expansion. t 4.6.1.1.2.2 7 Uncoupling Rod and Related Parts 1

 ;      Two means of uncoupling are provided. With the reactor vessel head removed, the lock plug can be raised against the spring force of approximately 50 lb by a rod extending up through the center of the control rod to an unlocking handle located above the control rod velocity limiter.                       i The control rod, with the lock plug raised, can then be lifted from the                          !

drive. If it is desired to uncouple a drive without removing the reactor pressure vessel head for access, the lock plug can also be pushed up from below. In this case, the pisten tube assembly is pushed up against the uncoupling ' rod, which raises the lock plug and allows the coupling spud to disengage the socket as the drive piston and index tube are driven down. The control rod is heavy enough to force the spud fingers to enter the socket and push the lock plug up, allowing the spud to enter the socket completely and the plug to snap back into place. Therefore, the drive can be coupled to the control rod using only the weight of the control rod.  ; i i ['* Q' i 4.6-5 Am. No. 56, (3/81) i

l ACNGS-PGAR h.6.1.1.2 3 Materials or construction Factors that determine the choice of construction materials are discussed in the following subsections. h.t.l.l.2 3 1 Index Tube  ! The index tube must withstand the locking and unlocking action of the enllet finge rs . A compatible bearing combination must be provided that is able to withstand moderate n'. salignment forces. Large tensile and column litdn are applied during scram. The reactor environment limits the choice o: ntterials suitable for corrosion resistance. To meet these varied requirements , the index tube in made from annealed, single-phase, nitrogen s t rengt hene d , uustenitic stainless steel. The wear and bearing requirements are provided by nitriding the complete tube. To obtain cuitable corrosion resistance, a carefully controlled process of surface preparation is employed.

• l h.6.1.1.2 3 2 Coupling Spud l

The coupling spud is aide of Inconel X-750 that is aged for maximum Id Waical strength and the required corrosion resistance. Bec ause misalignment tends to cause chafing in the semispherical contact area, the pa rt is protected by a thin chromium plating. This plating also prevents galling of the threads attaching the coupling spud to the index tube. l l h.6.1.1.2 3 3 Collet Fingers l Alloy X-750 is used for the collet ringers , which must function as leaf ( springs when cammed open to the unlocked position. Colconoy 6 hard facing l provides a long wearing surface, adequate for design life, to the area contacting the index tube and unlocking cam surface of the guide cap. i l I h 6.1.1.2 3 4 Seals and Bushings Graphitar-lb (or its equivalent) is selected for neals and bushings on the ' drive piston and stop piston. The material is inert and has a low friction [ coefficient when water-lubricated. Because some loss of Graphitar strength  ! i is experienced at higher temperatures, the drive is supplied with cooling water to hold temperatures below 250*F. The Graphitar-lh is relatively sof t, which is advantageous when an occasional particle of foreign matter l reaches a seal. The resulting scratches in the seal reduce sealing efficiency until worn smooth, but the drive design can tolerate considerable water leakage past the seals into the reactor vessel. I h.6.1.1.2 3 5 Summary All drive components exposed to reactor vessel water are made of austenitic stainless steel except the following: r l l (1) Seals and bushings on the drive piston and stop piston are Graphitar-14 (or its equivalent) O ; i 4.6-6 Am. No. 56, (3/81) , l

ACNGS-PSAR (2) All springs and members requiring spring action (collet fingers,

   /     )/      coupling spud, and spring washers) are made of Alloy X-750.
   \

U (3) The ball check valve is a Haynes Stellite cobalt-base alloy (or its equivalent). (4) Elastomeric 0-ring seals are ethylene propylene. r (5) Metal piston rings are Haynes 25 alloy. (6) Certain wear surfaces are hard-faced with Colmonoy 6. (7) Nitriding and chromium plating are used in certain areas where resistance to abrasion is necessary. (8) The drive pistca head, stop piston, buffer shaf t and buffer piston are rade of Armco 17 4 PH. (9) Certain fasteners and locking devices are made of Inconel X-750 or 600. Pressure-containing portions of the drives are designed and fabricated in accordance with requirements of Section III of the ASME Boiler and Pressure - Vessel Code. h.6.1.1.2.4 Control Rod Drive Hydraulic System

      ~,

The CRD hydraulic system (Figures 4.6-Sa, b) supplies and controls the

,v (        pressure and flow to and from the drives through hydraulic control units (HCU). The water discharged from the drives during a scram flows through                   !-

the HCUs to the scram discharge volume. The water discharged from a drive i during a normal control rod positioning operation flows through the HCU, l' I the exhaust head,_r, and is returned to the reactor vessel via the HCUs of nonmoving drives. There are an many HCUs as the number of control rod drives. , k.6.1.1.2.4.1 Hydraulic Requirements The CRD hydraulic system design is shown in Figures 4.6-Sa, b, and c, and , L.6-6. The hydraulic requirements, identified by the function they perform are as follows: (1) An accumulator hydraulic charging pressure of approximately 1750 to  ; 2000 psig is required. Flow to the accumulators is required only during scram reset or system startup.  ; (2) Drive pressure of approximately 260 psi above reactor vessel pressure is required. A flow rate of approximately h gpm to insert each control rod and 2 gpm to withdrav each control rod is required. (3) Cooling water to the drives is required at approximately 20 psi above

reactor vessel pressure and at a flow rate of approximately 0 34 gpm per drive unit.

J l L. 4.6-7 Am. No. 56, (3 /81',

AC f;GS-PGAll (h) ihe scram discharge volume is sized to receive, and contain, all the i I vater discharged by the drives during a scram; a minimtun volume of 3 34 cal. per drive is required (excluding the instrument volume ). 1 h.6.1.1.2.4.2 System Description  ; The CitD hydraulic system provides the required functions with the pumps, filters, valves, instrumentation and piping shown in Figures 4.6-Sa, b, and c, and descrited in the following paragraphs. Duplicate components are included, where necessary, to assure continuous system operation if an in-service component requires enintenance. h.6.1.1.2.h.2.1 Supply Pump One supply pump pressurizes the system with water from the condensate treatment system and/or condensate storage tanks. One spare pump is provided for standby. A discharge check valve prevents backflow through the nonoperating pump. A portion of the pump discharge flow is diverted through a minimum flow bypass line to the condensate storage tank. This flow is controlled by an orifice and is suf ficient to prevent pump dannge if the pump discharge is inadvertently closed. Condensate water is processed by two filters in the system. The pump suction filter is a disposable element type with a 25-micron absolute rating. A 250-micron strainer in the filter bypass line protects the pump when the filters are being serviced. The drive water filter, downstream of the pump, is a cleanable element type with a 50-micron absolute rating. A differential pressure indicator and control room alarm monitor the filter element as it collects foreign materials. h.6.1.1.2.h.2.2 Accumulator Charging Pressure Accumulator charging pressure is established by precharging the nitrogen accumulator to a precisely controlled pressure at known temperature. During scram, the scram inlet (and outlet) valves open and permit the stored energy in the accumulators to discharge into the drives. The resulting pressure decrease in the charging water header allows the CED supply pump to "run out" (i.e. , flow rate to increase substantially) into the control rod drives via the charging water header. The flow element upstream of the accumulator charging header senses high flow and provides a signal to the manual auto-flow control station which in turn closes the system flow control valve. This action maintains increased flow through the charging water header, while avoiding prolonged pump operation at "run-out" conditions. Pressure in the charging header is monitored in the control rocm with a pressure indicator and low pressure alarm. During normal operation, the flcv control valve rnintains a cnstant system flow rate. This flow is used for drive operation and drive cooling. O 4.6-8 Am. No. 56, (3/81)

ACNGS-PSAR k.6.1.1.2.4.2.3 Drive Water Fressure f% ('- ') Drive water pressure required in the drive header is maintained by the drive pressure control valve, which is manually adjusted from the control room. A flow rate of approximately 16 gpm (the suu of the flow rate required to insert 4 control rods) normally passes from the drive water pressure stage through eight solenoid-operated stabilizing valves (arranged in parallel) into the cooling water header. The flow through two stabilizing valves equals the drive insert flow for one drive; that of one stabilizing valve equals the drive withdrawal flow for one drive. When operating a drive (s), the required flow is diverted to the drives by closing the appropriate stabilizing valves, at the same time opening the drive directional control and exhaust solenoid valves. Thus, flow through the drive pressure control valve is always constant. Flow indicators in the drive water header and in the line downstream from the stabilizing valves allow the flow rate through the stabilizing valves to be adjusted when necessary. Differential pressure between the reactor vessel and the drive pressure stage is indicated in the control room. 4.6.1.1.2.4.2.4 Cooling Water Header The cooling water header is located downstream from the drive / cooling pressure valve. The drive / cooling pressure control valve is manually adjusted frou the control room to produce the required drive / cooling water pressure balance. /' 'h The flow through the flow control valve is virtually constant. Therefore, ( ,/ cace adjusted, the drive / cooling pressure control valve will maintain the correct drive pressure and cooling water pressure, independent of reactor vessel pressure. Changes in setting of the pressure control valves are required only to adjust for changes in the cooling requirements of the drives, as the drive seal characteristics change with time. A flow indicator in the control room monitors cooling water flow. A differential - pressure indicator in the control room indicates the difference between reactor vessel pressure and drive cooling water pressure. Although the drives can function without cooling water, seal life is shortened by long-term exposure to reactor temperatures. The temperature of each drive is indicated and recorded, and excessive tempertures are annunciated in the control room.  ; 4.6.1.1.2.4.2 5 Scram Discharge Volume The scram discharge volume consists of header piping which connects to each . HCU and drains into a dedicated instrument volume. The header piping is ' sized to receive and contain all the water discharged by the drives during a scram, independent of the instrument volume. During normal plant operation, the scram discharge volume is empty and vented to atmosphere through redundant open vent and drain valve. When a scram occurs, upon a signal from the safety circuit these vent and drain valves are closed to conserve reactor water. Lights in the control room s indicate the position of these valves. \ 4 N ,/ ' 1 4.6-9 Am. No. 56, (3/81)

ACNGC-PSAR Durind a scram, the scram discharge volume partly fills with water discharded fro- above the drive pistons. After scram is completed, the CRD seal leakage from the reactor continues to flow into the scran discharge volume until the discharge volume pressure equals the reactor vesnel pressure. A check valve in each HCU prevents reverse flow from the scram discharde header volume to the drive. When the initial scram signal is cleared from the reactor protection system (RPS), the scram discharge volune signal is overridden with a keylock override switch, and the scram dischrrde volute in drained and returned to atmospheric pressure. Hemote uanual switches in the pilat valve solenoid circuits allow the discharge vo h me vent and drain valves to be tested without disturbing the RPS. Closing the scran discharge volume valves allows the outlet scram valve seats to be leak-tested by timing the accumulation of leakage inside the scrau discharge volume. There is a separate scram discharde volume for the west and east banks of HCU's. Eight liquid-level switches activated by six transmitters evenly divided between the two instrument volumes, monitor the volume for abnormal water level. In addition four self-actuating switches, 2 attached to each volume provide a redundant, diverse monitor of volume water level. They are set at three different levels. At the lowest level, two switches actuate to indicate that the volume is not completely empty during post-scram draining or to indicate that the volume starts to fill through leakage accumulation at other times during reactor operation. At the second level, two switches produce a rod withdrawal block to prevent further withdrawal of any control rod when leakage accumulates to half the capacity of the instrument volume. The remaining four switches and four self-actuating switches are interconnected with the trip channels of the Reactor Trip System and will initiate a reactor scram should water accumulation fill the instrument volume. 4.6.1.1.2.4.3 Hydraulic Control Units Each hydraulic control unit (HCU) furnishes pressurized w .ter, on signal, to a drive unit. The drive then positions its control rod as required. Operation of the electrical system that supplies scram and normal control rod positioning signals to the HCU is described in Subsection 7.7.1.1 (Rod Control and Information System). The basic components in each HCU are: (1) manual, pneumatic and electrical valves; (2) an accuuulator; (3) related piping; (h) electrical connections; (5) filteru; and (6) instrumentation (Figures 4.6-5a, 4.6-Sb, k.6-5c, b.6-6, 4.6.7, and 4.6-8). The components and their functions are described in the following paragraphs. h.6.1.1.2.4.3 1 Insert Drive Valve The insert drive valve is solenoid-operated and opens on an insert signal. The valve supplies drive water to the bottom side of the main drive piston. 4.6.1.1.2.h.3 2 Insert Exhaust Valve 4.6-10 Am. No. 56, (3/81) O

t ACNGS-PSAR

                                                                                      ;
  ,em    The insert exhaust solenoid valve also opens on an insert signal. The        l l     }

valve discharges water from above the drive piston to the exhaust water

  \s_,/  header.

h.6.1.1.2.4.3.3 Withdraw Drive Valve The withdraw drive valve is solenoid-operated and opens on a withdraw si6nal. The valve supplies drive water to the to, of the drive piston. 4.6.1.1.2.4.3.4 Withdraw Exhaust Valve The solenoid-operated withdraw exhaust valve opens on a withdraw signal and  ! discharges water from below the main drive piston to the exhaust header. It also serves as the settle valve, which opens, following any normal drive novement (insert or withdraw), to allow the control rod and its drive to settle back into the nearest latch position. h.6.1.1.2.h.3 5 Speed Control Units t The insert drive valve and withdraw exhaust valve each have a speed control adjustment. The speed control unit regulates the control rod insertion and withdrawal rates during normal operation. The manually adjustable flow control unit is used to regulate the water flow to and from the volume beneath the main drive piston. A correctly adjusted valve does not require readjustment except to compensate for changes in drive seal leakage.

   <~s   4.6.1.1.2.h.3.6     Scram Pilot Valve Assembly
 'N  '#  The scram pilot valve assembly is operated from the RPS. The scram pilot valve assembly, with two solenoids, controls both the scram inlet valve and  ,

the scram exhaust valve. The scram pilot valve assembly is solenoid-operated and is normally energized. On loss of electrical signal , to the solenoids, such as the loss of external a-c power, the inlet port closes and the exhaust port opens. The pilot valve assembly (Figures 4.6-Sa and b) is designed so that the trip system signal must be removed r from both solenoids before air pressure can be discharged from the scram valve operators. This prevents inadvertent scram of a single drive in the . event of a failure of one of the pilot valve solenoids.  ! 4.6.1.1.2.h.3.7 Scram Inlet Valve The scram inlet valve opens to supply pressurized water to the bottom of the drive piston. This quick opening globe valve is operated by an internal spring and system pressure. It is closed by air pressure applied to the top of its diaphragm operator. A position indicator switch on this valve energizes a light in the control room as soon as the valve starts to open. h.6.1.1.2.4.3.8 Scram Exhaust Valve The scram exhaust valve opens slightly before the scram inlet valve, exhausting water from above the drive piston. The exhaust valve opens f) ( / faster than the inlet valve because of the higher spring setting in the valve operator. v i I 4.6-11 Am. No. S', (3/81) l i

ACNG3-PSAR h.6.1.1.2.4.3 9 Scram Accurulator The scram accumulator stores suf ficient energy to fully insert a control rod at any vessel pressure. The accum21ator is a hydraulic cylinder with a free-floating piston. The piston _sparates the water on top from the nitrogen below. A check valve in the accumulator charcing line prevents loss of water pressure in the event supply pressure is lost. During normal plant operation, the accunnlator piston is seated at the bottom of its cylinder. Loss of nitregen decreases the nitrogen pressure, which actuates a pressure switch and counds an alarm in the control room. To ensure that the accumulator is always able to produce a scram, it is continuously monitored for water leakage. A float-type level switch actuates an alarm if water leaks past the piston barrier and collects in the accumulator instrumentation block. 4.6.1.1.2 5 Control Rod Drive System operation The Control Rod Drive System (CRDS) performs rod insertion, rod withdrawal and scram. These operational functions are described in the following sections. h.6.1.1.2 5 1 Rod Insertion Rod insertion is initiated by a signal from the operator to the insert valve solenoids. This signal causes both insert valves to open. The insert drive valve applies reactor pressure plus approximately 90 psi to the bottom of the drive piston. The insert exhaust valve allows water from above the drive piston to discharge to the exhaust header. As shown in Figure 4.6-3, the locking mechanism is a ratchet-type device and does not interfere with rod insertion. The speed at which the drive moves is determined by the flow tnrough the insert speed control valve, which is set for approximately h gpm for a shim speed (nonscrara operation) of 3 in./sec. During normal insertion, the pressure on the downstream side of the speed control valve is 90 to 100 psi above reactor vessel pressure. However, if the drive slows for any reason, the flow through and pressure drop e. cross the insert speed control valve will dec esse; the full dif terential pressure (240 psi) will then be available to cause continued insertion. With 260-psi differential pressure acting on the drive piston, the piston exerts ar. upward force of 10h0 lb. h.6.1.1.2 5 2 Rod Withdrawal Rod withdrawal is, by design, more involved than insertion. The collet l finger (latch) must be raised to reach the unlocked position (Figure 4.6-3). The notches in the index tube and the ecllet fingers are shaped so that the downward force on the index tube holds the collet fingers in place. The index tube must be lifted before the collet fingers can be re lea sed . This is done by opening the drive insert valves (in the manner described in the preceding paragraph) for approximately 1 sec. The withdrau valves are then opened, applying driving pressure above the drive pisten and opening the area below the piston to the exhaust header. 4.6-12 Am. No. 56, (3/81)

ACNGS-PSAR Pressure is simultaneously applied to the collet piston. As the piston [ ) raises, the collet fingers are cammed outward, away from the index tube, by

 \s_,/ the guide cap.

J The pressure required to release the latch is set and maintained at a level high enough to overcome the force of the latch return spring plus the force of reactor pressure opposing movement of the collet piston; when this occurs, the index tube is unlatched and free to move in the withdraw direction. Water displaced by the drive piston flows out through the withdraw speed control valve, which is set to give the control rod a shim speed of 3 in./sec. The entire valving sequence is automatically controlled and is initiated by a single operation of the rod withdraw switch. h.6.1.1.2 5 3 Scram During a scram, the scram pilot valve assembly and scram valves are operated as previevete described. With the scram valves open, accumulator pressure is admitted under the drive piston, and the area over the drive , piston is vented to the scram disch ge volume. The large differential pressure (approximately 1750 psi, initially and alwa/s several hundred psi, depending on reactor vessel pressure) produces a large upward force on the index tube and control rod. This force gives the rod a high initial acceleration and provides a large margin of force to overcome friction. After the initial acceleration is achieved, the drive s continues at a diminishing velocity. This characteristic provides a high g initial rod insertion rate. As the drive piston nears the top of its

  '--  stroke, the piston reaches the buffe and the driveline is brought to a stop at the full-in position.

Prior to a scram signal, the accu =ulator in the HCU has 1750-2000 psig on the water side and 1750 psig on the nitrogen side. As the inlet scram valve opens, the full water side pressure is available at the CRD acting on a 4.1 in.2 area. As C'D motion begins, this pressure drops to the gas side pressure less line losses between the accumulator and the CRD. When the drive reaches the full-in position, the accumulator completely discharges with a resulting gas side pressure of approximately 1200 psig. The CRD accumulators are necessary to scram the control rods within the required time. Each drive, however, has an internal ball-check valve which allows reactor pressure to be admitted under the drive piston. If the reactor is above 600 psi, this valve ensures rod insertion in the event the accumulator is not charged or the inlet scram valve fails to open. The insertion time, howe 7er, will be slower than the scram time with a properly functioning scram system. The CRDS, with accumulators, provides the following scram performances at full power cperation, in terms of average elapsed time after the opening of the RPS trip actuator (scram signal) for the drives to attain the scram strokes listed: V 4.6-13 Am. No. 56, (3/81)

ACNGS-PSAR From Full-Out (Notch Position 48) To: Notch Position hk 28 12 Stroke (in.) 12 60 108 Time (eec) 0.28 0 91 1.620 h.6.1.1.2.6 Instrumentation 7.le instrumentation for both the control rods and control rod drives is defined by that given for the rod control and information system. The objective of the rod control and information system is to provide the operator with the means to make changes in nuclear reactivity so that reactor power level and power distribution can be controlled. The system allows the operator to manipulate control rods. The design bases and further discussion are covered in Chapter 7, " Instrumentation and Control System." 4.6.1.2 Control Rod Drive Housing Supports 4.6.1.2.1 Safety Objective The control rod drive (CRD) housing supports prevent any significant nuclear transient in the e rent a drive housing breaks or separates from the bottom of the reactor vessel. 4.6.1.2.2 Safety Design Bases The CRD housing supports shall meet the following safety design bases: (1) Following a postulated CRD housing failure, control rod downward motion shall be limited so that any resulting nuclear transient could, not be sufficient to cause fuel damge. (2) The clearance between the CRD housings and the supports shall be sufficient to prevent vertical contact stresses caused by therml expansion during plant operation. h.6.1.2 3 Description The CRD housing supports are shown in Figure 4.6-8. Horizonta] beams are installed immediately b' low the bottom head of the reactor vessel, between the rows of CRD housings. "'he beams are velded to brackets which are welded to the steel form liner of the drive room in the reactor support pedestal. Han6er rods, approximately 10 ft long and 1-3/4 in. in diameter, are supported from the beams on stacks of disc springs. These springs compress approximately 2 in. under the design load. The support bars are bolted between the bottom ends of the hanger rods. The spring pivote at the top, and the beveled, loose fitting ends on the support barn prevent substantial bending moment in the hanger rods if the support barc are overloaded. 4.6-1d Am. No. 56, (3/81)

ACNGS-PSAR 6 ss Individual grids rest on the support bars between adjacent beams. 3ecause [ ') a single-piece grid would be difficult to handle in the lit ited turk space l

 \s_/~    and because it is necessary that control rod drives, positica indicators, and in-core instrumentation components be accessible for inslection and maintenance, each grid is designed for in-place assembly or disassembly.

Each grid assembly is made from two grid plates, a clamp, and a bolt. The top part of the clamp guides the grid to its correct position directly below the respective CRD housing that it would support in the postulated accident. Wher the support bars and grids are installed, a gap of approximately 1 in. l at room temperature (approximately 70*F) is provided between the grid and the bottom contact surface of the CRD flange. During system heatup, this gap is reduced by a net downward expansion of the housings with respect to the supports. In the hot operating condition , the gap is approximately 3/h inch. In the postulated CBD housing failure, the CRD housing supports are loaded when the lower contact surface of the CRD flange contacts the grid. The resulting load is then carried by two grid plates, two support bars, four hanger rods, their disc springs, and two adjacent beams. The American Institute of Steel Construction ( AISC) Manaal of Steel Construction (Specificatior 'ar the Design, Fabrication and Erection of Structural Steel for Buildi- was used in designing the CRD housing support system. However, te vide a structure that absorbs as much f r~'s energy as practical without yielding, the allovable tension and bending stresses used were 90% of yield and the shear stress used was 60% of yield. ( ') These design stressen are 1 5 times the AISC allowable stresses (60% and 40% of yield, respectively). For purposes of mechanical decign , the postulated failure resulting in the highest forces is an instantaneous circumferential separation of the CRD housing from the reactor vessel, with the reactor at an operating pressure of lu86 psig (at the botton of the vessel) acting on the area of the separated housing. The weight of the separated housing, CRD and blade, - plus the pressure of 1086 psig acting on the area of the separated housing, gives a force of approximately 32,000 lb. This force is used to calculate the impact force, conservatively assuming that the housing travels through a 1-in. gap before it contacts the supports. The impact force (109,000 lb) l is then treated as a static load in design. The CRD housing supports are designed as category I (seismic) equipment in accordance with Section 3 2.  ; 4.6.2 Evaluations of the CRDS h.6.2.1 Failure Mode and Effmts Analysis This subject is covered in Appendix B. h.6.2.2 Protection from Commen Moc.e Failures The applicant's position on this subject is covered in Appendix B.

O) 4.6-15 Am. No. 56, (3/81)

ACNGS-PSAh h.6.2 3 Safctf Evaluation Safety evaluation of the control rods, CRDS, and CBD housing supports is dercribed below. Further description of control rods is contained in Section 4.2. h.6.2 3 1 Control Rods l 4.6.2 3 1.1 Materials Adequacy Throughout Design Lifetime l The adequacy of the materials throu6 h out the design life was evaluated in the mechanical design of the control rods. The primry mterials, Phc powder and Type-30h austenitic stainless stee?, have been found suitable in meeting the demands of the BWR environment. l h.6.2 3 1.2 Dimensional and Tolerance Analysis Layout studies are done to assure that, gisen the worst combination of part tolerance ranges at assembly, no interference exists which will restrict ! the passage of control rods. In addition, preopt -ational verification is l mde on each control blade system to show that the acceptable levels of operational performnce are met. I h.6.2.3 1 3 Therml Analysis of the Tendency to Warp i The various parts of the control rod assembly remain at approximtely the l same temperature during reactor operation, negating the problem of l l distortion or warpage. What little differential thermal grosth could exist i is allowed for in the mechanical design. A minimum axial gap is maintained between absorber rod tubes and the control rod frame assembly for the purpose. In addition, to further this end, dissimilar metals are avoided, i l h.6.2 3 1.4 Forces for Expulsion l An analysis has been performed which evaluates the mximum pressure forces ( vhich could tend to eject a control rod from the core. The results of this analysis are given in Subsection h.6.2 3 2.2.2, (Rupture of Hydraulic Line(s) to Drive Housing Flange). In summary , if the collet were to remain open, which is unlikely, calculations indicate that the steady-state control rod withdrawal velocity would be 2 ft/sec for a pressure-under line break, the l hiting case for rod withdrawal. 4.6.2 3 1.9 Punctional Failure of Critical Components The consequences of a functional failure of critical components have been evaluated and the results are covered in Gubsection 4.6.2 3 2.2 ( Analysis of Malfunction Pelating to Rod Withdrawal). h.6.2 3 1.6 Precluding Excessive Rates of Reactivity Addition 1 In order to chew that excessive rates of reactivity addition are precluded, analysis has been performed both on the velocity limiter device and the l effect of postulated critical component failures (Subsection 4.6.2 3 2.2). 4.6-16 Am. No. 56, (3/81)

ACNGS-PSAR h.6.2 3 1 7 Effect of Fuel Rod Failure on Control Rod Channel Clearances I A The CRD mechanical design ensures a sufficiently rapid insertion of control 2 ( ,/ rods to preclude the occurrence of fuel rod failures which could hinder reactor shutdown by causing sigr.ificant distortions in channel clearances. I h.6.2 3 1.8  !!echanical Damage In addition to the analysis performed on the CRD (Subsection h.6.2 3 2.2) and Subsection 4.6.2 3 2 3 (Scram Reliability) and the control rod blade, analyses were performed on the control rod guide tube (see Subsections 4.2 3 3 7 through 4.2 3 3 8 for these analyses). 4.6.2 3 1 9 Evaluation of Control Rod Velocity Limiter The control rod velocity limiter limits the free-fall velocity of the control rod to a value that cannot result in nuclear system process barrier damage. This velocity is evaluated by the rod drop accident analysis in I Chapter 15 ( Accident Analysis). h.6.2 3 2 Control Rod Drives 4.6.2 3 2.1 Evaluation of Scram Time The rod scram function of the CRD system provides the negative reactivity insertion required by safety design basis 4.6.1.1.1.1.l(1). The scram time gs shown in the description is adequate as shown by the transient analyses of ( Chapter 15 v 4.6.2 3 2.2 Analysis of Malfunction Relating to Rod Withdrawal There are no known single malfunctions that cause the unplanned withdrawal of even a single control rod. However, if multiple malfunctions are postulated, studies show that an unplanned rod withdrawal can occur at withdrawal speeds that vary with the combination of malfunctions postulated. In all cases, the subsequent withdrawal speeds are less than that assumed in the rod drop accident analysis discussed in Chapter 15 Therefore, the physical and radiological consequences of such rod withdrawals are less than those analyzed in the rod drop accident. h.6.2 3 2.2.1 Drive Housing Fails at Attachment Weld The bottom he9A of the reactor vessel has a penetration for each CRD location. A drive housing is raised into position inside each penetration and fastened by welding. The drive is raised into the drive housing and bolted to a flange at the bottom of the housing. The CRD housing material at the vessel penetration is seamless, Alloy 600 tubing with a minimum tensile strength of 80,000 psi, and Type-316L stainless steel pipe below the vessel with a minimum strength of 70,000 psi. The basic failure considered here is a complete circumferential crack through the housing wall at an elevation just below the J-weld.

 \s /

4.6-17 Am. No. 56, (3/81)

ACHGS-PSAR Static loads on the housing vall include the weight of the drive and the control rod, the weight of the housing below the J-veld, and the reactor pressure acting on the 6-in. diameter cross-sectional area of the housing and the drive. Dynamic loading results from the reaction force during drive operation. If the housing were to fail as described, the following sequence of events is foreseen. The housing would separate from the vessel. The CRD and housing would be blown downward against the support structure, by reactor pressure acting on the cross-sectional area of the housing and the drive. The downward motion of the drive and associated parts would be determined by the gap between the bottom of the drive and the support structure and by the deflection of the support structure under load. In the current design, maximum deflection is approximately 3 in. If the collet were to remain latched, no further control rod ejection vould occurl; the housing would not drop far enough to clear the vessel penetration; reactor water would leek at a rate of approximately 180 gpm through the 0.03-in. diametral clearance between the housing and the vessel penetration. If the basic housing failure were to occur while the control rod is being withd: avn (this is a small fraction of the total drive operating time) and if the collet were to stay unlatched, the following sequence of events is foreseen. The housing would separate from the vessel; the drive and housing would be blown downward against the CRD housing support. Calculations indicate that the steady-state rod withdrawal velocity would be 0 3 ft/sec. During withdrawal, pressure under the collet piston would be approximately 250 psi greater than the pressure over it. Therefore, the collet would be held in the unlatched position until ariving pressure was removed from the pressure-over port. h.6.2 3 2.2.2 Rupture of Hydraulic Line(s) to Drive Housing Flange There are three types of possible rupture of hydraulic lines to the drive housing flange: (1) pressure-under (insert) line break; (2) pressure-over (withdrawn) line break; and (3) coincident breakage of both of these lines. 4.6.2 3 2.2.2.1 Pressure-under (Insert) Line Break For the case of a pressure-under (insert ) line break, a partial or complete circumferential opening is postulated at or near the point where the line enters the housing flange. Failure is more likely to occur after another basic failure wherein the drive housing or housing flange separates from the reactor vessel. Failure of the housing, however, does not necessarily lead directly to failure of the hydraulic lines. If the pressure-under (insert ) line were to fail and if the collet were latched, no control rod withdrawal vould occur. There would be no pressure differential across the collet piston and, therefore, no tendency to unlatch the collet. Consequently, the associated control rod could not be withdrawn; but if reactor pressure is greater than 600 psig, it will insert on a scram signal. The ball check valve is designed to seal off a broken pressure-under line by using reactor pressure to shift the check ball to its upper seat. If 4.6-18 Am. No. 56, (3/81)

ACNGS-PSAR the ball check valve were prevented from seating, reactor water would leak [T to the containment. Because of the broken line, cooling water could not be ( ,)_ supplied to the drive involved. Loss of cooling water would cause no immediate damage to the drive. However, prolonged exposure of the drive to

                                                                                                  ;

tOmperatures at or near reactor temperature could lead to deterioration of material in the seals. High temperature would be indicated to the operator I by the thermocouple in the position indicator probe. A second indication would be high cooling water flow. If the basic line failure were to occur while the control rod is being withdrawn, the hydraulic force would not be sufficient to hold the collet open, and spring force normally would cause the col:et to latch and stop rod withdrawal. However, if the collet were to remain open, calculations indicate that the steady-state control rod withdrawal velocity would be 2 ft/sec. h.6.2 3 2.2.2.2 Pressure-over (Withdrawn) Line Break , The case of tne pressure-over (withdrawn) line breakage considers the complete breakage of the line at or near the point where it enters the housing flange. If the line vete to break, pressure over the drive piston would drop f rom reactor pressur e to atmoe ,,heric pressure. Any significant reactor pressure (approximately 600 psig or greater) would act on the bottom of the drive piston and fully insert the drive. Insertion would occur regardless of the operational mode at the time of the failure. After full insertion, reactor water would leak past the stop piston seals. This

  -     leakage vould exhaust to the containment through the broken pressure-over

[s) x_ - line. The leakage rate at 1000 psi reactor pressure is estimated to be 1 to 3 gpm; however, with the Graphitar-lh (or its equivalent) seals of the stop piston removed, the leakage rate could be as high as 10 gpm, based on experimental measurements. If the reactor were hot, drive temperature would increase. This situation would be indicated to the reactor operator by the drift alarm, by the fully inserted drive, by a high drive temperature annunciated in the control room and by operation of the drywell sump pump. h.6.2 3 2.2.2 3 Simultaneous Breakage of the Pressure-over (Withdrawn) and Pressure-under (Insert) Lines l For the simultaneous breakage of the pressure-over (withdrawn) and

• pr :ssure-under (insert } lines, pressures above and below the drive pisten  !

vould drop to zero, and the ball check valve vould close the broken pressure-under line. Reactor water would flow from the annulus outside the l drive, through the vessel ports, and to the space below the drive piston. As in the case of pressure-over line breakage, the drive would then insert (at reactor pressure approximately 600 psig or greater) at a speed dependent on reactor pressure. Full insertion vould occur regardless of ' the operational mode at the time of failure. Reactor water would leak past the drive seals and out the broken pressure-over line to the containment, as de scribed above. Drive temperture vould increase. Indication in the contr ol room would include the drift alarm, the fully inserted drive, the high a.-ive temperature annunciated in the control com, and the operation of the dz7vell sump pump.  ; \ i V 4.6-19 Am. No. 56, (3/81)

ACfiGS-PSAR l l l h.6.2 3 2.2 3 All Drive Flange Eolts Fail in Tension  ! Each CRD is bolted to a flange at the bottom of a drive housing. The h flange is velded to the drive housing. Bolts are unde of AISI blh0 steel, with a minimum tensile strength of 125,000 psi. Each bolt has an allowable load capacity of 15,200 lb. Capacity of the 8 bolts is 121,600 lb. As a result of the reactor design pressure of 1250 psig, the major load on all 8 bolts is 30,h00 lb. If a progressive or simultaneous failure of all bolts were to occur, the drive would separate from the housing. The control rod and the drive would be blown downward against the support structure. Impact velocity and suppnet structure loading would be slightly less than that for drive housing failure, because reactor pressure would act on the drive cross-sectional area only and the housing would remain attached to the reactor vessel. The drive would be isolated from the cooling water supply. Reactor water would flow downward past the velocity limiter piston, through the large drive filter, and into the annular space between the thermal sleeve and the drive. For worst-case leakage calculations, the large filter is assumed to be deformed or swept out of the way so it would offer no uignificant flow restriction. At a point near the top of the annulus, where pressure would have dropped to 350 psi, the water would flash to steam and cause choke-flow conditions. Steam would flow down the annulus and out the space between the housing and the drive flanges to the drywell. Steam formation would limit the leakage rate to approximately 8h0 gpm. If the collet were latched, control rod ejection would be limited to the distance the drive can drop before coming to rest on the support structure. There would be no tendency for the collet to unlatch, because pressure below the collet pisten would drop to zero. Pressure forces, in fact , exert 1435 lb to hold the collet in the latched losition. If the bolts fatied during control rod withdrawal, pressure below the collet piston would drop to zero. The collet, with 1650 lb return force, would latch and stop rod withdrawal. h.6.2 3 2.2.h Weld Joining Flange to Housing Fails in Tension The failure considered is a crack in or near the veld that joins the flange to the housing. This crach extends through the wall and completely around the housing. The flange material is forged, Type-316L stainless steel, with a minimum tensile strength of 70,000 psi. The housing rnterial is seamless, Type-316L stainless steel pipe, with a minimum tensile strength of 70,000 psi. The conventional, full penetration weld of Type-208 stainless uteel has a minimum tensile strength approximately the same as that for the parent metal. The design pressure and temperature are 1250 poig and 575*F. Reactor pressure acting or. the cross-sectional area of the drive , the weight of the control rod , drive , and flange, and ti. dynamic reaction force during drive operation result in a niximum tensile stress at the veld of approxinntely 5100 psi. If the basic flange-to-housing joint failure occurred, the flange and the attached drive would be blown downward against the support structure. The support structure loading would be slightly less than that for drive Am. No. 56, (3/81) 4.6-20

ACNGS-PSAR i housing failure, because reactor pressure would act only on the drive [' cross-sectional area. Lack of differential pressure across the collet (,_,h) piston would cause the collet to remain latched and limit control rod  ; motion to approximately 3 inches. Downward drive movement would be small; i therefore, most of the drive vould remain inside the housing. The pressure-under and pressure-over lines are flexible enough to withstand the small displacement and remain attached to the flange. Reactor water would follow the same leakage path described above for the flange-bolt failure, except that exit to the dryvell vould be through the gap between the lover end of the housing and the top of the flange. Water would flash to steam in the annulus surrounding the drive. The leakage rate vould be approximately 8h0 gpm. If the basic failure were to occur during control rod withdrawal (a small fraction of the total operating time) and if the collet were held unlatched, the flange would separate from the housing. The drive and flange would be blown downward against the support structure. The calculated steady-state rod withdrawal velocity would be 0.13 ft/sec. Because pressure-under and pressure-over lines remain intact, driving water pressure would continue to the drive, and the normal exhaust line restriction would exist. The pressure below the velocity limiter piston would drop below normal as a result of leakage from the gap between the housing and the flange. This differential pressure across the velocity limiter piston would result in a net downward force of approximately 70 lb. ' Leakage cut of the housing would greatly reduce the pressure in the annulus surrounding the drive. Thus, the net downward force on the drive piston

  ,g     would be less than normal. The overall effect of these events would be to

( ) reduce rod withdrawal to approximately one-half of normal speed. With a

 \._,/

560-psi differential across the collet piston, the collet vould remain unlatched; however, it should relatch as soon as the drive signal is removed. h.6.2 3 2.2 5 Housing Wall Ruptures This failure is a vertical split in the drive housing vall just below the bottom head of the reactor vessel. The flow area of the hole is considered equivalent to the annular area between the drive and thermal sleeve. Thus, flow through this annular area, rather than flow through the hole in the housing, vould govern 'eakage flow. The CRD housing is made of Alloy 600 seamless tubing (at *.he penetration to the vessel), with a minimum tensile  ! strength of 80,000 psi, and of Type-316L stainless steel seamless pipe  !' below the vessel with a minimum tensile strength of 70,000 pai. The maximum hoop stress or 9,000 psi results primarily from the reactor design r pressure (1250 psig) acting on the inside of the housing. If such a rupture were to occur,12 actor water would flash to steam and leak through the hole in the housing to the dryvell at approximately 1030 gpm. Choke-flow conditions vould exist, as described previously for the flange-bolt failure. However, leakage flow would be greater because flow resistance would be less, that is, the leaking water and steam would not have to flow down the length of the housing to reach the dryvell. A critical pressure of 350 psi causes the water to flash to steam. O)

 \u.-

4.6-21 Am. No. 56, (3/81)

ACNGS-PSAR There would be no pressure differential acting across the collet piston to unlatch the collet; but the drive would insert ac a result of loss of pressure in the drive housing causing a pressure drop in the space above the drive piston. If this failure occurred during control rod withdrawal, drive withdraval would stop, but the collet would remain unlatched. The drive would be stopped by a reduction of tre net downward force action on the drive line. The net force reduction wcald occur when the leakage flow of 1030 gpm reduces the pressure in the annulus outside the drive to approximately 540 psig, thereby reducing the pressure acting on top of the drive piston to the same value. A pressure differentia of approximately 71C psi would exist across the collet piston and hold the collet unlatched as long as the operator held the withdraw signal. h.6.2 3 2.2.6 Flange Plug Blows Out To connect the vessel ports with the bottom of the ball check valve, a hole of 3/h-in. diameter is drilled in the drive flange. The outer end of this hole is sealed with a plug of 0.812-in. diameter and 0.25-in. thickness. ' rull-penetration, Type-308 stainless steel weld holds the plug in place. The postulated failure is a full circumferential crack in this weld and subsequent blowout of the plug. If the weld were to fail, the plug were to blow out and the collet rem 'ned latched, there would be no control rod motion. Thera would be no pressi re differential acting across the collet piston to unlatch the collet. Reactor water would leak past the velocity limiter piston, down the annulus between the drive and the thermal sleeve, through the vessel Iorts and drilled passage, and out the open plug hole to the drywell at approximately 320 cpm. Leakage calculations assume only liquid flows from the flange. Actually, hot reactor water would flash to steam, and choke-flow conditions would exist. Thus, the expected leakage rate would be lower than the calculated value. Drive temperature would increase and initiate an alarm in the control room. If this failure were to occur during control rod withdrawal and if the collet were to stay unlatched, calculationt indicate that control rod withdrawal speed would be approximately 0.2h ft/sec. Leakage from the open plug hole in the flange would cause reactc. water to flow downward past the velocity limiter pistcn. A small differential pressure across the piston would result in an insignificant driving force of approximately 10 lb, tending to increase withdraw velocity. A pressure differential of 295 psi across the collet piston would hold the collet unlatched as long as the driving signal was maintained. Flow resistance of the exhaust path from the drive would be normal because the ball check valve would be seated at the lower end of its travel by pressure under the drive piston. O 4.6-22 A m, No. 56, (3/81)

l ACNGS-PSAR

;                                                                                                                                                       l

[ h.6.2 3 2.2 7 Ball Check Valve Plug Blows Out i { As a means of access for machining the ball check valve cavity, a 1.25-in.

diameter hole has been drilled in the flange forging. This hole is sealed

] witha. plug of 1 31-in. diameter and 0.38-in. thickrass. A full-penetration j veld, utilizing Type-308 stainless steel filler, holds the plug in place.

j. The failure postulated is a circumferential crack in this veld leading to a

                . blowout of the plug.

f If the plug _ vere to blow out while the drive was latched, there would be no 4 control rod motion. No pressure differential vould exist across the collet j piston to unlatch the collet. As in the previous 'ailure, reactor water j vould flow past the velocity limiter, down the annulus between the drive ! and thermal sleeve, throuh the vessel portp and drilled passage, through l the ball check valve cage and out the open plug hole to the dryvell. The

                . leakage calculations indicate the flow rate vould be 350 gpm. This i               . calculation assumes liquid flow, but flashing of the hot reactor water to

steam would reduce this rate to a lower value. Drive temperature would

rapidly increase and initiate an alarm in the control room.

I In the plug failure were to occur during control rod withdrawal (it would not be possible to unlatch the drive after such a failure), the collet j vould relatch at the first locking groove. If the collet were to stick, j calculations indicate the control rod withdrawal speed would be 11.8 j ft /sec. There would be a large retarding force exerted by the velocity limiter due to a 35 psi pressure differential across the velocity limiter

piston.

i i ( h.6.2 3 2.2.8 Drive / Cooling Water Pressure Cont. 1 Valve Closure (Reactor j Pressure, 0 psig) } The pressure to move a drive is generated by the pressure drop of 3 practically the full system flow through the drive / cooling water pressure j control valve. This valve is either a motor-operated valve or a standby j manual valve; either one is adjusted to a fixed opening. The normal

pressure drop across this valve develops a pressure 260 psi in excess of

] reactor pressure. j If the flow through tl e drive / cooling water pressure cont: 31 valve were to l be stopped (as by a valve closure or flow blockage), the drive pressure 4 would increase to the shutoff pressure of the supply pump. The occurrence l of this condition during withdrawal of a drive at zero vessel pressure vill ! result in a drive pressure increase from 260 psig to no mre than 2000

psig. Calculations indicate that the drive would accelerate from 3 in./sec j to approximately 5 In./sec. A pressure differential of 1970 psi across the l collet piston vould hold the collet unlatched. Flow would be upward, past j the velocity limiter piston, but retarding force would be negligible. Rod l movement would stop as soon as the driving signal was removed.

j 4.6.2 3 2.2 9 Ball Check Valve Fails to Close Passage to Vessel Ports I Should the ball check valve sealing the passage to the vessel ports be

l. dislodged and prevented from reseating following the insert portion of a lg '

drive withdrawal sequence, water below the drive piston would return to the a 4.6-23 Am. No. 56, (3/81) + - . - - . . -- - - - -- . - - _ - . . . _ _ - . - _ - -

ACNGS-PSAR reactor through the vessel ports and the annalus between the drive and the housing rather than through the speed control valve. Because the flow resistance of this return path would be lower than normal, the calculated withdrawal speed would be 2 ft/sec. During withdrawal, differential pressure across the collet piston would be approxinttely h0 pai. Therefore, the collet would tend to latch and would have to stick open before continuous withdrawal at 2 ft/sec could occur. Water would flow upward ptst the velocity limiter piston, generating a small ri tarding force of approximately 120 lb. h.6.2 3.2.2.10 Hycraulic Control Unit Valve Failures Various failures of the valves in the HCU can be postulated, but none could produce differential pressures approaching those described in the preceding paragraphs and none alone could produce a high velocity withdrawal. Leakage through either one or both of the scram valves produces a pressure that tends to insert the control rod rather than to withdraw it. If the pressure in the scram discharge volute should exceed reactor pressure following a scram, a check valve in the line to the scram discharge header prevents this pressure from operating the drive mechaniste. h.6.2 3 2.2.ll Collet Fingers Fail to Latch The failure is presumed to occur when the drive withdraw signal is removed. If the collet fails to latch, the drive tantinues to withdraw at a fraction of the normal speed. This assumption is rude because there is no known means for the collet fingers to become unlocked without some initiating signal. Because the collet fingers will not cam open under a load, accidental application of a down signal does not unlock them. (The drive must be given a short insert signal to unload the fingere and cam them open before the collet can be driven to the unlock position.) If the drive withdrawal valve fails to close following a rod withdrawal, the collet would remain open and the ( ive continue to move at a reduced speed. h.6.2 3 2.2.12 Withdrawal Speed Control Valve Failure Normal withdrawal speed is determined by differential pressures in the drive and is set for a nominal value of 3 in./sec. Withdrawal speed is maintained by the pressure regulating system and is independent of reactor vessel pressure. Tests have shown that accidental opening of the speed control valve to the full-open position produces a velocity of approxientely 5 in./sec. The CRDS prevents unplanned rod withdrawal, and it has been shown above that only multiple failures in a drive unit and in its control unit could cause an unplanned rod withdrawal. h.6.2 3 2 3 Scram Heliability High scram reliability is the result of a number of features of the CRD system. For example: (1) An individual accumulator is provided for each CHD with sufficient stored energy to scram at any reactor pressure. The reactor vessel 4.6-24 Am. No. 56, (3/81)

ACNGS-PSAR itself, at pressures above 600 pai, vill supply the necessary force t.-

 /'~N                            insert a drive if its accumulator is unavailable.
 \      \

(2) Each drive mechanism las i's own scram valves and a dual solenoid scram pilot valve; tht- bre, only one drive can be affected if a scram valve fails to ;en. Both pilot valve solenoids must be de-energized to ini , ate a scram. (3) The RPS and the dCUs are designed so that the scram signal and mode of operation override all others. l (h) The collet assembly and index tube are designed so they will not restrain or prevent control rod insertion during scram. ($) The scram discharge volute is monitored for accumulated water and the reactor vill scram before the volume is reduced to a point that could interfere with the scram. 4.6.2 3 2.h Control Rod Support and Operation 1 As described above, each control rod is independently supported and ' controlled as required by safety design bases. h.6.2 3 3 Control Rod Drive Housing Supports Downward travel of the CED housing and its control rod following the

      -                       postulated housing failure equals the sum of these distances: (1) the

[s x _,/ s

        )                     compression of the disc springs under dynamic loading, and (2) the initial gap between the grid and the bottom contact surface of the CRD flange. If the reactor were cold and pressurized, the downward motion of the control rod would be limited to the spring compression (approximately 2 in. ) plus a gap of approximately 1 inch.                                               If the reactor were hot and pressurized, the gap would be approximately 3/4 in and the spring compression vould be slightly less than in the cold condition. In either cose, the control rod movement following a housing failure is substantially limited below one drice " notch" movement (6 in.). Sudden withdrawal of any control rod through a distance of one drive notch at any position in the core does not produce a transient sufficient to damage any radioactive material barrier.

The CRD housing supports are in place during power operation and when the nuclear system is pressurized. If a control rod is ejected during shutdown, the reactor remains subcritical because it is designed to remain subcritical with any one control rod fully withdrawn at arv time. At plant operating temperture, a gap of approximately 3/4 in. exists between the CRD housing and the supports. At lower temperatures the gap is greater. Because the supports do not contact any of the CRD housing except during the postulated accident condition, vertical contact stresses are prevented. Inspection and testing of CRD housing supports is discussed in 4.6.3 2.1.

   'd 4.6-25                      Am. No. 56, (3/81)

ACNGS-PSAR 4.6 3 Testing and Verification of the CRDs 4.6 3 1 Control Rod Drives 4.6.3 1.1 Testing and Inspection 4.6.3 1.1.1 Development Tests

 'Ihe initial developrent drive (prototype of the standard locking piston design) testing included more than 5 Os scrams and approximately 100,000 latching cycles. One prototype was exposed to simulated operating conditions for 5000 hours. These tests demonstrated the following:

(1) The drive easily withstands the forces, pressures and temperatures imposed. (2) Wear, abrasion and corrosion of the nitrided stainless parts are negligible. Mechanical performance of the nitrided surface is superior to that of materials used in earlier operating reactors. (3) The basic scram speed of the drive has a satisfactory margin above minimum plant requirements at any reactor vessel pressure. (h) Usable seal lifetimes in excess of 1000 scram cycles can be expected. 4.6 3 1.1.2 Factory Quality Control Tests Quality control of welding, heat treatment, dimensional tolerances, material verification and similar factors is raintained throughout the manufacturing pr3 cess to assure reliable performance of the mechanical reactivity control components. Some of the quality control tests performed on the control rods, CRD mechanisms, and HCUs are listed below: (1) CRD mechanism tests:

a. Pressure velds on the drives are hydrostatically tested in accordance with ASME codes.

b. Electrical components are checked for electrical continuity and resistance to ground.

c. Drive parts that cannot be visually inspected for dirt are flushed with filtered water at high velocity. No significant foreign material is permitted in effluent water.

d. Seals are tested fer leakage te demonstrate correct seal operation.

l e. Each drive is tested for shim motion, latching, and control rod position indication.

f. Each drive is subjected to cold scram tests at various reactor pressures to verify correct scram performance.

4.6-26 Am. No. 56, (3/81)

ACNGS-PSAR (2) HCU tests:

    )      a. Hydraulic syatems are hydrostatically tested in accordance with the applicable code.

b. Electrical components and rystems are tested for electrical continuity and resistance to ground.

c. Correct operation of the accumulator pressure and level switches is verified.

i

d. The unit's ability to perform its part of a scram is demonstrated.

e. Correct operation and adjustment of the insert and withdrawal valves is demonstrated.

h.6.3 1.1 3 operational Tests After installation, all rods and drive mechanisms can be tested through their full stroke for operability. During normal operation, each time a control red is withdrawn a notch, the operator can observe the in-core monitor indir.ations to verify that the control rod is following the drive mechanism. All control rods that are partially withdrawn from the core can be tested for rod-following by inserting or withdrawing the rod one notch and returning it tv its original position, while the operator observes the in-core monitor indications. kJ ) To make a positive test of control rod to CRD coupling integrity, the operator can withdraw a control rod to the end of its travel and then attempt to withdraw the drive to the overtravel position. Failure of the drive to overtravel demonstrates rod-to-drive coupling intyrity. Hydraulic supply subsystem pressures can be observed from instrumentation 4 in the control room. Scram accumulator pressures can be observed on the nitrogen pressure gages. t 4.6.3 1.1.h Acceptance Tests Criteria for acceptance of the individual CRD mechanisms and the associatt i  ; control and protection systeem will be incorporated in specifications and test procedures covering three distinct phases: (1) pre-installation; (2) after installation prior to startup; and (3) during startup testing. The pre-installation specification vill define criteria and acceptable ranges of such characteristics as seal leakage, friction and scram  ; performance under fixed test conditions which must be met before the component can be shipped. l The after-installation, prestartup tests (Chapter lb) include normal and scram motion and are primarily intended to verify that piping, valves, o electrical components and instrumentation are properly installed. The test a fv) specifications will include criteria and acceptable ranges for drive speed, 4.6-27 Am. No. 56, (3/81) ;

ACNGS-PSAR timer settings , scram valve response times, and control pressures. These are tests intended more to docutent system condition rather than tests of performance. As fuel is placed in the reactor, the startup test procedure (Chapter 14 ) vill be followed. The tests in this procedure are intended to demonstrate that the initial operational characteristics meet the limits of the specifications over the range of primary coolant temperatures and pressuren from ambient to operating. The detailed specifications and procedures have not as yet been prepared but vill follow the general pattern established for such specifications and procedures in EWRs presently under construction and in operation. h.6.3.'.1 5 Surveillance Tests The surveillance requirements (SR) for the CRD system are described below. (1) Sufficient control rods shall be withdrawn, following a refueling outage when core alterations are performed, to demonstrate with a margin of 0.25% k that the core can be made suberitical at any time in the subsequent fuel cycle with the strongest opecable control rod fully withdrawn and all other operable rods fully inserted. (2) Each partially or fully withdrawn control rod shall be exercised one notch at least once each week. In the event that operation is continuing with three or more rods valved out cf service, this test shall be performed at least once each day. The weekly control rod exercise test serves as a periodic check against deterioration of the control rod system and also verifies the ability .f the control rod drive to scram. If a rod can be moved with drive pressure, it may be expected to scram since higher pressure is applied during scram. The frequency of exercising the control rods under the conditions of three or more control rods valved out of service provides even further assurance of the reliability of the remaining contol rods. (3) The coupling integrity shall be verified for each withdrawn control rod as follows:

a. When the rod is first withdrawn, observe discernible response of the nuclear instrumentation; and

b. When the rod is fully withdrawn the first time, observe that the drive vill not go to the overtravel position.

Cbservation of a response from the nuclear instrucentation during an attempt to withdraw a control rod indicates indirectly that the- rod and drive are coupled. The overtravel asition feature provides a positive check on the coupling integrity, for only an uncoupled drive can reach the overtravel position. O 4.6-28 Am. No. 56, (3/81) l

ACNGS-PSAR 1 (4) During operation, accumulator pressure and level at the normal  !

    /]           operating value shall be verified.

Experience with CRD systems of the same type indicates that weekly verification of accumulator pressure and level in sufficient to assure operability of the accumulator portion of the CRD system. (5) At the time of each major refueling outage, each operable control rod shall be subjected to scram time tests frot the fully withdrawn position. Experience indicates that the scram times of the control rods do not significantly change over the time interval between refueling outages. A test of the scram times at each refueling outage is sufficient to identify any significant lengthening of the scram times. h.6.3 1.1.6 Functional Tests The functional testing program of the CRDs consists of the five-year maintenance life test, 40 year design life test, 60 year extended design , life test, and improved single failure testing as described below: (1) Five-Year Maintenance Life Tests - This test is designed to demonstrate that the FSCRD could operate catisfactorily for a duty cycle equivalent to five reactor years of operation, with margin. The imposed duty cycles are given in Table 4.6-2. O

    .     (2) 40 Year Design Life Tests - The objective of thk tot is to verify

( that the FSCRD are capable of satisfying the design life and duty cycle requirements (Table 4.6-3) with all performance characteristics remaining within normal limits. (3) 60 Year Extended Design Life Tests - The objective for this test is to demonstrate margin with respect to life and duty cycle requirements by subjecting the test drive to an additional 20 'rs of simulated reactor life. In addition, the cyclic equiva; c of a five-year maintenance life is run at CRD operating tempe. ture greater than 350*F as part of the 20-year extended life test to demonstrate the FSCRD's ability to satisfy performance requirements during extended periods of high temperature operation. ( '4 ) The following tests with imposed single failures have been performed to evaluate the performance of the CRDs under these conditions: (1) Simulated Ruptured Scram Line Test (2) Stuck Ball Check Valve in URD Flange (3) HCU Drive Down Inlet Flow Control Valve (V122) Failure (4) HCU Drive Down Outlet Flow Control Valve (V120) Failure m (5) CRD Scram Jerformance with Valve (V120) Malfunction m e 4.6-29 Am No. 56, (3/81)

ACNGS-PSAR (6) IICU Drive Up outlet control Valve (V121) Failure (7) IICU Drive Up Inlet Control Valve (V123) Failure (8) Cooling Water Check Valve (V138) Leakage (9) CRD Flange Check Valve Leakage (10) CRD Stabilization Circuit Failure (11) IICU Filter Restriction (l?) Air Trapped in CRD Hydraulic System (13) ChD Collet Drop Test l i (1h) CR Qualification Velocity Limiter Drop Test I l Additional postulated CRD failures are discussed in Subsections l h.6.2 3 2.2.1 through 4.6.2 3 2.2.11.  ;

                          '4.6.3 2                                       control Rod Drive liousing Supports 4.6.3 2.1                                                               Testing and Inspection CRD housing supports are removed for inspection and maintenance of the control rod drives. The supports for one control rod can be removed during reactor shutdown, even when the reactor is pressurized, because all                                                                                                                                     j control rods are then inserted. When the support structure is reinstalled,                                                                                                                              !

it is inspected for correct assembly with particular attention to  ! l l rnintaining the correct gap between the CRD flange lower contact surface i and the grid. l h.6.4 Information for Combined Performance of Reactivity Control Systems This information will be supplied in the Allens Creak FSAR. h i 4.6 5 Evaluation of Combined Performance As indicated in Subsection h.6.4.2, credit is not taken for multiple  ! reactivity control systems for any postulated accidents in Chapter 15 i

                                                                                                                                                                                                                                  ?

4.6.6 Standby Liquid Control System I l 4.6.6.1 Design Bases  ! h.6.6.1.1 General Design Bases [ t b.6.6.1.1.1 Safety Design Bases I l l The Standby Liquid Control System shall meet the following safety design I l bases: O:, l 4.6-30 Am No. 56, (3/81) f B

                      ,      , _ , . _ - _ _ _ , _ _ - - _ - _ - - _ - , . . . , _ _ _ - - _                             _-_.-.___.m.  -

ACNGS-PSAR I a) Backup capability for reactivity control shall be provided, [ independent of norml reactivity control provisions in the nuclear reactor, to be able to shutdown the reactor if the normal control ever j becomee inoperative. b) The backup system shall have the capacity for controlling the reactivity difference between the steady-state rated operating l condition of the reactor with voids and the cold shutdown condition, j including shutdown cargin, to assure complete shutdown from the most  ; t tvactive condition at any time in core life. c) The time required for actuation and effectiveness of the backup  ! control shall be consistent with the nuclear reactivity rate of change f predicted between rated operating and cold shutdown conditions. A l fast scram of the reactor or operational control of fast reactivity [ transients is not specified to be accomplished by the system. l l d) Means shall be provided by which the functional performance capability j of the backup control system components can be verified periodically , under conditions approaching actual use requirements. A substitute solution, rather than the actual neutron absorber solution, can be injected into the reactor to test the operating of all components of the redundant control system. e) The neutron absorber shall be dispersed within the reactor core in sufficient quantity to provide a reasonable margin for leakage or imperfect mixing. , f) The system shall be reliable to a degree consistent with its ule as a special safety system; the possibility of unintentional or accidental shutdown of the reactor by this system shall be minimized. 1.6.6.2 4 Description The Standby Liquid Control System (see Figure 1.6-9) is mnually initiated 4 from the main Control Room to pump a boron, .eutron absorber, solution into the reactor if the operator believes the reactor cannot be shutdown or kept shut down with the control rods. Once the operator decision for initiation of the SLC system is mde, the design intent is to simplify the mnual process by providing two " key locked" switches. This prevents inadvertant injection of neutron absorber by the SLC system. However, insertion of the control rods is expected to assure prompt shutdown of the reactor should it l be required.

• I The " key locked" Control Room switches are provided to assure positive action from the min Control Room should the need arise. Standard power plant procedural controls are applied to the operation of the " key locked" control room switch.

The SLC system is required only to shut down the reactor and keep the reactor from going critical again as it cools. n\ (G 4.6-31 Am. h. 56, (3/81)

nUma-PSAR The GLC uystem is needed only in the inprobable event that not enough control rods can be inserted in the reactor core to accomplish shutdown and couldown in the normal manner, The boron solution tank, the test water tank, the two positive-displacement pumps, the two explosive valves, and associated local valves and controls are mounted in the Reactor Auxiliary 13uilding. The liquid is piped into the reactor vessel and discharged near the bottom of the core shroud from a mixing stand pipe so it r.1xes with the cooling water rising through the core (see Section 5.k, "heactor Vessel and Appurtenances," and Section 3 9.4, "Heactor Core Support Structures and Internals Mechanical Design"). The boron absorbs thermal neutrons and thereby terminates the nuclear fision chain reaction in the uranium fuel. The specified neutron absorber solution is codium pentaberate (H"213 100 16 10l!2 0). It is prepared by dissolving stoichiometric quantities of borax and boric acid in demineralized water. A sparger is provided in the tank for mixing, using air. To prevent system plugging, the tank outlet is raised above the bottom of the tank. At all times when it is possible to make the reactor critical, the GLC System chall be able to deliver enough codium pentaborate solution into the i reactor (see Figure 4.6-10) to ensure reactor shutdown. This is accouplished by placing codium pentaborate in the SLC tank and filling with l demineralized water to at least the low level alarm point. The colution can be diluted with water up to overflow level volume to allow for  ! evaporation losses or to lower the saturation temperature. l The saturation temperature of the recommended colution is 49 F at the low j i level alarm volume and approximately 59 F at the tank overflow volume (see , Figure 4.6-11). The equipment containing the solution is installed in a l room in which the air temperature is to be maintained within the range of TO F to '.10 F. In addition, a heater cystem maintains the solution i temperature at 75 P to ^5 F to prevent precipitation of the sodium pentaborate from the solution during storage. High or low temperature, or high or low liquid level, causes an alarm in the Control Root Each positive displacement pump is sized to inject the solution into the reactor in 50 to 125 minutes, independent of the anount of solution in the tank. The pump and system design pressure between the explosive valves and , the pump discharge is 1400 psig. The two relief valves are set slightly I under lh00 psig. To prevent bypass flow from the one pump in case of j relief valve failure in the line from the other pump, a check valve is installed downstream of each relier valve line in the pump discharge pipe. The two explosive-actuated injection valves provide assurance of opening , when needed and enuure that boron will not leak into the reactor even when the pumps are being tested. Each explosive valve is closed by a plug in the inlet chamber. The plug is circuuscribed with a deep groove so the end will readily shear off when pushed with the valve plunger. This opens the inlet hole through the plug. O, f f 4.6-32 Am. No. 56, (3/81)  ; l

                                                                                                                                                                                                           ;

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l l ACNGS-PSAR g The sheared end is pushed out of the way in the chamber; it is chaped so it I i vill not block the ports after release. The continuity of the explosive valve primer ignition circuit is monitored by measuring a trickle current through the primers. If either primer or primer ignition circuit becomes open-circuited, the continuity meter reads downscale. The SLC System is actuated by turning either of two keylocked switches on the control room conuole to the "RUN" position. The keys are removable only in the "0FF" position. Activaticn of divisional switch 1 or divisional switch 2 will open the asscciated divisional storage tank discharge valve, start the associated divisional pump, actuate the associated divisional explosive valve, and close one of two reactor water cleanup icolation valves. If division 1 is actuated, the inboard RWCU isolation valve vill close. If division 2 is actuated the outboard RWCU isolation valve vill close. A green light in the control room indicates that power is available to the pump motor contactor and that the contactor is open (pump not running). A red light indicates that the contactor is closed (pump running). If the pump lights, or explosive valve light indicates that the liquid nay not be flowind, the operator can immediately operate the other division switch which actuates the alternate divisional system. Cross piping and check valves assure a flow path througn either pump and either explosive

    '   valve. The local switch will not have a "stop" position. This prevents the separation of the pump from the control room. Pump discharge pressure is also indicated in the control room.

The storage and test tank overflow lines, both pump drip plate drains and pipe drain lines are to drain into equipment hubs located in the SLCS area floor. These hubs and a floor drain are connected to a connon header which drains into a 55 gal. drum located near the equipment access area. A locked - closed Glove Valve is to be located a minimal distance above the

storage container. The valve vill be nanually opened only during drainage. When the SS gal. drum is full, the valve vill be closed and drum will be replaced with an empty one. I Should either the test or storage tank over-flow, the riser system vill fill with liquid until it begins to flood the SLCS area floor and will be contained by concrete curb which surrounds the SLCS area. This area can then be drained through the floor drain by opening the aforementioned valve which will permit flow to the storage container. This would adequately minimize the possibility of sodium pentaborate contamination spread. Instrumentation consisting of solution temperature indication and control, solution level, and heater system status is provided locally at the storage tank. x 4.6-33 Am. No. 56, (3/81)

ACNGS-PCAR Table h.6-1 contains the procens data for the various codes of operation of the CLC. h.6.6.3 Cafety Evaluation The standby liquid control system is a special safety system and in maintained in a standby operational status whenever the reactor is critical. The system is expected never to be needed for safety reasons because of the large number of independent control rods available to shutdown the reactor. However, to assure the availability of the SLC System, two sets of the components required to actuate the system - pumps, motor-operated valves, and explosive valves - are provided in parallel redundancy. The system is desipled to bring the reactor from rated power to a cold shutdown at any time in core life. The reactivity compensation provided will reduce reactor power from rated to zero level and allow cooling the nuclear system to room temperature, with the control rods remaining withdrawn in the rated power pattern. It includes the reactivity gaina that result from complete decay of the rated power xenon inventory. It also includes the positive reactivity effects from eliminating steam voids, changing water density from hot to cold, reduced Doppler effect in uranium, reducing neutron leakage from bolling to cold, and decreasing control rod worth as the moderator cools. The specified minimum final concentration of boron in the reactor core provides a margin of -0.05 k for calculational uncertainties and assures a substantial shutdown margin. The specified minimum average concentration of natural boron in the reactor to provide the specified shutdown margin, after operation of the SLC System, is 660 ppm (parts per million). (See Figure h.6-12). Calcu lation of the minimum quantity of sodium pentaborate to be injected into the reactor is based on the required 660 ppm average concentration in the reactor coolant including recirculation loops, at 70 F and reactor normal water level. The result is increased by 25 percent to (825 PPM), to alllow for imperfect mixing and leakage. An additional 250 ppm is provided to accommodate dilution by the HHH System in the shutdown cooling mode. This concentration will be achieved if the solution is prepared as defined in Section h.6.6.h and maintained above saturation temperture. Cooldown of the nuclear system will require a minimum of several hours to renove the thermal energy stored in the reactor, cooling water, and acucciated equiprent. The controlled limit for the reactor vessel cooldown l in 100 F per hour, and normal operating temperature is approxirntely 550 F. l Une of the main condenser and various shutdown cooling systems requires 10 to 24 hours to lower the reactor vessel to room temperature (70 F); this is the condition of nnximum reactivity and, therefore, the condition that i ! requires the maximum concentration of boron. The specified boron injection rate is limited to the range of 6 to 25 ppm per minute. The lower rate assures that the boron is injected into the reactor in approximately two hours. This resulting reactivity incertion is conaiderably quicker than that covered by the cooldown. The upper limit injection rate assures that there is sufficient mixing so the boron does Am. No. 56, (3/81) 4.6-34

ACNGS-PSAR not recirculate through the core in uneven concentrations that could [} v possibly cause reactor power to rise and fall cyclically. The GLC System equipment essential for injection of neutron absorber solution into the reactor is designed as Category I (seiumic) for withstanding the specified earthquake loadings (see Chapter 3). The system piping and equiptent are designed, installed, and tested in accordance with requirements stated in Chapter 3 1 The SLC pystem is required to be operable in the event of a station power failure; therefore, the pumps, heaters, valves, and controls are povered from the standby ac power supply or de power in the absence of normal power. The pumps and valves are povered and controlled from separate buses and circuits so that a single failure vill not prevent system operation. The SLC system and pumps have sufficient pressure margin, up to the system relier valve setting of approximately 1400 psig, to assure solution injection into the reactor above the normal pressure in the bottom of the reactor. The nuclear system relief and safety valves begin to relieve pressure above approximately 1100 psig. Therefore, the SLC system positive displacement pumps cannot overpressurize the nuclear system. Only one of the two standby liquid control pumps is needed for system operation. If one pump in found to be inoperable, there is no immediate threat to shutdown capability, and reactor operation can continue during repairs. The time during which one redundant component upstream of the

     . explosive valves may be out of operation should be consistent with the following:      the probability of failure of both the control rod shutdown capability and the alternate component in the SLC system; and the fact that nuclear system cooldown takes several hours while liquid control solution injection takes approximately two hours. Since this probability is small, considerable time is available for repairing and restoring the SLC system to an operable condition while reactor operation continues. Assurance that the system vill still fulfill its function during repairs is obtained by demonstrating operation of the alternate divisional pump.

h.6.6.4 Inspection and Testing Operational testing of the SLC System is performed in at least two parts to avoid inadvertently injecting boron into the reactor. With both storage tank discharge valves closed and with the two valves in the test tank return line and the test tank discharge valve open, demineralized water in the test tank can be recirculated by locally starting e!.ther pump. The system can be functionally tested in the injection mode by opening the test tank discharge valve and closing the two valves in the test tank ret urn line. Both storage tank discharge valves are kept closed. The celected divisional system is actuated in the Main Control Room. The  ; selected divisional injection valve is fired and demineralized water is l pumped f rom the test tank to the RPV. The test tank contains enough , p demineralized. vater for 3 minutes of pumping. The test is performed during (s reactor shutdevn with the reactor vessel pressure at 0 psig and 125'F. l i 4.6-35 Am. No. 56, (3/81)

                 --.             -      .=.                    . .      =-

ACNGS-PSAR After functional tests, the injection valve shear plugs and explosive charges must be replaced and all the valves returned to their normal positions as indicated. After closing a local locked-open valve to the reactor, leakage through the injection valves can be detected by opening valves at a test connection in the line between the containment isolation check valves. Position indicator lights in the control room indicate that the local valve is closed for tents or open and ready for operation. Leakage from the reactor through the first check valve can be detected by opening the same test connection when the reactor is pressurized. The test tank contains demineralized water for approximately 3 minutes of pump operation. Demineralized water from the demineralized water system is available for refilling or flushing the system. Should the boron solution ever be injected into the reactor, either intentionally or inadvertently, then after making certain that the normal reactivity controls will keep the reactor suberitical, the boron is removed from the reactor coolant system by flushing for gross dilution followed by operating the reactor water

 , cleanup system. There is practically no effect on reactor operations when the boron concentration has been reduced below approximately 50 ppm.

The concentration of the sodium pentaborate in the solution tank is determined periodically by chemical analysis. h.6.6 5 Inst ruerntat ion The instruw2ntation and control system for the SLC system is designed to allow the injection of liquid poison into the reactor and the maintenance of the liquid poison solution well above the actuation temperature. A further discussion of the SLC instrumentation may be found in Chapter 7 h.6.7 References

1. J. E. Benecki, " Impact Testing on Collet Assembly for Control Rod Drive Mechanism 7BD 81hhA", General Electric Company, Atomic Power Equipment Department , November 1967 (APED-SSS5).

O 4.6-36 Am. No. 56, (3/81) l

l O O O TABLE h.6-1 STidirEY LIQUID CCNTROL SYSTD! OPEPATING PRESSURE /TD!PERTURE CCNDITIONS TEST MODES (a) Standby Circulation Injection Operating Mode (a) Test Test (b) Mode (a) PIPING Press. Temp. Press. Temp. Press. Temp. Press. Terp. psig (c) *F psig (c) *F psig (c) *F psi; (c) *F Makeup 70/1C20 Test Tank 70/100 Test Tank 70/100 StoraEe 70/110 , Pump Water (d) Static (d) Static (d) Tank (d) ! Suction Pressure Head (e) Head (e) Static l Head I $ j Pu=p Dis- Makeup 70/100 0/1220 70/100 70 Plus 70/100 (70 Plus 70/110 $ ! charge To Water Reactor Reactor y ! Explosive Pressure Static Static @

• Head

! , Valve Inlet Head) to @ l  ? 1220 - w N Explosive Reactor 70/100 Reactor 70/100 TO Plus 70/100 (70 Plus 70/110 Valve Outlet Stetic Static Reactor Reactor . To But Not Head to Head to Static Static Including 1150 (f) 1150 (f) Head Head) to First Isola- 1220 , tion Check , Valve I i ! First Isola- Reactor 70/560 Reactor 70/560 Reactor 125 (b) Reactor 70/560 ' ! tion Check Static Head (g) Static (g) Static Static (g) l Valve to the to 1150 (f) Head Head (b) Head to l Reactor 1150 (f)

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ACNGS-PSAR w TABLE h.6-1 (Continued)

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v (a) The pump flow rate vill be zero (pump not operating) during the standby mode and at rated during the test and operating modes. (b) Feactor to be a O psig and 125 F before changing from the standby mode to the injection test mode. (c) Pressures tabulated represent pressure at the points identified below. To obtain pressure at intermediate points in the system, the pressures

          'ab ' lated must be adjusted for elevation difference and pressure drop t'tseen such intermediate points and the pressure points identified below:

Piping Pressure Point Pump Suction: Pump Suction Flange Inlet Pump Discharge To Explosive Valve Inlet: Pump Discharge Flange Outlet Explosive Valve Outlet To Dit Not Including First Isolation Check Valve: Explosive Valve Outlet [~Nl First Isolation Check Valve To The Reactor: (d Reactor Sparger Outlet (d) During chemical mixing, the liquid in the storage tank will be at a temperature of 150 F maximum. (e) Pump suction piping vill be subject to demineralized water supply pressure during flushing and filling of the piping and during any testing where suction is taken directly from the demineralized water supply line rather than the test tank. (f) Maximum reactor operating pressure is 1150 psig at reactor standby liquid control sparger outlet. (g) 560 F represents nnximum sustained operating temperature. 4.6-38 Am. No. 56, (3/81)

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i ACNGS-PSAR I Table h.6-2 1 DESIGN DUTY CYCLE i (FIVE-YEAR MAINTENANCE LIFE TEST) l l s ' l Activity Cycles l Scram tests 27 l l Startup scrama 30 i l Operational scrams 48  ; Jeg cycles 4000 l l 1 k i Shim / drive cycles 175 I I l \  ;

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ACNGS-PSAR Table h.6-3 h0-YEAR DESIGN LIFE TEST O Activity Cycles Scram tests lho Startup scrams 160 i Operational scrams 300 l I 30,000  ! j Jog cycles l Shim / drive cycles 2,000 , ]  !

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