Text

Semiannual Operating Rept,May-Oct 1964
	ML20030A511
Person / Time
Site:	Big Rock Point File:Consumers Energy icon.png
Issue date:	11/27/1964
From:	Haueter R; CONSUMERS ENERGY CO. (FORMERLY CONSUMERS POWER CO.)
To:	;
References
	NUDOCS 8101090807
	Download: ML20030A511 (87)
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q; ' * -J / s Docket No. 50-lM C l-g f gu y g 1934;, [4 'p'ilo C m '\\ D e ^). /s i I!G"ic.?! '\\ 'S ~ 1 a, - H:L u;n:t ^ Report of Operation of Big Rock Point Nuclear Plant \\,, t;aEtutu 8 w....,_ 4 License No. DPR-6 p May 1,1964 Through October 31,196h %. c.D l /f-D - de f C Th s report, submitted in accordance with Paragraph 3 D.(3) of Operating License No. LPR-6 (effective May 1,1964), covers the first six-month period of operation of the Big Rock Point Nuclear Plant (Plant) under this license. I. SUMPMRY OF OF. RATIONS A. General At the beginning of the reporting period, the Plant was still shut down and in the final weeks of a scheduled outage to inspect the turbine-generator, rewind the generator field, and to load the re-actor core to 8h fuel bundles in preparation for increasing reactor power to 240 Mwt. Loading of the core to 84 bundles commenced'on May 1, 1964, and was completed on May 4, 1964. Necessary control rod drive testing and physics tests were completed by May 15 when the reactor head was installed. The generator was connected to the system on May 21 and power operation resumed. R&D testing, associated with power escalation to 240 Mwt hs outlined in 8.2.2.1 of the Technical Specifications), was started on May 22 and continued through May 30 without incident. On May 31, a reactor scram from about 69 Mwe occurred as a result of a spurious opening of the bypass valve. Plsnt operation and R&D testing resumed on June 1, 196h. On June 5, the +.arget Plant output of 75 Mwe was-reached. With the exception of the above-mentioned scram, all testing vas performed without incident. The remainder of the month of June was taken up with further tests on the 8h-bundle core. Flow-control tests utilizing the butterfly valves were performed with very gratifying results. R&D test results are summarized ~in "B" and "C" following, fh (h eCi

\\ s g Natural circulation tests were conducted late in June with the attainment of a maximum power of about 122 Mwt. Details of this testing (in answer to questions raised by DRL in a letter from E. G. Case dated Sept ?mber 1,1964) will be submitted in the near future. A second power scram occurrad on July 1,1964, during the routine 30-day exercising of the safety system, when a spurious trip on Channel No. 2 picoammeter coincided with i st operation of Channel No. 1 picoammeter trip circuit. Further testin6 Thowed that it was possible to induce a spurious operation of a second 'oammeter trip unit as a result of tripping the first unit. This et er. ult from the combination of a change in the load on a trip unit and ti marginal setting of the hysteresis potentiometer in the trip unit. A e procedure for setting the bysteresis control has been instituted. The Plant was shut down on July 13 1964, for a gamma scan of the 64-bundle core and reconstitution of a ih-bundle core for additional R&D testing at 60 Kw/ liter. More than, ' of the core was scenned with very good results. In general, peakin6 ' actors were found to be slightly lower than predicted. Loading of the hk-bundle core was completed in early August and the Plant was back on the line August 9 The 60 Kw/ liter power density target was achieved on August 18 with the kh-bundle core operating at 1050 psia. Core performance and stability tests were performed during the stepwise approach to 60 Kv/ liter. These tests demonstrated that the reactor was very stable and well within operating limits. On August 24, slight neutron flux oscillations were noted, and the next two days were spent trying to determine the exact cause. The reactor was shut down on August 26 when it appeared that the amplitude of the oscillations was increasing. Details of the testing performed at this time, and during the subsequent run, along with results and con-j clusions, are included in Appendix A. During the period August 26 through September 15, the Plant was shut down to inspect core internals in an attempt to find the cause of the flux oscillations. When nothing of significance could be found, the decision was made to resume operation with the 44-bundle core and per'orn further testing. The Plant was back on the line e -

1 o 3 September 15 The tests to determine the cause of the oscillations were essentially complete when this run was terminated on September 18 by a scram resulting from the spurious opening of the bypass valv<. Shortly following the scram, during a routine control rod drive coupling integrity check, two control rod drives showed evider.ce of galling and a third drive could not be withdrawn from the full-in position, indicating that a foreign object undoubtedly was lodged in the control rod drive nozzle extension. Since the oscillations had not reappeared, and the reactor had operated in a stable manner, it was decided to terminate the kh-bundle core testing. The Plant, therefore, was shut down for inspection, maintenance and for the removal of the foreign object from the control rod drive nozzle extension. (Details of the scram and the subsequent difficulties with control rod drives are contained in the TWX sent to the Commission on September 28, 196h.) During the shuffling of core internals to effect removal of the control rod drive, a total of chree bottom rollers from two con-trol rod blades was found to have become disengaged from the casting. Inspection of the remaining 30 control rods disclosed a number of lower rollers with excessive wear in the pin and/or pinholes. Examination of defective parts revealed that this excessive wear resulted from a mis-application of the pin design. New control rod blades are being fabricated. Sufficient control rods will be on hand, prior to the next start-up, to replace those control rods with rollers and pins showing excessive wear. While inspecting the control rod blades, two thermal shield hold-down studs were found to be fractured. Further examination reve. led four more fractured studs. At the end of the reporting period, investigations into methods of repair were underway. Details of the stud failures and the proposed repairs are included in Appendix A. During this reporting period, the reactor was brought critical 45 times, and was critical for a total of 1876 hours of operation. Total heat produced by the reactor was 321,259 Mwh(t). The turbine-generator was on the line for 1703 9 hours with a gross electric generation of lo3,h26 Mwh. Net plant electric generation for the period t was 96,2h3 Mwh.

3 e 4 B. R&D IWsting Bundle Core During May and June,1964, the Phase II recctor performance ~These tests c)vered the tests were completed for.the 84-bundle core. following range of operating parameters : Reactor Pressure 1250 and 1350 Psia Reactor Power 0 to 235 Mvt Recirculation Flow 3 5 to 12 5 Million Lb/Hr (Natural and Forced Circulation) Inlet Subcooling 8 to 34 Btu /Lb Two main objectives were accomplished in these tests. First, the power output was increased from 157 to 235 Mvt (50 to 76 Mwe). (successfully attained on June 5,1964. ) Secondly, steady-state and transient test data were obtained from the instrumented fuel assemblies under a variety of operating modes. The more important results of the teste are briefly summarized below: 1. Core Performance Tests 4 These tests consisted of obtaining steady-state measurements of flow rates, neutron flux, pressures, and temperatures in the nuclear steam' supply system. From these measurements, the important thermal-hydraulic parameters such as heat flux, kilowatt per foot of fuel rod, power density, and burnout ratio were computed. Table I summarizes the variations in these parameters at different rower levels. Some items worthy of note are : a. As the thermal power increased, the total recirculation flow decreased slightly from 12 3 to ll.9 M lb/hr. This was due to the gradual build-up of steam quality with the resulting in-s crease in two-phase pressure drop in the recirculation loop. i .b. It was necessary to raise reactor operating pressure frcm 1250 psia to 1350 psia when it was found that the turbine admission valves would hav; been vide open at about 73 Mwe at 1250 psia. c. The axial fJsx was flattened markedly because of the insertion of poison rods into the developmental fuel assemblies. L( 1-

-o 5 TABLE I CORE PERFCA W CE TEST RESULTS PHASE II POWER ESCALATION TESTS i ' Test Number 2 6 8 11 13 Plant Gross Power, Mwe ' 49 59 65 68 76 Reactor Power, Mwt 153 175 19h 2% 229 Recirculation Flow, M Lb/Hr 12 3 12.2 12.2 12.1 11 9 Subcooling, Btu /Lb 17 0 18.7 19 7 19 7 21.h Reactor Pressure, Psia 12k5 1245 1252 1258 1362 Core Average Exit Quality 0.041 0.049 0.057 0.064 0.076 MaximumHeagFlux, Btu /Hr-Ft 257,000 293,000 327,000 347,000 h3h,000 Fuel 3odPower,Kw/Ft 8.3 9.6 10.7 11.4 13 2 Core A'erage Power Densitf,Kw/L 29 32 37 39 43 Gross Peaking Factor-2.85 2.85 2.85 2.85 3 24 Hot Channel Exit Quality 0.11 0.13 0.15 0.16 0.19 Critical Heat Flux Ratio 2 94 2 59 2 32 2.18 1 56 I 4 )' + 4 4 .+s ,e .e c-m w,- -~ w-r- ->---w e-o

6 2. Pump Trip Tests Four pump trip tests were run with the 84-bundle core. Starting from various power levels, both recirculation pumps were tripped. The recirculation flow and reactor power were allowed to coast down to their natural circulation levels. The traces of recircu-lation flows and heat fluxes during the coast down transients are shown in Figure 1. The recirculation flows were measured with the in-core flowmeters in the instrumented assembly. The flow traces in Figure 1 show a slight trend with power l~ vel. As '.he reactor power is in-creased, the in-core flow initially drops slightly more rapidly, but levels out to a higher natural circulation flow. The heat flux also exhibits a slight trend to drop more rapidly as the reactor power is increased. 3 Flow Control Tests Flow tests were performed on the 84-bundle core to demonstrate that the reactor power can be controlled by throttling the recirculation flow through the core. The tests were performed by starting initially at rated recirculation flow and throttling the re-circulation flow in steps of 90, 80, 70, 60, and 55 percent of rated flow without movement of control rods. At each step, data were obtained for a plant heat balance, and data were also obtained from the in-strumented assemblies. The reduction of reactor power with recircula-tion flow is shown in Figure 2a. Another interesting way to display the data of Figure 2a is shown in Figure 2b. Here, Qo represents the reactor power when recirculatioa flow is rated, W ; and Q represents the reactor o power when the recirculation flow is some value, W, other than rated flow. Figures 2a and 2b indicate the following: a. The reactor power reduction is app. tmately linear with flow rate in the range of the flow reduction. As the natural circu-lation flow rate is approached (or, more appropriately, as the core pressure drop approaches the natural circulation driving head), internal flow paths may be set up in the reactor vessel through the leakage paths around the core. This would be ex-pected to cause the correlation in Figure 2b to become slightly ( nonlinear.

r T i b.. When the data in Figure 2a are' plotted in the normalized fashion of Figure 2b, one set of flow control data falls atop the other. - The linear correlating equation has the form S=a+bE Ro Wo The factor, a, in this equation has been shown to be largely a function of the doppler reactivity, with only slight dependence 4 on xenon reactivity and reactivity-in-voids. During one of the flow control tests (starting at 147 Mwt I and rated recirculation flow) a flux.vire was irradiated for approximately 1/2 hour at each of the flow steps in Hole No. 5 The resulting axial flux distributiens are shown in Figure 3 It is evident from the figure I that, with the exception of the fine structure on each of the axial pro-files, the axial flux shape remains essentially constant. This suggests that only a minor chan6e occurs in the steam void distribution during flow control operations. Additional flow and power information was obtained from the instrumented assemblies during the flow control tests as shown in Figure 4. The upper portion of Figure 4 shows the trend of exit quality in the. instrumented assembly during the flew control tests. During the low power test, no noticeable increase in exit quality oc-curred. For the higher power test, the instrumented channel exit quality apparently increased by about 3 percent in value as the flow was reduced from rated flov to 55 percent of rated flow. However, plus or minus 3 percent of the reading is s oproximately the uncertainty of the exit quality measurement. Therefore, it is doubtful whether a firm con-clusion can be made about a trend of exit quality frcm these data points. Shown in the second and third parts of Figure h are the instrumented assembly flow and power. Since these data points fall nearly on straight lines, it is concluded that there was no substantial flow redistribution and no substantial power redistribution during the flow control test.

The instrumented bundle radici power factor (defined as

.f the ratio of the instrumented assembly power to the average assembly power) ,--,-r y

a 8 i '{ ~ variation during the flow control tests is shown in the last section of Figure 4. Here, the trend of the data suggests that the radial power factor decreases with flow. However, the magnitude of the uncertainty prevents a firm statement from being made. The general conclusions and observations from the flow control tests reported above are: a. The reactor power changes approximately linearly with recircu-lation flow in a boilin6 vater reactor. This suggests that the recirculation flow may be profitably used as a means of con-trolling the boiling water reactor power. b. Provided the recirculation flow changes are made slowly enough, and the control rod pattern is not changed, only small changes in power distribution result. If the flow changes occur in less than the transit time of the steam voids across the core, as happens during loss of recirculation pumps, a transient redistri-bution ot flux vill result. 4. Load Following Tests F Tests were run to demonstrate the capability of controlling reactor power with recirculation flow varititions. The tests were run by putting the Plant on turbine governor cor.crol, and manually rejecting increments of power at varying rates. A reactor operator was stationed at the controls of the recirculation pump discharge butterfly valves with i a trace of reactor pressure measurement at his disposal. Thus, a manual feedback control path was obtained which used the signals of reactor pressure and rate of pressure change with time for control. The method of control was to manually throttle first one recirculation valve to its maximum closed position, and then the other. The variations of neutron flux and in-core flow ith recirculation flow were as expected. The maximum rate of electric load change which was ac-complished in the four load-following tests was h.5 Mwe per minute. However, this rate was not limited by plant equipment. Tbst data in- -dicate that smooth End rapid control of BWR's is available if automatic control equipment is used, together with flow control.

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9 .v-4 (~ C. R&D Testing - 4h-Bundle Core During portions of August and September,1964, Phase II reactor performance tests were completed for the 44-bundle core. The objective of this series of tests was raising the reactor core average power density to 60 Kw/1 (approximately 160 Mwt - successfully attained on August 18, 1964). The various tests consisted of one-pump trip test, steady-state thermal-hydraulic core performance tests and two control rod-oscillator tests. The results of these tests were as expected and confirmed the stability of the core. Table II presents the results of the core performance tests. Also shown in Table II are the four principal steps taken in the approach to 60 Kw/1. II. ROUTINE RELEASES, DISCHARGES AND SHIPMENTS OF RADIOACTIVE MATERIALS A. The gaseous radioactivity released from the stack during power 13 operation averaged 50 pc/see of N and 4 pe/see of fission gases. Based upon 1339 effective full power hours of operation during the re-13 porting period, this results in a total release of 2hl curies of N and 19 curies of fission gases during this reporting period. Gaseous release during periods when the Plant was shut down was negligible. B. During this reporting period, the liquid radioactivity releases to Lake Michigan, by way of the cit-culating water discharge canal, numbered 115 batches, with a total activity of 5 2 curies. Twenty-nine batches were released on a partially identified basic wherein at least 90% of the activity was determined to be a combination of CoS8 65, and Zn All other batches were released under unidentified limits. C. There were three off-site shipments of radioactive material as i follows: Shipment No. Date Transfer License No. Radioactive Material 1 8/13/64 DPR-6 to General Electric One In-Core Detector (APED SKM-54) Containing 90% Enriched Uranium (U 0 ) 38 2 9/15/64 DPR-6 to General Electric 266 pc Activated Company (Vallecitos 0017-60) Corrosion Products - Reactor Water and Crud Samples p 3 10/22/64. DPR-6 to General Electric 50 5 Curies - Activated (- Company (Vallecitos 0017-60) Stainless Steel Bolts and Miscellaneous Hardware

10 TABLE II CORE PERFORMANCE TEST RESOLTS PHASE II HIGH POWER DENSITY TESTS Test Number 31 34 35 36 Plant Gross Power, Mwe 39 5 46 49 51 Reactor Power, Mwt 11h 136 146 158 Recirculation Flow, M Lb/Hr 12.0 11 9 11 9 11.7 Subcooling, Btu /Lb 11.1 11.1 12 3 13 0 Reactor Pressure, Psia 1050 1050 1050 1050 Core Average Void Fraction 0 39 0.hh 0.h6 0.a' Maximum Heat Flux, Btu /Hr-Ft 310,000 371,000 463,000 444,000 Average Heat Flux, Btu /Hr-Ft 134,000 160,000 172,000 186,000 Core Average Power Density, Kw/L 42 51 55 59 Critical Heat Flux Ratio 30 25 2.0 1.6 ?

~ -~- y-11-c I l .III. RADIOACTIVITY LEVELS IN PRINCIPAL FLUID SYSTD4S ' A. Primary Coolant The activity levels in the primary coolant during this period of power operation ran as follows: j Min Avg Max -3 -2 -1 Reactor Water Filtrate ** pc/cc 3 x 10 1,o x 1o -1.8 x 10 -2 -1 -1

Reactor Water Crud ** pc/cc/Turb 3 3 x 10 2.h x 10 6 3 x 10

-5 4 -N l Iodine Activity *** pc/cc 1 x 10 1,1 x 1o 2 5 x'10 i B. Reactor Cooling Water System The principal radionuclides in the reactor cooling water 51 system were Na and Cr, which resulted from activation of the sodium [ chromate. inhibitor. (A portion of this water flows through the biological shield cooling jacket.) i Min Avg Max -5 -3 -1 Reactor Cooling Water ** pc/cc 9 7 x 10 3x: 0 1,1 x to l C.. Spent Fuel Pool f Radioactivity in the spent fuel pool is principally ~ 1 activated corrosion products going into solution from fuel and core components. During this six-month period'there has been a very sub-stantial amount of movement of fuel and components in and out of the j-pool. Min Avg Max -3 ] Fuel Storage Poo1** pc/cc 1 5 x 10 1 5 x 10 5 3 x lo-Iodine Activity *** pc/cc

Background

5 x 10~ 2 5 x 10~ IV. PRINCIPAL MAINTENANCE PERFDRMED A. The first-year routine inspection of the turbine-generator was performed-during the period March 30, 1964 through May 21, 196h. During this inspection, the first-stage row of blades was replaced when a crack in the root of one of the blades was found. Several bear-T ings and seals required some maintenance. The blading and steam paths

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12 C g of the turbine were generally found to be in good condition. The generator stator was rewedged after a number of slot-filler-pieces were found to be loose. Also, during.this outage, the generator rotor was revound to correct a condition of chorted turns that had existed since initial plant checkout. B. New seals were installed -in each of the reactor recirt sting pumps. In order to gain experience with new seal materials of potentially ~ longer life, No.1 pump seals were replaced with an aluminum oxide hard l surface material, while No. 2 pump seals were replaced with a tungsten } ' carbide hard surface material. The initial seals in both pumps had I utilized stellite. C. The six steam-drum safety valves were removed from the steam drum, j the springs changed, the seats lapped, and reset for pressures applicable to 1050 psia operation. These changes were necessary, since earlier operation during the period had been at 1250 and 1350 psia, while the kh-bundle, 60 Kw/ liter operation was at 1050 psia. D. . Considerable trouble shooting was performed on the main steam 1 bypass valve control system, and several suspected components were re-placed following the spurious opening of this valve. E. Following the scram on September 18, 1964, control rod drives in A-4 and E-3 positions were removed and replaced due to galling between the index tube and the upper guide sleeve. One index tube was replaced and the other cleaned up and polished. Both drives were cleaned up, ' reassembled and set aside as spares. Another control rod drive, in Position C-4, could not j be withdrawn due to a foreign object lodged in the control rod drive extension. In order to dismantle this drive, it was necessary to re-move the reactor vessel head as well as the fuel and support-tube-and-channel assemblies surrounding this drive. (The foreign object, the head of a 1/2-inch SS bolt, is believed to have originated in the steam drum during the steam drum preoperational cleaning and inspection.) F. The manhole in the steam drum was removed, drum internals inspected and the manhole replaced. The drum appeared to be very clean I -and all~ internals in good condition. 1 4 y, ,,,,,,-e + -,-,.--,....~5

13 11 k G. The steam drum safety valves were removed, springs changed and the valves reset in anticipation of resuming operation at 1350 psia. H. All scram ~ valve diaphragms have been replaced after the failure of an outlet scram valve diaphragm early in the period caused a drive to. insert slowly. Inspection of several other diaphragms at that time indicated sufficient wear to warrant replacing all diaphragms. t I. All 12 thermal shield hold-down bolts and the three alignment pins were removed from the reactor vessel thermal shield support pads j and shipped to General Electric, Vallecitos for examination. Inspection of the thermal shield, support pads, and alignment is ' virtually complete, with all components,:other than the bolts and pins, in good condition with no indications of wear, cracking or distortion. Details of the inspection, evaluation and planned repairs to the thermal shield hold-down system are included in Appendix A (attached hereto). V. CHANGES, TESTS, AND EXPERIMENTS PERIORMED PURSUANT TO 10 CFR 50 59 (a) A report of changes, tests, and experiments for the period of' August 31, 1963 through August 30,196h was sent to the Commission on i November 13, 1964. There were no other reportable items for the period August 30,196h' through October 31,196h except for a bypass valve oscillation test performed as part of the diagnostic test series following 'he detection of the neutron flux oscillations during the kh-bundle core l testing. The bypass valve oscillation test was utilized to determine the l exact pressure-to-flux transfer function. When utilized in conjunction with the reactivity-to-flux transfer function (determined from control rod oscillation tests), it provides the necessary information to obtain the reactivity effect of pressure changes. This test was performed at a power level of approximately 115 Mwt. Nothing unusual was detected during this test. The data derived from the test substantiated conclusions drawn earlier from other tests - that the reactor is inherently stable. VI. PERIODIC TESTING PERFDRMED AS REQUIRED IN THE TECHNICAL SPECIFICA'll - !S The following table shows the systems or functions which must be tested. periodically to comply with the requirements of the Technical' Specifications. The required frequency of testing and the { dates of testing are also listed: } j e r-es -.-w ~ ,,,,, - - =, s-ew .-,r-s-- -- - - - - - - - - - ~

lb System or Function Frequency of Dates Undergoing Test Routine Tests Tested Control Rods Continuous withdrawal and insertion Each major refueling 8-3-64 of each drive over its stroke with and at least quarterly 9 64 normal hydraulic system pressure. during periods of power Minimum withdrawal time shall be 23 operation. seconds. Withdrawal of each drive, stopping Each major refueling 8 64 at each locking position to check and at least quarterly 9-14-64 latching and unlatching operations during periods of power and the functioning of the position operation. indication system. Scram of each drive from the fully Each major refueling 5-6-64 withdrawn position. Maximum scram and at least quarterly 9-12-64 time from system trip to 90 percent during periods of power of insertion shall not exceed 2 5 operation. caconds. Insertion of each drive over its Each major refueling 5-15-64 entire scope with reduced hydraulic but not less frequently 8-6-64 system pressure to determine that than once a year. 9-10-64 drive friction is normal. Control Rod Interlocks Rod withdrawal blocked when any Each major refueling 8-3-64 two accumulators are at a pressure but not less frequently 8-31-6h below 700 psig. than once a year. 9-2k-6h Rod withdrawal blocked when two of Each major refueling 8-3-64 three power range channels read but not less frequently 9-14-6h below 5% on 0 - 125% scales (or than once a year. 9 64 below 2% on their 0 - 40% ccales) when reactor power is above the minimum operating range of these channels. Rod withdrawal blocked when scram Each major refueling 8-3-64 dump tank is bypassed. but not less frequently 9-14-64 than once a year. 9-24-64 Rod withdrawal blocked when mode Each major refueling 8-4-6h selector switch is in shutdown but not less frequently 9-13-64 position. than once a year. 9-24-64 [ Containment sphere access air-Six months or less. 8-15-64 locks leakage rate. w

15 System or Function Frequency of Dates Undergoing Test Routine Tests Tested Control Rod Interlocks (Contd) Post-incident spray system At each major refuel-5 64 automatic control operation. ing shutdown but not 8-7-64 less frequently than 8 64 once a year. Reactor safety system scram At each major refuel-5-1 6-64 circuits requiring plant shut-ing shutdown but not 8-7-64 down to check. less frequently than 9 64 once a year. Containment sphere isolation At each major refuel-5-1 6-64 trip circuits. ing shutdown but not 8-8-64 less frequently than 9-14-64 once a year. Resetor emergency cooling systems' At each major refuel-5-16-64 trip circuits (core spray system ing shutdown but not 8-8-64 and emergency condenser). less frequently than 9-2-64 once a year. 9-14-64 Liquid poison system component Two months or less. 5-15-64 check. 7 64 9-9-6L Test firing or explosive valves. One year or less. 5-15-64 The following instrument calibrations were performed on a frequency of one month or less: Reactor safety 'ystem circuits not re-quiring plant shutdown to check; air ejector off-gas monitor; stack gas monitor; emergency condenser vent monitors; process monitors; and, the area monitoring system. By Robert L. Haueter (Signed) Robert L. Haueter Assistant Electric Production Superintendent - Nuclear Consumers Power Company Jackson, Michigan Date: November 27, 1964 Sworn and subscribed to before me this 27th day of November 1964. (SEAL) Grace Warner (Signed) Notary Public, Jackson County, Michigan My commission expires February 16, 1968

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I dEEEEElE E DEI' AILED DISCUSSION OF RECENT OPERATING DIFFICULTIES AT BIG ROCK POINT The Six-Month Report of Operations of the Big Rock Point Nuclear Plant, of which this " Appendix A" is an attachment, briefly describes the recent operating difficulties experienced during the planned R&D Program testing with a small, kh-bundle core, operating at an average power density of 60 Kw/ liter. This " Appendix A" goes into greater detail, as well as presenting the analyses and conclusions with respect to the significance of these events. This Appendix is also intended to answer specific questions (re Division of Reactor Licensing " letter request" dated October 27,196+) with respect to the following: I. Neutron Flux Oscillations II. Main Steam Bypass Valve Malfunction III. Control Rod Drive Malfunction IV. Thermal Shield Hold-Down Stud Failures (

1 / I. NEUTRON FLUX OSCILLATIONS 4 1 L .I l I i

i ' ' = 1, II - I. NEUTRON FLUX OSCILLATIONS' LA. Summary On August 24', 1964, neutron flux oscillations were' noted by the operators of the Plant during operation at' a nominal 60 Kw/ liter - with the kh-bundle' core. Measurements indicated that the neutron flux -oscillations were occurring at a frequency of.l.3 cps with an amplitude. - varying-between 6% and 9%. These measurements, and data previously ob-tained during research and development' tests with.this core, showed that the neutron flux oscillations were not caused by hydraulic instability, i but were being driven by a mechanical vibration.- The Plant was shut down on August 26. Although the core was not-ccmpletely dismantled,. careful inspection of suspected components that might have caused the oscillations, showed that the reactor vessel internals were apparently intact. The ht~ bundle core was reconstituted . and further testing was performed during the week.of September 13 in order to determine the source of the oscillations. On September 18, the testing was terminated by a reactor scram. During the shutdown which followed, F further insl.ection of the reactor vessel internals revealed that 6 of the 12 studs holding the thermal shield had failed. An analysis of the failure of the hold-down studs was con-ducted and is presented in detail in Section IV. Tne prog. essive failure of these hold-down studs (all on one side of the thermal shield) produced a rotational mode of vibration of the thermal shield about the horizontal axis. Variations in the core bypass leakage flow of sufficient magnitude to cause the neutron flux oscillations at the frequencies observed can .be directly related to this mode of thermal shield motion. B. Stability of the High Power Density Core Two control rod oscillation tests were performed with h-bundle high power density cores at approximately the same conditions as occ rred during the neutron flux oscillation. The iirst test was on August 17, one week before the neutron flux oscillations were first ob-served, as a planned part of the research and development testir4;. The second test was on September 17 with the reconstituted hh-bundle core. ( - These tests confirmed that the two high power density cores operated with large stability margins, and were well damped. 4 m I ..m m.. -..s. .-n ,,.m .,,-,,_n,,- ,.m .,,,7 y.

e'- rz ^ 2 l-4 j ' Data are presented in Figures 3 and lo and are summarized in -the. table' below: Table I Predicted Vs Actual Gain and Phase Margins' for. High. Power Density Core Predicted Measured Test 37 Gain Margin 21 db 16 db i layp2st 17,1964 Phase' Margin 139 90 Test 49' Gain Margin 22 db 14 db ~ September 17, 1964 Phase Margin 139 84 L 3 Neutral Flux. Oscillations Gain Margin-22 db August 26, 196k Phase Margin 139 .The'significant results are that the gains and phase margins measured ~ during.the tests indicate a stable system with a high degree of damping. -Note that, in both tests the measured gain margins agreed within 2 decibels and the phase margins agreed within 6 degrees. The differences between the three predictions in Table I are explained as follows : 1. For Test 37, the void coefficient of reactivity was 5 3 cents per percent void. For Test h9, the void coefficient of reactivity was 4.8 cents per percent. void. The decrease of 0 5 cent per percent void for Test h9 compared to Test 37 resulted in an increase of 1 decibel in the gain margin for the prediction of Test 49, with no change in phase margin. 2. The prediction of the stability margin for the conditions occurring during the oscillations recognized that the butterfly valves in the recirculation flow loops were throttled, thus adding damping in the single-phase portion of the loop. This resulted in a slight reduction in the hydraulic feed-back response and a slight increase in the gain margin compared to the conditions at Test 37 j. 3 The control rod patterns and operating conditions for these three tests are presented in Table III of Attachment B. The test results described above showed stability margins which were smaller than the. predicted margins. When the analyses were l ~ -( run, they were. based on the assumed rated core bypass leakage flow. It is shown below that the actual bypass flows were larger because of the ~! n.-

3 ~ 'I ' failure of che thermal shield hold-down studs and subsequent movement of 1the therr21. shield. The differences between the predicted and measured stability-margins are, therefore, partially accountable by this effect. In additionZthe uncertainties in the measured gain and phase margins of ~ 2 db' and' *10 absorb some of the differences between the predicted and. l' measured quantities. C._ Analysis of Neutron Flux Oscillation When the reactor was s' ~.t.down on August 26,1964, it was concluded that the reactor system was not unstable but that the system was being driven by a mechanica' vibration which caused a re-activity variation of sufficient strength to result in the flux oscilla- ~ tions. The arguments which pointed to mechanical vibration are presented -in the following discussion. Data in support of these arguments are presented and discussed in A. above.and in Attachments A and B. 1. T.e control rod oscillation test indicated that the Plant was operatin6 with substantial stability margins. 2. As the reactor was being shut down, a fast picoammeter I recording was made at about 20 Mwt (5 Mwe) which showed that the neutron flux oscillations still continued. At this low operating power, the hydro-dynamic feedback.which would have been required to support spontaneous oscillations had virtually disappeared. This is primarily because the l' void coefficient (dk/k)f(ARg/Rg*) was lower by a factor of five at i 20 Mwt than at 150 Mwt. In this type of system, such a change in feed-back gain would have increased the damping and reduced or eliminated any natural oscillations. Since such reduction in power did not signif-icantly reduce the observed oscillatory response, the system was neces-j sarily responding to a driving function, probably as mechanically induced core flow variations. 3 The indication of flow oscillations by the in-core channel exit flovmeter during the neutron flux oscillations suggest,ed l two possibilities. First, that the neutron flux was causing flow oscillations. This required that the neutron flux oscillations have . sufficient magnitude to respond through the fuel as heat flux oscillations. Because the response of heat flux to changes in neutron flux are slowed ,( by'an' effective time constant of 6.5 seconds, any heat flux response to J

Rg =_ Core Bulk Exit Void Fraction

b variations of 13 cycles per second (cps) would have to originate from very much larger neutron flux variations than were experienced during the observed operation. During e control rod oscille. tion tests (Figure 3), the exit flow response confirms thic filtering effect by essentially no response at 1.25 cps. The second possibility was that driven flow oocillations were occurring through the core which caused tlie neutron flux oscillations. The flow oscillations could have occurred in one of two modes - either where the loop flow osci3]ates in series with the core flow (the so-called " loop mode" of flow oscillations), or where the flow was exchanged between the core and a flow path in parallel with the core (the parallel channel mode). The lack of indication of flow oscillations in the recircaLAtion flow measuring nozzles ruled out the " loop mode." Therefore, the flow oscillations evider.tly occurred in the parallel channel mode. It was suspected that a bypass flow path around the core was being alternately opened and closed by mechanical vibration of some component in the reactor vessel. 4. This latter notion was supported by three substantial pieces of information. First, the core pressure drop had dropped 0.4 psi with no accompanying change in recirculation flow. The only way this could cecur was for a bypass flow path to be opened around the core. Second, it was known that if the oscillations had been induced by hydro-dynamic instability, the natural frequency would have changed approximately 20% for a 20% reduction in flow (i.e., on a 1 to 1 basis). When the r?- circulation flow actually was throttled, the frequency of the neutron flux oscillation changed by almost a factor of 2 for a 20% reduction in flow. This suggested that the neutron flux oscillation frequency was a power series function of the flow rate, a condition which is characteristic of certain modes of flow induced vibration. Third, during the three days that the flux 1scillations were observed, the turbidity of the reactor water (while still very low) increased significently. This suggested the openirg of a new flow path and the flushing of crud from a previously more stagnant area. Accordingly, a search for abnormalities in the reactor vessel internals was conducted. Specifically, the major items inspected ( and found to be intact were as follows:

l ) 5 f. a. The steam baffles in the reactor vessel were properly : atered. b. All of the grid bars were in place and latched. c. All of the fuel hold-down assemblies were in place. d. All of the plugs in the unused fuel channels were in place. In addition, fuel, support-tu'e-and-channel assemblies, t sources, detuning stabilizers, control rods and orifices were examined for evidence of movement, unusual wear or other ancmalies, with nothing of significance noted. No evidence was found during this first'shu,tdown of mechanical. defects sufficient to fully explain the observed transients. Tb elimir. ate one possible contributor, an improved <ource hold-down and orifice was installed to further restrict a possible variable flow path at high-pressure drro conditions. D. Source of the Neutron Flux Oscillations The discovery of the failed hold-down studs for the thermal shield and the determination of the mode af failure (See Section IV) pro-vided the necessary additional clue to the neutron flux oscillations. Analysis of the other clues in light of the stad failures conclusively related the neutron flux oscillations to the hold-down stud failures. 1. The analysis of the stud failures *ndicated that just prior to shutdown on August 26, the 2 studs adjacent to the 6 failed studs were deflecting a maximum of 90 mils. Tnus, the thermal shie'.d was being lifted. at its highest point approximately 180 mils. The effect of the asymmetric loading o: the vibration of the thermal shield was to intro-duce a rotational hc3ree of freedom about a horizontal axis. The natural frequency of the rotntional mode of vibration was approximately 1 5 cps. Both tae side-to-side and the rotational modes of vibration demonstrate calculated natural frequencies in the range of the frequency of the neutron flux occillations. In the discussion below, it wi11 be shown that vibration amplitudes in the order of 1/h-inch could have caused variations of bypass flow of sufficient amplitude to cause the neutran ( flux oscillations. Tne actual vibrational amplitudes were believed to be

9 6 approximately 1/k-inch (based on the wear observed on the hold-down studs' after they were removed from-the reactor vessel). 2. .The -seal which prevents gross leakage of water around the thermal shield is a stainless steel washer (3/8"' thick by'approximately 8'10" OD). When the thermal shield rested on its support (as shown in Figure 1), the seal was loaded like a Belleville washer between the lower - edge of the. thermal shield and a protrusion in the reactor vessel vall. With the thermal shield lifting off its supports, the seal lifted off the protrusion. Under this latter configuration, the limiting flow area occurred in the radial distance between the OD of the seal and the ID of the reactor vessel. The tolerances between these two dimensions indicate that the most probable gap dimension was approximately 60 mils, but that the maximum dimension could be 90 mils and the minimum dimension zero mils. Steady-state calculations were made to determine the effect of opening gaps of various dimensions in the annular space between the seal and the reactor vessel wall. The results on the flow, pressure drop and void fraction are shown in Figure. 2. With reference to Figure 11, which shows the measured core pressure drop during operation with the 44-bundle high power density cores, the fellowing seque. ice of events can be postulated: a. On August 21, a decrease in the core pressure drop of about 0.2 psi was observed over a 24-hour period. This is associated with the failure of the first hold-down stud (s). Although fast records of neutron flux were not made on this date, the picoammeter records indicate that the intermodulation (similar to that shown in ~ Figure 7) may have been occurring at a greatly reduced amplitude. b. Late on August 23, another decrease in core pressure drop of 0.1 psi was noted. This is as-sociated with further stud failures. Soon there-(- after, on August 24,'the neutron flux oscillations were first observed on the picoammeter recorders in the control room.

17 - I

c...On August 25,:a third' reduction in core pres'ure

. drop of 0.2 psi occurred, associated with the further stud failures. The traces rhown in Figures 5 through 8 were taken on this date. d. When operation was resumed on September 13,1964, . vith' the reconstituted 44-bundle core, a further drop in core pressure drop was'noted. The core pressure drop at this time was about the same as that experienced at the end of the previous'run, although the recirculation flow vas higher. This is the equivalent of approximately 0.2 psi, and again could be associated with further stud failures. In any event, because conditions were again different from those existing at the time of the neutron flux oscillations, it is quite understandable that the oscillations did not recur. With

ach failure of hold-down studs, a shift in the average position of the thermal shield was occurring.

An average gap between the seal-and. reactor vessel wall of 40 mils agrees with the toti hange in core pressure drop of 1/2 psi. Moreover, the small change in total recirculation flow of less than 1% agrees with the observation of no change in recirculation flow made on the control room instrument (the measurement uncertainty of this instrument is estimated to be 1-3/4%). The results of control rod oscillation Test 37 indicated that a driving reactivity of approximately 5 cents was required to cause 'the neutron flux oscillations which were observed. Since.the void co-efficient of reactivity was 5 3 cents per percent void, a change in the average void fraction in the core of less than 1% was sufficient to cause the neutron flux oscillaticns. This relates to a change in gap dimension of approximately 30 mils at zero frequency (Figure 2). At 1 3 cps the change in gap dimension approximately one order of magnitude i greater was required since ' at this frequency.the void response to flow changes is about 1/10 as great as at zero frequency. 3 In summary, a direct relationship is established ( ' between the failure of the hold-down studs on the thermal shield and the neutron flux oscillatic observed at the Big Rock Point Plant in three ways:

_a

8 a. The apparent time of failure of the studs coincides with the start of the oscillation. b. The natural frequencies of vibration of the thermal shield occur at the observed frequencies of neutron flux oscillation. c. The observed amplitudes of vibration of the thermal shield are sufficient to result in reactivity changes which were required to drive the neutron flux. 1

AT'fACHMENT A DATA OBTAINED IURING hEJTRON FIUX OSCILLATIONS A. Transient Recordings The flux noise tracec showing the neutron flux oscilla-tion are presented in Figures 5 through 8. For reference, the record taken during control rad oscillation Test 37 f.s shown in Figure 3 and the flux noise trace *.aken immediately after ' n Figure 4. Also, the record taken during control rod oscillation Test 49 with the second h4-bundle core is given in Figure 10 and the accompanying noise trace in Figure 9 The results of the control rod oscillation tests are discussed in Attachment B following. The observations from Figures 5 through 8 are cited: 1. Figure 5 shows the neutron flux oscillations as they were occurring on August 26. The oscillations are at approximately 1 3 eps and at an amplitude of about 6 percent peak-to-peak. Both in-core and out-of-core (picoammeter) traces are shown. The in-core traces are all in phase and show the same amplitude. This is characteristic of a driving force that is uniform over the core. Picoammeter No. 3 signal lags the in-core signals by approximately 1/4 cycle." 2. Figure 6 shows the picoammeter signals together with various flow and pressure signals. Note that Picoammeter No. 2 signal is a clean cyclic disturbance, whereas Piconmmeter No. 3 signal is cyclic, but shows evidence of less regularity. The in-core outlet flovmeter shows a jagged, noisy signal which is characteristic of a malfunctioning electronic readout system. It is not possible to identify a cyclic disturbance of the same period as the neutron flux oscillation in this signal, although it may have existed. The core pressure drop and the reactor vessel water level signals in Figure 6 show cyclic variations with periods of 2 seconds and 4-1/4 seconds, respectively. These variations were observed in bo+h signals during the Plant start-up tests and during earlier research and development tests. They are caused by the differential pressure measure-( ment systems (i.e., pressure pulsations excite the large bellows and the

This was caused by electronic filters in Brush Recorder Channel No. 7 which introduced the phase lag.

2 y 1, If , fluidLmass in the pressure lines, which respond by oscillating at the system natural frequency. Direct evidence of this phenomena is seen in the difference between the signal characteristics of recirculation flow traces in Figure 6 and Figure.7 In the-former case the normal. plant flow transmitters were valued out of service and special fast response transducers were used alone. In the latter case the normal plant. flow transmitters vere in service in parallel with the fast trans ducers. It is evident'in Figure 7 that the natural frequency of the plant flow transmitter and fluid line of approximately 1/2 cps is being sensed by, and is controlling the response of, the fast transducers). There is no evidence of cyclic variations in the recircu- .lation flow signals in Figure 6 or in the riser pressure drop signal with the period of the flux disturbance. In the case of the riser pressure drop signal, an oscillation at'l.3 cps may have existed but was masked by the higher frequency disturbance. 3 Figure T shows the response of the neutron flux when l control rod C-3_vas inserted one notch. During this reactivity step, the power dropped about. 3 Mwt, or about 2 percent. Modulation of the l amplitude of the neutron flux oscillation is evident in Figure 7 Of i interest is the way in which the in-core traces and Picoammeter No. 2 i traces intermodulate. As the amplitude of No. 2 picoammeter diminishes, } the amplitude of the in-core flux traces expand and vice versa. The in-core outlet flow is seen to decrease when the con-4 trol rod is inserted. This is due to a reduction in the instrumented f fuel channel exit quality. 4. The traces in Figure 8 vere made after the recircula- ~ tion flow had been reduced to approximately 80 percent of rated, or to about 26,000 gpm. Several significant observations may.be seen in Figure 8. First, the frequency of the oscillation decreased to 0 7 cps. 0ther data obtained during the flow and power decreases are summarized in the following Table II: .p-1. 1 1 a . ~. ..,. _ ~. _, _ _ _ _. _. _,, _. _...,

= 2 3 TABLE II 4 -Plant Power Recirculation Flow-Frequency (Mwt) (Kgpm) (Cps) 50 31.0 1 30 49 31.0 1.25 '47 31.0 1.20 .45 29 2 0 90 41 26.0 0 70 25 26.0 0.68 25 31.0 1.10 This information indicated that the frequency of neutron flux oscilla-tions was most etrongly influenced by the recirculation flow, with slight dependence on the power level. It was. impossible to detect any trends in the amplitude of the oscillations due to the intermodulation ~between in-core and picoammeter signals. The second item of significance in Figure 8 is that the in-core outlet flow showed eviden7e of oscillating, even to the extent of displaying the harmonic distortion of the flux traces. This first occurred when the recirculation flow was reduced to 80% of rated, at which time the downscale spikes seen in Figure 5 abruptly stopped and the flow oscillations were observable. B. Related Observations The following observations were made prior to the shutdown on August 26, 1964: 1. The turbidity of the water in the primary hydraulic loop, while remaini, quite low, increased significantly between August 2k and August.6. The readings were: August 24 0.1 APHA Turbidity Units August 25 0 38 APHA Turbidity Units August 26 0.6 APHA Turbidity Units 2. During the high power density run with the first 44-bundle core, readings of reactor power, recirculation flow and core t pressure drop were taken by the plant operators at the start of each ( shift. The data are shown in Figure 11. Note that one flow change .,a

k s. -3 was made during the' run on August 20, accompanied by a core pressure drop reduction. Also note that core pressure drop reductions occurred' on August 21', 23 and 25 for which no flow changes were made and the-absence of an' increase in' pressure drop on September 13, even though recirculation flow was increased. 2. t 4 h i s i

ATTACHMENT B CONTROL ROD OSCII1ATION DATA This Attachment B presents the traces obtained during the control rod oscillation and flux noise tests related to the high power density operation. A discussion is also included on the sen-sitivity of the reactor system to changes in various operating para-meters. 1. Table III shows the reactor operating conditions during the two control rod oscillation tests and during the neutron flux oscillations. 2. Figure 3 is a portion of control rod oscillation Test 37G (1.25 cps). This trace was chosen because it most nearly fits the conditions during the neutron flux oscillation. The truces of flux response are relatively smooth with little tendency of harmonic distor-tion and no emplitude intermodulation. The amplitude of Picoammeter No. 2 signal it, similar to the amplitude seen in Figure 5 The notch vorth of the control rod during Test 37 was approximately 9 cents, indicating that the reactivity worth of the control rod motion was approximately 17 lines / 29 lines times 9 cents, or about 5 cents. This is also the approximate worth of the driving reactivity during the neutron flux oscillation. The oscillation of control Rod B-3 causes the No. 7 string of in-core ion chambers to respond with a greater amp.11tude than Pico-ammeter No. 2. This type of response was seen in other control rod oscillation tests and is indicative of a driving reactivity localized in the core. Note that there is only very slight, if any, indication that the in-core flows are respondir3 to the neutron flux oscillation. 3 Figure 4 presents the cteady-state flux noise record taken after Test 37 The flux and flow noise records show a random Gaussian noise pattern, characteristic of BWR's. The amplitude of the oscillation is about 4 percent peak-to-peak. At this point, there is no hint of the 1 3 cps oscillation seen later. The high frequency (9 cps) oscillation in Brush Recorder Channels 6, 7 and 8 was caused 4 by a malfunctioning chopper in the recording system.

2 I TABLE III REACTOR CONDITIONS PJRING CONTROL ROD OSCILLATION TESTS Test Number 37 49 Type of Test Control Rod Neutron Flux Control Rod Oscillator Oscillation Oscillator Date and Time August 17,1300 Hr August 26,1300 Hr September 17, 2100 Hr Reactor Vessel Pressure, Psia lOh7 1037 1041 Drum Pressure, Psia 1035 1020 1029 Reactor Power, Mwt 153 158 150 Plant Power, bNe 50 51.0 47 Recirculation Flow M Lb/Hr 12.2 11.7 12.1 Inlet Subcooling 13 6 13 0 13 5 Btu /L1 Inlet Temperature, F 539 539 539 Core Average Exit Quality 0.045 0.052 0.045 Core Pressure Drop, Pcid 10.7 10 3 97 Core Average Heat Flux Btu /Hr-Ft2 180,000 186,000 177,000 Core Average Power Density, Kw/L 55 59 56 Void Coefficient of Reactivity, p/(% Void) 53 53 4.8

Control Rod Pattern 17, 10, 17 23, 15, 23 17, 0, 17 SteamFlow,MLb/Hr 0 587 0.610 0 570
Rod positions shown are for Control Rods B3, C3 and D3 All other rods at Notch 23 4.

The control rod oscillations shown in Figure 10 are for Test 49, obtained with the reconstituted h4-bundle high power density core. Note that the notch worth for this test was approximately twice the notch worth for Test 37 Otherwise, the same comments are appropriate for both tests. 5 Figure 9 is the noise recording accompanying Test 49 ( The flux noise amplitude is again about h percent peak-to-peak, with no

3 indication of the 13 cps oscillation seen in the previous core. It is quite evident that the vibration of the thermal shield did not occur during the second 44-bundle high power density core. In-core ion Chambers 7A and 2B show very slight evidence of a flux. However, the frequencies of the two signals are not the same, nor is the obviously organized flux oscillation present over the whole core. The core pressure drop during the second hh-bundle core was 9 7 psi, with a total recirculation flow of 12.1 million pounds per hour. This compares to 10 3 psi at 11.7 M lb/hr for the first core. Clearly, misalignment cf the thermal shield had occurred between the two cores causing a larger leakage flow path and resulting in a lower core pressure drop. It also appears that the misalignment prevented recurrence of the vibration of the thermal shield. 6. The sensitivity of the reactor system stability margins to changes in various operating parameters is presented in the discussion below: a. The sensitivity of BWR systems stability to changes in recirculation flows depends on the method of changing flow. If recirculation flow reductions are obtained by increased valving in the recircu-lation loops, the stability margins increase. If recirculation flow reductions are made, either by removing one or more recirculation loops from service or by reducing the driving head of the recirculation pump by speed reductions, the stability margins decrease. No calculations were made for the kh-bundle core in reduced flow modes of operation. However, for the 84-bundle core the following calculations obtained (at 1250 psia reactor pressure) are :* r 4

A more detailed discussion of the mechanism may be obtained in GEAP-3971, VBWR Stability Test Report, June 1963

- 4 1 ~( Reactor: Power Recirculation Flow' Stability Margins (Mwt) (M Lb/Hr) Gain (db) Phase (Degrees) 196 12 5-20.8 108 -(Two-Loop Forced Circulation)' 189 6.99 13 5 85.6- -(One-Loop Forced - Circulation) 5.69 fl0.6 55.6 -189 .-(TVo-Loop Natural Circulation) l b. During Phase I of the research and development testing at the Big Rock Point Nuclear Plant,. control rod' oscillation tests were performed at 800, 1050, and 1500 psia. Within this range of pressure, a slight increase in the calculated stability margins was indicated as the pressure increased. The results of the predictions are shown in the following table :** .i. Reactor Power Stability Margins-Reactor - Pressure (Psia) (Mwt) Gain (db) Phase (Degrees ) 800 157 16 87 1050 157 17 5 91 1500 157 20 87 c. The sensitivity of the Plant stability margins to calculated power level for the L4-bundle high power density core operating at 1050 psia, and - rated recirculation flow is shown below: Reactor Power-Stability Margins (Mwt) Gain (db) Phase (Degrees) 160 21 3 139 i 123 21.2 136 4 i - {-_

- These 'results are published in GEAP-4567, Preoperational Power Stability Analysis of the Consumers Big Rock Point Plant, J. M. Case, Feb 1964.

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i II. MAIN STEAM BYPASS VALVE MAUUNCTION ) i l

II. MAIN STEAM BYPASS VALVE MALFUNCTION A. Description of Events On September 18, 1964, in the midst or preparations to conduct the last of the planned transient tests with the 60 Kw/ liter, 44-bundle core, the main steam bypass valve suddenly opened from the existing 6% value to 100%. The valve opening was initiated by a spurious electronic signal, followed by an unsuccessful attempt at manual re-closure and the bypass valve remained 100% open. The turbine admission valves, responding to reduced steam-drum pressure, immediately closed. This was followed at 53 seconds with initiation of the main steam iso-lation valve closure which required 50 additional seconds to complete. The conditions of the reactor immediately preceding the transient were essentially as follows: 1. Reactor Th ermal Power, Mwt 155 2. Peactor Pressure, Psia 10h5 3 Feed-Water Controller Manual L. Turbine Admission Valves Governor 5 Bypass Valve Position - % Open 6 6. Picoammeters and In-Cores - % 100 Fast response traces of certain variables, in addition to normal Plant records, are available as a result of the development test preparations underway at the time the bypass valve malfunction occurred. In the fol-loving evaluation of the sequence of events, these records play an important interpretive role. In addition to the actual measured data, a transient model of the complete thermal hydraulic loop was written for the Philco 2000 computer. The responses of additional variables of interest are presented as derived from that analysis. B. Evaluatf on of Transient L. The bypass valve opening at time "zero" resulted in a sudden increese in steam flow from the reactor and a corresponding negative presst re rate (Figure 1 attached at end of Section II). Tnis was later compensated as the turbine admiscion valves close due to the subnormal steam -drum pressure. The admission valves were rapidly closed by the Initial Pressure Regulator which was in service but backed off t slightly so that the governor was controlling. (Figure 10) The ~ yy w

l 3 2 i-( immediate effect of the markedly increased loss of steam was a steam-drum pressure decrease causing a lower pump suction pressurr. and a rapid change in the recirculation flow rate. This is shown in,he total re-circulation flow rate decrease (Figure 1) and also in the increase in core exit quality (Figure 3). The latter indication, however, was dependent on core power and vessel pressure as well as the recirculation -flow variation. As the initial rapid negative pressure rate diminished in magnitude with closure of the turbine admission valves, the recircula-tion flow decrease was arrested and even slightly increased. 2. Power changes as depicted by both picoammeter and in-core flux traces (Figure 2) were initially due to the recirculation flow f. decrease coupled with some flashing in the core. After about 2 5 seconds j the increased steam voids associated with the reactor vessel pressure decrease continued to force the power down, while the recirculation flow j-rate vent back up, and thus reduced the rate of power decrease during the next 5 seconds. t 3 After about 7 5 seconds the small changes in the temperature of the recirculation flow initially leaving the steam drum vere just beginning to reach the core inlet (Figure 5). The variations in drrm outlet subcooling were caused by: a. A decrease in the saturation enthalpy in the drum due to drum pressure decrease and resultant drum flashing, and b. A slight increase in feed-water flow resulting from the reduction in back pressure at the feed pump. ) ~No abrupt changes were introduced due to either of these effects (the feed-water controller was on manual control). Feed-water temperature was constant over the interval of interest, due to a 31-se: ond flow time. from feed-watt - heaters to drum. h. The core inlet enthalpy began to see these small changes after the downcomer flow delay of 7 5 seconds, and the neutron flux then responded slightly with the increases shown on the in-core trace over the next four seconds (Figure 2). Pico No. 3 indicated a change of power from about 37% to hT% of initial value during this interval. At 11 5 seconds 4 I -,, ~ .___m._, .m,,_

- l 3 r. i the power indication from both Pico No. 3 and in-core instruments again t j dropped off. This was attributed to increased flashing in the reactor vessel. 'The-reactor vessel pressure decrease by this time was great enough to be felt in the reflect'or region, with the boundary of saturated water, CX, having reached below the core and hence beginning to flash (Figure h). With.this voiding in the reflector, the CICs began to see additional leakage from the core, hence indicating falsely. high neutron flux levels. This reflector flashing effect began after about5secondsbutdidnotsignificantlyaffectdensity,/),inthe reflector until about 8 or 9 seconds (Figure h). Because of the low reflector flow time of 96 seconds, the cooler recirculation flow water was not.'elt outside the core during this time interval. 5 After about 13 5 seconds the recirculation flow began to recover. The core exit quality (Figure 3), which is measured by means of an instrumented bundle exit flowmeter, decreased correspondingly. This strongly indicated that power did not increase with flow, and was i borne out by the low value of in-core flux in this interval. The flashing in the core due to the reactor vessel pressure rate was ap-parently overcoming the positive effect of flow increase on power. . This increased flashing was now felt very strongly in the reflector, such that, at about 17 seconds, the CICs (Pico No. 2 and No. 3) picked up a high flux scram and the power dropped abruptly. The signal gain effect, (), (Figure h) due to decreased attenuation from core to CIC as density decreased, by analysis reached approximately a factor of five. This scram apparently occurred when.the actual core power was no greater than 35% of the initial value. 6. The analytical values of core peak heat flux and core j - peak fuel center temperatures (Figure 6) were always less than initial value throughout the transient. The upper curves are based on the pico-ammeter trace with the falsely high -indication of the last several seconds. 7 The effect of thermal stresses in the reactor vessel, resulting from this transient, have been calculated to have " consumed" ' the following percentages of the total 40 year fatigue life at the four i most sensitive points in the reactor vessel. ~w-< e ~m, + p m.,

i h Percent Fatigue Life Consumed i Normal 40 Years This Transient Inside Surface-Main Flange 1.00% 0.022% Poison Nozzle 4.47% 0.04% Control Rod Thimble Weld (Including Scram Stress) 75 0% 0.067% Outside Surface of Inlet Nozzle 0 33% 0.023% The key stress points in the reactor vessel are more sensitive than those in the steam drum, pumps and piping and the expected effect on the fatigue life of these latter components would be less than those listed above. No sig' ificant effect on fuel bundles or core internals would be expected since metal thicknesses are small and parts cooled from both sides. All reactor vessel internals and a portion of the fuel have been thoroughly inspected, and all components, other than the thermal shield hold-down studs, are in good condition with no deleterious effects noted. Consequences of the above series of events cannot be considered severe in any of the areas discussed. Puel temperature and heat flux were better throughout the transient than at the initial condition prior to bypass valve malfunction. The pressure and temperature transient imposed only minor stress effects on the reactor vessel in terms of the design life of the components. It compares roughly to the impact of one scram. Hence, this incident was worth two scrams in terms of fatigue effects seen at the control rod thimble veld. In terms of power and thermal hydraulic transients, the scram action was extremely conservative. The action of the picoammeter scram function,resulting from the voiding of the reflector region, in-troduces this unnecessarily large element of conservatism. C. Description of Curves The curves illustrating the transient response of the important variables are of several varieties. Those variables of significance only until the scram occurred are shown for the first 18 seconds. Those variables of interest throughout the transient are shown for 120 seconds. The curves were obtained as follows :

Figure 1 - Steam flow from the reactor ouddenly increased, going off scale on the Brush recorder. The transient, until about 50 seconds when it returned to scale on the recorder, was recon-structed by analysis of critical steam flow through the bypass valve, coupled with the reduction in turbine steam flow as the turbine admission valves closed. Recirculation flow is based on direct fast recordings from Pace transducers in each of the two recirculation loops over the full transient. Steam-drum pressure was not calibrated on the recorder at the time of the transient. A complete transient loop analysis was used to generate this pressure transf nt, based on known measured transients of core exit quality and recirculation flow, plus the above reconstructed steam flow transient. Figure 2 - The in-care trace (lcyar sensor C501) is from a slow speed chart recorder (not having beer, calibrated on the fast recorder at the time of the transient). This in-core string was part of the instrument probe installed in an instrumented fuel bundle in core Position 05-55 The maximum and minimum values during the transient are readily defined from these traces. The time re-lationship of the points was reconstructed, making use of the fast response recordings of pico numbers two and three. Twd other in-cores, not shown nere, were located in the same axial strir.g and demonstrated t he same general time relationships and transient amplituder. The next three traces are taken from direct piccammeter record-ings. Pico number one is expanded from a slow speed recording, utilizing the time relationships from the fast recordings avail-able for Channels 2 and 3 The effect cf the transient voids in the reflector is evident in proportion to the attenuation L of the signal that is seen by the respective CICs (Figure 4). Figure 3 - The reactor vessel pressure transient is derived from the transient analytical model. The feed-water flow, with its controller on manual, is shown as it varied due to reduction of the discharge pressure, which is directly related to the decreasing steam-drum pressure.

- ~.- - 6 i Core exit quality is essentially a measured quantity based on a fast recording of the instrumented assembly exit flow rate. Quality is directly related to thfe measured outlet flow re-fereneta to the measured recirculation flow. As shown, it is plotted as core average exit quality, assuming that changes of quality from this instrumented channel are representatf.ve of the changes in the average quality. Figure h - The location of the reflector saturation boundary is represented by ()(and is derived from the transient analytical model. This isalsotrueofthereflectoraveragedensity,$1whichaccounts for the flashing occurring above the saturation or boiling boundary as it moves down the reflector. Two curves are given representing the. density variation in the reflector region associated with CICs one and two (located near core center line) and for the region a 1 associated with CIC number three (located several feet above core center line). The density associated with chamber number i three is lower because this CIC is located where some steam t i voids are interposed between chamber and core in the full power condition. . Based on these densities as a function of time, the approximate change in attenuation factor for each channel is shown as it progressed through the transient. This demonstrates the reason for deviation of the three pico signals as the reflector flashing progressed. Figure 5 - This figure shows more than the 20 seconds represented in earlier i traces, going out to 120 seconds. The steam flow and recircula-i tion flow curves are repeated. The curve of drum water levels, L, is taken from a standard p plant circular chart recording. The magnitudes are readily discerned from this record, and the fast transient recording of recirculation flow is used as a guide in interpreting the time relationships for this variable as taken from the chart. The core inlet and outlet temperatures are presented as expanded + curves from point recorder data. About 7 5 seconds delay in j change of'the inlet teaperature is due to the flow time of the i i coolant from the steam drum to the core. m ,..~...__-m

T t 4I *. Figure 6 - The curves of heat flux and fuel temperature are calculated directly from measured reactor power (or indicated power). The TIGER (1) code was used to analyze this case, giving an accurate transient evaluation. The upper curves of both heat flux and fuel temperature are based on pico number three, which gives a falsely high repre-4 sentation of the real conditions existing. It is of interest, however, that even under these assumptions, heat flux and fuel a temperature fail to return to their original steady state values. The lower curves'are representative of actual conditions exist-ing and effectively show how the two variables are generally decreasing throughout the transient. 1 Figure 7 - These are the traces frcm the three in-core ion chambers in the instrumented probe which was in core Position 05-55, a . central position in the kh-bundle core. Note - that the maximum and minimum values during the transient are readily distinguishable. Figure 8 - These curves are traces from the in-core thermocouples in the instrumented probe at core Position 05-55 These traces pro-vide very accurate information as to the actual rate of change of reactor coolanc during the transient. The curves are ex-1 i panded and shown on Figure 5 Figure 9 - This figure shows the actual traces from the Log-N-Period Channels No. 4 and No. 5 Note - that these two channels exhibited the same characteristics as the picoammeter channels. Period indication on Channels No. 4 and No. 5, as evidenced by the safety channel scram recorders, approached but did not reach 10 seconds. Figure 10 -This curve shows the electrical output of the turbine-generator unit. Note - the very rapid drop in load as the Initial Pressure Regulator sensed the dropping steam pressure and closed the i turbine admission valves. D. Power Level, Power ' Sensors, and Reflector Density I'. Reactor power level indication is provided by five power t range instrument channels at Big Rock Point.. In each of these power (1}KAPL-2044, " TIGER II,' An IBM-704 Digital Computez Program: Temperatures From Internal Generation Rates," A. P. Bray and S. J. McCracken, May 1959 7 i +.. .m-m ._,._,_-,_,__-_,.m. ~ ,.m,

8 channels, the output of a gamma-compensated ion chamber (CIC) is fed into 4 a fast response flux amplifier. The CIC detector is placed near the in-side boundary of the concrete biological shield and responds to neutrons leaking from the reactor vessel, which is proportional to the core power level. Two of these five instrument channels are displayed on a loEarithmic scale with separate period indication, and three are fed into linear ampli-fiers: the "picos." The fission ent gy neutronc leaking from the core which eventually vill provide the power indication at the detector are attenuated both through the water surrounding the core and the reactor vessel ar d other structural material, and finally are thermalized in the material sur'ounding. the detector (and in the material of the detector to a smaller de62ee). This amount of attenuation from the core edge to the detector location is primarily due to the water surrounding the core. The density of this sur-rounding water is important, and small differences of density have marked effects on instrument channel reading as related to the actual power level existing in the core. During the initial operation of Big Rock Point, as in all APED reactors, the amount of attenuation due to water surrounding the core is measured. Therefore, increases of reactor temperature and corresponding water temperature increase and density decrease can be accounted for in the calibration of reactor poser as a function of chamber reading at all temperatures from cold to operating conditions. Initial heat balances are made considering the heat capacity of the reactor system, and final calibration of the system is performed with operating plant heat balances. Throughout the process, conservative limits of reactor power are applied until the final power heat balance has been performed. In a pressure loss transient, such as was observed recently, an instrument system error is introduced in a definitely conservative fashion. Neutron attenuation from the core edge to the chamber location is reduced, due to boiling in the water surrounding the core. Figure 4 shows this increased gain or reduced attenuation and the density changes. The attenuation due to water from the core edge to the chamber location has been estimated to be (for the 44-bundle core) in the range of 3 h 3 x 10 to 171o. depending on the chamber in question. The 44-bundle ( core, which is 8 fuel bundles vide in the NW-SE direction and 6 fuel

9 T bundles vide in the NE-SW direction, was loaded off center in the vessel, contributing two degrees of asymmetry. . Using the following approximate equation, the attenuation due to water density may be estimated: -X' + Chamber Reading = K {} P e f K is a proportionality constant of chamber and amplifier response versus flux; P is the flux of bundles at the edge of the core in g the direction of the chamber; and x is the water distance from t g core edge to chamber. The quantity )( ig the JL folding length of neutrons in water, typically 7 cm at 550 F and is inversely proportional to water density. 1 Due to the core asymmetrics, the water attenuation distances are significantly different for the three chambers. The offect of these differences was plotted as changes in gain with time, following the bypass valve opening (Figure 4). Channel 3 was located above the other two channels such that core-to-chamber neutron paths included some steam voids from the region above the reflector. Also the two neutron vindows located between the core and Channel 3 influenced its response. Hence, its initial density was slightly lower than for Chambers 1 and 2, which were at or S near the center line of the core (Figure 4). Gain effects are first 4 i seen by. Channel 3 at the boiling boundary moves down and boiling occurs i j at the top of the reflector. Apparent increase in power, thus indicated b3 the instruments, can occur without any real increase in actual power 1evel. Due to the fact that the in-core monitors are fission chambers, they are not susceptible to changes in gain er attenuation due to density variations. Once calibrated to rated power conditions by a heat balance, thr in-core monitors indicate reactor power for all - control rod configurt.tions and core power distributions which do not deviate markedly from the calibration conditions. Exaraination of Figures 1 and 3 show that the core exit quality was inveruely related to recirculation flow rate. This can occur' only if the povet response to flow and subcooling is compensated by a corresponding negative power response to flashing. Hence, the core voids were well behaved and did not introduce marked changes in power distribu-tion during the transient. s a ,+m... y. --e. ,-e -,.e-,.,,.-,-._-w-,v.-.,-_..-,,e

10 Since the in-core monitors did not see core power distri-i butions which differed radically from their calibration conditions, they gave a relatively accu / ate representation of actual core power throughout the transient. Moreover, all three records of in-cores located in an axial string show the same proportional variations during the transient, further supporting the consistency of the in-cores as a true representation of power (Figure 7). The transient analytical model by which reflector density, location of the saturation boundary, drum pressure, and reactor vessel pressure were calculated was written with three two-phase node 6: reactor vessel, risers, and steam drum. As input to the model, the actual measured transients of recirculation flow rate and core exit quality and steam flow were introduced. The transient of steam flow was based on measured values when these were on scale, and calculated from a steam blow-down model for the remaining portions. The two sources were found to be ccmpletely compatible. The other transient input was feed-water flow rate, which increases as shown in Figure 3 Feed-water temperature was assumed con-stant over the 18 second transient calculation. All tra two-phase nodes (reactor vessel dome, risers, and steam drum) are solved simultaneously by transiently considering the mass, volume, and energy equations of the individual nodes. Two-phase flow through the riser and through the steam separators into the steam drum is based on the Martinelli-Nelson correlation. Transient pressures in the reactor vessel, risers, and steam drum are calculated, as are the inlet and outlet flows of steam and water with respect to each of these nodes. Since recirculation flow rates are input from measured data, the single-phase portion of the loop was not calculated. Similarly, the core was not analyzed since the upper instrumented assembly flowmeter provides a direct measure of exit quality from the core, which would be the objective of the core model if used. Because of the interest in reflector density, this part of the loop was also evaluated. Again the transient mass, volume, and energy equations are solved to determine the void fraction and density as the saturation boundary moves down with pressure. Since the flow time through this region is 96 seconds, it was assumed that the temperature in the subcooled reflector was unchanged throughout the transient.

11 The transient heat flux and fuel center temperature were calculated by means of the transient code TIGER referenced earlier. Standard design criteria on thermal conductivity and heat transfer characteristics are assumed. Input to the code was in the form of measured power. Two cases were evaluated, one corresponding to indicated power as given by pico number three, and another corresponding to actual cLre power as given by the in-core chamber. E. Power-Subcooling Effects The effect of subcooling on thermal power during a transient of the sort experienced during the bypass valve malfunction has been evaluated. Of significance to this study are the transport times in the recirculation and feed-water loops: Loop Segment Transport Time - Seconds Feed-Water Heater to Drum 31.0 Drum to Vessel Inlet 75 Vessel Inlet to Reflector 36.0 Vessel Inlet Through Core 0.7 Through Reflector 96.0 Feed-water temperature changes are not felt in the steam drum until 31 seconds after the closure of the turbine admiscion valves (which also takes about h seconds). This cooler water is then mixed with the recirculation flow in the steam drum and transported to the core, requiring an additional 8 seconds. Hence, more than 40 seconds will have elapsed prior to any power response due to lowered feed-water temperature. Since the scram occurred at 17 seconds, this effect obviously played no role in the transient. Feed-water flow, on the other hand, increases gradually with the steam-drum pressure drop as illustrated in Figure 3 The effect is to additionally dilute the recirculation flow with more and more cool feedwater. For the transient experienced here, the contribution of this feed-water flow transient to core subcooling is less than 0 3 Btu per pound per second. At rated power for the core being considered, a change of this magnitude corresponds to less than one cent per second reactivity increase. Ignoring the compensating effect of voids due to pressure decrease, this reactivity rate would add approximately one percent power + ..e-y v

12 f per second after the cooler water reaches the core. Hence, in this transient, the contribution of subcooling to power is less than ten percent by the time the scram occurred. It should be poi:.ted out that the assumption of rated power and flow makes this analysis very con-servatively large. Another contribution to core subcooling, also delayed by the flow time from steam drum to core, is the gradual reduction in re-circulation flow enthalpy as the drum saturation enthalpy drops with pressure. The composite of these effects is shown by the core inlet temperature trace on Figure 5 The total effect of inlet subcooling on power is suggested in the power traces of Figure 2 at about 8 seconds (when this change first reaches the core). Somewhat masking this is the effect of flashing due to pressure decrease, and the dominating effect of power responding to the recirculation flow changes. F. Conclusions A thorough evaluation of all the data available, including post-transient analyses, lead to the following conclusions : 1. Actual core thermal po'.er, heat flux, and fuel center temperature decreased tnroughout the transient, with no succeeding maxima ever reachin6 the initial rated conditions. 2. The rapid pressure drop after the scram, and the sc-companying thermal transient, are relatively insignificant in terms of total fatigue life of the various structural components. 3 The effect of subcooling daring this transient was present,as expected, but contributed only minor perturbations to tne reactor power. h. The effects of reflector density give rise to indicated power level on ClCs which is substantially higher than real power in the core, and thereby very conservatively protects against high flux conditions during any transient during which voids are produced in the reflector. 5 No structural damage to reactor system components resulted from conditions encountered during this transient. In addition to minimum stress effects of the pressure induced thermal transient, the ( core and riser pressure drops were well behaved and generally were lower than initial values throughout the transient.

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The results of this analysis, therefore, confirm that all aspects were of minor consequence with regard to safety and mechanical integrity of the system. G. Main Steam Lypass Failure - Cause and Corrective Actions Just prior to the scram on Septeiber 18, 1964, the Plant was involved in running a series of tests designed to verify that the neutron flux oscillations (noted in the previous run of the kh-bundle core - 60 Kw/ liter) were not due to hydrodynamic instability. These tests, utilizing the control rod oscillator equipment and the Annin Turbine Bypass Valve,had been completed at approximately 115 Mwt. The control rod oscillator tests at 150 Mwt had been successfully completed and preparation was underway for testing at this power level utilizing the bypass valve. The Plant was at power operation (49 Mwe) with the kh-bundle core. The bypass valve was set at 5% open and connections were being made from a low frequency test oscillator to the bypass valve electronic servo amplifier. This would permit sinusoidal oscillation of the bypass valve-and evaluation of the void-to-reactivity feed-back mechanism. The test oscillator chassis was floating a few volts above ground potential and the signal lead was inadvertently connected to the servo amplifier before the unit was grounded. This simulated an opening signal to the bypass valve and the bypass valve opened 100%. The oscillator signal lead was immediately removed, but the valve did not respond and close (indicating a failure in a' component of the control system). Subsequent investigation indicates that the following sequence of events occurred: 1. The signal, imposed by connection of the test lead, quite likely initiated an opening signal to the valve. 2. The valve did not c?ose when the lead was removed due to sticking of the spool piece in the MOOG servo valve (verified by examination performed by the manufacturer). The malfunction of the MOOG valve prevented closure of the valve when the signal leads were removed. 3 The failure of the bypass valve to close when the 4 manual override switch was utilized was due to a second component failure (the flow-control valve installed downstream from the bypass valve power cylinder to limit the closing speed of the bypass valve) (Figure 11). -m-4 e r----

14 ( Examination of the valve after the incident disclosed a mechanical failure of the valve due to excessive back-pressure spikes from the drain line, which also drains other components. It was demonstrated that the piston of this valve could be displaced to a position which prevented flow through the valve, thus preventing closure of the bypass valve. We have concluded that the combination of the two separate component failures (outlined above) were responsible for the pressure blowdown experienced on September 18. The changes and improvements being implemented to prevent a recurrence of this event are : 1. A replacement flow-control valve has been purchased and installed. With the correction of the high back-pressure pulses on the valve (as noted on the next item), this valve should perform satisfactorily. However, a mechanical stop has been installed to that complete shut off of the valve is impossible. Also, the availability of other tyree of flow-control valves for this application is under investigation. 2. The drain line from the flow-control valve has been i enlarged, and tests have demonstrated a marked improvement in the back-pressure spikes experienced previously. 3 The rate of depressurization of the primary system dur-ing the transient was due in part to the additional steam flev capacity of the bypass valve above that specified. The power cylinder of the valve has been replaced with a new cylinder which reduces the stroke of the valve such that it will pass approximately 1 x 10 pounds / hour at 1450 psig. 4. The hydraulic oil in the system has been tested and found to contain metallic particles in the 10-100 micron size. Although this should not interfere with operation, the oil has been replaced with new oil filtered through a 3 micron clean-up system. The metallic particles in the oil are felt to be due to the wearing in of the new pumps and equipment. The oil will be sampled frequently to evaluate future trends. 5 The M0GG servo valves have been thoroughly tested and, in addition to the servo valve that failed, a second servo valve showed indications of a sticking spool piece. Both valves have been returned to the manufacturer for reconditioning.

15 6. A long-term test program is in progress and vill continue through the present outage. Prescribed teste are conducted each shift to check the electronic control system, the hydraulic system and tha nAnual closure function. In general, ope _ tion has been very smooth. 7 The decision has been made to replace the entire electronic contro) system. Specifications calling for a highly reliable solid-state system have been written and will be released to major con-trol manufacturers shortly for bids. The feasibility of utilizing coincidence circuitry is also being investigated. ( )

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A 4 ,+ a W i 5 j J III. CONTROL ROD DRIVE fRLFUNCTION s t e I ( l l .]

A a ~ - III. CONTROL ROD-DRIVES AND CONTROL RODS ~ A. Control Rod Drive' Malfunctions There have been two instances where galling between the i .index tube and the upper guide sleeve of control rod drives has been experienced. The first event occurred on February 20, 1963, and vas . attributed to foreign material ~ in the drive and/or reactor vcasel that I had not been flushed out during preoperational cleaning of the system. 4 The second event occurred on September 18,1964 when a two. drives showed signs of galling during coupling integrity checks 'immediately following'the scram due to the b, pass valve malfunct' ion. b At the time of this scram, the rate of pressure decrease was -causing ]- significant flashing of water in the reactor veshc] and it is not j unreasonable to-postulate that this affected the. flow paths such as to put any foreign material present in the reactor vessel into sus-pension. Some of~this material could then have been drawn into the drive screen crea during the scram, and particles bypassing the screen 1 could have initiated the galling. While this is a possible cause, we now tend to discount 1 it due to the complete absence of any detectable particles in the screen of either ' drive involved, or for that matter in the screen of the third drive or other drives recently inspected. From all observations, the 3 Big Rock Point reactor vessel and connecting primary system are clean with very low turbidity and crud levels. 1 J The crud in the reactor vessel, as in all boiling water j reactors, is very finely divided particles of soft oxides of copper, l iron, nickel, etc. There is no evidence that this crud has any delete-7.- rious effect upcn the operation of any movable parts inside the primary system. This is borne out by many years of operating experience in BWR's. A second and more plausible cause of the galling centers around water flashing during the pressure transient. It is quite con-1. i-ceivable that steam may have been drawn into the upper guide sleeve l during the scram and resulted in a lack of lubrication,-thus initiating ) the galling, j. / In any event, the onset of galling is evidenced by failure l of the drive to jog, and by erratic movement of the drive during normal i. 1 e w y e,,,. me,-w y y e----

2 I insertion and withdrawal. Since the progressive galling produces metal chips which can in turn interfere w!th proper collet action, movement of a drive suspected of galling will be limited to that necessary to deter-mine if galling exists. Should it be desirable to continue operation for a pericd of time with a known galled control rod, the procedures specified in Section 5 3 2 of the Technical Specifications will be followed. Al-ternate control rod patterns will be developed and used if this proves to be desirable in terms of permissible power level. B. New Control Rods The new control rods to be installed at Big Rock Point are described in a separate submittal to DRL requesting a slight change to the Technical Specifications to accommodate the new control rods. These control rods will utilize the same pin und roller concept, except 'ers will be slightly smaller in diameter and the that the stellite pins longer. The rollers are present only to reduce long-term wear on control rod blade sheaths and support-tube-and-channel assembly surfaces. The absence of several rollers would not be expected to significantly increase the movement of the rods. There is no evidence of significant " wobble" or " vibration" of the control rods, with or without rollers missing. There is also no evidence that vibration was a significant contributor to the pin wear experienced on the present blades.

i IV. THERMAL SHIELD HOLD-DOWN STUD FAILURES k 9

f l' IV. THERMAL SHIELD HOLD-DOWN S' IUD FAILURES A. Sumr.ry On October 11, 1964, operating personnel, in the process of inspecting reactor core components, detected 2 broken thermal shield' h'old-down studs. Subsequent investigations showed that k other studs + were also broken. The 6 broken studs and the 6 unbroken studs along with the 3 unbroken guide pins, were sent to General Electric- (Vallecitos) for physical and metallographic examinations. These examinations re-vealed that the studs suffered a fatigue type failure due to excessive movement of the thermal shield resulting from wear on the alignment pins. Detailed analyses readily associated the earlier detected neutron flux oscillations of the h4-bundle core with these stud failures. The proposed design solution to eliminate the problem is currently in process. B. Original Design The thermal shield is a large cylinder located in the. annular region between the reactor core and the reactor vessel. (See Figure 4 3A, Section h, FHSR) Functionally, it separates the inlet plenum and outlet plenum of the reactor vessel. The thermal shield rests on six brackets welded to its inside surface near the lower por-tion. The brackets, located at 60 degree stations around the thermal shield's circumference, rest on and are bolted to six supporting pads which are welded to the reactor vessel. Two one-inch diameter studs, approximately three feet long combined with sleeves hold each of the six brackets of the thermal shield to its mating pad. (See Figure 1) All six reactor vessel pads are tapped for two 1-inch diameter 8N studs. In addition, the reactor vessel pads at 60,180 and 100 were bored ftr a zero clearance fit with 1-1/2 inch diameter locating pins. The thermal shield brackets contain cleararce holes for the studs, which are machined elongated 3/8" in the radial direction to accommodate differential expansion of the thermal shield and reactor vessel. The pinholes in the thermal shield brackets are also machined elongated 3/8" in the radial direction but are close fitting on the locating pins in a circumferential direction. The studs holding the reactor vessel pads and thermal f-shield brackets together are 38" long and were loaded against sleeves m.- m 4-

4 2 n. .i of1-1/2inchschedule80 pipe. The sleeves are set in 3/16" deep 1 ~ counterbore h61es around the top of the elongated thermal shield bracket stud holes.' The three, equally spaced locating pins, are fixed in the reactor vessel pads and engaged in the elongated holes in the thermal shield. bracket permitting the differential expansion of the shield and reactor vessel but constraining the shield in the reactor vessel against lateral motion or rotation. In assembly, the studs were threaded into the reactor vessel pads approximately 1-3/4 inches, 8.1 the sleeve and nut were then added to the assembly. The nuts were tightened to a 500 inch-pound torque to limit the preload so that the upward force of the pressure differential across the thermal shield -vould unload the pads and facilitate the differential expansion. The long studs accommodate differential expansion by bending as the top of the stud moves with the sleeve and thermal shield while the bottom is fixed in the reactor vessel pad. With a 10 psi core differential pressure across the thermal shield, all grid openings plugged, and without fuel, the net upward lead on the bolts would be 86,60^ pounds which would produce a tensile stress of 13,000 psi at the stud thread root. With a h4-bundle core installed and operating at a 10 psi differential pressure, the net upward force was 54,000 pounds and the tensile stress at stud thread root was 8,200 psi. The bending stress in the stud due to the 1/16-inch radial differential expansion produces a stress of 6,000 psi. The annulus between the thermal shield and the reactor vessel is fitted with a seal ring constructed like a Belleville washer. This ring is sprung between the bottom of the thermal shield and a continuous protrusion or ledge around the inside circumference of the vessel to restrict core bypass leakage in this area. A differential pressure of 3 to 6 psi should unseat this ring and result in increased leakage. C. Analysis of Stud Failures The original design of the thermal shield hold-down was conceived to provide flexibility for differential thermal expansion between the thermal shield and the reactor vessel. Under the axial } loading of the pressure difference between the inlet plenum and outlet plenum, an elongation of the studs and a compression of the sleeves 1 . - ~ _ _, ~

3 i occurred. The original analysis indicated that, for a pressure dif-ierential of 3 to 4 pei, the loading.on 6 stud-and-sleeve pairs would be just sufficient to-cause the brackets to lift from their mating pads. With'a pressure differential of approximately 10 prei, such as experienced with the.'kh-bundle core, the brackets lifted approxi-mately 15 mils away from their mating pads. Under this configuration, a side-to-side mode of vibration of the thermal shield occurred. The I side-to-side mode had a calculated natural frequency of approximately 1.2 cps. The restraint on the lateral movement was a pin fastened to each of the pads between the studs which protruded up inside a slot on each of the brackets of the thermal shield. During the vibration of the thermal shield, the hold-down studs were bent as cantilever beams and completely reversin6 bending stresses were set up within the studs. Ultimately, 6 of the 12 studs failed. Failure of one or more of the hold-down studs resulted. in asymmetric loading of the remaining studs. In particular, the studs adjacent to the unsupported span of the thermal shield were put under the most severe axial load. The analysis indicated that, just prior to shutdown on August 26, the 2 studs adjacent to the 6 failed. studs were deflecting a maximum of 90 mils. Thus, the thermal shield was being lifted at its highest point approximately 180 mils. The effect of the asymmetric loading on the vibration of the thermal shield was to intro-duce a rotational degree of freedom about a horizontal axis. The natural frequency of the rotational mode of vibration was approximately 1 5 cycles per cecond. Both the side-to-side and the rotational modes of vibra-tion demonstrate calculated natural frequencies in the range of the freq.uency of the neutron flux oscillations. In Section I, Neutron Flux Oscillations, it was shewn that vibration amplitudes in the order of 1/4 of an inch could have caused variations of bypass flow of sufficient amplitude to cause the flux oscillations. The actual vibrational ampli-tudes were believed to be approximately 1/4 inch, based on the wear ~ observed on the hold-down studs after they were removed from the vessel. The key to understanding why the studs failed lies in i understanding how clearances developed in the guide pin bracket hole i .+.

t 9 b contact. During the early operation of this reactor (up to June 1963), excessive ~ vibration occurred in the support-tube-and-channel assemblies due to action of the recirculation water inlet flow. These vibrations were. subsequently reduced by the introduction of a flow diffuser to divert the inlet flow. Ho-ever, this early turbulence may have intro-duced the original oscillation modes to the thermal shield. The guide pins of the thermal shield retainer had, initially, a'Lero clearance fit. The pin contacted the elongated hole in the thermal shield bracket in a very small area (where it is tangent to the pin) on both sides of the pin. With relative motion, the resist-ance of.this small contact area of the pin to abrasive wear is very ~ slight so that continued oscillation of the thermal shield produced accelerated wear as more clearance built up. Elongated Pi Slot ~ l 4 } j Contact Area The wearing of the guide pin would take up the clearance between the studs, which are 1" diameter, in a 1-1/8" vide elongated thermal shield bracket hole. Thus, the added freedom of the guide pin allowed the studs to begin contacting the elongated holes of the thermal shield bracket. This can be predicted by totaling the clearances involved, the wear on the pins and stude and the wear of the elongated bracket holes. The sums of these measurements indicate that they were approximately equal for the pin and for the stud. As the wear on.the guide pins and studs increased, the allowable movement for the thermal shield increased. Low cycle fatigue has been indicated to be the cause of the stud failt es. Crud build-up i on the broken portions of the studs indicates that the breaks developed over a period of time that could run from several weeks to the total operating time of the Plant. ( The operation of the h4-bundle core with a 10 psi core pressure drop probably was sufficient to cause the completion of the stud failures. However, this was not the primary cause. 1 - I

._~ L D 5 l-D. Physical-Examinations Subsequently, the 12 hold-down studs and 3 guide pins ' were removed from the reactor vessel. The 6 failed studs were located ~ ' in the east ' sector of the reactor vessel. (0 - 180 from north - Figure 2). Physical examination, of the stude and guide pins, revealed flattened surfaces near the bottom of the pins a.a studs where' contact with the thermal shield bracket was made. l The amount of wear for each flattened surface of the pins and studs is shown in Table I. An attempt was made to correlate the flattened surfaces of the studs with their original positions in .the reactor vessel. However,'because.the shaft of the studs was covered' by a sleeve, it was very difficult after removing the studs to orient them as they had been during operation. The orientation of the flat areas can be' inferred from the wear on the elongated holes of the ther-mal shield bracket. This wear has been estimated from binocular obser-j vation and by measuring impressions made of these areas. The guide pins were pushed out of their respective loca- - tions, and the orientation of the wear areas was maintained by marking the top plate of the retainer assembly which was welded to the top of the pin. The. locations of the wear areas are in agreement with the pre ent relative location of the holes in the reactor vessel pad to b the holes in the thermal shield as indicated by impressions made at the bottom of the elongated holes. i Observation and measurement of the wear areas on the 1 guide pins and studs indicate an average of 0.090-inch total wear on the guide pins and an average of 0.063-inch total wear on the studs. 1 The diameter of the thermal shield was remeasured with all the guide pins in pl' e as well as after two pins had been removed. - Both measurements. indicated that the thermal shield has maintained its circular shape with no observsbie distortion. E. Metallurgical Analyses

1..The overall appearance of all studs (see Figures 4 and 5) was essentially the same whether they had broken or not. No

~ .= unusual corrosion was noted. All were reasonably straight with a .(; maximum bow of a few mils detected by comparison with a straight edge. 'l

o 6 Two wear areas were noted on the bottom end of each stud just above the threaded area and in some cases included a portion of the threads. Their height corresponded to the bracket thickness. The two wear surfaces for each bolt were *.ot necessarily of the same depth, ranging from 0.01 to 0.06 inch. (See Table I) The studs failed in the area between the top of the reactor vessel pad and the f1 st thread of the lower thread area with the root of the thread defining the plane of the break. 2. The failed ends are essentially in one plane with a series of arc-shaped ridges progressing across the failed surface at right angles to the planes of the worn areas. The broken surfaces range in color from light tan to a bright blue. The tan areas are near the circumference on one side and the b]ne areas are opposite. Cracks were noted in the root areas of several of the failed and unfailed studs. One stud showed a crack in the cap threaded area; other studs had cracks in the lower end. These cracks,7re also at right angles to the wear surface. Metallographic examination of a crack just below a broken surface shows it progressed essentially transgranularly. (See Figure 6) The main break also appears to be transgranular. Microhardness measurements show the material is in the annealed condition. Considerable tvinihg is noted in the grain structure. 3 The macroviews of the fracture surface point definitely to a fatigue-type failure. This type of fracture surface is one of the most familiar known to the metallurgical profession, and there is no other type failure which produces such a characteristic pattern. Usually, this type of failure develops over a period of time. The color shading of the fracture surface indicates that it developed over a suf-ficiently long period to havt several thicknesses of oxide. The arc-shaped edges are rather broadly spaced indicating that the fatigue failure was of the low to moderate cycle variety oc-curring in hundreds or thousands of cycles rather than the million cycle range. Two factors can contribute to a low cycle fatigue be-havior: high stresses and the water environment. The amount of twining indicates high stress levels although the lack of grain dis-( tortion indicates no excessive yielding. It has been shown that a

e T 'r.. 4 L{- water environment is more severe than an air environment for the propagation of cracks. The lack of grain distortion in the crack in-dicates that the crack propagation was cased by the presence of the 4 water. ~ F. Proposed Modifications to Thermal Shield Hold-Down Methods of retaining.the thermal shield under consi-deration at this time are referred to as the stilt and the yoke arrangements. 1. The stilt-type arrangement provides a leg which attaches to the inner side of the thermal shield and ties securely to the reactor vessel pad.- This will allow a flexible movement in the radial direction.to provide for thermal expansion but deters lateral movement. This design requires removal of the two gussets and the thermal shield bracket at each of the 6 locations. The bottom plate of the stilt will be bolted to the reactor vessel pad; and the top leg, which is a column construction witha rectangular cross section, will be bolted to the inside of the therrT' shiald. Exact dimensions of these pieces are not available at this time since the analysis and I design are currently in progress. This is the preferred fix at this time and will be followed unless installation requirements prove it to be impracticable. 2. The yoke-type arrangement, illustrated by Figure 3, consists of a stirrup device which will fit under tne reactor vessel [ pad, a 2-inch bolt and a clamping bar at the top of the thermal shield. The yoke is made of reactor grade 304 stainless steel and is self-aligning. T.e stirrup portion of the yoke device is fabricated from 1" plate (also 304 SS) which will be attached to the underside of the reactor ressel pad. Two retaining bars are located on the bottom of the stirrup to provide a self-aligning action on the single gusset 4 under the reactor vessel pad and the slanted rear side of the pad. A threaded ~ nut, which is held in place by a welded retainer, is used to attach the stirrup to the 2-inch bolt. The nut and its associated washer have spherical surfaces to aid in aligning the rod with its stirrup. The 2-inch bolt links through the clamping bar, which fits - over an are of the upper rim on the thermal shield. A hexagonal nut ,--p. ,,e --,.---ye t' "T-7+-+- Y -+ ' * --v-'*

C'+

.~ ~. 8 t and a spherical washer retain the clamling bar. The nut is prevented from loosening by means of a locking s eeve which is tack-welded to l 'the top of.the clamping ' car. An initial preload is applied by the use of a stud tensioner. The installation is planned at.the 0,120 and 240 f locations. TABLE I STUD AND GUIDE PIN WEAR i J St'ud Wear Areas (Inches) Stud Surface No. 1 Surface No. 2 Total

1 0.034 0.043 0.077 Failed 2

0.037 0.020 0.057 Failed 3 0.040 0.034 0.074 Failed' 4 4 0.025 0.035 0.060 Failed-i 5 0.031 0.003 0.034 Failed 6 0.020 0.017 0.037 Failed 7 0.020 0.047 0.067 cracked 8 0.027 .0.032 0.059 Okeh 9 0.051' O.020 0.071 Okeh 10 0.058 0.005-0.063 Okeh 11 0.061 0.010 0.071 Okeh 12 0.057 0.030 0.087 crachea Guide Pin Wear (Inches) j 60 0.040 0.045 0.085 Okeh 180 -0.050 0.040 0.090 0Feh 300 0.015 0.075 0.090 Okeh 2 l }

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ML20030A511

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ML20030A511

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