ML20085C590

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Rev 1 to Safeguards Rept for Saxton Core III
ML20085C590
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Site: Saxton File:GPU Nuclear icon.png
Issue date: 10/04/1991
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SAXTON NUCLEAR EXPERIMENTAL CORP.
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FOIA-91-17 NUDOCS 9110040026
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{{#Wiki_filter:. OhY [ J k SAFEGUARDS REPORT FOR SAXTON CORE 111 REVISION 1 7 \\. 4 4 9 ( team-rz r.nn

l TABLE OF CONTENTS Section Title Page 1.0 Introduction and Summary 1-1

1.1 Background

1-2 1.2 Progran Objectives 1-3 1.3 Linear Power objectives 1-3 1.4 Operating Conditions 1-4 1.5 General operating Procedures 1-5 2.0 Core Design 2-1 2.1 Mechanical Design 2-1 2.1.1 Core Loading 2-1 2.142 Fuel Assembly Design 2-2 2.1.3 Subassembly Design 2-8 2.2 Nuclear Design 2-10 2.2.1 Analytic Configuration 2-10

2. 7. 2 Power Characteristics 2-10

~ 2.2.3 Lif etime Characteristics 2-12 2.2.4 s eactivit/ Characteristics 2-12 2.3 L+rmal-Hydraulic Design 2-15 2.3.1 Gent ~al 2-15 2.3.4 Thermal cid Hydrauj'c Design Criteria 1-15 2.3.3 1hermal and Hydraulic Analysis 1 16 2.3.4 Thermal Analysis of Fuel Rod 2-20 2.4 Fuel Perfo mace.e Evaluation 2-22 g 2.4.1

Background

2-22 2.4.2 ~xpected Fuel Performance of Saxton Core Ill Rods 2-23 3.0 Safety Analysis _ 3-1 3.1 General 3-1 3.2 Loss of Coolant FJow 3-5

3. 3 Loss of Coolant 3-7 Appendix A - Saxton Failed Fuel Monitor System Appendix B - Summary Report on Buckling of Saxton, Core 11 Fuel Assemblies and Prevention of Buckling in Core 111 Appendix C - Change Reports 17 to 23 Appendix D - Parts of WCAP-3850-3 Which Describe Experimentally Observed Strain in CVTR Puel Rods Appendix E - Rod Withdrawal and Steam Line Break Accident Analysis for Saxton Core 1135 MWt Operation 4-11 I

LIST OF ILLUSTRA'? IONS rigure Title 2.1-1 Saxton Core III Assembly Configuration 2.1-2 Alpha Numeric Location Identification Within Assembly 2.1-3 Saxton Reactor 9 x 9 Loose Lattice + Load Follow Fuci \\ssembiv 2.1-4 Saxtsa Reactor Plant 9 x v '.cose Lattice Tie Rod Assembly 2.1-5 saxton Reactor Plart Loose La?tice Plutonium Fuel Rod Assembly 2.1-6 Saxton Reactor Plant Water Fil.ed Tube Assembly 2.1-7 Saxton Reactor Plant 9 x 9 Load Follow ila Rod Assembly 2.1-8 Subassembly 503-4-31 2.1-9 S ubass embly 503-4-33 2.1 .0 Subassembly 503-4 25 2.1-11 Sub assembly 503-4-32 2.1-12 Subassembly 503-4-34 2.2-1 Burnups and Enrichments in Load-Follow and Loose-Lattice Assemblies Used in Core III Nuclear Design 2.2. Relative Peak Rod Linear Power as a Function of Saxton Core III Operation at 23.5 MWt. 2.2-3 e 'icipated Power Operation of Saxton Core III 2.2-4 L.

cal Boron Concentration vs. Saxton Core III Operation 2.2-5 Saxton Core III Temperature Coefficient at Full Power as a Function of Boron Concentration 2.2-6 Saxton Core III Zero Power Temperature Coefficient (Beginning i

of Life) t 2.2-7 Rod Power Characteristics of Loose Lattice Assemblies 2.2-8 Rod Power Characteristics of Load Follow Asscublies 2.2-9 Inverse Boron Worth as a Function of Saxton Core III Boron Concentration 2.3-1 Core Radial Power Distribution 2.3-2 Design Nuclear Radial Power Factors in Thermally Controlling Core III Assemblies 2.4-1 Variation of Loose Lattice DNBR With Single Rod Clajding Strain ) 3.1-1 Saxton Loose Lattice Fuel Melt Limit 3.1-2 Flux Depression and Melting Temperature Variation with Pellet Burnup 3.2-1 Saxton 28 MWt Operation Loss of Flow Accident (Flow Coastdown Curves) 3.2-2 Saxton 28 MWt Operation Loss of Flow Accident (MG Set Inertia) 3.2-3 Saxton 28 MWt Operation Loss of. Flow Accident (No MG Set Inertia) (Fraction of initial Value) 3.2-4 Saxton 28 MWt operation Loss of Flow Accident (No MG Set Inertia) 3.2-5 Saxton 28 MWt operation Loss of Flow Accident (No MG Set Inertia) (W3-DNB-R) 111 j

- --.... - =_ _... - [ LIST OF ILLUSTRATIONS (Cont'd) Figure Title l 3.3-1 Comparison of Calenlated and Measured Heat Transfer Coefficients 3.3-2 Saxton - Double Ended Cold Leg Break - Loose Lattice Rods S axt on - 0.173 2 Break Loose Lattice Rods 3.3-3 Saxton.0375 Ft{t Break loose Lattice Rods 3.3-4 3.3-5 Saxton - Double Ended Cold Leg Break - Load Follw Rods 3.3-6 Saxton - 0.173 Ft2 Break load Follw Rods - Zirc Clad 3.3-7 Saxton .0375 Ft2 Break Load Follw Rods - Zirc Clad 3.3-8 Saxton Double Ended Cold Leg Bra.ak - Load Follow Rods 3.3-9 Sr,xton - 0.173 Ft2 Break - Load Follw Rods - S.S. Clad 3.3-10 Saxton .0375 ft2 Break Load Follw Rods - S.S. Clad I I iv .--,v.= r. ,.y..t u e--r-- ce - ,v wr==*

L7ST OF TABLES Table Title 2.1-1 Core Location and Composition of Loose-Lattice Assemblies 2.1-2 Core Location and Composition of Peripheral Assemblies (Core 1) 2.1-3 Initial Core Location and Composition of Saxton Core 111 Subassemblies 2.2-1 Summary of Saxton Core III Transient Analysis Physics Parameters 2.2-2 Installed Reactivity and Minimum Boron Requirements for Core 111 2.3-1 Engineering Hot Channel Factors 2.3-2 Thermal and liydraulic Design Parameters 2.3-3 Thermal and liydraulic Design Parameters Worst Design Conditions 2.4-1 Summary of Power RatinFs and Burnup Expected for Core 111 A b W 4 't V / R

l SAXTON CORE Ill

1.0 INTRODUCTION

AND StHMARY The Saxton Core III configuration, similar to previous cores, contains 21 fuel assemblies. The central nine consist of seven plutonium loose-lattice assemblies which were reconstituted f rom the Core 11 assemblies and two new enriched UO load-follow assemblies. The remaining twelve 2 outer assemblies are eleven partially depleted U0 assemblics from Saxton 2 Core I and one unitradiated UO assembly. The increased lattice pitch 2 of the loose-lattice assemblies extends the burnup capability of the Zircaloy clad plutonium fuel by taking advantage of the reactivity increase associated with the increased water to fuel ratio.

1.1 BACKGROUND

The Saxton Reactor was designed to investigate areas of interest in the development of pressurized water reactors. It has an extensive I in-core instrumentation system for measuring outlet temperature, flow rate, and pressure drop. In addition, fission detectors are used in conjunction with analysis to detemine nuclear power distributions throughout the core. An extensive experimental follow program was carried out during the operation of two previous cores in this reactor. The detailed information obtained during the design and operation of these cores provides the basis for the nuclear design of Saxton Core III. Saxton Core I Saxton Core I was composed of 21 UO fuel assesnblies with an enrichment 2 of 5.7 w/o U-235. The core operated for a total of 8630 megawatt days and accu:nulated a core average burnup of 9700 MWD /MThi. i 1-1 h

1 Saxton Core 11 Saxton Core 11 contained a partial core loading of plutonium fuel. Nine plutonium fuel assemblies were installed in the center of the core with 12 uranium fuel assem111es installed on the periphery. A series I of critical experiments van carried out using both fuel types before the core was installed in the Saxton reactor. Included in these measurements I were critical configuratioas at a lattice pitch equivalent to that of the loose lattice region of Saxton Core III. Extensive hot-tero and cold-tero power measurements were made at the beginning of life and periodically during Saxton Core II operation. At power measurements of various core parameters have also been made i periodically and the critical boron concentration as a function of depletion has been monitored throughout operation. Com;.arisons of analysia with experiment show good agreement. 1 ( During the course of the operations follow work a number of basic improvements were also made in the methods of analysis of plutonium fueled system. l In particular, a detailed PDQ-7 depletion calculation was carried out l to analytically simulate the actual operation of Saxton Core II. These l calculations are experimentally verified by the good agreement between the calculated and measured critical boron concentrations as a function of core life. The analytic results-from this calculation include a detailed burnup and power distribution for every plutonium fuel rod. This detailed distribution provided the basis for the selection of fuel-rods for further irradiation in Saxton Core III. f 1-2 -ye-.p.... e.g-.y---,gy-..-yyyy-.w.ysyms,. -%,,p.q. - ymyy q-y9 ,y.,,pg, p 4.g.- m. 9--pq< -?w g, g w- ,,,3 g,.g. y#%,y g.,.ymi,g.9q7p eqg,+%g, eg ..mg- ,,y g e++qp 4.-

1.. PROCW1 OBJECTIVE The objectives of the Saxton Core 111 irradiation program are to: 1. validate f uel element design code predictions, including determination of pove'./burnup f ailure limits; 2. demonstrate performance capability of Zircaloy clad oxide fuel elements over a broad spectrum of burnups and power levels; and 3. obtain depiction cha racteristics and transuranic isotope generation data for high burnup, mixed exida fue). Because the Saxton Core 11 mixed oxide fuel roda were designed for relatively low peak burnups and operation at power densities < 16 kw/ft, there is a significant risk of failuri of certain of these rods in Core III. By careful selection and placement of these rods in the ~~ loose-lattice assemblics, it is possible to control their burnup and 4 operating power levels and thus permit power /burnup limits to be established while operating safely and in full compliance with the reactor license Technical Specifications. 1.3 LINEAR POWER OBJECTIVES The linear power objectives are an expected peak kv/f t of 21.2 in the loose-lattice assemblies and 17.6 in the load follow assemblies. The corresponding denign linear poser including a conservative combination of the design uncertainties are 24.0 kw/f t and 19.9 kw/f t in the two type assemblics, respectively. These design linear powers are the basis for the analysis for Core 111 and will be achieved at a design core power less than 28 MWt. [ 1-3

As in Core II 35 MWt operation, the nominal inlet temperature during full power operation is 480'I'. The 19.9 kw/f t design linear power of the two load-follow assemblie s is 4% higher than the design value for Core II 35 MWt operation. The increased linear power of the rods in the loose-lattice assemblies resu'ts from the use of water-filled tubes in alternate rod positionsj whit.h by reducing the heat addition to a coolant. channel Jed increases the margin to DNB. 1.4 OPERATING CONDITIONS Core III operating conditions are select ed to maintain fuel temperature below center malt and the minimum DNB raitio greater than 1.3 at the control and protection system reactor trip setpoint conditions thus protecting the fuel for anticipated transient conditions. The fuel and moderator temperature coefficients and kinetic parameters for Core III are intermediate to those of previous cores. There f ore, Core III represents no extrapolation fro:n previous operation except for the increased peak linear power ratings. Operation is restricted to prevent center melt and to maintain margin to DNB. As in Core 11 35 MWt Operation, the reactor coolant pump is operated from the motor-generator set to benefit from the increased pump inertia. The operating pressure for Core III is increased from 2200 psia to 2250 psia to more closely simulate conditions in current PWR's. Approximately 50 fuel rods operate within 10% of the peak linear power in the seven loose-lattice assemblies (design peck linear power 24 kv/f t) and approximately 80 fuel rods operate within 10% of the peak linear power in the load-follov assemblies (design peak linear power 19.9 kw/f t). The Core II mixed oxide fuel rods used in the loose-lattice assemblics i were designed for lower peak burnups and lower linear power than they will experience in Core III. As a result certain of these rods may A fail due to excessive clad strain and internal gas pressure at the high linear power and burnups to be achieved in Core III. f ( 1-4

The probable mode of f ailure will be short cracks or local blis ters having no significant " ballooning" which could restrict coolant flow or af fect adjacent rods. The location of the !vidual fuel cods in the loose-lattice assemblics have been judiciou oclected to obtain power /burnup comkinations favoring the longest fuc. :1fetimes and those likely to f ail early are located in the center removabic subassembly to permit easy access. In addition to daily sampling and periodic operation of the letdown / charging syste:n radiation detector for monitoring coolant activity, an experimental f ailed fuel monitor has been installed at Saxton. The system utilizes 7 the pressure drop across the steam generator to circulate reactor coolant by a gamma detector. The system is described more fully in Appendix A. 1.5 GENERAL OPERATING PROCEDURES i &ixton Core III will be operated such that the lowest power associated with the followings three design conditions will not be exceeded at any time during Core III steady state operacion: a) a design peak linear power of 24.0 kw/f t in the pu0 -UO 1 se-2 2 lattice assemblies as detennined by power distribution measurements, f or b) a design peak linear power of 19.9 kw/f t in the unirradiated UO 2 i fueled load-follow assemblies as determined by measurements ar { c) 28 FNt I \\ The power associated with the above limit (a) ; t expecten to be 24.9 Ff4t and is expected to be controlling irc n the beginning of Core 111 life to approximatsly 1200 effective full power hours. During this ( i 1-5 E

v i period the power will be slowly increased to 25.8 kWt which is the power associated with limit (b) above. Af ter this the core power will continue to increase to a maximum of 26.5 MWt to maintain the design linear power of 19.9 kw/f t in the load-follow assemblics. A peak pellet burnup of approximately 55,000 MWD /HTM will be achieved f af ter 5000 equivalent full power hours (ETPH * "1.5 MWt) operation. In order to simulate Plant Load Follow Operation from a materials standpoint, cycling between approximately 100% and 40% of full power will be performed. The load follow operation will be carried out in a manner which will not result in a higher peak power in the core than that imposed by the design limits. The node of operation will be based on the following / s 1. Achieving a minimum of 1000 cycles in Core III. [ 2. Existing operational limitations with regned to poder loading [ and unloading. I 3. Core III lifetime. ( 1-6 .imi. mum _-____.--

2.0 CORE DESIGN 2.1 MEC11ANICAL DESIGN 2.1.1 CORE LOADING The 21 main fuel assemblies (9 x 9 rod array) in Saxton Core 111 are made up of seven loose-lattice assemblics, two load-follow assemblies, one peripheral untrradiated UO assembly, and eleven U0 as s emblics y 2 f rom Core 1 arranged as illustrated in Figure 2.1-1. The seven loose-lattice assemblies (described in greater detail below) consist of four 36-rod assemblies, one 35-rod assembly, one 34-rod assembly, and one 32-rod assembly which accommodates, in addition to the 32 fuel rods, a 4 rod removable subassembly. The rods in the loose-lattice assemblies are composed of irradiated rods removed f rom Core 11 ascemblies and loaded into new assemblies containing water-filled hollow tubes at every other rod location. 1 Each of the two load-follow assemblies (described in g,reater detail below) will cont.ain 60 fuel rods with design variations '.n fuel poliet diameter density and internal gas pressure. These assemblies have the same pitch as the standard section 72 rod assemblies. Fuel rod pitch und cross section of the 11 peripheral Core 1 Lasemblies remain identical to those shown in Figure 203.1 of the Saxton Final Safeguards Report. The un:.rradiated peripheral UO assembly contains pelletized, enriched y UO, Zirca; y-4 clad fuel rods (described in greater detail below). 2 The remainin3 fuel loading will consist of six control rod followers which were in Cores I anti II, four removable fuel subassemblies in the peripheral UO; fuel assemblies and the special Irshaped assemblies whic9 were in the peripheral plot positions in Cores I and 11. I 2-1 s)

2.1.2 FUEL ASSEMBLY DESIGN P 1. Overall Construction g3 The construction of the. 9 center assemblies and of the unitradiated peripheral UO assembly remains essentially the same as that af 2 Cores 1 and II, except that these assemblies have been reinforced as iFcribed in Appendix B. No change has been made in the } overall dimensions of the fuel assemblies. Two locations in each of the 21 main feel assemblics in Core III (A-5 and E-1 in Figure 2.1-2) will be used for either flux thimbles, secondary [ source rods, removable is.el rods, removable cladding tes t apecimen assemblies, removable water-filled tubes containing spot veld test specimens or removable solid Zircaloy baro. In the text, removable fuel rods, tubes, etc. are defined as those which can be removed without removing the top nozzle f rom the fuel assembly. A non-removable fuel rod, tube, stc. is cefined as one which cannot g be removed frem the fue3 assembly with ut first removing the top .io z zle. The ce.nter nine assemblies dif fer f rom previous asseelies { in thet the top mozzle is removelle. = [ Each rzmov2ble top nozt is fastened to the assembly by thre" stainless steel tie rods and hald by capci.-e nuts with integral locking cup washers. The tie roda consist of Type 304 stainlesn L steel tubing, 0.391 inch diameter by 0.015 inen wall thickness and hse 304 stainless.tcel special end plugs. The bottom end plug ex each tia.od is welded into the bottom notzle and the rod passes throagh the grids in the same manner as a fuel rod. The top end plug consists of two square cuctions with a threaded section at the top end. The threaded section protrudes through the top no::zie and the captive nuts are torque loaded to 100 in-lbs. One of tha square sections fits into the nozzle eL plate and prevents rotation under torquing. The other square section i shoulders on the nozue end plate thus fixing the fuel assembly overall length and presenting overloading the can. I (1) Suzmary Report on Euckling of Saxton Core 11 Fuel Assemblies and Prevention of ' Buckling in Core III. (See Appendix B) 2-2 i -f _. _ j

.,,.fE a Ctw The top of th: u-' - -^ assembly has a specially prepared reinforced edge which fits into, and is heI/ by, the top nozzle plate. Each can assembly and top nozzle is a matched pair to ensure proper alignment. The fuel assemblies, with top and bottom nozzles, are depicted in F' re 2.1-3. 2. Loose-Lattice Asremblies Table 2.1-1 presents the quantity of each kind of rod, tube or test, which cccupies the 72 locations in each of the loose-lattice assembles. The pitch between fuel rods in the loose-lattice assemblies is 0.820 inches. Removable rod locations E-1 and A-5 (See Figure 2.1-2 for an illustration of the location ident!!1 cation within an assembly) in Table 2.1-1 are shown in Figure 2.1-3. The removable. rod locations in Figure 2.1-3 are vacant because these locations are filled af ter the assembly is fabricated. The loose-lattice tie rods referred to in a foot note in Table ( 2.1-1 will be open to the primary coolant permitting a ready entry and exit of water via two 1/16 inch diameter holes near the bottom of the clad and two 1/8 inch diameter holes near the top of the ludding. The loose-lattice tie rods are shown in Figure 2.1-V.

  • in There are two arrangements of tie rods in the loose-lattice assemblies, a 7}pe A and Type B ioose-lattice assembly as shown in Figure 2.1. 3.

In the Type A assembly the tie rods are located in the B-2, B-8 and H-2 locations within the assembly. In the lyrpe B arrangement, the tie rods are located in the A-2, D-9 and J-4 assembly rod locations. { 2-3

The locations of the f uel rods and water-filled tubes in the Type A the 7~ype B assemblies in order assemblies are the reverse of those 1 to maintain a uniform fuel rod pitch between adjacent loose-lattice as r emblics. As a resul*. the tie rods are in dif f erent locations in the two types of assemblies. The 7 loose-lattice assemblics consis t of f our 36-rod Type A, one 35-rod Type B, one 34-rod Type B, and one 32-rod Type A assembly. The 32-rod --~ + Type A assembly contains,in addition to the 32 rods, a removable subassembly with 4 fuel rods arranged on a 0.758 inch pitch. The louse-lattice assemblies have been strenghtened by spot welding 0.026 inch thick angles between the can and 6 stainless s+ eel water-filled tubes as illustrated in the cross section of the assembly shown [ in Figure 2.1-3. The one inch long clips previously used to f asten the halves of the.an have been replaced by full length clips in the spans between the grids. The grid springs have been reset to give a nominal 6.5 lbs. contact force compared to the 15.5 lbs, force previously used. The combined ef fect of the charges produces a :.9 'ety f actor greater than 1.5 between the forces generated in the assembly and tne buckling strength of the cans. A detailed explanation of the buckling is given in Appendix B. The loose-lattice fuel rods, which are described in the Saxton Core II Plutonium Project Safeguards Report, were selected f rom previously irradiated plutonium rods removed f rom Core 11 fuel assemblies. Figure 2.1-5 depicts a typical plutonium rod. The rods were inspected prior to being loaded into the new assemblies. ( -4 a

r ds The general basis for acceptance of irradiated Pu0 -UO2 2 was satisfactory performance of a rod in Core II and an absence of ancealous conditions as determined by visual and dimensional inspections at the time of reconstitution. Leak testing of the reconstituted loose-lattice assemblics served as a further check on fuel rod integrity. The water-filled tubes, referred to in columns 4 and 5 of Table-2.1-1 and which occupy approximately every other rod location in the loose-lattice assemblies, are shown in Figure 2.1-6. They are f abricated from Zircaloy-4 tubing having the same physical properties as Zr-4 cladding and are filled with end plugs. Each water-filled tube has two 1/16 inch diameter holes at the bottom and two 1/8 inch diameter holes the top of the tubing to allow the flow of primary coolant at through ths rd. Forty-two of the water-filled tubes will be hydrised to surious levels as described in Change Report No. 31 in Appendix C. L The loose-lattice region of the core contains one L-assembly at core location E-2. This L-assembly contains 9 pelletized ( UO fuel r ds clad with type 348 stainless steel. 2 3. Load-Follow Assemblies At the beginning of Core III life, core locations E-3 and i C-3 vill be occupied by load-fcllow assemblies 503-18-3 and 503-18-1 respectively; however, at mid-life the location of these assemblies will be reversed. These two assemblies each contain 60 fuel rods arranged on a nominal pitch of 0.580 inches. They have the same overall dimensions as Core II type fuel rods and contain pelletized 002 fuel, f tw enrichments (9.5 or 12.5 v/o U-235). The d 2-5 y% ug j m--- +' ~

pellet densities range between 8;.5 and 94.5 percent of theoretical ~ l density and have nominal pellet to clad gaps of 5.5 to 9.5 mils. The rods in these assemblies t.re described in greater detail in Reference (2). Each of the two load-follow assemblies also contain 4 water-filled Zircaloy tubes (item 5 in Figure 2.1-3) and one stainless steel vnter-filled tube (iteiu 9 in Figure 2.1-3). These tubes are illustrated in Figure 2.1-6. The tie rods in each load-follow assembly contain filler material. Two contain inconel (item 8 in Figure 2.1-3) and one contains stainless steel filler (item 7 in Figure 2.1-3). 'jhe purpose of filler material is to control local power perturbacions in the assemblies. The tie rods are illus trated in Figure 2.1-7. The load-follow assemblies were also strengthened as described in Appendix B. A solid square stainless steel bar, with two angles spot welded between the bar and the can, is used in 5 place of two stainless steel clad fuel rods in opposite corners of the assembly. An angle. 0.05 inches thick, has been spot welded to the inside of the can at the center of each long span between grids. Full length angle clips are also used between the edges of the can halves as was done with the loose-lattice assemblies. In each assembly the removable rod locatiot. A-5 (See Figure 2.1-2) will be occupied by a solid Zircaloy bar (as described in Change- ) Report 23, in Appendix C). Removable red location E-1 will be occupied by a flux thimble tube in Assembly 503-18-1 at core location C-3 and by a cladding test specimen assembly 4 (as described in Change Report No. 22 in Appendix C) in Assembly 503-18-3 at core Leation E-3. As with the loose-lattice assemblies, locations A-5 and E-1 are blank in Figure 2.1-3 since these locations are filled af ter the assembly is_ fabricated. ( (2) WCAF-7219-L Rev. 1. Addendum to Saxton Core III License Application (Westinghouse Confidential), April 29, 1969. 2-6 4 x m 1 -)

I 4. Core I Assemblies Eleven previously irradiated assemblic; from Core I are loaded k into Core III periphe;al locations. The contents and core locations of these assemblics is given in Table 2.1-2. The pitch and cross section of these assemblies remains identical te that shown aan on Figure 203.1 of the Saxton Final Safeguards Report. An examination of the thermal gradients across these assemblies shows that, the worst thermal gradient would produce a bow of approximately 0.015 inches over the length of the assembly. 1 This bow would always be of an elastic nature and therefore no buckling of the assembly would occur. 5. Unirradiated Peripheral Enriched UO Aseembly 2 \\ One of the 12 peripheral assemblies will be unitradiated. The assembly is 503-10-7 and is located at the.B-2 core Ic.ation. This assembly contains 68 Zircaloy-4 clad fuel rods containing pelletized UO fuel arranged on a 0.580 inch pitch. As explained in Appendix B 2 Reference (1), this Assembly was strengthened similarly to the load-follow assemb'fes. However, this acaembly er tains removable solid Zircaloy bars at both the E-1 and A-5 removable rod locations; whereas, the load-follow assemblies each contain only one solid Zircaloy bar. 6. Mid-life Inspection At approximately mid-life of Core III, a detailed fuel inspection will be carried out. A minimum of 4 assemblies (two load-follow and two , loose-lottice) will be removed from the core and given a visual inspection with a periscope and T.V. camera. The top nozzles will be removed and a sample of both fuel rods end water tubes will be examined. The sample size will be dependent upon the results obtained by examinatic. of a minimum of 5 fuel rods and 5 water tubes. In ( additio., some fuel rods and water tubes will be removed _for a more exteasive examination. 2-7 hs

l l 2.1.3 SUBASSEMBLY DESIGN l l l In general, the removable subassemblies are similar in construction to the main fuel assemblies except that the can is of 0.019 inch thick pe: forated stainless steal and the rods are arranged on a 0.536 inches pitch. On insertion of a subassembly into a fuel assembly, he end plates of the subassembly are compatible with the surrounding nozzle plates so that the coolant encounters a flos path similar to that in a fuel assembly. l The location of the' subassemblies in Core III at initial start-up and l the individual suMssembly comoosition is given in Table 2.1-3. The location of the petipheral subassemblies in the core may be switched during Core III life. The arrangement of fuel rods, water-filled tubes, flux thimtie tubes, etc. within each subassembly is illustrated in Figure 2.1-8 to 2.1-12. I Subassembly 503-4-31 which is initially in the N-1 location is illustrated l 4 l in Figure 2.1-8. This subassembly contains four previousl: irtidiated Pu0 - UO Zr-4 clad fuel rods, a flux thimble in the center rod location 2 2 and 4 water-filled Zr-4 tubes in the four corner locations. Subassembly 503-4-33 is initially in the. N-2 location and is illustrated in Figure 2.1-9. This subassembly contains 2 hydriding_ effects test fuel rods which contain fuel having an nyerage D 0 content of 120 ppm 2 I (see Change Report 18 in. Appendix C) and 2 high pressure creep rods at 1915 psia initial pressure (see Change Report'19 in Appendix C). This subassembly will also contain a flux thimble in the center locatim and 4 mter-filled Zr-4 tubes at the four corner locations. S Anssembly 503-4-25, which is initially in the N-3 -location, is illustrated in Figure 2.1-10. This subassembly contains 5 non-remevable. fuel rods : lad with 304 s.s., enriched to 5.7 w/o U-235. In addition, this subassembly cM. mins 4 removable fuel rods clad with Zr-4, enriched to 5.7 w/o U-235 { with various internal pressures. 2. -,,. . ~, _ _ = - _.. ~,._._ .. ~....

t Subassembly 503-4-32 which is initially in the N-4 location is illustrated in Figure 2.1-11. This subassembly contains 4 non-removable fuel rods clad with Zr-4, enriched to 12.5 w/o U-235; 2 low pressure creep test rods (see Change Report 19 in Appendix C); a flux thimble, and 2 irradiated Puo ~ 2 UO, Zr-4 clad rods. 2 Subassembly 503-4-34 which is initially in the.'-5 location and described in Change Report 17 in Appendix C is illustrated in Figure 2.1-12. This subassembly contains 2 hydriding effects test fuel rods (described in Change Report 18 in Appendix C) and 2 materials compatibility test rode (described in Change Report 17 in Appendix C). k 1 2-9

e n 4= - m TABLE 2.1-1 CORE IDCATION AND OrtPOSITION OF 11MSL-LATTICE ASSEffBLIES

  • Assocla-Re:novable Rnd Locat ion E-1 and A-5****

Assembly Assembly Ntsnbe r Ntsaber of Zr Clad ** of Water-Filled Tubes Sub-ted L-Spot Puo,-00 6econdary Serial Core Puo -10 Not Assembly Assembly Assembly Flux Weld 01 adding 7 Hydrided Hydrided 7~ location Serial Thimble Test Test Ir-4 Clad Nuclear y 2 Number Specimens Specimens Fuel dods Source ( Number location Zr-4 Clad Fuel Rods 503-17-2 D-2 34 (1-VP) 2 25 B none none 1 (E-1) 1 (A-5) 503-17-8 C-2 36(2-VP) 27 A none none 2 503-17-5 C-4 36(4-VP) 25 A none none 1 (E-1) 1 (A-5) 303-17-4 D-3 32 20 A N-1*** none 2 503-17-6 D-4 34(2-VP) 10 17 B none none 2 503-17-3 E-2 36 25 A none 50 3-3-2 1 (E-1) 1 (A-5) 503-17-9 E-4 36(1-VF) 25 A none none 2 I' All of these assemblies contain 3 stainless steel water-filled tie rods, 6 water-filled stainless steel tubes to which 0.028 inch thick angles are spot ' welded. (a total of 72 rod locations). See thange Report 21 in Appendix C for d4tta11ed explanation h is subassee M y 1ccation contains eubassembly 303-4-31 which is composed of 4 water-filled tubec; 4 Puo - U0, Zr-4 clad fuel y 7 tods and a fotx thimble. See Figure 2.1.4 ' T :. planation of alphanumeric location designation E-1 and A-5. t VP denotes number of Puo - UO, Zr-4 clad fuel rods which have vipac fuel. y y

%WI TABLE 2.1-2 CORE LOCATION AND CG1 POSITION OF PERIPHERAL ASSDiBLIES (CORE I) Number of Pelletized Removable Rod Location E-1 and A-5 UO Assembly Assembly Stdinless Associeted* Subassembly Pelletized UO Flux Secondary Serial Core Steel Clad leAssembly Location StainlessStehl Thimble Nuclear Number Location Fuel Rods Serial Number Clad Fuel Rods Source '~ 503-2-4 B-3 61 503-3-5 N-2 2 503-1-12 C-1 70 503-3-1 none 1(A-5) 1 503-2-2 c-5 61 503-3-3 N-3 2 503-1-16 D-1 70 none noo' ' 1(E-1) 1 1 1(E-1) 503-1-11 D-5 70 503-3-4 none 503-2-6 E-1 61 503-3-7 N-4 2 503-1-5 E-5 70 none none 2 503-1-9 F-2 70 503-3-10 none 2 503-1-17 F-3 70 none none 1(A-5) 1 503-2-5 F-4. ~61 503-3-9 N-5 2 503-1-8 B-4 70 503-3-6 none 1(E-1) 1 348 stainless steel clad fuel rods. Each L-Assembly containe 9 pelletized UO2 l

s=m I f TAELE 2.1-3 INITIAL CORE IECATION AND COMPOSITION OF SAXTON CORE III SUBASSEMBLIES Eumber of Number of Number of Number of ' Subassembly Initial Flux Thimble Number of Pellettred Pelletized Pelletized Pelletized Serial Subassembly at Water-Filled 5.7vioUO Rods 5.7 12.5w/oUO ruo -UO Miscellaneous ErCladkods w/o UO 7,r Clad Rbds Rodk fr$m Numbe r Core Center Tubes S.S.Cfad Core II Location ( 503-4-31 N-1 Yes 4 4 thigh burnup) 503-4-33 N-2 Yes 4 2 hydride-120 ppm

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2.2 NUCLEAR D'. SIGN 2.2.1 Analytic Configuration The burn-ups and enrichments in load-follow and loose lattice assemblies use-in Core III nuclear design are shown in Figure 2.2-1. The loose-lattice plutonium fuel region which is made up of seven reconstit2ted assemblies is subdivided into zones distinguished by the amount of burnup the rods rece.ived during their previous operation in Core II. At alternate positions in the loose-lattice, water-filled Zircaloy tubes are installed to provide additional moderator while maintaining the desired flow character-istics. Three of these alternate positions in each assembly contain water-filled stainless steel tie rods. The two UO 1 ad-follow assemblies are installed on the flats of the 2 center nine assemblies (See Figure 2.2-1) and contain veintions in the U-2.34-W enrichment. The purpose of these enrichmente variations is to increase the number of rods that operate near the peax power rat g of 19.9 kv/ft. These two assemblies are to be interchanged at an intemediate point in core lifc. To avoid uceeding an imposed power limit when the assemblies are reversed, the enrichment patterns are identical. As previously stated in the Mechanical Design cection (page 2.1-6), the tie rods in the load-follow assemblies contain filler material to control local power distribution in these as.semblies. Further, each load-follev assembly contains 4 water-filled Zircaloy tubes and 1 vater-filled stainless steel tube as an additional means of improving the assembly power distribution characteristics. In addition, variations in fuel density and pellet dimensions are abss used to flatten ans. pcuer. 2.2.2 Power Characteristics Power distributions throughout the expected life of Saxton Core III were determined using a LEOPARD-PDQ-7 analysis sequence. The calculations show t 2-10 2

that the peak linear power in both load-follow and loose-lattice assemblies s will decrease with burnup with the loose-lattice tods burning down at a faster rate than the load-follow rods. Figure 2.2-2 shows the relative change in linear power calculated for each fuel type. In order to maintain the peak linear power for as long as possible, the reactor power vill be increased as the peaks burn down. Tigure 2.2-3 summarizes the anticipated power operation and the resulting effect on the peak linear power for each fuel type. As shown in this figure the peak linear power is maintained constant for approximately the first 1200 hours operation by increasing the core power f rom 24.9 to 25.8 Wt. At thattime the peak linear power in the load-follow des;. fuel has reached its limit. Thereafter, the core power is increased at a reduced rate holding the peak linear power in the load-follow rods at the design limit. Because of the dif ference in burnup rate the linear power in the loose-lattice rods reduces with increased buruup. At the end of the design life the core power reaches approximately 26.5 Wt and both load-f ollow and loose-lattice rods will have operated at their respective design peak linear powers for a significant part of the total operating period. The actual power operation sequence will be based on power measurements at the beginning of.$.ife and periodically during operation. It is apparent from the operating sequence summarized in Figure 2.2-3 that the M values of core power (26.5 Wt), loose-lattice design linear power (24.0 kw/f t), and load-follow design linear power (19.9 Kw/f t) will l not all occur simultaneously during the operating of Saxton Core III. A small margin l.5 W t,in total core power was included to insure that the j i design peak linear power in each fuel type could be reached even if the local l l peaking factors were less than those anticipated. Therefore, the thermal-I hydraulic and transient evaluations were made at a reactor power level of 28 Wt and with the design peak linear power in each fuel type i.e. the design linear pow 2r of 24.0 kw/ft it. the loose-lattice assemblies and 19.9 kw/ft in the load-follow assemblies. The design core and assembly power distributions are given in Figure 2.3-3. The power characteristics of the fuel rods in the peak 1cose-lattice and load-follow assemblies are shown in Figures 2.2-7 and 5 2.2-8. 2-11

r 2.2.3 Lifetime Characteristics The lif etime available in Saxtou Core III was determined using the LEOPARD-PDQ-7 analysis sequence and supporting one-dimensional PANDA calculations. The anticipated boron concentration requirement as a function of lifetime is summariced in Figure 2.2-4. 2.2.4 Reactivity Characteristics The reactivity characteristics of Saxton Core III were determined by means of both one-and two-dimensional calculations. The analytic procedure used was consistent with that producing gcod agreement between analysis and experiment in two previous Saxton cores. Core III transient analysis physics parameters are listed in Table 2.2-1. Moderator Temperature Coefficient The moderator temperature coefficient was determined in a series of one-dimensional radial PANDA calculations. The calculations were made at the beginning and end of the expected core life with an appropriate range of boron concentration. The results are summarized in Figures 2.2-5 and 2.2-6. Doppler Coefficient Doppler and power coefficient calculations were carried out using one-dimen-sional radial PANDA calculation. The latest methods for calculating fuel ~ temperature were used and an effective resonance temperature was determined by cultiplying this calculated temperature by an empirical factor determined l l from a correlation of power coefficient data from previous Saxton cores. Fuel temperature variations resulting from differences in power, fuel rod . design and fuel type were included. l l f 2-12

.=. The doppler coef ficient is governed primarily by the amount of U-238 present. However, since Core II also contained a small quantity of pu-240, the doppler coefficient in this core was r11ghtly nore negative than that of Core I. Fuel (including both U-238 and Pu-240) will be removed to form the loose lattice region in Saxton Core III. However, the aionne of Pu-240 present will actually be higher than that initially installed in Core II because of the buildup of this isotope during Core II operation. The end result is that the doppler and power coefficient as calculated for Core III is within the range of that determined for Core II. Control Rod Worth The worth of control rods in Saxton Core III was determined by using PDQ-7 two-dimensional calculations. A total bank worth of 18.4% Ak/k was deter-mined which is an intermediate value to that of the two previous cores. The most reactive stuch rod was calculated to be worth 6.4% Ak/k. The bank worth with the maximum worth rod stuck withdrawn is thus 12.0% for Core III as compared to 11.7% for Core II. The worth of boron as a function of concentration was determined using one-dimensional PANDA calculations. The results are summarized in Figure 2.2-9. I Steam Break Calculations and Minimum Boron Requirement i Adequate boron concentration will be maintained during Core III lifetime i l to insure at least a 1% shutdown by rods for the worst possible primary system cooldown that could result from a steam break. Hence the core will l not return to critical as a result of this accident. Detailed calculations were performed to determine the minimum. allowable boron concentration - versus core lifetime. Table 2.2-2 lists the installed reactivity for various Saxton Ccre III conditions and specifies the minimum boron requirement to prevent return to critical, as a result-of a steam break cooldown, by 1% Ak/k. In computing the maximum reactivity addition on cooldown, it was assumed f that a temperature drop to 212*F occurred. The reactivity addition corres-ponds to the integral of the temperature coefficient curve for zero ppm, 2-13

from 495'F to 212*F. Note that the teuperature coefficient is most negative when there is no boron in the coolant. The difference in the maximu:n reactivity addition between the HFP and HZP, BOL cases is just the power defect (approximately 1.2% Ak/k). Conclusions The calculated response of Saxton Core III is intermediate to that of two previous cores. It represents no extrapolation from previous operation except for the higher peak linear powers. Even then, this higher linear power is reached at a lower total core power (28 MWt) than that reached in Saxton Core II (32 MWt). e ( 2-14 ,e_< .*--,,-,-r,-- ,,,.-.,-w..,n-, .4 ,s,, 4 m---,,

~ TABLE 2.2-1

SUMMARY

OF SAXTON CORE III TRANSIENT ANALYSIS PHYSICS PARAMETERS Parameters Core III, 0.0003 Delayed Neutron Fraction, 6,gf Prompt Neutron Life 1.5 x 10' see -5 Doppler Coefficient -1. 6 x 10 &/k'F Modcrator Temperature Coefficient at _4 110*F, 1250 PPM, All Fods Out +.73 x 10 ak/k/*F Moderator Temperature Coefficient at ,4 l H.F.P.,1215 PPM (495'F) -1.0 x 10 ok/k/*F i Moderator Temi.erature Coefficient at End of Life, Hot Conditions (495'F, ,4 Zero PPM Boron) -3.0 x 10 ak/k/'F -6 Pressure Coef ficient at 80'F (BOL) -1.0 x 10 ak/k/ psi l i i Pressure Coefficient at Hot Conditions -6 l (495'F, BOL,1215 PPM Boron) +1.9 x 10 ak/k/ psi Void Coefficients at 80'F i (if 00 PPM Boron) +0.08% ak/k/% void Void Coefficients hot operating (495'F, 1215 PPM Boron) .25% ak/k/% Void Total Control Rod Bank Worth, H.F.P. (495*F, 1215 PPM) 18.4% ak/k Bank Worth With Rod #5 Stuck l (495*F,1215 PPM Baron) 12.0% ak/k l l f l

TABLE 2.2-2 INSTALLED REACTIVITY AND MINIMUM FORON REQUIREMENTS FOR COR Minimum Boron Installed Worth of Uncontrolled Maximum Reactivity Concentration Reactivity Control Reactivity Addition on J ore Condition Cooldown (at 0 ppm) bquirement for: % AK/K Bank With K = 0.99

  • I

% AK/K Rod Stuck Stuck Rod Stuck Rod % AK/K I AK/K Steam Break 30 330 Hot Zero Power (BOL) 11.3 12.0 .3 3.0 0 50 12.0 -3.7 4.2 7.3 Hot Full Power (BOL) Equi Xe 0 0 1.3 12.0 M 3.5 Hot Zero Power -77 5000 Hrs

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e.% I L-Reviced 3/70 APPENDIX A t Failed Fuel Monitor System g ntroduction , A f ailed fuel monitor is installed as an experimental system at the Saxton g checkout and calibration procedure and anticipated eactor. The general layout, erformance are described below. D Fuel element failure is indicated by the increase of gamma activity reau tin However, it should be noted that ? rom fission products in the reactor coolant.

hsra are other sources of gamma radiation such as:

a.

Background

b. N-16 activity Other radioactive isotopes such as corrosion products c. ')escription of System The system utilizes th'e pressure drop across the steam generator to circulate reactor coolant in a bypass loop which consists of about 210 feet of stainless steel tubing, a heat exchanger, a radiation monitor, a remote operated flow control' valve and a remote reading flow meter (Figure 1). The radiation monitor consists of a coil of stainless steel tubing which To reduce the background radiation the coil is shielded surrounds two GM detectors. with 4" of lead. The detectors have an operating mnge from 1 mr/hr to 30 R/hr. The heat exchanger maintains reactor coolant below 140 F to prevent damage A thermoccuple is provided for monitoring the temperature at to the GM detectors. ths detectors. O The remote operated flow control valve and remote reading flow meter permit >0 flow to be adjusted to provide delay time of about 40 seconds. 3 the reactor coolant This delay time provides for sufficient decay of the N-l'6 activity in the coolant )3 full power. to allow proper operation of the system while the reactor is at p The readout system consists of two independent ratemeter channels and a The two channels and their 3 recorder which can record either of the channels. ).) l associated electronica are shown in Figure 2. Checkout and Cab ration The two channels have received a functional checkout and their operating The range, counts / minute vs. mr/hr using a Co-60 source, has been determined. background level in counts / minute for zero power operation has been obtained and a flow rate of 0.42 gpm in the bypass loop has been established as being optimum l l for the system. A:-

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2.3 THERMAL-lWDRAULIC DESIGN s 2.3.1 GENERAL A detailed evaluation of the thermal and hydraulic characteristics of the proposed Saxton Core III configuration has been conducted. The thermal and hydrculic margins are identified. It is demonstrated that sufficient margins exist such that safety limits as defined below, are not exceeded during steady state operation and reactor transients. The maximum fuel temperatures have been conservatively estimated for both the highest power load-follow fuel rod and highest poi er loose-lattice fuel rod. At no time during reactor operation will center melting occur in the core. 4 Extensive calculations have been performed to investigate the hydraulic. conditions to be expected during the Saxton Core III operation. The minimum DNB ratio within the core is greater than 1.30 at the maximum reactor control set point conditions. 2.3.2 TilERMAL AND INDRAULIC DESIGN CRITERIA Acceptable Saxton Core III operating conditions require that several thermal and hydraulic criteria be met. The design criteria and design methods which were imposed and employed were the same as used in the..,4yais of Core II 35 MWt operation. They include: l (1) A DNB ratio for all fuel rod channels of greater than 1.30 at the reactor limiting control set point conditions. (2) Center melting of the fuel is not permitted during normal steady i state operation or anticipated transient conditions. l- '( 2-15

Evaluations were conducted at both nominal and overpower conditions using the THINC code.( 2.3.3 THERMAL AND HYDRAULIC ANALYSIS Best estimate X-Y power distributions are shown in Figure 2.3-1. These values are increased by an 8% nuclear uncertainty before being used in the thermal and hydraulic analysis. Figure 2.3-2 shows the detailed rud by rod power distribution for the limiting loose-lattice assembly (1)-2) and the limiting load-follow assembly-(E-3). These power distributions include an 8% nuclear uncertainty factor. Hot Channel Factors The total hot channel factors for heat flux and enthalpy rise are defined as the maximum-to-core average ratio of these quantities. The heat flux factors consider the local maximum at a point _(the "het spot"), and the enthalpy rise factors involve the maximum integrated value along a channel (the " hot channel"). Each of the total hot channel factors is the product of a nuclear hot channel factor describing the neutron flux distribution and an engineering hot channel factor to allow for variations from design conditions. (1) Chelemer, H. Weisman, J., Tong, L.S. "Subchannel Thermal Analysis or_ Rod Bundle Core", WCAP-7015,. January, 1967. 2-16 l

(a) Heat ylux (Fq) 4 The total hot channel factor is the product of a riuclear hot channel factor describing the nuclear power distribution effects and an cngineering hot channel factor to allow for variations from nominal design conditions. The engineering heat flux factor includes the variations in fuel rod diameter and fuel pellet diameter, density, enrichment and eccentricity. Tables 2.3-1 and 2.3-3 summarize the heat flux hot channel factors for Saxton Core III. (b) Enthalpy Rise (FaH) The enthalpy rise in the hot channel is calculated using THINC. For convenience, the total enthalpy rise factor can still be considered as the simple product of several subf actors. In actuality, the interaction of various effects such as power i distributions, mixing, and flow redistribution cannot adequately be described by simple factors. Parametric 9tudies were used to determine the subfactor attributed to statistical variations in pellet diameter, density, and enrichment, and rod diameter, pitch and bowing. These variations result in an 8% increase in the enthalpy rise of the hot channel (1 + 3c = 1.08). Possible inlet flow maldistribution effects are taken into acccunt by reducing the flow to the fuel assembly containing the hot channel by 7%, based on experimental data. Flow redis-tribution, including the effects of local boiling, and mixing are calculated using THINC. The net result of'these effects is a decrease in the enthalpy rise of the DNB limiting hot channel by 8%. Tabit 2.3-1 summarizes the breakdown of engineering hot channel factors in the DNB limiting channel. 2-17 ~

Because of naheated rods, the total enthalpy rise cannot be considered as a simple product of nuclear factors and engineering factors. Enthalpies at the exist of the hot channel are calcu-lated by THINC and are reported with the nuclear hot channel factors in Table 2.3-3. Departure From Nucleate Boiling The evaluation of the Saxton Core III, DNB conditions have been made using the W-3 correlations ( }. The W-3 DNB design minimum value of 1.30 has been chosen statistically to insure a 95% probability that DNB will not occur with a confidence level of 95%. For fuel channels adjacent to the assembly enclosure, the presence of the unheated wall affects the amount and degree of coolant mixing inside the channel. The effect of the unheated wall on the DNB ratio has been considered (modified W-3 correlation used)( in the design analysis using experimental data. The minimum DNBR during overpower conditions is 1.30, as insured by the reactor trip setpoints (See Section 3.1). Thermal and Hydraulic Design Parameters ( Detailed THINC analyses were conducted for the various limiting thermal and hydraulic conditions in the core. The following overall limits apply at-nominal steady state operation: l 1. 28 MWt 2. 19.9 Kw/ft in load-follow assemblies ( 3. 24.0 Kw/ft in-loose-lattice assemblies l l l (2) L.S. Tong, " Prediction of Departure from Nucleate Boiling for an l Axially Non-Uniform Heat Flux Distribution", J. of Nuc. Energy l - Vol. 21 pp. 241-248,-(1967)._ (3) L. S. Tong, et. al., " Critical Heat Flux-on a Heater Rod on the Center of Smooth and Rough Square Sleeves, and in Line Contact with an Unheated l Wall", ASTM 67-WA/ilT-29 (1967). y 1 2-18 s .Q ,,s..---,e, ,w- ,-,_w,s..,.--s.,. ,-e n -+,... -.n.y ,,,,,,,4 n._,. ,y,,-

Note that these 11 nits do not occur simultaneously (See Section (1.5). The 24.0 Kw/f t limit occurs at 24.9 MWt at the beginning of life; whereas, the peak fuel rod power in the load-follow assemblics (19.9 Kw/ft at 25.8 MWt) does not occur until later in life (s1200 EFPH). Table 2.3-2 presents the thermal and hydraulic characteristics for Core III at its design power of 28 MWt. Table 2.3-3 presents the thermal and hydraulic characteristics of the various types of assemblies and sub-assemblies, each at their worst design conditions. Initially Core III power (24.9 MWt) is limited by the design linear power in the loose-lattice (24.0 Kw/ft) which occurs in the loose-lattice assembly at core location D-2. Further, the Core III hot channel is in the assembly at core location D-2 which has a minimum steady state DNBR=1.81. At this time all other assemblies in the core have lower linear powers and higher DNBR's than the assembly at core location D-2. \\ Af ter about 1200 EFPH, Core III power is limited by the design linear power in the load-follow assembly (19.9 Kw/f t) and occurs in the load-follow assembly at core location E-3, which has a minimum steady state DNBR=1.75. That assembly which limits core power changes from loose-lattice to load-follow after about 1200 EFPH, because the peak rod linear power in the loose-lattice burns down at a faster rate than that of the load-follow assemblies. The core hot channel after about 1200 EFPH is also.in the load-follow assembly at core location E-3. All other assemblies in the core at this time have lower peak linear powers' and higher DNBR's.. Core power will not exceed a nominal 28 MWt. Further, that nominal power at which the measured power distribution combined with analysis gives 19.9 Kw/ft in load-follow assemblies, or 24.0 Kw/ft-in loose-lattice assemblies will not be exceeded. ( l 2-19 l l l- .,-,.-,-n- . _.. _. _.., ~.., _ _, _,,

- ~_- - ~ _ -.. -... The worst condition for the peripheral fuel assemblies occurs at a core power of 28 MWt. The assembly at core location B-3 has the highest peak linear power (14.9 Kw/f t) and a minimum DNBR=3.30. If the subassembly having the worst local flow conditions (Materials Compatability Subassembly) were put in the highest power location (assembly location B-3), the minimum DNBR would be reduced to 2.35. In summary: 1. The minimum DNBR occurs in the load-follow assembly at core location E-3. 2. The highest peak linear power occurs in the loose-lattice assembly at core location D-2. 3. Feripheral subascemblies have higher DNBR's and lownr peak linear powers than the asse... lies at core location E-3 or D-2. 2.3.4 THERMAL ANALYSIS OF FUEL ROD The characteristics of the two basic types of fuel rods which will be exposed to the highest thermal conditions during Core III operation have been established. The power and burnup history which were employed for Cote III operation are shown in Figure 2.2-3. The thermal conditions' and character-istics are summarized in Table 2.3-2 with the temperature estimates being provided below. Fuel Central Temperature The maxionna fuel central temperatures for the core at the design conditions has been estimated to be 4540*F for the loose-lattice rods and 4650*F for the load-follow rods. At the maximum overpower conditions of 112% power a maximum central fuel temperature of 4880*F is predicted for the loose- -lattice assemblies. This-is approximately 120*F below the fuel melting I temperature predicted for this mixed oxide fuel. At this 112% overpower 2-20 l l _.,_ ~ ___ _ _.. - _ _

condition, the highest power load-follow fuel rod (22.2 Kw/f t) is expected to have a maximum central temperature of 4900*F. This is also below the melting temperature for UO at the equivalent fuel burnup. 2 i Fuel Clad Temperature At the design fuel power condition of 28 MWt about 150 of the loose-lattice fuel rods (60% of total loose lattice rods) can be expected to operate with a mean cladding temperature at some point on the fuel rod greater than 700*F. All of the 120 load-follow fuel rods are expected to operate with some small fraction of their cladding at a temperature greater than 700'F, at the maximum 28 MWt reactor operating power level. At the design basis power condition i (all of the uncertainties in the power calculation considered to be maximum at the same time) the mean cladding temperature at the peak loose-lattice power condition (24 Kw/ft) would be approximately 725'F. At the comparable load-follow condition (19.9 Kw/ft) the maximum mean cladding temperature would be 720*F. These temperatures were used in the eva!uation of mechanical limits. \\ l i ( 2. , _,, ~,._.. _ _..__ _ _.

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

h TABLE 2.3-1 Engineering Hot Channel Factors Pellet Diameter, Density F Enrichment, and Eccentricity 1.045 9 Rod Diameter, (Pitch and Bowing) Pellet Diameter, Density Enrichment 1,08 E F R d Diameter, Pitch and Bowing 4H Inlet Flow Maldistribution 0.92 Flow Redistribution and Mixing s E 1,00 Fg, Total l l l I l ( y -ew,,- ,r. ,-w,v-w- ,.r'*,r+cie v - gn- ~w w - vr-,.-

-.... -.... ~ - -. -.... _.. - TABLE 2.3-2 Thermal and Hydraulic Design Parameters

  • Total Core Total Heat Output 28.0 MWt 6

Total Heat output 95.56 x 10 Btu /hr Heat Generated in Fuel 97.4% System Pressure - Nominal 2250 psia System Pressure - Minimum - Steady State 2200 psia 6 Total Flow Rate ** 3.21 x 10 lb/hr 6 Effective Flow Rate for Heat Transfer 2.73 x 10 lb/hr Flow area for Heat Transfer Flow 2.41 ft Average Velocity Along Fuel Rods 6.28 ft/sec , Coolant Temperatures Nominal Inlet 480 F Maximum Inlet Including Instrument l Errors and Deadband 485 F Average 9(se in Vessel 26.0 F Average Rise in Core 27.0 F Average in Vessel 493.0 F l' Average in Core 493.5 F Heat Transfer i 1 Active Heat Transfer Surface Area 2 l of Fuel Rods 375.3 ft Average Heat Flux 219,400 Btu /hr-ft Average Thermal Output 6.58 kw/ft-Maximum Clad Surface Temperature at Nominal Pressure 657.4 F -(

  • These thermal and hydraulic design parameters supercede all those previously reported for Core III.
    • At 63 cycles

% LL 2.14 THERMAL AND ifTDRAULIC DESIGN f'ARXtETERS WORST Ut.SlCN O'NDITIONS lead-Ts%1ow loore-ta t t ice loose-Lat t ice Peripheral Assembly Peripheral Assembly Peripheral Assembly 6 Assembly st f.est.=bly at Assembly at at Core location F-4 at Core at Cor e location B-3 j Case Location Core locat8en Core Location Containing Materials location Containing Materiata Center Core Region E-3 D-2 D-3 Compatability Suba gesbly s B-3 Campatebility Subassemb1L F Heat Flux Hot Channel Factor 3.28 4.10 3.0 1.84 2.26 2.58 4 F Nuclear Radial Tsetor 2.23 2.95 2.80 1.32 1.62 1.86 F ' Nuclear Axial Factor 1.35 1.33 1.33 1.33 1.33 1.33 l Nominal Outlet Enthalpy in Hot onannel, stu/lb 542.5 528.1 524.9 510.4 5.2.7 618.9 l i Saturation Enthalpy at Minimum Steady State Freasure, Stu/lb 695.0 693.0 695.0 695.0 695.0 695.0 l Maximum fleet Flux l Btu /hr-ft 662,300 800,200 761,900 402,800 495,200 633.600 Maximum Thermal Output, kw/ft 19.9 24.0 22.9 12.1 14.9 19.0 Hinimum W-3 DNB Ratio at 1002 Power Nominal Condition 1.75 1.81 2.09 3.44 3,30 1,9n Core Power at which mininnan W-3 DNB Ratio occura, MWt 25.8 24.9 24.9 28,0 28.0 28.0 As a result of more refined calculations, it has been determined that the nazisam linear power for the wteriale competability subassembly in any Core III peripheral cubassembly location will be 19.0 kw/f t rather than the 16.1 kw/f t reported in Change Reports 17 and 18 (See Appendix B) and that the hot channel factors are somewhat higher than those given in Change Report 17

1 B C D E F l -l I, l 0.54 0.60 0.56 --l Y .n 9 4 1.82 0.54 _9 0.59 1,77 2.14 e a 1 a { 4 l 1 A ij

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.66 3 .79 1.92 1.92 L. 2.50 1 l I t ,l 3, _ >l 3 l l y, _ I 0.53 1.63 2.11 1.70 0.62 l '~ _4 m ~ s. n n< q \\ w m, y i l 0.49 O.62 0.47 l ~ n l I CORE RADIAL POWER DISTRIBUTION Figure 2.3-1

_~ DESIGN NUCLEAR RADIAL POWER FACTORS IN THERMALLY CONTR01. LING CORE III ASSDiBLIES i (Values include 8% uncertainty factor) 1.84 1.86 1.88 1.88 1.94 1.97 2.01 2.02 2.03 2.10 2.15 2.17 2,17 2.24 2.24 2.27 2.27 2.24 DESIGN PEAR ROD 2.36 2.37 2.33 2.50 2.50 2.48 2.33 . 59 2.56 2.49 2.67 2.76 2.70 2.52 2.40 2.95 . 2.90 l LOOSE-LATTICE ASSDiBLY AT CORE LOCATION D-2 l 2.24 1.96 1.80 1 2.16 2.24 2.26 2.21 1.15 2.03 1,64 2.21 2.23 2.25 2.04 1.87 1.59 2.24 2.16 2.22 2.25 2.11 2.03 1.86 1.65 DESIGN PEAK ROD KW RT = 19.9 2.23 2.20 2.16 2,06 1.92 1.85 2.24 2.15 2.09 2.19 2.22 2.09 2.02 1.87 1.65 2.25 2.22 2.20 2.20 ' 1.97 1.85 1.63 2.28 2.22 2.22 2.25 2.13 2.00 1.66 l { 2.24 2.28 2.22 2.17 2.00 1.75 1.65 LOAD FOLLOW ASSEMBLY AT CORE LOCATION E-3

2.4 TUEL PERTORMANCE EVALUATION 2.4.1 BACf4ROUND The objectives of the Saxton Core III irra.diation program are tot 1. valiJcte fuel element performance predictions, including determination of pwer/burnup failute limitsi 2. demonstrate performance capability of Zircaloy clad oxide fuel elements over a 1d spectrum of burnups and pwer 2evels; and 3. obtain depletion characteristics and transuranic isotope generation data for high burnup, mixed oxide fuel. Because the Sa.xton Core 11 mixed oxide fuel rods were designed for relatively low peak rod average burnups and operation at power linear pwer densities < 16 kw/f t, there is a significant risk of failure of certain of these rods in Core III.( ) By careful selection and placement of these rods in the * >ose-lattice assemblies, it is possible to control their burnup and operating power IcVels and th is permit power /burnup limits to be established while operating safely and in full compliance with the reactor license Technical Specifications. The Core III loose-lattice region has been designed for a peak linear power of 24 kv/ft. Design analyses predict clad strains and internal gas pressures which exe.eed design limits early in Core III life if rods with highest prior burnup are operated within 10% of the peak power. Thereforo, it is necessary to ret.trict the highest power rods in Core III to those of relatively low prior burnup (j 18,000 MWD /HTM). Furthermore, the high burnup rods will be limited to operation at lower linear powers in order to achieve acceptable lifettmes. The high burnup rods which are likely to fail first will be located in the center removable subassembly to permit easy access should removal of these rods become necessary. ( (1) WCAP-3385-52, Saxton Plutonium Program, Mechanical and Thermal-flydraulic Design of Partini Plutonium Core, December 1965. i I 2-22 l l l

2. . 2 EXPECTLD TUEL pERp0RMANCE Op SAXTON CORE Ill RODS The operation of all rods in Saxton Core III has been analyzed and the results compared with the following pVR fuel element design criteria: 1. Cladding strains less than 1%. 2. Interns'

casure less than external (system) pressure (except for certain land-follow rods).

3. Clad stresses less than 0.2% yield strength at operating conditions. 4 No center melting of fuel. DNB ratio of at Icast 1.3. The latter two limits are considered in Section 2.3 of this document. Table 2.4-1 summarizea the projected power-burnup combinations for all rodstobeirradiatedinthecenternineassemh,hicsofSaxtonCoreIII. It includes ' aoth beginning-of-life and end-of-life burnups as well as I peak linear powers expected for fuel rods in the seven loose-lattice assemblies. The table alt.o shows the predicted peak linear power and burnup tor the rods in the two load follow assimblies. The analysic of expected fuel performance is based on current design procedures and material properties data. Among the uncertainties considered in the hot channel interpretation of the code predictions were the nucicar and ( engineering f actors.' Core Il power history, projected Core III power levels, e and the dimensional uncertainties (diametral gap, fuel density and plenum size). The fuel rod performance was generally analyzed on a "most probable" basis; however, the studies aleo evaluated the ef fect of " worst combinations" of dimensions, and projected power history, i Analysis of the four rods to be located in the center removable subassembly indicates a high probability of failures.ggg}g_in Core 111 life because of large expected cladding strains (> 21) relative to the 1% strain design criterion and high internal gas pressure. Any fcilures in these rods would provide a better indication of performance capabilities of other ( 2-23 l J

- - _. - ~. l l rods in the experiment, i.e., help predict when f ailures are likely to occur and confirm the mode of failures. The design codes also predict failvecs in some of the loose-lattice rods outside the subassembly due l I to cetesolve internal pressur at moderately high burnups.

However, such failures are expected to be statistical, i.e., will not all occur at one time, because of the broad spectrum of burnups and linear powers represented and statistical variations in diametral gaps, densities, etc.

l Previous observations of high power CWR and Saxton test rods indicate that f ailures will occur as either: 1. short circumferential cracks associated with clad ridging at pellet interfaces, or 2. local clad blistering with short, randomly oriented cracks. In either case, cladding strains will be small with no significant " ballooning." Thus, any such f ailures will not restrict coolant flow or have significant ef fecs on adjacent fuel rods. In addition, evidenet. to date on defected rods shows that the cracks will not propagate to produce catastrophic fuel rod failures. Furthermore, the nature of the defects (short ruptures) would limit the activity release and thus permit continued operation of the reactor with a limited number of failed rods. Since intentional operation j to failure is planned for the loose-lattice fuel, thermal-hydraulic analysis of the flow blockage required to reduce the steady-state DNBR at 112% power to a lower limit of 1.3 has been performed. Figure 2.4-1 summarizes these results and indicates that diametral strains would have to be >35%. Therefore, loose-lattice c1&dding failures are unlikely to result in violation of the DNBR lower limit of 1.3. Cladding strains in failed rods are generally found to be small (<5%) compared to the 16-35% required to produce a steady-state DNBR of less than 1.3.(2) ( )WCAP-3850-3, Post-Irradiation Examination of CWR, Fuel Assemblies. August,1968 (See Appendix D for appropriate pages). 2-24 i

In sumary, the Core III f oradiation testin6 is expected to provide valuable information in the valf.dStion of fuel element design procedures. t j a h t t a '2-25 5 --.. -. a -

TABLE 2.4-1 6 J SLHMARY OF POWER RATINGS AND btT10WPS EXPECTED TOR CORE III Expected Peak Pe ak Pe ak Group Linear

  • uwer No. of Initial Burnup.

Expected EOL Burnup, No. kw/ft Rods MWD /KTM MWD /MTM y 20 4 > 32,00% 55,000 LL-2 21.2-18.7 47 16,000-19,000 39,000-45,000 LL-3 18.7-it.3 43 19,000-21,000 39,000-44,000 LL-4 17.4-14.3 57 21,000-23,000 38,000-44,000 LL-5 14.3-10.6 78 23,000-27,000 36,000-44,000 LL-6 10.6-8,5 20 27,000-32,000 36,000-45,000 Total 249 LT-1 17.6-15.8 60 0 19,000-21,000** LF-2 15.8-14.1 24 0 16,800-19,000** LF-3 14.1-12.3 12 0 14,700-16,800** LF-4 12.3-8.8 4 0 10,500-14,700** Total 120 These four rods ure in the center removable subassembly.

    • These values assume constant exposure at the peak linear power i

indicated. Since the two load follow assemblies are to be l interchanged midway through Core III life, the peak burnup values for the high linear power rods will be slightly reduced from the tabulated i values. L f LL - Loose-Lattice LF - Load Follow (

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3.0 SAFETY ANALYSIS 3.1 GENERAL A no fuel melt limit has been inposed for the high linear power loose-lattice assemblies during Core III operation. The effect of this limit on the permissible fuel rod operating conditions is illustrated in Figure 3.1-1. The figure shows the calculated center fual melt limit in terms of linear pow r density and peak fuel pellet burnup. This relationship has been ialculated for the loose-lattice fuel rods considering: 1) The dependence of the mixed oxideb melt temperature with fuel burnup. 2) The reduction in fuel center temperature, as estimated using the Laser Computer program, due to flux depression.( } 3) The burnup dependent fuel rod center temperature due to irradiation i and time induced changes in the thermal and physical characteristics (i.e., pellet-clad gap conductance, fuel swelling, clad creep, fission gas release, etc.). Figure 3.1-2 provides the fuel melt temperature and flux depression fuel burnup relationships which were employed to generate the Saxton fuel i melt limit curve shown in Figure 3.1-1. Also included !in Figure 3.1-1 are four representative curves showing the design-overpower limits for different initial peak pellet burnups. These curves in combination with the proposed power variations (Figure 2.2-3) have been used in selecting power density /burnup combinations and hence, the fuel rod locations in the loose-lattice assemblies to prevent fuel center melt during Core III operation. 1 Poncelet, C.G., " Laser - A Depletion Program for Lattice Calculations Based Upon MUlT and THERMOS", WCAP-6073, April 1966. t i 3-1 l ~..... .~

1 It should be noted that 12.5% of the peak pellet power density in Figure 3.1-1 is nuclear and engineering uncertainty. In addition the peak pellet 1s evaluated at the 112% overpower condition. There is a high degree of assurance that actual fuel temperature would be less than indicated by the design overpower limit curves because in the above approach the nuclear uncertainty, the engineering uncertainty, and the trip accuracy uncertainty have been combined additively rather than statistically. The overpower trip setpoint for Core III operation will be set for a power no greata than 5% above the lowest power associated with the following three conditionst fueled loose-lattice assemblies a) a design prek of 24 kw/ft in the pu-002 as determined by power distribution measurements, or fueled load b) a design peak of 19.9 kw/f t in the unitradiated U02 follow assemblies as determined by power distribution measurements, or l c) 28 Ma't other trip setpoints are as follows: a) Hot leg trip cetpoint 511*F* b) Low pressure trip setpoint 2125 psia 6 l c) Low flow trip setpoint 3.05 x 10 lbs/hr d) Low M-G set frequency trip setpoint 60 cycles /sec

  • The setpoint for reference power level of 28 MWt is $11*F.

For initial operation at 24.9 MWt the hot leg trip setpoint will be reduced to 508'F. ( 3-2 -v, ,e;e---- -r e

The n.inimum DNB ratio occurs in the 3oad follow assemblies. The sbeve reactor trip setpoints ensure that reactor trip will be initiated prior to reaching a minimum DNB ratio of 1.30. The transient resultira in the minimum DNB ratio is the loss of flow incident which is re-analyzed in detail in Section 3.2. The rod withdrawal accident results are nearly identical to that presented for Core II 35 MWt operation. (See Figure VI-I Appendix E), only ansturptions e and f were modified since the maximum overpower reactor trip is 112% for Core III versus 114% for Core 11 and that the limiting value of the ~ cioderator temperature coefficient is -0.5 x 10 ak/*F for Core III versus -2.0 x 10 Ak/*F for Core II. The reduction in the trip setpoint has the stronger effect, since the reactivity feedback is mostly dependent on the fuel temperature coefficient and only very slightly dependent of the moderator temperature coefficient. As a net result, the nuclear flux peak is now slightly below 120% and the hot spot heat flux is below the 114% reported for Core II (See Figure VI-I in Appendix E). Thus, the minimum DNB ratio is higher than the 1.31 DNB ratio reported for Core II. This is a direct result of the higher steady state DNB ratio of Core III (1.75) versus Core II (1.52). At the minimum DNB point during the transient, there is still a 3.5% power margin to a DNB ratio of 1.3. Hence, the core is adequately protected against the rod withdrawal accident. The loss of coolant accident is re-analyzed since the higher fuel temperature and heat flux in the peak power density region of the core during Core III operation would result in increased clad temperature transients. The steam line break accident has been analyzed previously in Core II 35 MWt Safety Analysis Report, the appropriate pages of which are in Appendix E. Although the linear power deasity of the peak rods are considerably higher in Core III, the steady state DNB ratio (1.75) for Core III is approximately 15% higher than the Core II (35 MWt) steady state DNB ratio (1.52). As a result, the control and protection system will trip the 3-3

reactor and terminate the first phase of the steam break accident with 3 greater margin of DNB than the one reported for Core II (35 Mk't). s even 4 Af ter the reactor trip is initiated, the principal concern for this accident is tbc possibility of the core returning to significant pwer as a result of loss of shutdwn margin resulting from the primary cooldwn and the negative moderator coefficient. As discussed in Section 2.2, the worth of the Saxton control bank is sufficient to maintain a 1% shutdwn margin / considering the maximum possible reactivity addition from cooldwn and ,[ the highest worth control rod stuck out of the core (see Table 2.2-2). ) Hence, there would be no return to pwer. l I k 3 s S _ _, ~ -. -..

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3.2 LOSS OF C001. ANT " LOW The loss of flow analysis presented in the Cort II 3$ MWt Safety Report has been repeated for Core III operating conditions. The transient W-3 DNB ratio was evaluated only for the load-follow assemblies since the loose-lattico assemblies were not DNB limiting. Other assumptions and initial conditions were as follows: a) Reactor poser level is at 103% of nominal rating (28 MWt) as result of possible calorimetric errors. b) Inlet temperature is 485'r allowing 5'T for instrumentation error. c) Priuary pressure is at 2230 pria allowing 50 psi for instrumentation error. d) Maximum expected absolute value of the fuel temperature coefficient (Doppler) is: -1.0 x 10-5 ak/*F. 4 e) Minimum expected absolute value of the moderator temperature coefficient is -0.5 x 10'0 ak/'F. f) Scram delay of 1.1 see (0.5 seconds due to instrumentation and 0.6 seconds due to rod motion in a region of small effectiveness). g) The hot spot heat flux is evaluated for the maximum fuel gap (9.5 milu, cold). As shown in Figure 3.2-2, the largest fuel to clad gap results in the highest heat flux response after reactor scram. h) A negative reactivity insertion rate of 2.22 x 10 ok/ seconds (reactor trip). l 4 1 ( l l l 3-5 t

The flow coastdown curves with and without MC set inertia are shown in Figure 3.2-1 and $dA'also y.iven f n the Core 11 Safety Analysis Report. The flow coastdown with inertia is also presented in Figure 3.2-1 and was obtained f rom recent experimental measurements at the Saxton Plant. Figures 3.2-2 and 3.2-3 show the neutron flux response, the average and the hot spot heat flux responses with and without MG set inertia. With MG set inertia a minimum DNB ratio of 1.52 occurs at approximately 1.5 seconda as shavn in Figure 3.2-4. Therefore, there is adequate margin to DNB and fuel damage will not result from this accident. With out inertia, the minimum DNB ratio is about 1.14 and clad damage could be expected in the hot channel. However, as shown in Figure 3.2-5, the W-3 DNB ratio decreases below 1.3 only for channels within 10%.of the hot channel power. For the Saxton Core 111 power distribution, this amounts to about 80 fuel rods of the load follow assemblies. Hence, f or the very unlikely conditions that no inertia is availabic, the core damage will be limited to approximately 60 fuel rods. This is slightly better than similar conditions analyred for Saxton Core Il at 35 MWt, where approximately 107 rods were found to reach DNB. l l l ( 3-0

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t SAXTON 28 MWt OPERAT20N LOSS OF TLOU ACCfDENT i TOTAL SCRAM DELAY: 1.1 SECONDS -1,x30~5{AK/K/'F 4 /K/'F 5 DOPFLER REACTIVITY C0 TTICIENTt . 5 x 10-MODERAT0B PEACTIVITY COEFFICIENT: TRIP REACTIVITY: 3% TUEL GAP: 9.5 MILS i c.c j I i + 1 i 4 j } I i i l I l_ i I i i l' . ;...p..,._a ..._ J 4 _j ...,.,. a... l. l j l-1 I t '/ i + o.0 r c l 1...{ l -r-, k 3.- ; 7y, __ _L'.... l 1 L_._l i i l i I l I/ i I I r-'l ----l i g . g.._

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3.3 JfSS OF COOLANT l ( The loss-of-toolant accident has been re-evaluated for Saxton Core III. The assumptions and analytical techniques used are as followst 1. The moderator teeperature coefficient has been re-evaluated. A less negative coefficient of -0.5 x 10 ' (ak/k)/*F is now used in the analysis rather than the value of -2.0 x 10~4 (ak/k)/'F l previously reported. 2. The heat transfer models in the LOCTRA-R2 core thermal analysis code are the aamn as those described in Section 4 of the report i "Saxton toss-of-Coolant Accident Prevention and Protection" except that stabic film boiling immediately following the occurrence of DNB at 0.5 seconds, is no longer assumed. Instead, post.DNB heat transfer coefficients during the transition and stable film boiling phases of blowdown are now calculated using an empirical correlation developed by Westinghouse, from steady state heat transfer data. r It has been compared with experimentally determined transient data recently obtained as part of the Westinghouse Flashing Heat Transfer research and development program.(1} A comparison of the measured heat transfer coefficients obtained from the transient blowdown data and the coefficient calculated with the empirical correlation is presented in Figure 3.3-1. It is concluded that the heat transfer coefficient during the transition and stable film boiling phases of blowdown can be conservatively estimated by the recently developed Westinghouse empirical correlation. 3. The two load follow assemblies contain both Zircaloy and stainless steel clad fuel rods.( ) 4. Reactor power level is 28 HWt. 1 \\ l l l 3-7 l _m. .... _,. -.. ~, .w .r ._.r,,..,,,,,-

6 The specific cases analyzed and a summary of the results are as follows: 1 Total % Core Total % Zr-H O 2 Bre ak,S,i_re Clad Melt Reaction Double Ended Break - 1.28 f t 8.9 12.5 Intermediate - 0.173 f t 1.6 7.5 Surge Line - 0.0375 ft 0.0 1.6 Figures 3.3-2 through 3.3-4 show the clad temperature transients for the high power density Zircaloy clad rods located in the Loose-Lattice assemblies. Presented in these figures are the clad temperature transients for rods operating at various fractien of the design peak linear heat rate of 24 kw/ft. Figures 3.3-5 through 3.3-10 show the clad temperature transients for the Zircaloy and stainless steel clad rods located in the load follow casemblies. The design peak linear heat rate in these assemblies is 19.9 kw/ft. s Since the stainless steel clad rods located on the core periphery are operating at approximately 33.1% of the peak linear heat rate of 24 kw/ft, the peak clad temperatures exhibited by these rods would be much lower than those presented in Figures 8 through 10. The clad melt is 2imited to 8.9% for the double-ended break with no melt occurring for the surge line break. REFIRENCES 1. Farman, R. F., Cermak, J. O., " post DNB Heat Transfer During Blowdown", WCAP-9005, October, 1968. Westinghouse Proprietary Report. 2. Helehan, J. B., Addendum to Saxton Core III License Application. WCAP-7219, July 16, 1968. Westinghouse Confidential Report. L 38 4 r -, ww .-e,

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y p.bea2_J.me.,+_ ~. -m -smeAm.44._m (Aa p% d. 4 en a a w _.w,4 a 4. APPENDIX A FAILED FUEL HONITOR SYSTEM A-1

APFDDIX A ,(' Failed Fuel Monitor System Introduction A failed fuel monitor is installed as an experimental system at the Saxton Reactor. The general layout, checkout and calibration procedure and anticipated performance are described below. Fuci element failure is indicated by the increase of gama activity resulting from fission products in the reactor coolant. However, it should be noted that there are other sources of Samma radiation such as: a.

Background

b. N-16 activity c. Other radioactive isotopes such as corrosion products Description of System The system utilizes the pressure drop across the steam generator to circulate reactor coolant in a bypass loop which consists of about 210 feet of stainless steel tubing,.a heat exchanger, a radiation monitor, a remote operated flow control valve ( and a remote reading flow meter (Figure 1). The radiation monitor consists of a coil of stainless steel tubing which surrounds two GM detectors. To reduce the background radiation the coil is shielded with 4" of lead. The detectors have an operating range from 0.01 mr/hr to about 1 r/hr. The heat exchanger maintains reactor coolant belov 140"F to prevent damage to the GM d.etectors. A thermocouple is provided for monitoring the temperature at the detectors. The remote operated flow control valve and remote reading flea meter permit the reactor coolant flow to be adjusted to provide delay time of ab ut 40 seconds. This delay time provides for sufficient decay of the N-16 activity in the coolant to allow proper operation of the system while the reactor is at full power. The readout system consists of two independent ratometer channels and a recorder which can record either of the channels, l'he tve channels and their associated electronics are shown in Figure 2. Checkout and Calibration The two channels have received a functional checkout and their operating range, counts / minute vs. mr/hr using a Co-60 source, has been determined.- The ~ background;1evel in counts / minute for zero power. operation has been obtained;and r = L _ f a flow rate of 0.k2 gpm in the bypass loop has been established as being optimum for the system. t -4 e

il Revised

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3/70 fatem Characteristics esign Pressure Primary 2500 psia. Secondary 150 psia O ,j tesrials Primary: 1s", Type 316 Stainlesa Steel Tubing Low, primary, maximum .73 gpm D O low Control Valve remote, air operated fail open, Valve Hammel-Dahl Masonellan Valve ~ Positioner Sentry Equipment Corp. sat Exchanger Shell 150 psi, tube 2485 psi Design psi 350 F, 6800F 0 Design temp 225 psi, 4880 psi Test psi

700F, 700F Test temp Tubing:

3/8" O.D. x 0.035 wall )ils: Radiation Monitor, Experimental 0.D.. 3/8 inches 3/8 inches wall 0.065 0.035 i length 9.5 13.5 turns 24 34 /P Cell Foxboro Integral Orifice Orifice diameter 0.159" 0-42 sph .M. Tube - Phillips Model 18509 Operating voltage 500 to 650 volts R ngs 1 mr/hr to 30 R/hr l g e I e 4 4

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l [ t i 'GM1 GM2 n- % h-b-$RADIATIONMONITOR e S g - a' { ALARM PRE COAX CABLE AMP COUNT RECORDER . RATE METER COAX CABLE JUNCTION BOX JUNCTION BOX o o o CONTAlt MENT VESSEL REACTOR CONTROL ROOM. EXPERIMENTAL INSTRUMENT ROOM PRE AMP RECORDER-COUNT COUNT -/ N. RATE RATE -E'N ./ /. METER #1 METER #2-r ~ ' -FAILED FUEL MONITOR. SYSTEM - ELECTRICAL: ' FIGURE A-2 A l l

a. ,.a_ 4 Ae m.. --s ,#ueA _..Mr,i Jt.,>_ 4_ Ja a2,a4,.- 1. k APPENDIX B SUM 1ARY REPORT ON BUCKLING OF SAXTON CORE II FUEL ASSEMBLIES AND PREVENTION OF BUCKLING IN CORE III i B-1 h a

SUFDIARY Buckling of the cans on some Saxton Core II-assemblies was observed after Core II operation. The extent of the buckling of the Core II cans and the design modifications used to prevent buckling in Core III are herein reported. The buckling of the central plutonium assemblies was caused by frictional forces between the_ grids and fuel rods arising from differential thermal expansion between the stainless steel assembly can and Zircaloy fuel cladding. The buckling of the peripheral uranium dioxide fueled assemblies was caused by thermal gradients across the assembly and was limited to those assemblies with the Core II type grid design. The major modifications to the loose lattice assemblies to prevent buckling during Core III operation consist of: (a) reduction of the grid f riction loads through resetting of grid springs; and, (b) stiff ening of the can structure through the use of full length clips between can halves and replecement of six Zircaloy water tubes by stainless steel water tubes with angle braces welded between the tubes _and cans. To prevent buckling in the load follow assemblies during Core lli operation, the assemblies have been modified by reducing the grid friction loads through resetting of grid springs and stiffening _ of the can structure through: (a) replacement of fuel rods in two corner locations-by square stainless steel bars with angle braces welded between the bars and cans; and, (b) angles welded to the inside of the can between fuel rods on the long sides of the can. None of the buckled assemblies will be reused in Core III. f i B-2

~ l l l i 1.0 SUL1MAfJ,0F, BUCKLING j i l 1.1 Core,IJ, Plutonium,Assemblieg l t Buckling was observed in eight of the nine central plutonium assemblies. The buckling appeared to be of a random nature with l no apparent pattern or consistency. However, the four corner assemblies of the square pattern formed by the central nine assemblies appeared, in general, to have the worst buckling. t The maximum lateral deflection of the buckles were estimated by 1 I visual inspection to be 0.06 to 0.08 inches. The center span between the second and third grids of the plutonium assemblies experienced the worst buckling with the greatest frequency, the frequency and severity of buckling decreasing towards the end of the assembly. The direction of the buckling (toward or away from the fuel rods) appeared to be completely random and the severity of buckles independent of direction. s i Rub marks, which were observed on several assemblies, could be ( attributed to handling or contact with spacer bars in the spent fuel cask during shipment. However, in at least one case it is definitely concluded, based on the appearance of the marks, that the rubbing occurred in the core and resulted from interference with a control rod assembly. The single plutonium assembly which did not exhibit buckling contained eighteen stainless steel' clad rods. The effect of l stainless steel clad fuel rods would be to reduce the friction load exerted by the Zircaloy rods and increase the effective j strength of the can (the stainless rods being put into comprension l as well as the can during differential thermal expansion). The one plutonium assembly which contained eight stainless steel clad rods was found to have only minor buckling. By analysis, this number of stainless rods is insufficient to prevent buckling of the assembly but the reinforcing effect evidently did reduce (_ the extent of buckling in this assembly. B-3

Core,II,UO, Assemblies 1.? 2 Buckling was observed on three c! the seven Core II design UO2 # <4 assemblies. In this case, however, the buckling was minor in ature and generally was restricted to the upper spans on the sides of the assemblies facing the center of the core. No buckling was observed on any of the Core I design assemblies used in Core II. A detailed summary of the buckling is given in Table 1 and Figures 1 and 2 show some typical buckles. Figure 3 shows a core r cross section indicating buckled assemblies. l I \\ \\ l B-4

t TABLE 1 Core Buckling Serial No. Tyype Location Observed 503 l-7 SS Clad UO Rods ID No 2 503-1-19 SS Clad UO Rods 3F No 2 503-1-10 SS Clad UO Rods SD No 2 503-10-6 SS Clad 00 R ds 1C Yes Minor 2 503-10-2 SS Clad UO Rods 2B No y 503-10-3 SS Clad UO R ds 2F Yes Minor 2 503-10-4 SS Clad UO Rods AB No 2 503-10-5 SS Clad UO Rods SE No y 503-12-2 Zr Clad Puo -UO Rods 2C Yes 2 y 503-12-5 Zr Clad Puo -UO Rods 2D Yes 2 2 503-12-4 Zr Clad Puo -U0) Rods 2E Yes 2 503-12-3 Zr/SS Clad Puo -UO Rods 3C ~ No g 2 503-12-6 Zr Clad Puo -UO Rods 3E Yes 2 2 503-12-7 Zr Clad Puo -UO Pods 4C Yes 2 2 503-12-1 Zr/SS Clad Puo -UO Rods 4D Yes 2 2 503-12-8 Zr Clad Pu0 -UO Rods 4E Yes 2 2 503-7-1 SS Clad UO Rods 1E No 2 503-2-3 SS Clad U0 Rods 3B No 2 503-11-1 SS Clad UO Rods 4F No 2 503-11-2 SS Clad UO Rods SC Yes Minor 2 503-13-1 Zr Clad pug -UO Rods 3D Yes 2 2 l B-5

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ksf' {/? .\\. t l ', \\ {'- \\. %_. s ]k.', i,)$% L ..., ' 'l' D \\ s s. _s e i s'AN ,N' jf M 'r\\< , g( / f s /N / l 's- /_. M N / s s. i M ,g a'd' --~~ / l ~~ ~ ~..... . M.' / \\ 1 P ;.t w m s atsu,.uwe w \\ l' 80Ctaf9e & VtlM stat frv9.t l TICURE 3: Saxton Core Cross-Section Showing Buckled Fuel Assembly Locations "B" l B-7 s Q

i 2.o geysE_gy_pyggp3yg The following possible loading methods were examined as buckling modes for the assemblies: A. Externally applied loads on the assembif due to: 1. Shipping and handling 2. Interference with reactor internals b. Loading generated _ internally to the assembly through: 1. Frictional effec;s during differential expansion 2. Temperature differentials across the assemblics 2.1 Externally,3pplied, Loads, Evidence of buckling due to thcse types of loading would have been the collapse of the fuel assembly end spans between the nozzles and end grids. The required end loading would also have had to have been of such magnitude that the top nozzle hold down springs would have collapsed. Examination of the assemblies showed no evidence of cither of these conditions. In addition, analysis showed that the conditions in-the reactor during handling operations, which would be necessary to produce i this magnitude of loading, could not be realistically predicted. It was concluded, therefore, that the buckling did not result from externally applied loads. 2.2 Ig3pyga}}y_ggggyagef_[ gags 2.2.1 Plutonium, Assemblies (Zircaloy Cladding) Calculations show that the buckling of. these assemblies' occurred due to dif ferential expansion on initial heatup., The buckling, however, would be of an elastic nature at that point, disappearing on subsequent cool down, except for some small amount of permanent set resulting from ( B-8 l l l l -

i l l relaxation due to irradiation. The inrge buckles developed j by a ratchetting mechanism through a number of full l temperature cycles of the core, the buckles growing by the i additional permanent set occurring with each cycle. l l A f ull temperature cycle is from cold shutdown, through hot ( operating temperature back to cold shutdown conditionr., l l Examination of the reactor's thermal history, shows that only 1 five such cycles had occurred prior to the mid-life detailed observation of three fuel assemblies. Estimations of the i permanent set indicate that only small deformations would have been present at this stage; this is probably the reason that the buckling was not observed. Subsequent to this observation, eighteen additional cycles occurred which account for the large buckles observed at the end of life. 2.2.2 E92_ES9lfd_Assggb}igs (Stainless Steel Cladding) Although the UO2 assemblies which exhibited buckling were of the Core II design with six point contact grid support, the buckling in the assemblies is not attributed to frictional loading. In these assemblies, both the assembly can and fuel cladding are stainless steel. Therefere, any 1 j friccional loading in the can caused by differential expansion j between the rods and can at operating temperatures would be small and would result from tensile ctresses in the can. In addition, buckling was only observed on the hot side of the cans and not randomly distributed around all sides as I would be expected uith axial friction loads. ( It appears, instead, that the observed buckling in these assemblies resulted from a combination of thermal gradients across the assemblies _and the resistance to bowing exerted l by the rod bundles in the grids used in the Core II-design. 1 l B-9

The six point contact support used in the Core Il grids provide an effective built-in condition for the rods at each grid location and thus resist bowing of the assemblies through restraining moments on the rods. If the rmal gradients sufficient to produce bowing in an unrestrained conaition were present in the Core II assemblies, the restraint offered by the grids could result in compressive buckling stresses on the hot face of the ansemblics. This would not be the case with Core I design assemblies where the grids provide a four point support for the rods and little restraining tement. Examination of the core temperature distributions based on power distributions during Core II operation showed four assembly positions where thermal gradients would be sufficient to cause buckling in Core II design assemblies and one location which was marginal. Of the four locations, one was occupied by a Core I assembly which exhibited no buckling. Two of the remaining locations were occupied by Core II assemblies which did exhibit buckling. The Core II assembly occupying the fourth location showed no buckling. In this case, the actual average temperature of coolant flowing through the assembly is in question. From instrumentatlan in the 3 x 3 test assen.bly (503-4-29) which was suspended in the 9 x 9 assembly at this location, coolant temperatures approximately 20*F below expected were indicated. Because of channeling effects through the 3 x 3 assembly, the indicated temperature would be basically the discharge temperature from the 3 x 3 assembly and would reflect the effect of deleted fuel rods in the assembly. However, the low temperatures in the 3 x 3 would also tend to reduce coolant temperatures in the 9 x-9 assembly and thus reduce temperature gradients across the assembly since these are directly related to coolant temperature. Although the temperature effects cannot. he accurately predicted, it would appear that buckling did not ( occur because of reduced coolant temperatures. B-10 l l 2 -~rr- =- J w + - - e

s The last of the three Core II assemblies which exhibited buckling was in the marLinal location where the thermal gradients were not sufficiently high to predict buckling. The buckling in this case, however, was very minor and localized and could possibly have resulted from a local weakness in the can (a thin ligament or out of flat condition). l l [ t ( B-11 l S m e m m

d I 30 U221f!SoI!9F_9f_99BE_!!! e!!!U!h!E!_I9_fBEYEEI_!!'95!!EE The center fuel assemblies in Core III were of Core 11 design and contained Zircaloy water tubes and/or Zircaloy clad fuel. Therefore, they would experience high friction loads due to differential expansion between the stainless steel assembly can and the Zircaloy rods and would be expected to buckle. To prevent buckling during Core III operation, the Core 11 assembly design required modification. There were two possible approaches to assembly modification to prevent the bucklir.g: 1. Increase the stiffness of the can sufficiently to withstand imposed loads. 2. Decrease the friction load by reducing the normal force applied to the fuel rods by the grid springs. A combination of both approaches was used. The spring contact force was reduced to the minimum which would not risk fretting of the fuel rods. However, this minimum contact force would still be sufficient to cause buckling. Therefore, the can was also strengthened. 31 h2989_ hat 3}ge,63sem)}}gg The fuel assembly has been strengthened by replacing six Zircaloy water tubes by six stainless steel ones and spot welding 0.028 inch thick angles between the can and the tubes. The one inch long clips previously used to fasten the two halves of the enclosure l have been replaced by full length clips in the spans between grids. A cross-section of a repaired loose 16ttice assembly is shown in I l Figure 4. The grid springs have been reset to give a nominal 6.5 lbs contact-force compared to the 15.5 lbs force previously.used. The combined effect of the changes produces a safety factor of 1.5 between friction forces generated by the grids and the buckling strength of l the cans. r k B-12 .,.. - ~ =

t?ESTlHGHOUSE ELECTRIC CORPOR ATION 1 ( ) r % ( ~ ?, v v J v J ~- 0 0 0 0 0 0 0 0 C 0 0 0 0 0 0 0 0 0 0 0 0 0 0 ^ sD O O O O O O O L 0 0 0 0 0 0 0 0 L O 0 0 0 0 0 0 0 0 O O Cl O O GIDOQ FIGURE 4: Cross-Section of Repaired Loose Lattice Assembly -(

  • B-13' S

e --woe -e n,-r, 'e ,n,-, w-w ,,-me v + ~ -rm+<

i \\ 3.2 kgap,[g}}gy, Aggggb}igs A slightly different method has been used for the load follow assemblies to obtain the required strength. A solid square stainless steel bar, with two angles s' pot welded between the bar and the can, is used in place of two stainless steel clad fuel rods in opposite corners of the assembly. A full length angle 0.05 inch thick has been spot welded to the inside of the enclosure skin at the center of each long span between grids. Full length angle clips are also used between the ends of enclosure halves as was done with the loose lattice assemblies. A cross-section of a repaired load follow assembly is shown in Figure 5. The MAPI assembly to be used in the periphery of the core will be similarly treated. The same buckling strengtn factor has been achieved for these assemblies as in the loose lattice.

3. 3 Peripheral,UO,Assemblles 2

Because of the large thermal gradients across these assemblies in Core III, only Core I type will be used. An examination of the thermal gradients across the assemblies in the whole core shows that, although no buckling will occur, the worst gradient will produce a bow approximately 0.015 inch over the length of the assembly. This will always be of an elastic nature and will cause no interference problems. 3.4 _Thergal, Hydraulic, Considerations, The modifications used for both the loose lattice and load follow assemblies have not compromised the thermal-hydraulic performance in Core III operation. For both types of assemblies the minimum DNB ratio will not be below the current limit (1.30) specified in the operating license for the Saxton reactor. ( B-14

y bNs 1 WESTlHGHOUSE ELECTRIC CORPOR ATION i J' s 'O O L O 0 0 0 0 0 0 0 0 0 C O ~ r. r., r) [3 [- ~ / a u. %/ I ( ( O O C 0 0 0D 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 000 J_ 0 0 0 0% FIGURE 5: Cross-Section of Repaired Load Follow Assembly ( B-15

1 A With the modifications discussed herein buckling is not anticipated in either the new or irradiated assemblies used in Core III. Assuming buckling did occur, however, up to 0.060 inch inward deflection of the can surfaces can be tolerated without danger of exceeding the minimum a'llowable DNBR. Deflections of this magnitude in the outward direction could also be tolerated without

problems, i.e.,

binding of control rods would not occur. The fuel assemblies will be inspected at mid-life for any evidence / of can buckling before reinsertion for continued reactor operation. In addition, the control rod scram times are normally checked every six months as'a check against gross buckling. 3.5 Nuclear, Aspects The modifications proposed will have no effect on the operation of the load foli,w assemblies. However, the addition of stainless water tubes in the loose lattice assemblies will reduce the power i output of these assemblies by approximately 3%. This may be alleviated slightly by increasing the nominal power output from 26 MWT to approximately 26.3 FRCE without seriously affecting any thermal hydraulic or nuclear margins. 3.6 Health,and, Safety The modifications to Core III assemblies of Core II design will prevent buckling and present no hazard to the health and safety of the public. The assemblies will be given a detailed examination for buckling at mid core life. ( B-16 i w g e-y----e-- 9, c,.-

I i ,a APENDIX C CHANGE REPORTS 17 to 23 i ( C-1 .___________i________

Docket No. 50-146 DPR-4 Change Report No. 17 page 1 of 3 Pages t Change Report No. 17 MATERIALS COMPATIBILITY TEST 1. DESCRIPTION OF Cl!ANGE A subassembly designated Test Assembly No. 503-4-34 uill contain either a primary or an alternste materials compatibility test in which case it will be designated 503-4-34A. The subassembly containing the primary mater'.als compatibility test is illustrated in Figure 1 and the primary test rod in Figure 2. 1ne subassembly containing the alternate materials cor patibility test is illustrated in Figure 3 and the alternate test s ctions are illustrated in Figure 4. The subassembly for the primary test will contain two fuel rods in addition to the two materials compatibility test rods; whereas, the subassembly for the alternate test will contain no fuel rods. The rods which will be with the primary materials com-patibility test will be clad with Zircaloy-4 having a nominal wall thick-ness of 23 mils and contain uranium dioxide (UO ) pellets, uniformly en-g 2 riched to 12.5 w/o U-235 and 89.5% to 94% theoretical density, a. Primary Test Design The primary design illustrated in Figures 1 and 2 consists of a Zircaloy-4 tube to which are attached 10 stainless steel sleeves, 2.425 inches long, equally spaced over the tube length. The sleeves are mechanically attached to the Zircaloy tube by bulges. Each stainless steel sleeve in the subassembly grid area has welded to it four inconel spring fingers which hold the tube in place by pressing against a stainless steel cylinder which is welded into each grid of the test subassembly, b. Alternate Test Design The alternate design is illustrated in Figures 3 and 4. In this design the Zircaloy tube test sections containing stainless steel l' t C-2

Docket No. 50-146 DPR-4 Change Report No. 17 Page 2 of 3 Pages 4 sleeves are notched and held between the subassembly grids by the notches. As in the primary test design, the sleeves are mechanically attached to the Zircaloy tube by bulges. Section B-B of Figure 3 is an illustration of the cross section of the alternate design subassembly showing the arrangement of the test sections in the assembly. The outer two cylindrical cross sections are part of the grid structure while the inner two are test sections. 2. PURPOSE OF CHANGE The purpose of the change is to allow substantiation of the expectation of negligible crevice corrosion between stainless steel and Zircaloy-4 in an operating PWR environment. 3. SAFETY CONSIDERATIONS The parameters pertinent to safety considerations for the primary ma-terials compatibility test design are presented in Table I for the as-sembly containing the tests and two 12.5 w/o U-235 enriched fuel rods. The location of the fuel rods and materials compatibility tests in the assembly are shown in section B-B of Figure 1. It'can be seen in Table I that the calculated values of specific power, heat flux, DNB ratios, and fuel rod and rod surface temperatures arc well within safe limits. Further, it'is highly unlikely that the ma-terials compatibility test should fail in such a'way as to interfere with the heat transfer-from the two fueled rods. It is highly unlikely in the extreme that a failure of an alternate test section could_cause an unsafe condition since the subassembly containing the backup alternate test will contain no; fuel rods. 3 C-3 0

Docket No. 50-146 DPR-4 Change Report No. 17 Page 3 of 3 Pages t 4 HEALTil AND SAr0TY It is our ennelusion that the health and safety of the public will not be endangered by this change. I l l .e C4

-.-__-~_ . _ =, l IABLE I T}lERMA1, AND IIYDRAULIC PAPAMETERS TOR MATERI ALS COMPATIBILITY l TESTS IN ASSEMBLY CONTAINING TWO 12,5 w/o U-235 ENRICllED RODS T ( Mwt.) 480*F inlet Maximum Linear Pvver Density 16.1 kw/ft Maximum Surface lleat Flux 540,000 ETU/ilr. ft.2 Ilot Channel Factors Fq 2.43 T 4.01 g3, Minimum DNB ratios 100% Power (2250 psia T 480'F) 2.45 inlet 112% Power (2075 psia T 490'F) 1.93 g,3, l Mean Clad Temperature at maximum ourface heat flux 100% Power 711*F 112% Power 719'F Clad Stresses

  • Maximum (10,000 psi Clad Strain l

l Maximum <0.5% l l Fuel Center Temperature 100% Power (16.1 kw/f t) 3750'F 112% Power (18.1 kw/ft) 4120'F (

  • Pressure tensile stress, hot, operating end-of-life l

l C-5

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_ _ - - = _. - -.. - - . -. _ -.. _ -. ~,. -. Docket No. 50-146 DPR-4 Change Report No. 18 Page 2 of 2 l' ages 4 Approval to operate rods similar to these at a maximum calculated operat-ing stress of 22,000 psi has been previously given in Change No. 30. Fuel rt.ds containing 30-75 ppm moisture have previously been successfully operated at Saxton. It is our opinion that this change does not represent any unreviewed safety question. 4 IIEALTil AND SAFETY It is our conclusion that the health and safety of the public will not be endangered by this change. l l l C-11

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DPR-4 Change Report No. 19 Page 1 of 2 Pages A Change Report No. 19 TEST FUEL RODS - CLAD CREEP BEHAVI0R 1. , DESCRIPTION OF CHANGE Four test fuel rods will be inserted in Saxton Core 111 periphoral test subassemblics. Two of the test rods will be high pressure creep rods and two low pressure creep rods. The four rods are clad with Zircaloy-4 having a nominal thickness of 23 mils and contain uranium dioxide (UO ) 2 pellets, uniformly enriched to 12.5 w/o U-235, having less than 30 ppm 1120 and are helium filled. The high pressure creep rods contain 94% theoretical density fuel pellets and have a 10 inch gas plenum where as the low pressure creep rods contain 89.5% theeretical density fuel pel-lets and have a 2 inch gas plenum. 2. PURPOSE OF CHANGE The purpose of the change is to assess the effect of irradiation on clad creep behavior. 3. SAFETY CONSIDERATIONS Parameters pertinent to safety are given in Table 1. The maximum calculated operating stress in the rods is 16,000 psi which is much less than pre-i viously approved stress levels. The stress level in the low pressure creep rods waa not determined since it will be much less than that in the high pressure creep rods, which is well within acceptable limits. Rods similar l to the low pressure creep rods were previously approved in Change No. 22, 24 and 29. The high pressure creep rods will be similar to rods previously approved in Change No. 30, the principle dif ference being that the high pressure creep rods will operate at a 6,000 psi lower stress level than those of Change No. 30. It is our opinion that this change does not represent any unreviewed safety l ( question. C-14 i

DPR-4 Change Report No. 19 Page 2 of 2 Pages ( 4 HEALTil AND SAFETY lt is our conclusion that the health and safety of the public will not be endangered by this change. 4 l l t l I ( C-15

O l i l TABLE I SAFETY PARMTETERS FOR CREEP f3ST RODS Test Rod Type Peak Power Initial Pressure cold Operating Stren Level

  • kw/ft psla psi High Pressure Creep

<16.1 1915 <16,000 Low Pressure Creep <16.1 290 negligible compressive stress a j l l i

  • Pressure tensile stress, hot, operating, end-of-life

. - - - ~..... - - - - - -.. -.. Docket No. 50-146 DPR-4 Change Report No. 20 Page 1 of 2 Pages l l Change Report No. 20 Cor@OSION TEST JJ Z1RCALOY-4_ SPOT WELDS AND BI-METALLIC SANDWICH i l 1. DLECEIPTION QP CHAMCE -n~ n-Test sectione containing candwich spot velds as shown in Figure 1 vill tse suspended ver'.' ally in four removable water filled tubes as shown in Figure 2. .he tubes containing the spot veld tests will replace four standard water tubes in two of the central nine Core III assemblies. The loc.ation ut the four removable water tubes which contain the tests arc shown in the following table. L Assemb1v Core Location Tube Location in A eembiv* $03-17-8 C-2 E-1 and A-5 503-17-9 E-4 L-1 and A-5 The following taolo is a tabulation of the parts shown in Figures 1 and 2. Part No. Title Materig Figure 1 1 Top Eng Plug Zircaloy-4 2 Retainer Zircaloy-4 3 Pin Zircaloy-4 l 4 Retainer Zircaloy-4 5 Retainer Zircaloy-4 l 6 Retainer Zircaloy-4 7 Clip Stainless Steel 304L 8 Clip Inconel 718 Figure 2 9 Water Tube Zircaloy-4 10 Bottom End Plug Zircaloy-4

  • See Figure 3 for explanation of assembly location designation.

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Docket No. 50-146 DPR-4 Change Report No. 20 Page 2 of 2 Pages l 2. PURPOSE Or CHANGE The purpose of the change is to permit the evaluation of corrosion and hydriding ef f ects on spot welded joints in an operating pres.surized l water reactor environrnent. l 2. SATETY CONSIDERATIONS The test sections containing bi-metallic sandwich and spot welds will be suspenued vertically in the water tube by being welded to the upper end plug. Coolant access will be through 0.125 inch and 0.062 inch diameter bleed holes in the upper and lower sections of the tube respectively. The bleed holes are such that it is highly iniprobable that material large enough to be detritnental to the reactor operation could get from inside the tube into the coolant. It is our opinion that this change does not represent any unreviewed safety question. 4 HEALTH AND SAFETY It is our conclusion that the health and safety of tht. public a111 not be endangered by this change. l C-18 4

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l l Docket No. 50-146 DPR-4 Change Report No. 23 Page 2 of 3 Pages i i ruel rods at restricted coolant channels were replaced by solid Zircaloy i bars to maintain thermal-hydraulic safety mergins and to minimize the disturbance of the power distribution in surrounding fue3 rodo. The following table is a tabulation of the thermal and hydraulle design parameters which still pertain to Saxton Core III operation. Total Core i Total Heat output 28.0 MWL 6 Total Heat output 95.56 x 10 Stu/hr Heat Generated in Fuel 97.4% System Pressure - Nominal 2250 psia System Presaure - Minimum - Steady State 2200 paia 6 Total Flow Rate

  • 3.21 x 10 lb/hr 6

Effective Flow Rate for Heat Transfer 2.73 x 10 lb/hr l Flow area for Heat Transfer Flow (unit cell) 2.2 ft Average Velocity Along fuel Rods 6.83 ft/sec l Coolant Temperatures l l Nominal Inlet 480 F Maximum Inlet Including Instrument Errors and Deadband 485 F Average Rise in Vessel 26.0 F l Average Rice in Core 30.5 F Average in Vessel 493.0 F Average in Core 495.2 F Here Transfer Active Heat Transfer Surface Area of 2 l Fuel Roda-376.2-ft 2 Average Heat Flux- - 220,400 Btu /hr-ft Average Thermal Output 6.62 kw/ft Maximum Clad Surface Tenperature at Nominal Pressure 657.4 F-t. .At 63 cycles l-j C-33 e rs v -en-- >,,t-w..r-~$w,-,-*yr*-retv-w*e- ---

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~. Docket No. 50-146 DPR-4 Change Report No. 23 I i Page 3 of Pages i I i i L Loose Load Lattice rollow l ,A_s sembly As seinbly ( U0 -Pu0; UO l Center Core Regign 2 2 i Fq Heat Flux Ilot Channel ractor 3.65 3.01 1 F Nucleat Radial Factor 2.62 2.10 T Enthalpy Rise llot Channel gg Factor 2.94 2.35 NUC Y Nuclear Axial Factor 1.33 1.38 g Nominal Outlet Enthalpy Hot Channel 528.3 Btu /lb $40.9 Btu /lb Saturation Enthalpy at Minitoum Steady State Pressure 695.0 Btu /3b 695.0 Btu /lb 2 2 Maximum Heat Flux 799,000 Btu /hr-ft 662,300 Bru/hr-ft Maxistum Thermal output 24.0 kw/ft 19.9 kw/ft W-3 DNB Ratio at 100% Power l Nominal Conditionu 2.0 1,7$. 1 ( l It is our opinion that this change does not represent-any unreviewed safety question. 1 4 IIEALTil AND SAFETY lt is our conclusion that the henith and safety of the public will not be endangered by this change. l. ~ C-34 Li

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en re tr WE571HGHOU$E EL ECTRIC CORPOR ATION i, lull 1cngth anyle as e.hcr n in Location of Zircaloy Bar in Figure 5 of Heterence (1) Unit ra llated l'eripheral UO3 As s eni-bly an<i Load rollov Asscebly gm ~ ~5 =% D'O O O OOOO e O QO O O O OO DO OO 00 O O ~ J 0 00 00 0 0.. ~ \\ 0 0 O 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 ,0 0 0 0 0 0 0 0 g O O o C aO Ocm - = - - location of Zircale-y Bar in Full length angle a9 shown in Univradiated Peripheral t'02 Figure 5 of Refeuneo (1) Assembly FIGURE 2 LOCATIONS OF SOLID ZIRCALOY BARS WITilIN ASSEMBLIES { (1) Summary Report on Buckling of Saxton Core 11 Fuel Assefablies and Prevention of Buckling in Core 111. C-36 l

l l l l i APPDIDIX D PARTS OF WCAP 3850-3 WilICil DESCRIBE EXPERIMENTALLY OBSERVED STRAIN IN CVTR FUEL RODS The ridging described-in the following pages is at most a fev-mils while blisters of up to 8 mils were-observed. This corresponde. to <$%. l l r 9 l l t k~ 4 D-1 .-e,-~, .w -e p. v -c.n,,- ,--,N we y ,r,.. -, v r -,,w,-,,--.N.++A'+-r, .,w-ww.r-+-w,+ ua.n,-n,'-w---- -,w-, ,--,-.c,,-.w,, w w.m.v v, w w, m

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3. 4.1.1. L Gama Swming, Fuc t hole E 3. ?C1 ad U. EC Failed fuel rod 53.861 and sister rod $3.862 were scanned for gross gamma

[ activity in the Transfer Building pool after wire spacers had been removed. The results, averaged over scans taken 90' apart, are given in Figures 3-3 and 3-4 Peak-to-average activity was 1.17 f or rod 53.861 and 1.16 for rod $3.862. Integrated activity for rod 53.862 was about 5.5 percent lower than the failed rod, which can be accounted for by the combined uncertainties of the instru-mentation (!$ percent accuracy) and hot channel f actors (for example, the flux skew f actor, F, is 1.025 f rom critical experiments *). 4

3. 4.1. 2. 3 frofilometry, Length Ncasuremente Typical profilometer traces are shown in Figures 3-5 and 3-6 for rods 53.861 and 53.862.

The scan for the latter rod shows no anomaliesi average diameters over the fuel bearing pertion of the rod f all between 0.4871 and 0.4883 inch. Fuel rod ovality (as estimated by comparing adjacent minima and maxima on the helical trace) varies between 0.2 and 240 mils. At no point along the rod do the measured ditmeters lie outside fabrication specifications of 0.486/0.489 inch. Failed rod 53.861 is characterized (Figure 3-5) by clad ridging, in teatrast to the relatively smooth trace from its sister rod. The height of the ridges is nominally 1.0 - 1.5 mils, and the spacing between ridges is approximately the length of a-fuel pellet 0.650 inch. Areas of ridging generally are con-centrated within 25 to 58 inches from the bottom of the rod. Average diameters f all within 0.4870 and 0.4885 inch, but localized high ovality (up to 3.2 mils) or ridging result in extremes of 0.4860 and 0.4902 inch.. The measured post-irradiation length of rod 53.862 was 103.094 inches. A pre-irradiation measurement is not available for this rod 6 although the rod had been identified as being within fabrication specifications (102.851/102.967-inches). Using the nominal dimension, the length increase during irradiation was 0.12 percent. No 12ngth measurement 'eas made on f ailed rod-53.861 because of the defect at the upper end plug.

  • CVNA-28 7, CVTR Fuel Management, pp. 125-134 D-2 t

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. -.. - _ - ~. -.. ~ - - - - - - - -. -.... -. = -. l 3.4.:.1.4 he:it.r Leak Teeting s The bottom (^.100 inch) section of fuel rod 53.861 beneath the def ect was leak tested with the apparatus described in Section 3.3.2 and Figure 3-1. No sec-ondary leaks were detected in the fuel tube or lower end plug. The minimum detectable leak with the helium test is estimated as 4 x 20'O cc/sec. (STP). Fission gas collection and analysis f rom sister rod $3.862 appeared normal and no further leak testing was considered necessary. 3.4.1.2 Special Test Assembly G-4

. 4.1. E.1 Vicual Examisation, Failed Rod 83.831 Axial locations and orientations (with respect to the wire spacer hole in the lower end plug) of the blisters on the f ailed rod are shown in F.igure 3-7.

Photomacrographs of the blisters at 14 and 16-1/2 inches from the bottom were presented in W(.\\P-3850-2, p. 42, Dec. 21,1967. The 14-inch blister was ini-tially observed as def ective, lut the 16-1/2-inch blister may have been damaged during post-irradiation handling. As measured by the profilometer, the 14-inch blister (largest) was s7.8 mils high and extended axially over approxi-mately 600 mils of the clad. Crud deposition was very light and unif orm on the f ailed region and probably was not a factor. A large ridge (Section 3.4.1.2.3) was observed 83 inches l from che. top of the rod. In addition, there were indications of other ridged areas on the rod and sister rod 83.832, but the small magnitude of che individ-ual ridges made identification impossible, even with the aid of the profilometer scans. 3.4.1.2.2 Gamma Ecanning, Fuct Rode 83.831 and SL.832 Failed fuel roo 83.831 and sister rod 83.832 were scanned for gross gamma ac-tivity in the Transfer-Building pool.- The-axial distributions arc presenced in Figures 3-8 and 3-9. Peak-to-average activity is 1.15 and 1.16 for rods 83.831 and 83.832, respectively. The integrated activity-of rod 83.832 is about 2.7 percent higher. The curves shown here have not been normalized f or r 'k D-3 hw ,,, - - - ~ - -

,. -,~.-.-

decay with the G-3 rods (Figures J-3 and 3-4) and the discrepancy in activity levels reflec ts the dif f ering power and decay histories of the two assemblies (Table 3-1). 3.4.1.: 3 Profilometry, Length Reasuremente The surface profile of failed rod 83.831 is characterized by blisters and rids-ing. Figures 3-10a and 3-10b show the trace in the vicinity of large clad blis-ters at 16-1/2 and 17-1/2 inches from the bottom, and also show a typical ridg-ing pattern at 144 inches. Regions of regular ridging (N1 mil ridge height) exist over most of the fuel-bearing zone of this rod, along with several larger (s2 mils), discrete ridges. A particularly prominent ridge (%4 mils) was ob-served at the approximate top of the f uel column, 83 inches f rom the botten. Average clad diameters (excluding local blistered and ridged regions) range between 0.4884 and 0.4907 inch. This rod is characterized by areas of high ovality, particularly in the plenum region (sB3 to 100 inches) where the oval-ity exceeds 4 mils. Sister rod 83.832 also showed areas of regular (%1 mil) ridging and larger, discrete ridges similar to the f ailed rod. Again, high ovality (55 mils) oc-curs in the plenum zone, and within the bottom 412 inches of the rod (2 - 4 mils). Average rod diameters vary between 0.4866 and 0.4878 inch over the fuel-bearing region. Post-irradiation length measurements (10.002 inch) for the G-4 rods were: rod 83.831 (f ailed), 103.160 inch, rod 83.832, 103.149 inch. No explicit pre-irradiation lengths were measured f or these rods, but caking the nominal length of 102.909 inch, the length increases are 0.24 percent for rod 83.831 and 0.23 percent for rod 83.832. Taking tha maximum length allowed under the f abrica-tion specifications. the increase f or these rods is 0.18 percent or greater. 3.4.1.2.4 Helium Leak Testing The defective region of rod 83.831 (9 - 18 inches) was sectioned out and the remaining bottom and top pieces were tested for secondary leaks. No leaks were found. i 1 D-4 _ -.. - _,. ~ ~

s APPENDIX E ROD WITl!DRAWAL AND STEAM LINE BREAR ACCIDENT ANALYSIS FOR SAXTON CORE II 35 MWt OPERATION i l l l l ( r.

VI. ACCIDENT ANALYSIS A. GENERAL Tha increased power level (35 MWt as compared to the 23.5 MWt licensed operating power) and the corresponding changes in operating conditions and instrument settings require that some of the inci-dents previously reported (in the Saxton Final Safeguards Report and in the Safeguards Report for the Saxton Reactor Partial Plutonium core II) be re-evaluated. Information in Section 502 of the Final Safeguards Report relating to the possibic causes of incidents and the safeguards provided applies to the incidents analyzed for this report and will not be repeated. B. REACTIVITY INCIDENTS 1. , Rod Withdrawal Incident An urcontrolled rod withdrawal is assumed to occur from an electrical or mechanical failure in the nuclear inst;umentation and control systems or by oper:itor error. In this unlikely event, an electrical interlock ensures that only one of the two control rod groups would be withdrawn. Assuming that the most reactive control rod group is withdrawn at its maximum rate (1.5 inches per minute) in its maximum worth region, the reactivity addition rate is limited to less than 7.2 x 10-5 ask/sec. In Section VI of the Safeguards Report for the Saxton Reactor Par-tial Plutonium Core II, rod withdrawal transients are presented l for cold and hot suberitical, and full power operation initial j conditions. These analyses were based on a conservative high insertion rate of 2.5 x 10*4 Ak/sec and illustrate the effec-tiveness of the overpower trip in terminating a rod withdrawal transient. Startup from the hot or cold suberitical condition is of less consequence than rod withdrawal from power, so rod withdrawal from power is re-analyzed. The principal change affecting the transients starting from subcritical is the increase in the overpower trip setpoint from 115% of 23.5 MWt to 107% of 35 MWt. The additional energy generated and, hence, the increase in fuel temperature in reaching the higher trip level is not sig-nificant. It should be noted that the analyses presented for rod withdrawal from subcritical in Section VI of the Safeguards Report for the Saxton Reactor Partial Plutonium Core 11 are highly conservative in that no credit is taken for either the j startup rate trip (set at 2 decades / minute) or for the reduc-tion in overpower trip setpoint (to 5 MWt) during zero power I operation. VI-1 (E-2) ~

._-.~_ Thc transient for rod withdrawal from power is shown in Figure VI-1 based on the following conservative conditions a) Initial power level is 193% of the nominal (3! FNt) power to allow for calorimetric errors. b) Initial primary coolant pressut e is at its minimum value of 2150 psia to aalow for instrument errors. Initial coolant inlet temperaturia is at it.s maximum value of 485'F allowing for instrument error. d) Minimum expected absolute value of negagive fuel tempera-ture (Doppler) coefficient: -1.0 x 10' Lk/'F. e) Minimumexpectedabsolutevalueo{21k/*F. negative moderator tem-perature coefficient: -2.0 x 10~ f) Reactor trip initiation due to overpower at 114% of the 35 MWt power (7% over the 107% reactor trip setpoint to allow for instrumentation errors). With the nuclear flux and hot spot heat flux peak at 120% and 114% of their nominal full power values respectively, the mini-num DNB ratio (calculated using thi. W-3 correlation) is 1.31 and occurs 3.7 seconds after initiation of the incident. This indicates that the power level reacter trip protection would prevent core damage in the im;.robable event of an uncontrolled rod withdrawal. l 2. S cam Line Break Incident l Rupture of a secondary plant steam line is reflected into the primary system as a Jtep load increase and results in a decreas-ing coolant inlet temperature. The negative roderator temper-ature coefficient causes reactivity and power to increase. For 'arge breaks, the control capability of the plant is exceeded und the reactor protection system will automatically initiate a reactor trip by either the overpower or low pressure condition. Following trip, heat extraction exceeds heat generation and the coolant temperature decreases further. Due to the negative mod-erator temperature coef ficient, the shutdown margin is reduced until heat removal is terminated.. The consequences oa a steam line break for the proposed 35 MWt Operating conditions are not changed from those reported ( V1-2 (E-3) e a .-g e ,--~~gn,-. - -n

previously in Reference (1), (2) and (3). The lower secondary temperature, 420*F for 35 MWt operation as compared to 490' for 23.5 MWt operation) will result in a smaller steam mass flow rate through the break and will reduce the rate of cool-down in the primary system Since the average coolant temper-ature is reduced to 500'F lor the proposed operating condi-tions, the moderator temperature coefficient is less negative than that assumed in the reference analyses. Both of these effects will result in a less severe reactivity transient. The effect of the increased power level is to increase decay heat following reactor trip which reduces slightly the primary sys-tem cooldown. The operating conditions for 35 MWt result in less reduction in shatdown margin during blowdown of the steam generator contents. C. MECHANICAL INCIDENTS 1. Loss of Flow Incident A loss of coolant flow could result from loss of electri c-power to the reactor coolant pump motor or froa mechanical fail-ure in the pump motor or coupling between the pump motor and pump. A mechanical failure causing sudden seizure of the pump motoc is not considered credible. Following losa of coolant flev, coolant temperature will increase because of reactor trip e } circuit delays which allow continued power generation while flow coastdown is occurring. If the heat generation is net termina-tad rapidly enough to prevent DNB, clad failure can result. Power generation during a loss of coolant flow iteident is tr-f minated by an automatic reactor trip initiated from eithe

c. low vcitage signal on the reactor coolant pump bus, a lcw ?*

cy signal on the MG set control or from a low flow signo In u.e event ot a loss of the streion 480 V auxiliary

=r.rical system, reactor coolant pump coastdown is hxtended (see Section IV) by automatic switching which transfers the field supply of the generator to the station battery. This extended coastdown (1) Saxton-Final Safeguards Report (2)

Saf eguards Report for Phase I of Saxton Nuclear Experimental Corpor-ation Five-Year Research and Development Program (December 1961). (3) Saxton Nucletr Experiment Corporation - Safeguards Report for the Saxton Reactor Partial Plutonium Cork II (March 1965). ( VI-3 (E-4)

f ~ c. FUEL TEMPERATURE COEFFICIENT = 1.0x10 " ak/*F MODERATOR TEMPERATURE COEFFICIENT = -2.0x10' ok/'F l TRIP LEVEL 114% Reacter Trip 120 Initsated l )l l <1 1 100 NUCLEAR TLUX i I 80 m HOT SPOT HEAT FLUX w a 60 PRIMARY COOLANT - PRESSURE s .2170 t u W m 40 .2160 m;$ \\ mm U* 2150 m .g i 20 0 1 -i i 0-5 10- ' 15 _ TIME, SECONDS. -CONTINUOUS-ROD WITHDRAWAL.- (REACTIVITY INSERTION RATE =(2.5 x'10~' Ak/ SECONDS)- e 1 FIGURE VI-1 ,VI-8. v (E-5) 1 -Q ._,-.,__._._..__..._.._..__._...-._..-_._..._._.a}}