ML19340A533

Technical Info for Amend to License DPR-2 as Amended,First Reload Fuel for Dresden Nuclear Power Station

ML19340A533
Person / Time
Site:	Dresden
Issue date:	01/27/1961
From:	Elliott V, Clint Jones, Sege G GENERAL ELECTRIC CO.
To:
References
GEAP-3586, NUDOCS 8008250810
Download: ML19340A533 (59)

Text

{{#Wiki_filter:2 / Commonwealth Edison Company General Electric Company - - " ~~; m ' ms?= 3 :=gt z p p =y~ J ..........J.s O' .w1 s 2La. UUe i TECHNICAL INFORMATION FOR AMENDMONT TO LICENSE DPR-2 AS AMENDED FIRST RELOAD FUEL FOR THE DRESDEN NUCLEAR POWER STATION By: Commonwealth Edison Company General Electric Compa y j<_<,'/<, ~E$ e.i yo January 27, 1961 / /g A<- / lh 'w 008250h/O .3

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/Af& TECHNICAL INFORMATION FOR AMENDMENT TO LICENSE FOR DRESDEN (FIRST RELOAD FUEL) ~ A? .s Submitted by: / ~ Safegu[ard Evaluations $1 C. R. ones (3 gp M Approved by: G. Sege - MaWager ~ '., Safeguard Evaluations Design Engineering br% k va V. A. B)1ibtt - Manager Design Engineering January 10, 1961 ATOMIC POWER EQUIPMENT DEPARTMENT GENERAL ELECTRIC COMPANY San Jose, California

i ] TABLE OF CONTENTS P,_ age INTRODUCTION 1 PURPOSE AND SCOPE 1 RELATION TO PREVIOUS SUBMITTALS TO THE AEC 1 SECTION I REVISIONS TO APPENDIX "A" OF LICENSE NO. DPR-2 2 SECTION II PHYSICAL DESCRIPTION OF DRESDEN CORE II 9 1. Mechanical Design 9 2. Thermal and Hydraulic Chr racteristics 19 3. Nuclear Characteristics of Type II Fuel 21 4. Nuclear Characteristics of 12 Advanced Fuel Assemblies 41 5. System Stability SECTION III SAFETY ANALYSIS 48 1. System Performance during Control Rod Run-out 48 2. Sudden Step Insertion of Reactivity 49 3. Mechanical Failure of Control Rod Drive 49 4. Accidental Valve Closure on Steam Line 50 5. Fuel Cladding Failure 50 6. Loss of Recirculating Pumps 50 7. Refueling Error 51 8. Fuel Assembly Placed in Wrong Position 52 9 Group Removal of Control Rods at Startup 52 10. Introduction of Cold Water 53 11. Maximum Credible Accident 53 12. Justification for Revision to Temperature Coefficient 54 REFERENCES 55

INTRODUCTION PURPOSE AND SCOPE This report is an appiication for an amendment to License DPR-2, as amended, for the operation of the Dresden Nuclear Power Station. The purpo.se of the report is to request revisions to Appendix A of License No. DPR-2, as amended, which will provide for the installation and operation of 112 new fuel assemblies that will be used to replace fuel assemblies now in the Dres-den Core. This fuel, together with the remaining original fuel assemblies, will comprise the Dresden Core II. The report consists of three sections, as follows: Section I, entitled " Revisions to Appendix A of License DPR-2," contains the revisions required to permit the installation and operation of Dresden Core ~1 and defines the conditions under which new fuel of different design may eeplace fuel of current design and continue to operate safely. Section II provides a description of the mechanical design and physical properties of the new fuel and core, including a discussion of the system dynamic s. Section III is devoted to a discussion of the safety of Dresden Core II and a re-evaluation of the safeguard analysis of the Power Plant. Sections II and III are not intended to constitute " Technical Specifications" in the sense of the lic ensing regulations (10CFR, Part 50, Section 50. 36). Although the primary concern of this report is the central core loading of 100 new stainless steel clad fnel assemblies, it is anticipated that it may be desirable to load the new fuel in s.he r patterns. In this event, the definition of the proposed cores and a discussion of the safety implications will be submitted in a separate report and further Commission authorization requested. RELATION TO PREVIOUS SUBMITTALS TO THE AEC This document does not supersede any of the previously submitted documents concerning the Dresden Nuclear Power Station. The intent of the document is to provide the auditional inieraation necessary for the evaluation of a new and improved fuel design which will be used ; s the first reload fuel for the core cur-rently operating. The amendment and analysis presented herein assumes the prior or concurrent approval of the proposed changes to Appendix A of License DPR-2 filed with the Commission on January 5,1961.

! SECTION I REVISIONS TO APPENDIX "A" OF LICENSE NO. DPR-2 A. Applicant requests revisions of Section "B Design Features" of Appendix "A" of DPR-2, as follows: 1. In item "2 Nuclear Core," delete the entire table starting with " Max-imum core diamete r..... " and ending with "...... fuel as sembly 488. " and substitute therefor the following : "a.. Dresden Core I Maximum core diameter (circum-scribed circle) 129 in Maximum active cold fuel length 107-1/ 8 in s Maximum number of fuel assemblies Type 1 488 b. Dresden Core II Maximum core diameter (circum-scribed circle) 129 in Maximum active fuel length - cold 111-15/16 in Maximum number of fuel assemblies by types TypeI 488 Type II 100 Type PF-1 through PF-12 (one each) 12 For fuel type number definitions, see Table I. Maximum number of all fuel types 488 Fuel assemblies of TypeII will be located in the central region of the core. Fuel assemblies of Type PF-1 through PF-12 may be located in any position in the reactor, provided each such assembly is separated by at least four Type I fuel assemblies from any other such assembly or any Type II fuel assembly. I Fuel assembiies of Type I will occupy the remaining positions to complete the core configuration. The reactor may be operated at any power up to and including rated power with any number of the various types of fuel assemblies in-stalled provided the maximum number and location are within the limits specified above. "

. 2. In item "3. Fuel," delete the entire paragraph starting with "Each fuel assembly..... " and ending with "..... O. 710 in. " and substi-tute therefor the following : "a. Dresden Core I Fuel Each fuel assembly consists of 36 vertical fuel rods (except for occasional special assemblies such as instrument bearing as-semblies which will have 34 or 35 fuel rods plus the thimbles of approximately the same dimension of the clad fuel rods), each of which is made up of solid cylindrical pellets of uranium di-oxide, enriched in ure.nium-235 to a maximum of 1. 5%, and clad in zircalcy-2. Each fuel rod is composed of four sepa-rately clad segments. Pertinent fuel design parameters in the cold condition are: Fuel pellet diameter

0. 498 in Fuel pellet length Regular fuel pellets

0. 625 in Segment end pellets

0. 500 in Outside diameter of cladding

0. 567 in Cladding wall thickness 0,030 in Fuel pellet density averaged over a fuel segment, minimum 94% of theoretical Cros s-sectional center-to-center distance between fuel rods

0. 710 in b.

Dresden Core II Fuel Each fuel assembly consists of vertically-positioned, rod-type fuel elements. The physical properties of each assembly are given in Table I. The number of fuel rods are given for a reg-ular assembly. Several assemblies, however, have been de-signed such that rods may be replaced with instrumentation. The minimum fuel peller density averaged over a fuel segment is 94% of theoretical for all fuel assemblies except P.F-7, PF-8, and PF-9 which is 90% of theoretical."

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. 3. In item "9 Safety System, " delete the paragraph starting with " Rated local power..... " and ending with "..... the point monitored. " and substitute therefor the following: "The maximum local heat flur. is given in, item 3 of Section E, entitled Power Operation. The in-core local power monitors will be used to corroborate the calculated shape and absolute value of the core power distribution. This power distribution will be used, together with appropriate analytical techniques, to determine the limiting thermal characteristics of each type of fuel assembly. " B, Applicant requests revisions of Section "E Power Operation" of Appendix A of DPR-2, as follows: 1. In item "1 Power Test Programs, " delete the entire section starting with "Following succes sful completion.... " and ending with ".... radiation survey. " and substitute therefor the following: "a. Dresden Core I Follow ing successful completion of the initial loading and critical testing prog tam, the power test program will be carried out. This program will include the following:

i. A series of ter,ts at low power, and low-to-rated pressure, with studies of stability, reactivity, and power distribution as they are affected by variations in reactor pressure, temperature, voids, control rod positions, and transient poisons; ii. A series of tests at various power levels up to rated power, at low-to rated pressure, with studies similar to those indicated in item E.1. a. i; iii. Studies of power transient effects, simulated equip-ment failures, and process mishaps; and iv.

Radiation surveys. b. Dresden Core II The power testing for the second core will be limited to careful observation of reactivity, power distribution, and stability as indicated by the reactor instrumentation during the normal gradual approach to power. "

. 2. In item "3 Determination of Maximum Reactor Power, " delete the entire section starting with "The ' rated' or operational.... " and ending with ".... phase of operation. " and substitute therefor the following: " a. Dresden Core I The " rated" or operational thermal power of the reactor is not, at least during initial power opention, a fixed vclue. Lower than equilibrium fuel cycle power pro-duction may be necessary initially, due to the possibilities of: i. Less than full load of 488 fuel assemblies in order to meet shutdown margin criteria; and ii. Higher peaking factors initially than at equilibrium fuel cycle. The maximum reactor power is, consequently, defined as that thermal power at which the maximum heat flux for any fuel rod is reached. This maximum heat flux will never 2 exceed 350,000 Btu /ft -hr, based on calculations and ex-perimental data. The peak rated heat flux and resulting rated reactor power are then set to 80% of their maximum values and, as indicated in item B.9, the high neutron flux scram setting will be no higher than an indicated 120% of the rated reactor power. However, in no case will the high neutron flux setting be allowed to exceed an indicated reactor thermal power of 782 hiw (125% of the planned operational power of the full loaded core). Within the limitations on reactor power and heat flux set forth in the paragraph above, the reactor will be operated in such fashion as to: 1. Maintain at all times a burnout heat flux margin of at least 1.5 at the point closest to burnout in the hottest channel in the core based on a uniform steam quality over the cross section of the channel; and ii. Be always well within the bounds of stability, as evi-denced by the operation itself and any experimental data produced during the " Power Test Program" phase of operation. I

. b. Dresden Core II The maximum reactor power for Dresden Core II is de-fined therefore as that thermal power at which the max-imum heat flux for any fuel rod is reached. This maximum heat flux, based on calculations and experimental data, will never exceed the following values in Btu /(hr)(sq ft): Fuel Type I 350,000 Fuel Type II 445,000 Fuel Type PFI through PF4 425,Ov0 Fuel Type PF5 through 9 415,000 Fuel Type PF10, 11 and 12 475,000 The peak rated heat flux and resulting rated reactor power are then set to 80% of their maximum values and, as in-dicated in item B. 9, the high neutron flux scram setting will be no higher thap an indicated 120% of the rated re-actor power. However, in no case will the high neutron flux setting be allowed to exceed an indicated reactor thermal power of 782 Mw (125% of the planned operational power of the fully loaded core). The reactor will be operated within the above limits such that a burnout margin of at least 1.5 will be maintained in each type of fuel closest to burnout in the hottest channel in the core based on a uniform steam quality over the cross section of the channel." In item "5 Reactivity Limits, " delete the entire paragraph starting 3. with "The moderator temperature.... " and ending with ".... to confirm this. " and substitute therefor the following:

. "The limitations relevant to the effect of moderator temper-ature on reactivity are as follows: (i) The positive coefficient will be less than Ix10~4 Ak/k*F at 70 *F; (ii) The maximum total reactivity addition will be less than one dollar, starting at 70 *F; and (iii) The maximum moderator temperature at which the co-efficient ceases to be positive will be less than 500 *F. These limitations apply to any critical assembly achieved by withdrawing rods in the section of the core (as determined by analysis) which has the greatest positive coefficient. A com-bination of experiment and analysis will be used to show that the above conditions are met."

9-SECTION II PHYSICAL DESCRIPTION OF DRESDEN CORE II !f 1. Mechanical Dealgn When it is desired to replace Type I Dresden fuel assemblies, the new fuel assemblies will be of a different design. The changes in design have been found by analysis and experiment to be advantageous. The fuel channel dimension of the new fuel remains unchanged, thus maintaining complete interchangeability of fuel assemblies. The basic fuel pellet for most of the fuel elements, new and old, is a sintered, solid cylindrical pellet of about 94% theoretical density. The diameter is slightly less than 1/2 inch, and the length is slightly less than 3/4 inch. The exact pellet diameter varies with the fuel design and is given in Table I, Section I, hereof. The pellets are enclosed by a stainless-steel or zircaloy jacket which forms a fuel element about nine feet long. Three of the advanced fuel elements, PF -7, PF-8, and PF-9, have been designed around a new fabrication technique in which a homog-enous matrix of UO2 Powder is worked externally within the cladding to produce a continuous fuel rod of about 90% theoretical density. Each fuel rod is to be placed in a square array of six rows of six fuel rods, seven rows of seven fuel rods, or eight rows of eight rods. The desired geometry is maintained by means of a square channel about 4-1/4 inches inside dimensions with spacers at several positions along the length which also serve to minimize deflection and vibration. The three fuel types are defined as follows: Type I The first Dresden core fuel consisting of cylindrical zircaloy-clad fuel rods in a "6x6" matrix. Details of the fuel assem-bly are shown in Figure 1 and Reference 1. Type II A fuel of a new and improved design con-stituting the primary reload of about 100 elements. Details are shown in Figure 2. Type PF-1 through PF-12 Fuel using advanced fabrication techniques. A total of 12 experimental fuel assemblies are planned. Details are shown in Figures 3 and 4. The location of the fuel and the control rod position in the core are defined by the numbering system shown in Figure 5.

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. s Three typical core configurations are shown in Figures 5, 6, and 7, the exact location depending principally upon the status of the first Dresden Core fuel. The fuel assemblies of type PF-1 through PF-12 will be located at several positions in the core, spaced so that the performance of any such assembly will not significantly affect nuclear or thermal performance of the entire core. A. Fuel Tyne II Each of 100 fuel assemblies constituting the primary Dresden reload will consist of seven rows of seven nonsegmented fuel rods except, as stated below, one rod will be segmented as required for positioning spacers. The rods, clad with 304 stainless steel, are held in position at the top and bottom by stainless-steel tie plates. Between the tie plates, rod po-sition is maintained by four wire-type spacers, spaced equally along the length of the fuel assembly. The four spacers are held in longitudian1 position by a seg-mented fuel rod which is located in a position where the re-sulting end plugs do not present any power peaking problem. Nine rods of each fuel assembly are loaded with thoria pellets. The water-to-fuel ratio is 2.67:1. Details of the design are given in Table I, Section I, and Figure 2. B. Experimental Fuel Assemblies PF-1 through PF-12 Experimental fuel assemblies PF-1 through PF-12 incorporate fuel designs resulting from recent engineering developments in rod-type fuel. Figures 3 and 4 show the essential design fea-tures of the new fuel. The same general design for fuel Type I applies to the 12 ex-perimental assemblies with the following variations: A removable fuel rod design is used. a. b. All fuel rods are suspended from the top tie plate. c. " Corner" fuel rode in the fuel assemblies will contain either a ThO2 - UO2 mixture, ThO2 - UO2 - Er2O3 mixture, or variable enriched UO2 mixture to reduce power peaking in the corners of the fuel assembly ad-jacent to the control rod blade corner. d. Erbium-oxide (Er2O ) will be incorporated in fuel as-3 semblies PF-1, PF-2, and PF-10 as a burnable poison which will allow additional burnup without increasing the reactor control requirements. TYPE PF.1 THROUGH PF-12 26 OR TYPE I FUEL 000 000 0 2s no0o 0 Ci o i.__ 24 ODOCnCDOOOD 0000 21 n _u_ n._.o o n G o n r._. n i _.i i : n n n_ o U n_ n n_ n n r_ o n OOCOno 2. 20 ! i t i I i. I-i! I l. l i 1 ' l l n r n_" h [ b- - -- = hh h i b O h ~ ~ ~ [ ~ ~ ~~ ] bi bn 19 ~ 8 ~7-ocorn_g g a.g.a n_ o n_ n n n__ n.n_ i8 O _ h h_ =C h h,N.__ _ I [ I [ C O l_~ .__ _i. [ ] [ _I '_ 17 -- 'O.sb '.', h l-l ) j l ll ! 16 il l ' a f,;,$D >"3..E+ ] F- ,,_,.} I.,,,_,n,,,,,_I iJ-M c :< .c - _.,,i ; .=

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il 1 F l-l-l1 l ll r le i 4 r 7,. j:2 ] ] __. ] i 1 l__._ _ m' 1 4 i _{ m'y,}'t f s; lU]I -._ {> 10 '. l m Ijl 09 lm ! _,m E C j 'j 'CQ= :, jj(m q m _. l I! l _4 d I i C l U{ i i! ! i 3 g.:: _ .__ _I__.-, ] -, l_ [ _ lI .i ji 08 lj 3 fd,y i 6 h f r []' r-- [ j 6 ' ' i i l Llj U Lq L) p1 07 06 l] l[] l l ! i iIL_Li. Yhl] _- _ _ _ il _ h 05 i m i_ ] U 04 l' i ll l t ][ ]! l l ' { _*l 03 l! ll - - -l1 _ _ -{ R 02 i i 1 l, - L_ Ot L TYPE 11 FUEL ONLY A B C D E F G J K 51 52 53 54 55 56 57 58 59 60 61 62 63 64 65 06 67 68 69 70 71 72 73 7 FIGURE 7 New fuel rod manufacturing processes, such as swaged e. powder (fused), swaged-over ground pellets, and swaged-over unground pellets, will be incorporated in nine assem-blie s, PF-1 through PF-9, using stainless-stael cladding. f. As indic.ted below, some ext . mental assemblies are de-signed with relatively thin stainless-steel cladding which is not self-supporting when exposed to reactor pressure but, in-stead, depends on internal support from UO2 Pellets and internal springs. The fuel assemblies are designed for an average exposure of 15,000 MWD /T. Based on experimental data, the design fission gas release rates for 15,000 MWD /T exposure are as follows: a. Swaged powder fuel rods - 100% gas release, with the assumption that 10% of the fuel volume is available for fission gas. b. Regular pellet, and swaged-over ground or unground 46% fission gas release with pellet-type fuel rods no free volume within the pellet for fission gas. Each fuel assembly incorporates some of the design features listed. The special features of each fuel assembly are given below. All of these improvements in design have been tested in the VBWR, GETR, or General Electric's Heat Transfer Test Facili.y at San Jose, Califo rnia. PF-1 itainless steel clad swaged-over unground pellets The design utilizes thin-clad stainless steel which reduces the parasitic absorption of neutrons in stainless steel. The swaged-over pellet proc ess consists of filling an oversized tube with either ground or unground pellets and then passing this tube through a swaginJ machine which reduces the outside diameter to a specified dimension. The assembly uses the burnable poison Erbium Oxide (Er2O) 3 in all fuel rods, as well as Thoria in the corner rods. The water-to-fuel ratio is 2. 9:1. PF-2 stainless steel clad swaged-over ground pellets The design is the same as PF-1 except lower-enriched UO2 corner rods are used instead of Thoria corner rods, and ground rather than aground pellets are used in the fuel rods. The wate r-to-fuel ratio is 2.9:1.

- 19 PF-3 stainless steel clad swaged-over ground pellets The design is similar to PF-1 except burnable poisons are not used, and ground rather than unground pellets are used in the fuel rods. The water-to-fuel ratio is 2.9:1. PF-4 stainless steel clad swaged-over unground pellets The design is the same as PF-3 except unground rather than ground pellets are used in the fuel rods. The water-to-fuel ratio is 2.9:1. PF-5 8e PF-6 stainless steel clad swaged-over unground pellets The design is similar to PF-4 except thicker stainless-steel clad fuel rods are used. The different clad thickness will provide a means for comparing the operating charac teristics of the thin-clad versus the thick-clad fuel rods. The water-to-fuel ratio is 2.91. PF - 7, PF-8, and PF-9 stainless steel clad swaged fused powder The design utllizes a low cost fabrication process. The swaged powder process consists of filling a tube with " fused" UO2 Powder, g' plugging the ends of the tube, swaging the tube down to the re-quired diameter, removing the end plugs, welding permanent end plugs in place, and then performing the necessary quality control to insure that a sound fuel element has been manufactured. A i UO, density of 907 is obtained in this process. The water-to-fuel i ratio is 2.9:1. 1 ] PF-10 zi rcaloy - 4 clad with lower enriched UO corner rods 2 { The "8x8" assembly is a non-segmented continuous rod design, ] using conventionally loaded UO2 Pellets, as shown in Figure 4, i with Erbium Oxide and low enrichment UO c rner r ds. The Z J water-tc-fuel ratio is 2. 4:1. j PF-11 and PF-12 zircaloy - 4 clad with thoria corner rods i The design is similar to PF-10 except burnable poisons are not used, and thoria corner rods are used to replace the lower-en-riched UO c rner r ds. The water-to-fuel ratio is 2.4:1. 2 2. Thermal and Hydraulic Characteristics l The objective of the Thermal and Hydraulic design of the new fuel is to extend the heat generation capability beyond that already established for Dresden without compromising the integrity of the cladding. Two potential causes of cladding failures are the. effect of high temp-erature on the mechanical properties of the cladding material and increased fission gas pressure resulting from molten fuel.

. Considering the mechanical properties of the cladding material and its ability to withstand internal pressure from fission gas released during the fuel lifetime, the following cr teria have been established: Fuel Type I - 2 Maximum heat flux 350,000 Btu /ft hr Maximum fuel temperature at design overpower 4900*F Minimum burnout heat flux margin

1. 5 The fuel design in the first Dresden Core has demonstrated its ability to meet the thermal and hydraulic performance require-ments listed above.

Fuel Type II - Maximum heat flux 445,000 Btu /ft2 hr Maximum fuel temperature at d design overpower 4900*F g Minimum burnout heat flux margin

1. 5 Fuel Type PF-1 through PF 2 Maximum heat flux 425,000 Btu /ft hr Maximum fuel temperature at design overpower 4900*F Minimum burnout heat flux margin

1. 5 Fuel Type PF-5 through'PF 2 Maximum heat flux 415,000 Btu /ft hr Maximum fuel temperature at design overpower 4900*F Minimum burnout heat flux margin

1. 5 Fuel Type PF-10, 11, and 12 -

2 Maximum heat flux 475,000 Btu /ft hr Maximum fuel temperature at design overpower 4900*F Minimum burnout heat flux margin

1. 5 The maximum heat flux is based an the central fuel temperature, therefore the actual maximum heat f1' x in the reactor may be limited u

by the burnout -atio before the heat flux listed above is reached. The maximum heat flux actually existing in the reactor will be limited by either the central temperature or the burnout ratio, whichever is smaller.

, Relative to the first core, a higher allowable power density may be achieved in the central region of Dresden Core II using the new stain-less-steel fuel assembly design. This is accomplished by increasing the number of fuel rods in a fuel assembly from 36 to 49 which de-creases the diameter of the fuel pellet and increases the surface area, thus reducing the rise in temperature across the fuel pellet and allow-ing more heat to be removed per element at the same surface heat flux. The new fuel design also eliminates the segmented fuel rod which in-creases the heated surface area and eliminates the " hot spot" factor resulting from flux peaking in the space between fuel pellets. In ad-dition, the new fuel has a single fission gas zone in a relatively cool region which results in a reduction in fission gas pressure. A further increase in the allowable power density is achieved by re-ducing the burnout ratio, which permits increasing the peak surface heat flux. This reduction is justified by: (a) A large number of heat transfer tests conducted by APED, making possible the accurate pre. diction of the actual burn-out limits, and (b) The excellent correlation of predicted power distribution-and measured power distribution leading to less uncertainty in the ability to predict the location and magnitude of the hottest channel. 3. Nuclear Characteristics of Type II Fuel A. Introduction From he standpoint of its nuclear characteristics, the Dresden Core II has a gross similarity to the first Dresden core. It re-mains a light-water-moderated, heterogeneous, low-enriched, oxide-fueled reactor. The first and second cores are similar in that they both contain a large number of critical massee and, hence, are able to sustain various power and flux shapes. The local nuclear characteristics of the reactor are closely coupled to the fertile and fissionable materials, fuel rod diam-eter, fuel rod pitch, cladding material, moderating character-istics and control capabilities. The variation of the nuclear characteristica with time depends upon the particular choice of the above parameters. Nuclear characteristics of the new fuel have been developed with maximum burnup and compatibility with the present fuel as fun-damental design criteria. The following is, therefore, a discussion of the various nuclear properties as they are related to the individual fuel assembly and to the core as a whole.

= . 1 a. General Description The active fuel height is 110 inches, and the equivalent cylin-1- drical core diameter is 129 inches. ' The central zone of Core II will be about 60 inches in equivalent diameter and contains fuel of Type II. b The core average void content at full power is about 18% of the active coolant in the core center and about 20% in the outer regions depending on control rod patterns. The spatial void 'ations in the first core (Core I) result in greater differ-ences in neutron slowing down properties than differences inherent between fuel Types I and II. The control system is identical with that in Core I and consists of 80 cruciform blades of stainless steel containing 2 weight-per cent natural boron. The decrease in neutron thermal diffusion lengths in a lattice with stainless-steel cladding used with Type II fuel results in a lessening of control strength of the control system in relation to the effect of the system with Type I fuel in regions where Type II fuel is used. The control strength is not affected in those regions of the core with Type I fuel. Control of Type II fuel is discussed in detail below. Details of the mechamcal design and materials composition of all fuel types are presented in proposed amendments to Appendix A of DPR-2. The fuel rod diameter and number of fuel rods per assembly were chosen to achieve a water-to-fuel volume ratio of 2.67:1 in the Type II fuel. This water-to-fuel volume ratio gives the system inherent safety bij virtue of the negative oper-ating void and temperature coefficienti with minimal penalty in life of the fuel. The water-to-fuel ratio within the flow channel is 1. 65 : 1. This region is significantly undermoderated, and the I formation of steam voids inside the channel thus causes a de-crease in reactivity. Comparing the Type I fuel with a water-to-fuel retIs i of 2.1:1 to the Type II fuel, the additional enrich-I ment required in a steel lattice to overcome the increased neutron absorption in the clad, unless otherwise compensated fo r, would lead to a poorly moderated system. Thus, Type II fuel assemblies have been designed to provide a greater water-to-fuel volume ratio than a zircaloy-clad lattice in order to j achieve near 7.timum moderation. The choice of this optimum water-to-tuel ratio also leads to similar coefficient character-istics for both fuel types. i i l

-o.. i t The channels surrounding each fuel assembly may be zircaloy or stainless steel. The stainless-s+. eel channels to be used initially with some assemblies serve as a tem-porary poison to meat control requirements. They can be replaced with zircaloy channels after some fuel depletion - in order to increase reactivity later in reactor life. The use of the stainless-steel channels as temporary poison permits the use of sufficient enrichment to achieve an f average discharge exposure of approximately 12,000 MWDt in the Type II fuel. Nuclear characteristics associated with the individual UO 2 2 uel rods of Type II assemblies are given in Table f and ThO 2. The reactivity associated with the thoria rods is reduced, due to the decreased enrichment. The nuclear character-istics of Type II fuel bundles with zircaloy or stainles s-steel channels are shown in Table 3. The characteristics of the assembly are not appreciably dif-ferent from those of Type I fuel. The fast fission effect is, i however, decreased because 18% of the rods are thoria which has very little fast fission. The neutron lifetime remains characteristic of.a thermal reactor, and has a value of ap-proximately 5x10-5 seconds. The effective delayed neutron fraction is the same for Type I and Type II' fuels in the cl.an condition (0.0071). With burnup, the effective delayed neutron fraction will become smaller as U-233 and plutonium are pro-duced. Howeve r, these two isotopes have about the same effective delay fraction so that in the equilibrium core Type I and Type II fuels have essentially the same effective delay fraction. In the hot operating condition the most probable neutron energy is about 0.06 ev. The conversion ratio in j Type II fuel is less than in Type I fuel. This is predom-inantly a result of the increased absorption in the stainless-i steel clad.

A 4 G T ABLE 2 Nuclear Characteristics - Type II Fuel Rod Hot Hot Cold No Void 20% Void ,68*F 546*F 546*F UO Fuel Rod 2 c 1.03 1.03 1.04 p 28 . 81 .76 .72 1 p .73 .66 . 61 f .82 .83 .84 i 3 1.79 1.78 1.77 2 I~ cm

2. 0

3. 9

4. 9 cm 35 51 67 K.

1.105 1.006 0.945 ThO Fuel Rod 2 e 1.00 1.00 1.00 02 p .87 . 81 .78 p 9 .66 total f .82 .84 .85 i 3 1.36 1.34 1.34 2 2 L cm

1. 9

3. 7

4. 6 c m

35 51 67 K. O.940 0.802 0.752 1 i i

1 . TABLE 3 Nuclear Characteristice - Type II Fuel Assembly Hot Hot Cold No Void 20% Void _68'F 5 46F 546*F Zircaloy Channel Assembly C 1.02 1.03 1.03 P total .82 .77 .75

1. 9fres

.09 . 12 . 13 6 'I Ithermal

1. 27 1.34 1.36 2

2 L cm

2. 5

5. 3

6. I 2

7 cm 34 52 63 K. 1.138 1.152 1.146 Stainless-Steel Channel Assembly C 1.02 1.03 1.03 P total .82 .77 .74 6 71 fres .09 .12 . 13 E9 fthe rmal

1. 13 1.20 1.21 2

2 L cm

2. 1

4. 6

5. 2 2

7 cm 33 50 60 K, - 1.018 1.038 1.028

B. Reactivity and Control Requirements Sufficient control capability must be available in the reactor to meet operational reactivity requirements. The reactivity which must be controlled is composed of the following contri-butions: shutdown margin, reactivity in voids and maneuvering, cold-to-hot defect including Doppler, xenon, samarium, and burnup. The control strength in the Type I andType II fuel is not affected significantly by the presence of other fuel because of the short neutron diffusion length, except where the control rod pattern distorts the flux and changes the relative irrpor-tance of the control rods. This effect is present in any con-trolled reactor and is decreased when the strength of the distortion or control decreases as is the case in the Type II fuel. Table 4 shows the control strength for Type II fuel assemblie s with zircaloy or steel channels. In the cold clean condition, the control rod strength with zircaloy channels is 10% Ak/k and is du:reased to 8% Ak/k with steel channels. The effect of steel channels on reactivity worth cf a fuelbundle ranges from

10. 5% Ak/k in the uncontrolled fuel to 8% Ak/k in tha controlled situation. Hence, the exact number of the steel channels u ied as a temporary poison, on some of the fuel bundles, will de-termine the effective average control strength. After the fuel has been irradiated for about 2000 MWD /t, some of the temp-orary steel channels can be removed and replaced with the zircaloy channels to provide additional reactivity to sustain burnup.

Table 5 shows the reactivity requirements that must be bal-anced by the permanent and temporary control system for the Dresden Type II fuel in the most reactive condition. The condi-tions are those which would prevail for a reactor fully loaded with Type II fuel with half stainic.ss-steel and half zircaloy chan-nels on the toel assemblies. Fcr a core partially loaded with Type I fuel, the shutdown margin will increase because of the stronger control strength. In the hot or hot operating condition, the control strength increases as shown in Table 4. In this situation, the shutdown margin is very large for both the Type.I and Type II fuels and the power shaping ability of the control sys-tem remains about the same as in the original core. The control system and the auxiliary steel channels have been designed so that the reactivity in the reload fuel will remain about 0. 01 Ak subcritical even if the strongest control rod is

~ l, l ccmpletely inoperable and removed from the core. The reac-tivity worth of the strongest rod depends on the size and flux l distribution in the surrounding unrodded region, as well as the locations of the other control rods. t I-The strongest control rod worth exists when a single control rod is withdrawn from smallest possible fuel configuration which is i critical. The most significant situation in the use of Type II fuel exists when a. rod is withdrawn from a single control rod cell with the reactor just critical' and when all of the Type II fuel has zircaloy channels. The maximum rod worth under these hypothetical con-i ditions is 0.02 Ak for the cold clean core. Configurations using steel channels result in a lower rod worth since a larger region is required for criticality. It is expected that the reactor will have the steel channel assemblies loaded in a relatively homogeneous distribution throughout the Type II fuel zone. Thus, the fuel with zircaloy channels will not occupy large regions of the core when in the fresh condition. In addition, the reactor is normally operated so as to avoid withdrawing con-trol rods in localized regions of the core so that,. in general, the. -worth of any control rod is considerably less than the above value. 1 The maximum rod worth in Type I fuel is about 0.03 Ak for a mi: I imum critical region in the cold case. 7 ] The total reactivity worth of all control rods in the hot condition { for Type II fuel with zircaloy channels is 0.18 Ak, and 0.14 Ak, j for steel channels. Comparing Type I and Type II fuel, the shorter thermal diffusion length of Type II fuel results in a decrease in rod l worth and, therefore, a reduction in maximum local reactivity in-i sertion rate due to control rod removal. 1 i i 4 } l l

. TABLE 4 Control Type II Fuel Flow Hot Hot Channel No Voids 20% Voids Type Cold (68 'F) (546 *F) (546 'F) k. Ak/k k. Ak/k k. Ak/k Zircaloy Rod in 1.0216 9674 9450 .102 .160 .175 Rod out 1.1382 1.1516 1.1461 Stainless Steel Rod in 9397 9005 .8768 .077 .133 .148 Rod out 1.0185 1.0383 1.0286

- TABLE 5 Effective Reactivity Requirements Cold Type II Fuel Cold to Hot (including Doppler) t. 005 Ak Hot to Operating with Voids

0. 010 Ak Xenon and Samarium O. 035 Ak Burnup and Maneuvering

0. 086 Ak Shutdown Margin O. 030 Ak Total Control Required 0.160 Ak Temporary Poison (50% Stainless Steel Channels)

O.060 Ak Control Rods 0.100 Ak Total Control Available 0.160 Ak

C. Power Distribution a. Gross Power Distribution The nuclear characteristics of the Core II are space de-pendent and closely coupled to the nonuniform distribution of steam voids and, hence, moderating properties. The agreement between operational data and calculations for the first core (Core I) has demonstrated that it is possible to predict a control rod pattern which will result in a sat-iefactory gross power shape. Analysis of power distribution data from Core I has demon-strated good agreement between the measured power shapes, including local effects, and the ones calculated prior to the Dresden Reactor operation. Similar success in predicting the required loading of steel channels and control rod pat-i terns to obtain a satisfactory gross power shape for the second core (Core II) is expected. Figure 8 shows a one-dimensional radial power distribu-tion which might be obtained in Core II with a typical control rod pattern. Calculations and information obtained at shut-down will be used to develop data of this type which will be used to guide. the selection of permissible control patterns, fuel schedules, and orifice locations. l 'The operating power distribution will be monitored through j the use of in-core ion chambers supplemented by wire ir-radiation data and periodic gamma probing the core during shutdown periods for refueling. Based upon these data and the predicted power shape, future fuel distributions and con-trol rod patterns will be arranged to provide the optimum power distributions for maximum fuel exposure. b. Local Power Distribution The relatively high enrichments used in stainless-steel-clad lattices tend to increase local power peaking in the fuel rodslocated along the water gaps between the fuel bundle s. In order to reduce the local power peaking across a fuel bundle caused by the water gaps, an improved lattice design is utilized. Nine thoria rods are placed within the bundle to reduce the power peaking. The peak-to-average

, 2.0 l 1 1 I I I I I l 1 i l l l 1 1 I I f l \\ y l I I.5 1 i I I I I I I I I I I I I l O I I i x g l l 1 I a l l E l I d I i l Q l O.5 g I I I I I I TYPE IE FUEL TYPE I FUEL I e 18 l ie l Ij l ie I i l l I 0 O 20 40 60 80 10 0 120 14 0 16C RADIUS IN cm l RADI AL POWER DISTRIBUTION FIGURE 8

. power in'the Type II fuel is 1. 22, compared to a value of

1. 45 for the Type I fuel. Figures 9 and 10 show the ar-rangement of fuel rods within the fuel bundle and the peak-to-average power distribution for each rod within the fuel bundle. The two figures show the position of control rod, zircaloy, or steel channels. In both of these distributions, the void formation is considered uniform throughout the region within the channel, and no voids exist in the water exterior to the channels. The calculation is based on an average void content of 20 percent by volume of the water contained within the channel.

D. Reactivity Coefficients The water-to-fuel ratio of 2.67:1 for Type II fuel was established such that the core is slightly undermoderated. Since both lattices were designed to operate close to the optimum, the coefficients in Type I and Type II fuels are similar for the oper-ating condition. Thus, for a given change in conditions, each fuel in the reactor will have similar response characteristics. However, the coefficients cannot be matched exactly for all conditions, because they are affected by fuel burnup, xenon distribution, and control rod patterns. Details of the Type II fuel coefficients are given in the following paragraphs. Similar data for Type I fuel are given in references 1 and 2. a. Temperature Coefficient The reactivity response of the reactor with temperature is dependent upon the position of the control rods. When the control rod is withdrawn, the space it occupied is filled with water and an increased flux peaking occurs in the water gap. The insertion of the control rod depresses the flux in the water gap region, displaces moderator, and shifts the local iultiplication to a more undermoderated condition. The presence of a control rod, without modera-tor diaplacement, would yield a more negative coefficient since the thermal diffusion length increases with tempera-ture resulting in a greater fraction of the neutrons being absorbed in the rod. The local temperature coefficient is thus expected to vary as the control rods are repositioned during the operation of the reactor. In a manner similar to that which causes the loss of neu-trons to control rods with increased temper ature, a nega-tive component to the temperature coefficient exists due to I

. Cinnnel Type Zirculoy Temperature Sh6 F Void Fraction 20% Pressure 1015 poi 1.03 1.16 1.06 1.03 1.04 1.09 1.2 Th09 95 1.03 93 90 91 97 1.09 Tho2 1.19 98 .88 .85 .85 91 1.ch S 1.18 97 .87 .84 .85 90 1.03 93 1.00 91 .87 .88 93 1.06 20, 1.00 1.10 1.00 97 93 1.03 1.16 Tho2 1.15 1.00 93 1.18 1.19 95 1.03 0 Th0 Th0 Th0 O 2 2 g 2 Water Control rod nosition (Water) 3 "ircaloy Channel All fuel rods are UO unlecs indicated to be Th0 2 2 FIGURE 9 POWER DISTRIBUTIOli III FIEL ASSDIBLY WITII A ZIRCAIDY CIIANNEL

. 1 Channe1 Type Statnicno Steel 'thmperntitre Sh6"P Void Fraction ?W Prassitre 1015 pst [oo 1.13 1.05 1.on 1.02 1.07 1.1 Tho2 92 1.03 95 92 93 97 1.07 Tho2 1.17 99 91 .88 89 93 1.02 3 1.17 99 91 .88 88 92 1.02 91 1.02 94 91 91 .95 1.05 I no, j 98 1.10 . 1.02 99 99 1.03 1.13 I 20, 1.12 98 91 1.17 1.17 92 1.00 0 Tho Tho Tho Tho 2 2 2 2 Water Control rod position plater) } Stainless Steel Channel All fuel rods are Uo unless indicated to be Tho 2 2 FIGURS lo POWER DISTRIBUTION IN FUEL ASSEMBLY WI'HI A STAIllLESS STSEL CIIANNEL i [ r r e- -c

• e

. increased leakage with temperature from local regions with k,, greater than unity. During steady-state operation, the leakage from various regions of the core aust be such that the effective multi-plication is everywhere equal to urity. Since the migra-tion area increases with temperature, the leakage will produce an additional negative effect (when the migration length increases) in regions of the. core where local re-activity is greater than unity; i. e., uncontrolled regions or fresh fuel regions, and a positive effect in regions where local reactivity is less than one; i. e., controlled regions, fuel with steel channels, or highly depleted fuel. Because the power density and, hence, the importance of a region will be highest in those regions of greatest re-activity, the temperature coefficient in these regions is more representative of the core coefficient. As the re-activity decreases due to fuel depletion, the local leakage effect will likewise decrease so that the change in migration length is of lesser importance. The tempera-ture coefficient will become less negative toward the end of the operating period. The amount of plutonium and U-233 produced in type II fuel a much less than the plu-tonium produced in type I fuel. This decreased conversion reduces the tendency for the temperature coefficient to become less negative with burnup in the type II fuel. As discussed above, the coefficient depends upon the number and location of control rods, channels, and local multiplication. The sign of the temperature coefficient at ambient conditions is not critical in a BWR in which (!) the coefficient is negative in the operating condition, (2) the void coefficient is negative for any critical array at any condition, (3) it is not possible to increase the moderator temperature rapidly from ambient conditions without changing the fuel temperature. A justification for the revised moderator temperature coefficient is given in Section III of this report. As shown in Table 6, the temperature coefficient is nega-tive at all temperatures above room temperature for minimum critical configurations with zircaloy channels on the fuel assemblies. With half the fuel assemblies in stainless-steel channels and arranged in a minimum critical array, a small positive moderator temperature coefficient exists at room temperature but becomes nega-tive well below the operating conditions. The total re-activity change from ambient to the peak value is cal-culated to be about 30 cents.

. 4 An effect not included in the table is the insertion of control rods which tends to make the coefficient sub-stantially more negative. TABLE 6 Temperature Coefficient Ak/k/ *F Type II Fuel 140 *F 330*F 496*F Zircaloy Channels - 2. 0 x10~ -1.1x10 -2.1x10 ~ 50% Zircaloy Channels +1.4x10-5 - 5. 3x 10 -5 - 1. 2x 10 - 4 50% Steel Channels b. Void Coefficient The void coefficients were calculated for various condi-tions of the core characteristic of uncontrolled regions because it is expected that voids would first be formed in such regions. The void coefficients for fuel assemblies with zircaloy channels and steel channels are given in Table 7. As in the case of the temperature coefficient, the configuration is assumed to be just critical with no control rods inserted. The void coefficient is more nega-tive than the values shown in the table if control rods are inserted as will be the case during normal operation. c. Doppler Coefficient An important reactivity coefficient from a safety stand-point and one that does not change with fuel irradiation is the Doppler coefficient. The Doppler coefficient is a prompt in that it occurs instantaneously with an increase in the fuel temperature and n'eutron removal without time delay for heat transfer to the moderator. The Doppler broading is sufficiently large to induce a negative reactivity of approximately 3% in a fast tran-sient. The Doppler coefficient is greatest at low temperatures and decreases

, TABLE 7 Void Coefficient Ak/k/% Void Type II Fuel Zircaloy Channels Tempe rature 68'F 546 *F 546*F 546"F Void Fraction 0 0

0. 3

0. 5 Ak/k/% Voids

-1.4x10-3 -1.6x10-3 -1.9 10-3 -2,3x10-3 Steel Channels Ternpe rature 68*F 546*F 546*F 5 46 *F Void Fraction 0 0

0. 3

0. 5 Ak/k/% Voids

-6.5x10-4 -8.4x10-4 -1.Ox!0 -1.2x10 -3 -3 50% Zircaloy Channels 50% Steel Channels Tempe rature 68'F 546*F 546*F 546*F Void Fraction 0 0

0. 3

0. 5 Ak/k/% Voids

-1.lx10-3 -1.2x10-3 -1.4x10-3 -1.7x10-3

. as the fuel temperature increases. Thus, the Doppler coef-ficient is greatest for the cold startup and decreases at the higher fuel temperatures attained during operation. Table 8 shows the reactivity coefficient due to the Doppler effect j for the cold, hot, and hot operating situations. The reac-tivity effect is dependent upon the resonance escape prob-ability and increases with increasing prc'aability of capture in the resonances. The Doppler effect is given for each fuel type, and their combined reactivity effects in the lattice are also listed. The thoria rods used to obtain an improved power distribution also have the advantage of increasing the Doppler coefficient for the lattice. d. Effects of Pressure on Reactivity The magnitude of a pressure surge in the Dresden reactor is minimized by the pressure regulator, and experience at the Dresden plant indicates that during normal operation the pressure variations are held to less than 1 psi, which results in a relatively small effect on reactivity. Dresden Core II will have virtually the same response to pressure changes as Core I. E. Irradiation Dependence of Core Characteristics As the fuel irradiation in the first Dresden core continues, minoi changes in the core characteristics take place. The local characteristics of Core I have changed slightly due to the depletion of U-235 and production of pluto-nium. The dependence of the reactor characteristics with time is expected to be only slightly altered with the loading of the Type II fuel as semblies. Because the steel-clad fuel Type II has increased parasitic ab-sorption and increased enrichment, the conversion ratio is re-duced compared to Type I fuel. Hence, the slope of the reactivity decrease with time is greater and the tendency for the tempera-ture ceefficient to become less negative is likewise reduced. A change in the power distribution may be assn 1ated with the change in control rod pattern. The change in power distribution causes a new void distribution which, in turn, affects the neutron spectrum and isot ape buildup. Considerable flexibility exists throughout the fuel cycle to change the control rod pattern. The control rod pattern can be rotated periodically; that is, those

TABLE 8 Doppler Coef ficient Type II Fuel Reactivity Dopple r Coefficient Fuel Moderator Coefficient Due to Doppler Temperature Temperature J_ g / *F } b /'F 'F 'F I dT kdT UO2' Fuel Rods 68 68 -0 94x10-4 -0 14x10-4 1300 546 -0. 51x10-4 -0 10x10-4 1300 546-20% voids -0. 51x10 -4 -0.11x10-4 ThO2 Fuel Rods 68 68 -3.2 x10-4 -0. 35 x10-4 1300 546 -0.96x10-4 -0 16x10-4 1300 546-20% voids -0 96x10-4 -0. '17x10- 4 Type II Fuel Assemblies 68 68 -0.18 x 10 - 4 1300 546 -0.11x10-4 1300 546-20% voids -0.' 12x10 -4 4

40 control rods which are out' of the core are inserted, and those j which are in are withdrawn so that a more uniform fuel burnup will take place. The gross power and void distributions can also be' altered by the fuel reloading and discharge schedule, as well as by changes in the orifice pattern. In addition to the gross power shifts with irradiation due to con-trol rod motion and fuel burnup, local effects take place within the fuel bundles. Those fuel rods adjacent to the water gaps are in a region of high neutron flux and consequently are irradiated to higher exposures than the average rod in the bundle. The thoria compensated corner rods flatten the power across a fuel bundle, and any power peaking in the fuel bundle will decrease with lifetime due to the more rapid burnout of the fissionable material at the peak, where the flux is higher. Fuel irradiation changes the isotopic composition in a fuel rod in such a manrer that the power distribution through the fuel rod is more strongly concentrated at the outside surface, improving the heat transfer characteristics of the fuel rod. This effect is caused by the large resonance absorption in U-238 and Th-232 on the surface of the fuel rods which produces a higher concentration of plu-tonium or U-233 at the surfr.ce of the fuel rods compared to the inner regions of the rod. As previously discussed, t'te reactivity coefficients become less negative with exposure primarily due to control rod withdrawal from the core. The reactivity coefficients, previously dis-cussed, were based on a configuration without control rods. The increase in control rod worth with an increase in void or temp-erature is an important effect which contributes to the reactivity coefficients. Hence, the coefficierts for the initial core with control rods inserted will be substantially more negative than - given in Section 5. 4. As the fuel is depleted and control rods and stainless-steel chan-nels are removed, the values shown in Tables 6 and 7 for the zircaloy channels will be more descriptive of the core coeffi-cients, except for slight variations due to burnup. The presence of plutonium and U-233 also sffects the temperature and void coefficients of the core. However, because of the enrichment level, the buildup of plutonium and U-233 in the fuel will have a negligible effect on the coefficients. The buildup of U-233 in the thoria rods located along the water gaps will have a tendency to overcome any positive effect due to the putonium buildup in the

. central portion of the bundle, because U-233 does not have a large, low-energy resonance. In the inW1 nperating condition, the effective delayed neutron fraction is about O. 007 including the contributions from U-238 and the Th-232 fis sions. '1he delayed neutron fraction in the reactor will change with irradiation and will have a value of about O. 006 in the irradiated fuel due to the smaller delayed neutron fractions characteristic of plutonium and U-233. F. Liquid Poison System The worth of the liquid poison is reduced slightly compared to,the first Dresden core. The worth in the Dresden Core II with 50% stainless-s' teel channels is estimated to be about O.17 Ak during operation. This value is well within the O.15 Ak minimum re-quired by the license. In addition, k,, cold for TypeII fuel with zircaloy channels is about 1.14, which indicates that even if the c:.,re was incorrectly loaded with all zircaloy channels, the shut-down margin would still be met. The micimum poison worth during refueling is about O. 047 Ak. This value is within the 0. 04 Ak minimum specified by the. licence .and well in excess of the maximum worth of a single fuel assembly (about O. 02 Ak) or the maximum worth of a single rod (about O. 02 Ak). The change in the liquid poison system affected by Dresden Core II does not, therefore', alter the design function of the system. 4. Nuclear Characteristics of 12 Advanced Fuel Assemblies A. General Description The nuclear characteristics of the 12 advanced fuel assemblies, PFI through PF12, are essentially the same as the Type I and TypeII fuel assemblies. Two variations in nuclear properties which have been made, how-ever, are the use of burnable poisons in a slightly enniched fuel to aid in the control of long-lived cores and the use of thorium fuel slightly enriched with U-235. The thorium fuel rods, placed near the intersection of water gaps, reduce the normal water gap power peaking. The advanced assemblies will be placed in a position in the reactor where it can be shown by analysis that the power

-- - = 1

• i generated in the element is well below its thermal limits, separated such that elements will have a minimum effect locally i

and on the core as a whole. Each' advanced assembly will be separated by at least four 'Iype I fuel assemblies, insuring com-plete domination of overall core characteristics by the Type I and 'Iype II fuel. The tempe.~ature and void coefficient of the experimental fuel will, therefore, have virtually no effect on the transient behavior of the entire core. This method of load- { ing is very conservative because the nuclear properties of the j advanced assemblies are sufficiently similar to fuel Types 1 and II so that loading in any configuration would not be expected to result in undesirable consequences. From a nuclear standpoint, tl:e twelve elements may be divided into seven different types resulting from variations in cladding material, cladding thickness, and burnable poison or thorium c ont ent. The detail physical characteristics of each element are ~ given in Table I, Section I, and the seven groups of different nuclear properties are given in Table 9 l B. Reactivity and Control Effects Comparing the values of k. given in Tables 3 and 9, the maxi-mum variation is calculated to be about a093 Ak greater in fuel types PF-7, 8, and 9, than k. of fuel Type I. To estimate the change in reactivity when the experimental fuel assemblies are placed in the core, a series of cell calculations were performed assuming an experimental assembly is placed at the center of every group of 32 uncontrolled, fresh type I fuel assemblies. This analysis indicated that the maximum re-activity change should be about +0. 3% Ak. An analysis was also performed assuming that the assemblies are inserted in a controlled region The maximum reactivity i i change for any of the advanced assemblies resulting from this analysis, without control rods mserted in such region, was es-timated to be about +0.6% Ah Because of the small number and dispersion of the advanced as-semblies, these small changes in reactivity will not appreciably affect the steady-state or transient performance of the reactor. C. Power Distribution Effects The 12 dispersed advanced fuel assemblies will not appreciably effect the gross-power shape of the core. The power produced

- 43 TABLE 9 i Nuclear Characteristics 12 Advanced Fuel Assemblies

1. Cf (therrnal)

Fuel Type T L2 k,,

• Ea (the rmal) f PF 1 68.5

6. 2 1.195

.0602 .0837 PF 2 68.3

5. 9 1.155

.0626 .0872 PF 3, 4 68.3

6. 3 1.225

.0615 .0909 PF 5, 6 69.7

6. 2 1.195

.0637 .0916 PF 7, 8, 9 68 9

6. 8 1.233

.0571 .0846 PF 10 72.2

8. 0 1.146

.0511 .0746 e, PF 11, 12 72.0

7. 3 1.163

.0532 .0784 -s =.

• k,, hot, average voids with Zr-2 channels 4

i ) 1

4 within the element will, however, be slightly higher than the surrounding element as a result of the slightly higher fission Cross section. 1 The local effect of the advanced fuel assembly on the power distribution has been studied using the same cells described in the discussion of reactivity given above. The cases repre-senting the greatest power change are given.in Figure.ll. Analysis of the thermal capability of the advanced assemblies in-dicates the following maximum power output at the overpower condition, consistent with a burnout factor of 1.5. l TABLE 10 Maximum Power Advanced Fuel Assemblies Type Maximum Power (Mw/assy) PF 1, 2, 3, 4 3.36 PF 5, 6 3.41 PF 7, 8, 9 3.68 PF 10, 11, 12 4.51 Using the results of the cell calculations to determine the ratio of average power in the advanced fuel to the average power in the surrounding (type I) fuel, the maximum permissible power of the region in which the fuel may_ be placed without reaching the the rmal limits, can be determined. The analysis is based on an average power per fuel assembly at 25% overpower of 1.73 Mw in the 452 element core. The results of this analysis are given in Table 11, below: TABLE 11 Permissible Power in Regions Surrour. ding Advanced Fuel Power Level PF 1-12 Max (Mw/assy) Pewer Level Type I Type Power Level Type I of Type I Average Core Power PF 1, 2, 3, 4

1. 37 2.45 1.42 PF 5, 6 1.35 2.53 1.46 PF 7, 8, 9

1. 33 2.77 1.60 PF 10, 11, 12 1.20 3.76 2.17 4

1.4 l.3 63 @1.2 w2 FUEL TYPE FUEL TYPEI-25% CONTROL PFI 4ei.i x 1.0 ~ .90 l0 20 30 40 DISTANCE FROM CENTER OF FUEL ELEMENT (CM) LOCAL POWER DISTRIBUTION i 1.4 1.3 6 5 .2 1 E FUEL TYPE I-25% CONTROL d 21.1 FUEL TYPE PF7,8 0R 9 m 1.0 .9 0 .O 20 30 40 OlSTANCE FROM CENTER OF' FUEL ELEMENT (CM) LOCAL POWER DISTRIBUTION FIGURE 11

• J The first column is the ratio of the power lev,-1 of the advanced fuel to the power level of the surrounding Type I fuel. The second column is the maximum power capability of the surrounding Type I fuel element, and the final column shows how much above the av-erage power the Type I fuel surrounding the experimental assembly may be operated without reaching the thermal limits of the exper-imental fuel. The table shows that the fuel Types PF1 through 4 are the most sensitive to burnout, but the power distribution given in Figure 8 shows that the maximum relative power in the Type I fuel is about 1.2 which, on the basis of burnout, would not restrict the location of the advanced fuel.

It should be pointed out, however, that the power distribution and the relative power given above are for illustrative purposes and do not necessarily represent the actual power distribution that will be chosen for operation. The primary purpose is to show a rep-resentative design case which illustrates the method and degree of conservatism used in determining the position and operating char-acteristics of the various fuel assemblies. 5. System Stability The operation of the Dresden Nuclear Power Station has adequately demonstrated the ability of the power plant to respond stably to a large variety of operating conditions. During. startup, rod oscil-lating tests, pressure. regulator and system response tests all in-dicated that a large margin from any instability at all operating conditions exists. The current improvement in fuel design, to be used as the central core region, will have only a very minor effect on the over-all system stability. Of the numerous mechanical thermal and nuclear parameters affecting stability of Core I, only two are changed. First, the new Type II fuel rod is slightly smaller in diameter which leads to the conclusion that heat generated within the fuel will reach the coolant slightly faster than would be expected for the previous de-sign at the same power density. The resulting change in thermal response is reflected in the " fuel time constant" which is 10 seconds . 'for the new (Type II) fuel and 13.5 seconds for the Type I fuel of the first Dresden core. From the dynamic analytical standpoint, this represents a change in gain of less than two which, from the degree of stability experienced during Dresden testing, is insignificant. From the practical standpoint, by comparison, the VBWR has operated stably with a mixed core of both plate and rod fuel assemblies which had time constants of about 1/10 second and 13 seconds, re spe ctively.

This variation in time constant is therefore insignificant compared to the change that could be tolerated without approaching instability. Second, it is planned that the new improved fuel will be operated at a slightly higher power density than the fuel of current design. The ac-cumulative effects of higher power, less flow restriction, and larger entrance nozzle lead to a slightly higher velocity of the coolant in the fuel channel. The time required for a unit volume of coolant to travel through the co:te, i.e., " sweep time, " will decrease slightly, tending to increase the core stability. The conclusien is, therefore, that the effects of the fuel time constant 4 will tend to be compensated by the decrease in sweep time. Both ef-fects, however, are so slight that no detectable change in stability is expe cted. I 4 1 J .-w,- N -, _, - - - ~ r--

l . SECTION III SAFETY AN ALYSIS The remarkably stable operation of the Dresden Nuclear Power Station, under numerous artificially-induced transient conditions during startup and subsequent power <- ration, is experimental evidence that the rela-tively minor change. uesign of the new fuel will not affect the stability of the core. In addition, it has been demonstrated by the stable operation of the mixed core (rod and plate) fuel elements of the VBWR that even elements of a relatively large difference in thermal response may be operated together without affecting the inherent stability of the boiling water reactor. This experimental ev dence, the large body of knocledge previously ob-tained from the Boran experiments, SPERT, and EB WR, together with the qualitative analysk discussed in part 3 of the former section, lead to the conclusion that Core II in its entirety will respond to changes in external demand in the came stable fashion as the first core. The entire assembly - Dresden Core II - may be developed progressively. That is, the first replacement fuel may be most or all of the experimental assemblies, PF1 through PF12, and two of the Type IIassemblies described in Section II. The remaining new fuel assemblies would not be loaded until the next regularly scheduled reactor shutdown. The safety evaluation is con-sidered applicable to this mode of operation for the following reasons: 1. In any core configurmon, the experimentai fuel assemblies will be operated well within the thermal limits specified in Appendix A of DPR 2. 2. The over-all performance of any coie configuration steady-state or transient is not appreciably affected by a few separated fuel assemblies. The fundamental evaluation contained herein is, therefore, direc.ted towa rd a re-examination of several of the accidents affecting fuel temperature and reactivity originall and amendmentsiI).y presented in the Preliminary Hazards Summary Report The objective of the evaluation is to show, by compar-ison, that Dresden Core II does not compromise the conservatism inherent in the original Dresden design. Following the review of accidents is a justification of the proposed license revision concerning the moderator temperature coeffici6nt. 1. System Performance during Control Rod Run-out If the operator runs cut one control rod at a time in a random fashion, one after another, the maximum rate of reactivity insertion reached will be about 0.02%& per second for Dresden Core I. i i

L. aec 49 Under the same accident conditions, the maximum reactivity insertion rate for the central region of the Dresden Core II will be less than 0.02% & per second for the cor e loaded with stainless-steel and zircaloy chan-nels in the configuration described in Section II of this report. The previous analysis which represents a limiting value for Dresden Core II showed that, if the operator pulled out control rods steadily to increase reactivity beginning at ratad power, the flux would rise gradually to scram level and safely shut ,wn the plant. 2. Sudden Step Insertion of Reactivity The effect of increases in reactivity in steps of 0.4 to 0.6% ah in Dresden Core I was the subject of several computer studies. The results showed that even if a scram should not occur, the transient would settle out very quickly (about two seconds) at a slightly higher power with no serious system effects and that even larger step changes in reactivity would not produce much more serious results. The mechanisms which limit the severit of the accident are the prompt 28 increase in neutron absorption in the U by Doppler broadening and the effects of void formation. Assuming the same step increase in reactivity locally within the SS-clad fuel region of the Dresden II core, the same nuclear mechanism would limit the severity of the transient. A change to the original fuel design - which ia of interest in this evaluation is the increase of Doppler broad-ening of the corner ThO2 rods. The resonance integral and total res-onance capture of Th232~is slightly less while the Doppler broadening is greater than that of U238 The ThO2 rods produce, therefore, a slight increase in over-all Doppler reactivity coefficient. Considering the effects of voids, the slightly shorter time constant of the SS-clad fuel, compared to the original Zr-clad fuel, and the effect of the in-creased Doppler coefficient resul.t in a slightly lower flux peak in the SS-clad fuel. None of the effects examined cause a substantial change in response of the fuel to the assumed accident condition. All tend, howeve r, to re-duce the severity of the accident. 3. Mechanical Failure of Control Rod Drive The control rod worth in the stainless-clad fuel region of the new core will be less than that of the first core. This change will tend to require a greater number of rods removed to attain critic ality and is, therefore, at least equivalent to the measure of safety provided in the original design.

j. I i' 4. Accidental Valve Closure on Steam Line The response of Dresden Core II to the accidental closure of valves i on the primary steam line will not change significantly from the i original analysis. The resulting core void reactivity and flux excursion will not exceed those of Dresden Core I. The maximum calculated fuel temperature in the central core region is.les s than 5000 *F. Some center melting in the hottest fuel rod may occur, but no cladding failures are expected. 7 i 5. Fuel Cladding Failure i An evaluation of Dresden Core Iindicated that about 4000 fuel element segments could be leaking at a rate of 10-6% of the noble gas activity per second without exceeding the expected average maximum permissible stack emission rate. This analysis is equally applicable to the new fuel with the exception that the rods are not segm-nted so that 1000 fuel rods would result in approximately the same release as 4000 fuel segments. j Although this appears to be an undesirable increase in the effect of l a single failure, the factor of 4 decrease in the number of welds in-l creases the reliability adequately to compensate for the change in effert. 4. Relt. ant to this discussion are the 937 " corner rods" containing ThO2 (900.n Type IIfuel assemblies, and 37 in the experimental fuel assem-blie s). Assuming the same distribution of fission products resulting from fission of U233 and U235, the criterion for evaluating the change in hazard resulting from including Th02 corner rods in the new fuel design is the retention of fission products by ThO. The best avail-2 able experimental evidence indicates that the fission gas retention of ThO2 is superior to that of UO. In addition, the thermal conductivity 2 is ne 2 times grea - than that of UO and, unlike UO, ThO2 has only 2 2 one oxide state, henc*, it is stable in an oxidizing atmosphere. 3 These physical properties indicate that the central temperature would be slightly lower in a Th02 rod than a UO2 rod for the same surface l heat flux and that cladding failure will result in release of a smaller quantity of fission gas than from a regular rod. 6. Loss of Recirculating Pumps i j Assuming that all four recirculating pumps are tripped out, the SS-clad i Type.II fuel central region of the core, operating at a higher average power relative to the surrounding region, will reach a burnout ratio of 1.4 with natural. circulation at rated power, and 149 with natural circulation as tripped l from rated power. i'

I - 4 Undar currently planned operating conditions, the remainder of the core (Type I fuel and'lype PF-1 through PF-12) will be operated at a slightly lower average power and slightly higher peak-to-average - power relative to the total core average. The resulting effect is to subject feiver Type I fuel assemblies to the condition described in the original analysis. The analysis indicates, t!.arefore, that clad-ding failure should not occur in either fuel. 7. Refueling " r.,or The refueling accident evaluated in References 1 and 2 was re-evaluated for Dresden Core II, assuming the following conditions: j a. The entire core, less one center Type II stainless-steel-clad fuel assembly, has been loaded incorrectly with all zircaloy channels in the central reload region. b. The control rod adjacent to the vacant fuel position is with-drawn, causing the reactor to become just criticah The source level instrument flux counter fails or the operator c. fails to observe increase in count rate upon withdrawal of the rod. d. A new Type II fuel assembly is inserted in the vacant fuel po-sition at the maximum design rate for the hoist. ( m 12 inch / sec.) The period scram circuitry fails, e. f. The moderator temperature is 68 'F. The maximum reactivity rate reached during this hypothetical accident is 0.00287 Ak/sec. The resulting center fuel temperature will be about 4600*F. Some center melting in the hottest fuel rod may occur, but no cladding failures are expected. The reactor is automatically shut down by the high neutron flux scram circuitry, with some localized boiling occurring. In addition to the mechanical and procedural measures presented in the original accident analysis of Dresden, which are designed to pre-vent the accident, the core must also be incorrectly loaded with all Zr channels. Even in the unlikely event that all the accident conditions were simul-- taneously satisfied, the incident would not result in release of fission products to the reactor or the enclosure. 4 y., _..__,e.

. 8. Fuel Assembly Placed in Wrong Position It is' planned to operate the central region of the core (type II fuel as-semblies) with no orifice. If, by error, a type I fuel assembly with a small diameter orifice were placed in the central core region, the fuel rods could overheat and rupture. If the ruptu re occurred during the approach to full power, the change in power which. occurs relatively slowly and the longitudinal and radial <ak-to-average temperature within a fuel bundle would cause a pro-gressive failure of the fuel tubes. Within about two minutes of the first failure and 1;efore any stack re-lease, the off-gas monitoring system should indicate a relatively large increase in activity and will continue to increase with increasing power. The extent of the damage that could result from this hypothetical ac-cident is, therefore, limited to the relatively gradual rupture of a few fuel rods. 9 Group Removal of Control Rods at Startup The worst startup accident was re-evaluated for the Dresden core. The principal assumptions used for the purpose of analysis were: A minimum slightly subcritical size for the central $egion of the a. core resulting from withdrawal of three control rods which produce 4 the maximum rod worth in the remaining rod. b. Removal of the fourth rod at its maximum design rate. Coincident failure of the period scram circuitry. c. d. Reactivity insertion starts with the reactor power at 10-12 times rated power (equivalent to the spontaneous fission source level). e. A moderator temperature of about 68 'F. f. Dresden II Core loaded with SS and Zr channels in a " checker-board pattern, " and with fuel dencribed in Section II. The calculated reactivity rate resulting irom this hypothetical accident is ~0.0019 ok/k per second. The maximum hot spot temperature should be about 4 '50 *F. Under the conditions described, no melting of the fuel or clad would be expected. By comparison, the first Dresden core could reach a calculated maximum reactivity insertion rate of ~0.0029 Ak/k per second with a maximum hot spot temperature of about 500C *F. e .,n e.

-=_... 53 - The reduced rod worth and the slight improvement in Doppler coef-t ficient contribute to a reduction in the temperature ris e. In either i case, Dresden I or Dresden II Cores, it is not expecte d that fission 4 products will be released to the coolant. i 10. Introduction of Cold Water The worst " cold water accident" previously examined for the Dresden I core involved the rapid introduction into the reactor, water in a cold j coolant loop. l The accident 'was re-examined for the Dresden II core assuming: One recirculating pump is shut down and its discharge valve a. clo sed. 1 b. Water at 109 *F is contained in the loop. The reactor is at 85% of rated power (the maximum permissable c. power with three recirculation pumps operating). d. The operator fails to allow the coolant loop to warm up by natural circulation through the bypass valve before starting the recircu-lating pump. The maximum temperature of the fuel following the hypothetical accident would be less than 4700 *F. Even though the hypothetical accident, as shown above, will not result in serious consequence, the p(2) are such that the probability of such an rocedures outlined in " Operating Proce-dures and Emergency Plans" event is very remote. 11. Maximum Credible Accident i i The " maximum credible accident" for the Dresden plant is conceived as i following a hypothetical instantaneous complete severance of one of the bottom inlet lines to the reactor while the reactor is in a " hot" standby condition. Fellowing the accident, the fuel starts on a rising temperature transient, eventually reaching the melting temperature of all the fuel. Comparing the former analysis of tne first Dresden core with the cur-rent analysis,--it may be observed that: The total release of energy resulting from the accident has not a. i been changed and the probability of a stainless-steel-water j reaction is even less than the previously evaluated zirconium-i water reaction, and l b. A.relatively minor change, resulting from the higher power density I in the central region of the new core and the changes in fuel geometry,_ i ,_,--__,z..-._-___ m . _ ~ m.,~

1 54 - i is that some SS cladding reaches melting temperature in the cen-tral region of t te core slightly faster, and some Zr cladding reaches melting temperature slightly slower, than would have l occurred in the previous core. In summary, the changes in the core will not affect the total stored energy within the reactor vessel or the applicability of the original argument regarding simultaneous nuclear excursion or metal-water i reactions. There is, however, a relatively minor change in fission product release rates. But since the primary factor, i.e., power, l has not changed, the small increase in release rates will not appreci-ably affect the radiological aspects of the maximum credible accident. 12. Justification for Revision to Temperature Coefficieg } From a safety standpoint, the positive temperature coefficient of the reactor is of interest only if a mechanism exists which could lead to temperature transient, slow or fast, which is not self-limiting beicw operating conditions or a very fast temperature transient in which high fuel temperatures are reached before equilibrium conditons are es-1 tablished. Accidents resulting from the first condition are avoided by the re-quirement that the temperature coefficient becomes negative at lees than saturation temperature at rated pressure. A maximum initial temperature coefficient of 1x10~4 Ak/k'F and a maximum total reactivity addition of one dollar ok are specified in the j license amendrcent to limit the rate at which an incident may occur. 3 In order to verify the selected limits, several transient studies were performed on an analog computer. In principal, the studies simulated starting the reactor from a cold condition with an initial reactivity in-4 1 put and plotting the time-temperature history of the fuel and moderator. The results showed that, within a wide range, the initial reactivity input had a relatively small effect on the final fuel temperature. For the an-alog model, it was assumed that the water was stagnant, which accel-erated ils temperature effecte by the recirculation loop transport time. It was also as su. ned that the water within the channel had a positive temperature coefficient which does not include the time required to heat the water outside the channels. Both these assumptions tend to cause the results to be highly conservative. The results of one of the computer studies, shown in Figure 12, represents the temperature-time h: story of the reactor using the expected moderator-reactivity-l temperature input with reactivity step insertions of 20d and $1.20. The safety of the system is assured, within the limits of the definition, by virtue of the Doppler coefficient, the relatively large time constant of the fuel, the huge heat capacity of the water, and the fact that the transient is self-terminating before reaching operating temperature. 4 e om c g g g q' 1 A o :.) o m 100 % N N N N N O b N N M 55E5 D M I E m =$ M,i N.ENM A D C t!ii Ilji Idi ilfi I. H...U '. I F. @l 17 l,lI I!I .i.i*.n ill!ili ii.iili'diliiiiiii4F. fi#liatt-H*iifrii ai:Th W. F_iWM..u.3..ii.i.~ill.i.iJ'f.i.t.l.i.l.l. t":.i L.i4i7s. i t . a.................._...:...m.... .p ~. m 7 !ll! !in iHi i!H l!h E ti[ il ljHiH5filf 78: TFE t.55fff!!!'!ibij ?.f...jhlifliipirMi' HAlisiihmHijiiti a. !I!! Hl! Fl! ili! HF I!h W1 iWfit!1m lin DW EU!iii4!!'iii!Li Em i-siin~nt..:mt smA M.fpihilii itiifi# ili! Uffiii! j%i#ilii i

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56 - It should be observed that the moderator coefficient within the channels, for Dresden Cores I and II, has been shown by analysis to have a nega-tive temperature coefficient. If the moderator is heated by the fuel, the transient is, therefore, limited by the rate at which the temperature of the water outside the channels rises. The proposed revision to the temperature coefficient provides a meas-urable criteria with the necessary operating margin without compromising safety. S l T

57 - REFERENCES (1) GEAP-1044 " Preliminary Hazards Summary Report for the Dresden Nuclear Power Station, " by G. Sege, May 1, 1957. (1) GEAP-3009 "An-3ndment No. I to Preliminary Hazards Summary Report for the Dresden Nuclear Power Station, " by D. P. Ebright, May 1, 1958. l (1) GEAP-3076 " Amendment No. 3 to Preliminary Hazards Summary Report for the Dresden Nuclear Power Station, " by J. L. Murray, December 23, 1958. (2) GEAP-2071 " Operating Procedures and Emergency Plans for the Dresden Nuclear Power Station, " by D. P. Ebright, January 25, 1958. 4 J l i t l -}}