ML19322A756
| ML19322A756 | |
| Person / Time | |
|---|---|
| Site: | Oconee  |
| Issue date: | 04/01/1967 |
| From: | DUKE POWER CO. |
| To: | |
| References | |
| NUDOCS 7911210779 | |
| Download: ML19322A756 (600) | |
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,-' DUKE POWER COAIPANY A AIEND.\iENT NO.1 TO APPLICATION FOR LICENSES DOCKET NOS. 50-209 50-270 Additional information submitted in response to selected questions contained in l letter of Dr. Peter A. h! orris, Director, Division of Reactor Licensing, to Applicant dated N11rch 23,1967.
'Ihe answers to the following questions are submitted as Supplemenc 1, dated April 1, 1967:
1.1 3. 4 8.1
- 1. 2 3. 6 3. 3
- 1. 3 4. 2 8. 4
- 1. 5 4.4 8. 5
- 1. 0 4. 5 S. 6
- 1. 7 4. 0 9.1 2.1 4. 8 9. 2
- 2. 2 7.1 9. 3
- 2. 3 7. 2 9. 4
- s. 2. 4 7. 3 9. 5
- 2. 0 7. 4 9. 0
- 2. / /.6 9. 7
- 10. 1
.0.0 0 0' M 5 e
'N Dockets 50-269 and -270 s_,) Supplement No. 1 April 1, 1967 QUESTION Supply specific ASME Code vessel classifications for all components, 1.1 including heat exchangers, in the systems which handle reactor coolant.
ANSWER Listed below is the Code classification for those components in the auxiliary systems which fall under the jurisdiction of the ASME Code. The pressure containing parts of those components serving an engineered safeguards function will be radiographed and all welds will be of tha full penetration design.
- i. High Pressure Injection and Purification System
- a. Letdown Cooler - ASME Section III-C
- b. Seal Return Cooler - ASME Section III-C
- c. Purification Demineralizer - ASME Section III-C
- d. Letdown Storage Tank - ASME Section III-C
- 2. Chemical Addition and Sampling System
- a. Reactor Coolant Sample Cooler - ASME Section III-C
- 3. Low Pressure Injection and Decay Heat Removal System
- a. Decay Heat Removal Cooler - ASME Section III-C
[--) b. Borated Water Storage Tank - ASME Section VIII \- / c. Core Flooding Tanks - ASME Section III-C
- 4. Spent Fuel Cooling System
- a. Spent Fuel Cooler - ASME Section III-C
- b. Spent Fuel Coolant Demineralizer - ASME Section III-C
- 5. Waste Disposal System
- a. Quench Tank - ASME Section III-C
- b. Deborating Demineralizer - ASME Section III-C
- c. Reactor Coolant Bleed Holdup Tank - ASME Section III-C
- d. Miscellaneous Waste Holdup Tank - ASME Section III-C
- e. Spent Resin Storage Tank - ASME Section III-C
- f. Waste Neutralization Tank - ASME Section III-C
- g. Waste Gas Decay Tank - ASME Section III-C h, E"aporator Condenser - ASME Section III-C
- i. Waste Evaporator - ASME Section III-C
- j. Evaporator Condensate Demineralizer - ASME Section III-C
- 6. Reactor Coolant System See Table 4-9, page 4-33, o f the Preliminary Safety Analysis Report.
Those components not covered by the ASME Code, such as valves and piping will be designed and fabricated to meet the requirements of ASA B 16.5 or MSS SP 66 specifications and ASA B 31.1, respectively. 7~ (( 1.1-1 (4- 1-67 ) g] 356
\
1
Reactor coolant piping is designed to ASA B 31.1 and inspected to re-quirements of Section III, Class A. Pressure containing castings will be radiographed to meet ASTM E-71. The pressure containing parts of all pumps will be liquid penetrant tested in accordance with Appendix VIII of Section VIII of the ASME Code. In addition, all pressure containing welds of all engineered safeguards pumps will be fully radiographed. In summary, those com-ponents not covered by the ASME Code will be fabricated and inspected to requirements equivalent to those imposed by the Code. O l l d
\' %
ii 1.1-2 (4- 1-67 ) 357
\ Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION Your calculations indicate tha t xenon oscillation might occur in 1.2 this core. Please describe the method by which xenon oscillations would be controlled should they occur. ANSWER Initial studies of the Oconee ?. ore, where realistic fuel tempera-tures are generated by thermal-nuclear iteration, indicate no instability at any time during the life cycle. These results are encouraging, but until more detailed analyses are comple:ed, it will be assumed that axial xenon oscillations are possible, azi-muthal oscillations are unlikely and radial oscillations will not occur. To illustrate a control procedure, an anal sis was made of an oscillating core without the stabilizing effect of the fuel temperature (Doppler). (If not controlled this produces a diver-gent oscillation.) Partial control rods , having a 3-f t long poison section, were moved up and down about the mid-plane of the core to offset oscillatory power shifts. This procedure was successful in limiting the bounds of the oscil-
~N lation and with more precise rod movement over shorter time periods should maintain the resultant power peaks within allowable limits.
If further analysis substantiates the assumption that oscillation will occur, this or some similar procedure will be developed during the detailed design. More detailed results of the stability analysis of the core are presented below, followed immediately by the methods section con-taining the details and limitations of the threshold and diffusion theory calculations employed. The closing section outlines an overall approach to the solution of the stability problem in regard to additional detailed calculative programs as well as a method for the correction of unbalanced power distributions. (a) Summary of Results
- 1. Threshold Analysis In the threshold analysis (l) axial, azimuthal and radial oscillations were investigated for beginning of life, flattened and slightly dished power distributions. The results of note are as follows :
- a. For a fixed dimension, the tendency toward spatial xenon oscillation is increased as the flux increases.
- b. For a fixed flux, the tendency toward spatial oscil-lation is increased as the dimension of the core in-creases. -
0- . m([v 5 b UL 1.2-1 (4-1-67) ~0000 358
- c. The large size of current PWR designs permits an ade-quate xenon description using 1-group theory.
- d. Flattened power distributions are more unstable than normal beginning-of-life dis tributions. Dished power distributions are even worse.
- e. In a modal analysis of the core, modal coupling can be ignored. In addition, the core is not large enough to permit second-harmonic instability.
- f. A large, negative power coefficient tends to dampen oscillations. If this coefficient is sufficiently large, oscillations cannot occur regardless of core size or flux level. Current designs have a substantial negative power coefficient.
- g. The critical diameter for azimuthal oscillations is larger than the critical haight for axial oscillations.
- h. The core is not large enough to excite radial oscilla-tions.
- i. Examination of the diameter, height and power coeffi-cient f ar this design indicates that oscillations should not occur at the beginning of life with un-flattened power distributions. However, there exists a finite probability of oscillations at some later time, since core depletion tends to flatten the power distribution.
- j. The period of oscillation (25 to 30 hours) is long enough to permit easy control of the cscillations,
- k. The modal analysis of this core toaard the end of the initial cycle (with about 80 per cent flatness) showed that axial oscillations are poss .ble , azimuthal oscil-lations are unlikely and radiai oscillations will not occur.
- 2. Depletion Analysis Diffusion-depletion calculations coupled with heat trans-fer equations were employed to investigate further the axial stability of the core since the analytical study indicated that this was the most probable mode of oscilla-tion. The results follow:
- a. Axial instability did not occur at any time during the initial cycle. An average fuel temperature of 1,400 F was maintained during the cycle.
- b. The threshold for axial instability near the end of the initial cycle was found to coincide with a core average fuel temperature of 900 F.
Diffusion theory was also used to examine the problem of controlling the system with rods if the stabilizing power Doppler was not present. The following was concluded : L
- a. Partial control rods are quite adequate in controlling axial oscillations. These rods have 3-ft-long poison 1.2-2 (4-1-67) 0000 357
sections which are moved up and down about the mid-
\ plane of the core to offset oscillatory power shif ts.
- b. Detailed power profiles will be available to the operator as output from the instrumentation. The large period of the oscillation will allow partial rod movement such that axial power peaks are held well within allowable limits.
(b) Methods l 1. Threshold Analysis The threshold analysis is described in 3 . 2. .2. 2 3 of the Preliminary Safety Analysis Report, including Tables 3-8, 3-9 and 3-10.
- 2. Depletion Analysis Core-averaged quantities were used in the analytical ana-lysis. For a more comprehensive investigation, it is desirable to study xenon oscillations with diffusion-depletion programs including heat transfer. Such calcu-lations, which include the important local temperature effects, allow the designer to look for xenon oscillations under actual operating conditions. For these reasons, the Babcock & Wilcox Company LIFE depletion program was
] '
modified to include axial heat transfer. The equations and iteration scheme are outlined below.
- a. The average fluid tempe: 2ture for each axial core region is computed from a previously known power density distribution as follows :
(1) Z out AT t = (T out ~ in i D (Z) M Z in where: AT = temperature change in region "i" t PD(Z) = power density in 7 direction Zg ,Z = region "i" boundaries and 3 (2) C = AT eore (Z) 0 , where H = active fuel height. Equation (1) is solved for Yout of, region "1." Since Tin is known from core inlei: conditio'ns, the average fluid temperature is definei,1 as follows: J ! .. ... o 4 y , 1.2-3 (4-1-67) i 0000 360 4 ffu a ,1; L i
"~r - - - . .-e. - ___w ,
() T +T
=
"t i" T i fluid 2
- b. The newly cot 9 uter-region-averaged ".uid temperatures are used to compute new fluid densities. These fluid densities are then used to adiust the. number densities for water and soluble poison. Local or bulk boiling are not permitted ,
- c. The average fuel temperature for each axial core region is then computed from the average fluid temperatures and power densities.
(4) _ T fuel
=
_f+Tfluid PD where PD = coverage power density of region "i" and K is defined by
} fuel - fluid core K= _
PD core
- d. After the new fluid temperatures , moderator densities and fuel temperatures are obtained, these quantities are used as new LIFE input to obtain a new power dis-tribution until either a convergence criterion is met or a specified number of iterations is made.
This analysis used an exact solution in that the spectrum was recalculated for each zone (11 axial zones described the reactor) for each iteration at every time step. This included the effects of the moderator coefficient. This LIFE package was used to determine the ef fects of the uncertainty in the power Doppler on the stability of the core. The uncertainty in the Doppler was more than com-pensated with a reduction in fuel temperature of 500 de-grees. The core was analyzed with core average fuel tem-peratures of 1,400 F and 900 F. Attached Figure 1 com-pares the cyclic response of these two cases following the 3-ft insertion and removal (after two hours) of a 1.2 per cent.o k/k rod bank near the beginning of life. The third curve on the graph depicts the behavior of the core if the heat transfer equations were not included in the calcula-tion. Attached Figure 2 shows the same comparison toward the end of life. It is easily verified that the 900 F fuel temperature case approached the threshold condition for axial oscillation in this core. On the basis of the information presented, it can be said that for a realistic fuel temperature this core does not exhibit axial insta-o(. ' - 1.2-4 (4-1-67) 0000 361
r 1 bility at any time during the initial cycle. The 1-D model was used to determine a method of controlling the core without taking into account the stabilizing effect
< of the power Doppler. Normally, this would produce a divergent oscillation as shown in Figure 3. A study was completed wherein a 1 per centak/k rod bank with a 3-ft-long section of regular control rod material was success- , fully maneuvered to control the core after a perturbation j of the power shape at a point about 3/4 of the way through ! cycle 1. The controlled results are also shown in attached Figure 3. The minimum rod motion was one foot, and the time step employed was 4.8 hours. More precise rod move-ment over shorter time periods would produce a much smoother power ratio curve. This control mechanism appears quite adequate for the intended use.
(c) Conclusions . 4 ' Instability in the radial or azimuthal mode is not expected since the diffusion theory study showed that the core is stable throughout lifetime and the L/D ratio is 1.1. The results are encouraging, but until additional analyses are completed, it will be assumed that axial xenon oscillations are possible. Consequently, rod motion will be used to compensate for un-O balanced power distribution as indicated by the instrumenta-tion. Work is under way to provide a 2-dimensional depletion program < which allows nuclear-thermal iterations. A detailed quantita-tive analysis of core stability and control procedures, employ-ing either partial or normal control rods, is to be undertaken with the new program. (d) References (1) Neu'aold , R.J. , Xenon Oscillation, BAW-305,1966. (Also Ref.16, Section 3, PSAR)
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l l ! l l a g 0'000 362 !O i 1.2-5 (4-1-67)
O O FIG, 1. EFFECT OF FLIEI, TEM PEllATilitE (DOPPl Elt) ON XENON OSCil.l.ATION O q lieginning of 1.ife w. Power Itatio Taken 36" From Top and Bottom of Active Fuel Case 1 - No Temperature Iteration T el = 1400 F Case 2 - Temperature Iteration with T fuel = 1400 F Case 3 - Temperature Iteration with T = 900 F f el 2.8
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sO . ,o .2 - i C O 0 - 3 I I l LM 0 1 2 3 4 5 Time, days Q W: Osc i l la t ion ini tia ted a t T = 2 days Figure 1, Question 1.2 (4-1-67)
O O O - FIG. 2. EFFECT OF FUEL, TEMPEllATUltE (DOPPLElt) ON XENON OSCII,1,ATIONS Near End of Life Power Itatio Taken 36" From Top and 130ttom of Active Fuel O. Case 1 - Temperature Iteration with'T fuel
= 1400 F o i Case 2 - Temperature Iteration with T = 900 F g f el 2.4
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Figure 3, Que st ion 1.2 (4-1-67) 1
l O Dockets 50-269, -270 and -287 Supplement No. 1 April 1, 1967 Revised May 25, 1967 QUESTION Discuss the use of aluminum components in the primary system from 1.3 the standpoint of experience with these components in service and state the criteria to which these components will be designed and fabricated including corrosion and fit-up considerations. ANSWER Evaluation of the relative advantages of stainless steel and alumi-num has resulted in a decision to use stainless steel in all aux-iliary system components and piping except for certain tanks listed below. Boric Acid Mix Tank Borated Water Storage Tank Condensate Test Tank Reactor Coolant Bleed Evaporator Feed Tank Reactor Coolant Bleed Holdup Tanks Aluminum is used in the auxiliary systems tanks as noted above where the operating temperatures and pressures are low, i The major application of aluminum is in those auxiliary systems [/( sms containing boric acid. Published literature (1)(2) indicates that aluminum is e mpatible with borated water. In the course of The Babcock & Wilcox Company contract work on Indian Point Unit No.1 for the Consolidated Edison Company of New York, Inc., a series of corrosion tests were made on various reactor materials in boric acid solutions. Anodized and nonanodized alumi-num types 6061 and 6063 were exposed in a 1 per cent boric acid solution at 200 F for a total exposure time of 1000 hours. Some of the specimens were completely submerged in the corrodent while others were half submerged. All of the specimens gained weight with the corrosion rate varying from 2.9 to 33.6 mg/dm /no. 2 There was no evidence of pitting or any other evidence of selective attack. Anodizing the specimens resulted in an apparent increased corrosion. Anodized specimens corroded at a rate of 19.7 to 33.6 mg/dm 2 /mo, and the unanodized specimens corroded at a rate of 2.9 to 9 mg/dm2/mo, A corrosion rate of 10 mg/dm2 /mo, converted to more familiar terms, would only amount to a metal thinning of 0.00018 in./ year. If the aluminum components are to be painted following installation, no lead base paints will be used. 1
. 1 Experience at the B&W Test Reactor in Lynchburg, Virginia, for the l past five years has indicated no problems stemming from the joining of stainless steel and aluminum.
n ;. , 1.3-1 (Revised 5-25-67)
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- 4,
O ASME code design techniques similar to those for steel components will be used in the design of these components, and the allowable stress values are those presented in these codes. O l (1) Chemical Engineers Handbook, Section 21, " Materials of Construction." (2) Reynolds Metals Company publication, Aluminum in Modern , l Architecture, Section 8.1, "Use of Aluminum in Piping Systems." { , 1.3-2 (Revised 5-25-67) 4
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O m- Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION Discuss the containment tests which will be performed initially and 1.5 over the service life of the containment which will assure that at least the specified 50% of containment leakage will exit through filters via the penetration room. ANSWER Presently, there are no practicable means of directly measuring leakage exiting exclusively through the penetration rooms to the filters. Until cuch a technique is developed, the adequacy of the safeguards designed to control the release of radioactivity to the environm'ent under emergency conditions can be established by the criteria outlined below. These criteria vill limit the effects of a release under emergency conditions to less than the ef fects associated with a Reactor Building leakage rate of 0.50 per cent per day, one half of which leakage is processed through the pene-tration room ventilation sys tem. This condition will prevail if: (a) The total Reactor Building leakage rate is 0.25 per cent per day or less, demonstrated by accual test; or (b) The net leakage rate (the difference between the total Reactor Building leakage rate and the sum total of the individual leakage tests of each testable penetration terminating in or vented to the penetration room) is 0.25 per cent per day or less. The integrity of the Reactor Building would be initially established and periodically demonstrated by retesting the building (1) in accordance with test frequencies required for reactor buildings with leakage rates of 0.50 per cent per day and (2) consistent with the above criteria. Notwithstanding the above requirement and criteria, it is intended that testable penetrations would be periodically tested to maintain their high degree of leak tightness. It is believed that the specified 50-50 split is conservative, and a greater percentage will occur through the penetrations, and hence l be processed through the penetration room ventilation system. De-sign of the penetration room ventilation filters .:111 be based upon ' all leakage via penetrations. Experience, as summarised in the ". S. Reactor Containment Technology Handbook," indicates that Reactor Building leakage is much more likely at penetrations than it is through the liner plates or weld joints. Although always within technical specifications on leak rate, all leaks round and corrected at CVTR were at penetrations and isolation valves, and there has never been any indication of a measurable leak in the. liner. This has been substantiated by The Babcock & Wilcox Company's experience with the N. S. Savannah's high pressure containment, the Indian Point Plant, and the B & W test reactor in Lynchburg which contains a low pressure steel containment. No post-operational Reactor Building
] \-
leakage rate tes t in the United States is known to have indicated
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a significant amount of leakage through the liner plates or weld O seams. Further, Duke Power Company's power plant experience with carbon steel tanks, boiler casing, breechings, and other plate weldments shows that leakage is much more probable at penetrations and connections than elsewhere. l O 0000 6' l.5-2 (4 67 )
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i e Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 l QUESTION Discuss the frequency and type of maintenance likely to be per-
. 1.6 formed on the hydro plants. What is the time required to restore the hydro plants to operation for these various types of main- ; tena:997 This should include maintenance that might be performed on the penstocks.
ANSWER The frequency, type of mainteaance, and time to restore at least one hydro unit to service in the Keowee Station are expected as
! follows:
(a) A general inspection of the penstock will be made sometime j during the first year of operation. Inspection crews can
; vacate the penstock and return the units to service upon j two hours notice. This will be followed by similar inspec-i tions scheduled at ten-year intervals.
I (b) It major repairs are required to the penstock necessitating installation of equipment and/or scaffolding, it is expected
]
that the penstock can be refilled and the units available for service upon notice of two to six hours depending upon the I repairs being made. Satisfactory experience with a penstock built in 1918 of riveted steel plate indicates that repairs ( to the concrete lined Keowee penstock are improbable over the life of the unit. (c) The penstock will be unwatered for annual simultaneous inspec-tion of the two hydro waterwheels. The units can be made available for service upon two hour notice. (d) Major repairs to a* hydro waterwheel will require unwatering of the penstock. Repairs will be scheduled for only one wheel at a time in which case the other wheel can be available for service upon two hours notice. An example of a major repair is rewelding cavitation damage which is expected with a fre-
; quency of seven to ten years.
The above estimates are based on experience at our Bridgewater Hydro Station where two units are connected to a common tunnel-penstock. We expect the Keowee units to give improved service due to modern methods and materials which should decrease the frequency of making repairs in the penstock and to the units. Inherently, hydro units are low maintenance pieces of equipment that are highly reliable in service and require very little time to restore them to service from a maintenance condition. After completion of Oconee Unit 2, maintenance and inspection on hydro
- l units will not be scheduled to coincide with seasons of heavy sys-4 tem loads or with scheduled outages of either of the Oconee '.
nuclear units. 1.6-1 (4-1-67) 0000 i7
1 l l l l Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION Please provide a discussion of how the larger water gap and thinner 1.7 thermal shield in this proposal affect, as compared to currently licensed plants , (1) the neutron irradiation, and (2) the thermal stresses in the pressure vessel wall. ANSWER (1) Neutron Irradiation Calculations have been performed in connection with the Oconee reactor pressure vessel design to determine the relative effects of varying the baffle and thermal shield thicknesses on the neutron flux (>l Mev) at the vessel wall. These calculations were performed with the P1 option of the P3MG1 code (l) using 34 fast neutron groups. The results showed that the neutron flux level at the ves-sel wall is dependent, for the most part, on the total metal and water thickness between the core and the vessel. However, there was some variation in fluxes depending upon the particular config-uration of steel-water laminations. Also, the gain in neutron attenuation by replacing water with steel diminishes somewhat with increasing steel thickness. In general however, the results showed that for total steel thick-(' '\' nesses in the range of 3 to 6 in., 1 in, of steel in place of 1 in. of water would reduce the neutron flux above 1 Mev by about 30 per cent. In pure water the calculations showed that the neutron flux 3 would be reduced , on the average, by a factor of 6 in 6 in. of water. Based on the above analysis a comparison has been made of the neu-tron attenuation in the Oconee vessel with those in San Onofre, Turkey Point No. 3 and 4, Indian Point No. 2 and Girna. The total distance oetween the core and the reactor vessel in the Oconee unit is 21 in. This provides from 1.5 to as much as 5.75 in. more water between the core and the vessel than in the other reactor units. For neutrons above 1 Mev it was found that this additional distance would provide additional attenuation ranging from a factor of 1.1 to 5 times greater than that in the other PWR's considered. (2) Thermal Stresses The gamma heating in the reactor vessel is produced by primary gammas from the care and by secondary gammas originating in the core liner , barrel, thermal shield , and the vessel itself. In the Oconee design the major portion of the heat is generated by gamma rays from the core and by secondary gamma rays from the core liner and barrel. (% Since the gammas from each of these sources must penetrate the ( ,) thermal shield to reach the vessel, the vessel heating. rate is-dependent on the thermal shield thickness. ' 3S . . . 1.7-1 (4-1-67) 110
,0000
~ -- -
r or designs which employ thicker thermal shields , or in which O
.ternals are to be exposed to higher neutron fluxes, gamma rays originating in the thermal shield or in the vessel itself may govern the vessel heating rates. Since gamma rays from these sources would have to penetrate only portions or none of the ther-mal shield to reach the vessel, the vessel heating in such cases would be less dependent on thermal shield thickness than in the Oconee design.
A comparisca was made between the gamma attenuation provided by the water and metal in the Oconee vessel and that in other PWR's by assuming that, in each design, the vessel heating was dependent on the gamma ray attenuatior provided by the thermal shield. This approach would be conservative since, as noted above for some de-signs, gamma sources other than those attenuated by the thermal shield may contribute appreciably to the vessel heating. The results of the comparison showed that the dif ference in gamma attenuation between Oconee and other PWR's ranged from negligible difference to a factor of 5.3 less for the Oconee design. The maximum steady-state stress resulting from gamma heating in the vessel has been calculated to be 3,190 psi (tension). This is a relatively low value, and no problems are anticipated from thermal stresses in the reactor vessel wall. O Re ferences , (1) Bohl, H. , Jr. , et al, P3MGl, A One-Dimensional Multigroup P-3 Program for the Philco-2000 Computer, WAPD-TM-272. O 0000 11 i' 1.7-2 (4-1-67)
Dockets 50-269 and -270 l Supplement No. 1 April 1, 1967 QUESTION Provide a drawing indicating the location of all areas within the 2.1 site boundary which will not be owned by Duke Power and those that will oe leased or otherwise used for purposes other than power generation. State the control that will be exercised by Duke Power over these areas. ANSWER This information has been added on Figure 2-2 (Revised 4-1-67). Duke will own all land within the site boundary except for two tracts: 1) a small rural church plot of 4.6 acres, the church being of historical interest and having no active membership, and 2) 9.8 acres of the United States Government's Hartwell Reservoir property located in the floodplain of that project. Duke will contracturally arrange to control all activities in-cluding exclusion or evacuation of personnel and property from these tracts in case of necessity. These arrangements will be in the nature of covenants running with the land and will be binding upon present and future owners of these tracts for so long as the Oconee Nuclear Station is in operation. Three residences to be owned by Duke and shown on Figure 2-2 will be leased to the occupying families at their request for use as O single family residences, with the provision that they will imme-diately evacuate the exclusion area upon notification by Duke, and with full control of other uses of the property to be vested in Duke. O- *M ' - 2.1-1 (4- 1-67 ) l
j l i I i j Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 l ! QUESTION Estimate the expected transient population around the future Lake i 2.2 Keowee as a result of summer cottages, boat access and any commer-I cial activities. I ANSWER It is expected that Lake Keowee's 300 mile shoreline will be fully developed by 1985 at which time the estimated transient population during a summer holiday or week-end will be 7500.
; This estimate is based on development of lakeside lots, use of I access areas and growth of commercial activity to support this j expanded recreation.
j There will not be any summer cottages, boat access areas or commer-cial activities within the site boundary. 1 I i j l < I l l 1 l l t I I i i f i t .I I i 0000 13 !O 2.2-1 (4-1-67 )
Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION Locate the water intake for the town of Seneca with reference to the 2.3 reactor and also indicate the distance to che proposed intake point on Lake Hartwell for the city of Anderson and the towns of Clemson and Pendleton. Provide stream flows, travel times, and estimated dilution to these intakes. Estimate the length of time that these municipalities could suspend use of these intakes. ANSWER The water intakes, existing and future, for Seneca, Clemson-Pendleton and Anderson are shown and located relative to the site on attached drawing dated 4-1-67. Due to discharge of liquid effluent into Keowee tailrace below Lake Keowee, no Oconee effluent would reach the Seneca intake located on . Lake Keowee. The estimated streamflows are: Incremental Total Incremental Total Keowee to at Keowee to at Clemson Clemson Anderson Anderson At Keowee Tailrace Intake Intake Intake Intake O Cfs Cfs Cfs Cfs Cfs Minimum instantaneous, no hydro units operat-ing - 30 54 ' min. ) 84 (min. ) 170 (min. ) 200 (min.) Annual average - 1100 357 (avg. ) 1457 (avg. ) 2010 (avg. ) 3110 (avg. ) Minimum monthly aver- i age - 646 Lowest yearly aver- i age - 858 l The estimated travel timcs computed on the basis of reservoir volume divided by the mean flow between Keowee and each intake are tabulated below. The estimated dilution is also shown based on d_scharge to the Keowee tailrace of 60 gph of liquid effluent as outlined in 11.1.2.5.1 and uniform mixing in the reservoir between Keowee and intakes. At Clemson Intake At Anderson Intake Mean Travel Mean Travel Har twell Volume Flow Time Volume Flow Time At El. 654 Ft3 Cfs Days Dilution Ft3 Cfs Days Dilution Min. Flow 675x100 57 *137 1:3.8x104 920x107 115 *926 1:9x104 Avg. Flow 675x106 1288 6.1 1: 6. 5x105 920x107 2105 50.5 1: 14x105 Hartwell , g At El. 625 L Min. Flow 260x106 57
- 53 1: 3.8x104 429x107 115 *432 1:9x104 Avg. Flow 260x106 1288 2.3 1: 6.5x105 429x107 2105 .23.6 ,1:14x105 y
4
; 2.3-1 (4-1-67 ) ,
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- Minimum flow would never have duration of more than a very few days, O
and hence, dilution at intakes would always be in excess of values shown. The intake at Clemson has and the intake at Anderson will have open-ings '. several elevations to permit selective intake of water if mixing is not uniform with depth. The new Anderson intake will be loc:ted on an arm of the reservoir that is fed by a separate tribt.tary. Any discharge to the Keowee tailrace would not be expected to flow upstream on this arm of the reservoir to reach the intake. The estimated length of time that municipalities could suspend use of intakes based on volume of water in clearwells and elevated tanks and amount of usage is - Seneca - day Clemson-Pendleton - 1 days Anderson - has present intake on another watershed and new intake on Hartwell would supplement this. There-fore, in case of necessity, Anderson could obtain water from existing intake and supplement from clearwells and elevated storage for several days. O I l l l l 2.3-2 (4-1-67 ) 0000 15
g-'s k-u Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 i QUESTION Discuss the reasons for discharging liquid radioactive waste into 1 2.4 the tailrace of the hydro plants rather than into Lake Keowee. In this regard provide the following information: (a) What is the effective transit time and dilution factor from the plant discharge canal through the lake to the intake canal and how would these be af fected by various flow conditions in the rivers? (b) What are the
, corresponding factors between the discharge canal and the tailrace of the hydro plants? (c) How will the flow through the hydro station be affected by low ilow in the rivers feeding the lake?
ANSWER Discharge to the tailrace can provide under Duke Power Company control the maximum immediate dilution of effluent whenever required as com-pared to all other alternatives. The maximum hydro discharge during operation is expected to be 19,800 cfs, and the maximum cooling water discharge to Lake Keowee is expected to be 3150 cfs for two units. a) The effective transient time between the intake and discharge canals is 6.9 days ~when reservoir is at its maximum drawdown elevation 775 and 9.9 days with reservoir at its average elevation 793.5. These times are conservatively calculated based on no Keowee hydro station or spillway discharges and no I inflow to reservoir, which would substantially increase effective transient time. For this calculation, both units were assumed to i be operating with condenser cooling water flow of 3150 cfs. Long-term dilution of Oconee ef fluent to the plant discharge canal is dependent upon inflow to the reservoir rather than flow in the discharge canal. With an average reservoir inflow of 1100 cfs, the dilution factor (for 60 in PSAR 11.1.2.5.1) wouldbe1:4.9x10gphOconeeeffluentasnoted b) Dilution of 60 gph into the tailrace would be by a factor of 1:1.35x104 with the hydro station shut down and minimum outflow
. of 30 cfs leakage; a factor of 1:4.9x105 for the long-term I
average discharge of 1100 cfs; and a factor of 1:8.9x10 6 obtain-able by operating the hydro units and discharging 19,800 cfs. Dilution factors in the ta11 race resulting from putting Oconee effluent into the discharge canal will be the same as indicated 3 in (a) above. The transient time between the discharge canal and the Keowee j tailrace would be extremely short when both plants are operating. When only Oconee is operating, the normal maximum transient time would be twe to three days based on the expected Keowee hydro station operation, t c) Reservoir storage will be used to supplement natural inflow dur-fN ing low flow periods'for power generation at Keowee. " Based on
\-)
0000
~~ ' 16 2.4-1 (4 67 )
v c; .g
'4 a computer run of a 10-year hydrograph which includes the driest period of record, the flows through the Keowee hydro station are:
858 cfs - Minimum yearly average ^ 646 cfs - Minimum monthly average 30 cfs - Minimum instantaneous Corresponding dilution factors of 60 gph effluent may be obtained by multiplying above flows by 448. O I O: 4 2.4-2 (4- 1-67 ) 17 0000
i, I ( I i
- Dockets 50-269 and -270 Supplement No. 1
$ April 1, 1967 4 1 QUESTION Describe the scope of the preoperation and postoperation environ-2.6 mental monitoring program including the type and frequency of sample collection. i ANSWER The environmental monitoring program is designed to establish environmental radiation levels and detect any changes which may , occur. Sampling points have been or will be chosen both on site and off site at distances up to 10-15 miles from the station, generally in prevailing wind directions and near population centers. ! The samples collected will include the following: (1) Water (streams , wells , lakes and rain) (2) Airborne particulate material (suspended and settlement or . fallout) l (3) Soil and silt t (4) Vegetation , (5) Milk ( (6) Fish and animals In addition radiation levels will be measured at various locations. The gross alpha and gros.- beta activity of the samples will generally be measured and specific radionuclides will be identified when appropriate. Special analyses, such as determination of Sr90 i in milk, will be performed by an outside laboratory. l I Since the program bas not been completely developed, the frequency 1 of observations at iampling have not been finally determined, but it is planned tt determine integrated radiation doses at monthly intervals. Water and airborne particulate samples will be collected weekly and radioactivity determined. Other samples will , be collected and analyzed on a frequency to be determined. l l l 'N . i ' i
.000( 18 2.6-1 (4-1-67) 1 1
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1 Dockets 50-269 and -270 . Supplement No. 1 l April 1, 1967 QUESTION It is our consultants' tentative opinion that the maximum hypo-l 2.7 thetical earthquake should be about 0.10g for those Class I structures which are founded on bedrock and 0.15g for any Class I structures located on overburden. Please provide your structural design criteria for the maximum hypothetical earthquake. 1 ANSWER The structural design criteria for the maximum hypothetical earth- ! quake is 0.10g and 0.15g for Class I structures founded on bedrock and overburden respectively. The Reactor Building seismic design will be based on dynamic analysis using response spectra curves based on the design earthquake ground acceleration and will be checked for a safe and orderly shutdown for the hypothetical earthquake ground acceleration. i Additional design criteria are being presented in the answers to questions 8.1, 8.4 and 8.5. d 1 i i i i 'I i I
. 'o' 0000 19' Os_ /
2.7-1 (4-1-67 )
O V Dockets 50-269 and -2'O Supplement No. 1 April 1, 1967 QUESTION Provide a description of the methods used to calvulate core void 3.4 fractions. ANSWER The Babcock & Wilcox Company's computerized void pregram uses a combination of Bowring's (1) model with Zuber's(2) correlation between void fraction and quality. The Bowring model considers three different regions of forced convection boiling. They are: (a) Highly Subcooled Boiling In this region the bubbles adhere to the wall while moving upward through the channel. This region is terminated when the subcooling decreases to a point where the bubbles break through the laminar sublayer and depart from the surface. The highly subcooled region starts when the surface tempera-ture of the fuel reaches the surface temperature predicted by the Jens and Lottes equation. The highly subcooled region ends when (1) T sat -Tbulk
- y O* where: Q = local heat flux, btu /hr/ft2 i
l = 1.863 x 105 (14 + 0.0068p) V = velocity of coolant, ft/sec p = pressure, psia The void fraction in this region is computed in the same manner as Maurer(3), except that the end of the region is determined by Equation (1) rather than by a vapor layer thickness. The nonequilibrium quality at the end of the region is computed from the void fraction as follows: (2)
- 1 1+ -1 where: xd = n nequilibrium quality at end of region 1 ad = void fraction at Tsat - Tbulk " "O V
p = fluid component density, Ib/f t3 l n U 00_0.3 _ 20
, , 3.4-1 (4-1-67) h gp \
(b) Slightly Subcooled Boiling O In this region the bubbles depart from the wall and are transported along the channel (condensation of the bubbles is neglected). This region transcends the point where the thermodynamic quality is zero. In general, this is the region of major concern in the design of pressurized water reactors. The nonequilibrium quality in this region is computed from the following formula: (3) P h x*=xy+ g (o - $3 p)d:: 6hg g(1 + e) "d where: x* = nonequilibrium quality in region 2 hgg = latent heat of vaporization, btu /lb 1
= fraction of the heat flux above the single 7,
phase heat flux that actually goes to pro-ducing voids O gp = single phase heat flux, btu /hr/ft2
& = mass flow rate , Ib/hr Ph = heated perimeter, ft z = channel d is tanc e , ft The void fraction in this region is computed from:
(4) x* Co x* + pg /pg (1 - x*) + Af PE EEC f-m . PfL . where: g = acceleration due to gravity, ft/sec2 gc = constant in Newton's Second Law = 32.17 lb m ft
?
lb f sec~ Co = Zuber's distribution parameter Af = the flow area, ft2 e = the surface tension Equation (4) results from rearranging equations found in reference (2) and assuming bubbly turbulent flow in determin-ing the relative velocity between the vapor and the fluid. Zuber has shown that Equation (4) results in a better predic-tion of the void fraction than earlier models based on empirical slip ratios. 3.4-2 (4-1-67) 0000 21 ,
.y.
J (c) Bulk Boiling l In this region the bulk temperature is equal to the satura-tion temperature, and all the energy transferred to the fluid results in net vapor generation. Bulk boiling begins when 1 the thermodynamic (heat balance) quality, x, is greater than the noncquilibrium quality, x*. The void fractica in this l region is computed using Equation (4) with the thermodynamic
! quality, x, replacing x i
l k i l .I i REFERENCES : (1) Bowring, R.W. , Physical Model, Based on Bubble Detachment , and Calcu-lation of Steam Voidage in the Subcooled Region of a Heated Channel, HPR-10 OECD Halden Reaktor Project, December 1962.
; (2) Zuber, N. and Findlay, J.A., Average Volumetric Concentrations in Two Phase Flow Systems, presented at the ASME Winter Meeting, 1964. To be published in the ASME Transactions.
(3) Maurer, G.W., A Method of Predicting Steady-State Boiling Vapor Frac-tions in Reactor Coolant Channels, Bettis Technical Review WAFD-BT-19. 1
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0 0 0 3. '2 2 3.4-3 (4-1-67) 4 ! I v
i r Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 yUESTION Please provide information on current burnout experimental studies l 3.6 with multirod geometries and nonuniform heat generation for the
; configuration and service conditions of the proposed reactor.
] ANSWER The Babcock & Wilcox Company has and will continue to pursue an ambitious program of critical heat flux testing. Under previous single channel critical heat flux programs sponsored by B&W and EURATOM, B&W obtained data on tubular or annular test sections of 72 in. in length. Data were obtained for the following condi-tions: 20 0 ATs = 150 ) l 1,000 0 P ; 2,000 ; 0.5 x 106 ; g ; 2.5 x 106 where: AT3 = inlet subcooling, F P = pressure, psia O- G = mass velocity, Ib/hr-ft 2 The following axial heat flux shapes have been tested where a P (power) ~< heat flux. (a) Uniform Heat Flux (b) Sine Heat Flux (P/P) max = 1.396 @ 507. L (c) Inlet Peak Heat F1.ux (P/P) max = 1.930 @ 25% L (d) Outlet Peak Heat Flux (P '?T max = 1.930 @ 75% L The critical heat flux testing program also included bundle geometry tests. The test section length was 72 in, with a uniform heat flux. A total of 513 data points were obtained covering the following con- , ditions: j 0 0 STg 0 250 1,000 '= P = 2,400
~
]
! 0.2 x 106 c = 3.5 x 106 i .
l The geometry of this section c >nsisted of nine pins of 0 420 in. Q ' t^ 3.6-1 (4-1-67) Ougt)g. 23
diameter on a 0.558 in. square pitch. Analysis of che last data of this set is in process. Following the uniform bundle critical heat flux tests, a program of two additional nonuniform 72 in heated length tubular tests was undertaken to obtain data for peaking conditions more closely related to the Oconee design. The additional flux shapes being tested are: (a) Inlet Peak Heat Flux (P/P) max = 1.65 @ 28% L (b) Outlet Peak Heat Flux (P/P) max = 1.65 @ 72% L These tests,"still in progress, will cover approximately the same range of pressure , mass flow and 2T as the uniform bundle test. In all tests performed recently, or planned in the future, the test section geometry was designed to-be as nearly identical to the B&W pressurized water reactor design as possible. The range of vari-ables tested, ie, pressure, mass velocity and inlet subcooling, has been selected in such a fashion that full coverage of operating variables is assured. Tests conducted before mid-1966 are all typical of PWR operating conditions although the representation is not as exact as it has been since that date. In addition to critir.a1 heat flux testing, B&W has in progress a test using a test section which is an exact mockup of the bundle critical heat flux test section. This test section is instrumented to provide information on interchannel mixing, flow distribution, velocity distribution and pressure drop. While this test is not a critical heat flux test, it will provide information of significant value in predicting the thermal performance of parallel rod cores. 3.6-2 (4-1-67) ..., I t, ; L
0) V Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 l 1 QUESTION Please resubmit a revised version of Figure 7-2, incorporating 4.2 your present intentions relating to the design of the nuclear instrumentation and protection systems. ANSWER Figure 7-2 revised 4-1-67 and filed wi th Amendment No. I shows our present intentions with respect to the design of the nuclear instrumentation and protection sfstems. I l i I v) 1 l l h 00.00. 25 4.2-1 (4-1-67)
i Dockets 50-269 and -270 i Supplement No.1 ~ April 1, 1967 QUESTION Please submit a schematic diagram (similar to the format of Fig. 4.4 7-2) showing your proposed three-wire d.c. system. Please include ! a failure analysis which shows that no single fault within this g system (e.g., short, ground, failed breaker, faulted charger . . . , j etc.) can preclude the actuation of protection and safeguards de-
- vices under accident conditions.
1 ANSWER The answer to this question is being submitted as revisions to the 3 Preliminary Safety Analysis Report dated April 1,1967. Figure 8-3, 125/250 Volt DC System and 120 Volt AC Vital Power System, covers the schematic diagram requirement.
- Revised 8.2.2.5, 125/250 Volt DC System, and 8.2.2.6, 120 Volt AC Vital Power Eusses, pages 8-4, 8-4a, 8-4b, 8-4c and 8-4d revised 4-1-67, cover the description and failure analysis requirement.
1 1 l O T 1 0000 26, D
- s_ /
! 4.4-1 (4-1-67)
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Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION Does the design of your protection system conflict in any way 4.5 with the proposed IEEE Standard for Nuclear Power Plant Protection Systems? If so, please state reasons justifying your position. ANSWER The protection systems will be designed to meet the proposed IEEE Standard for Nuclear Power Plant Protection Systems. l l O
-}
4.5-1 (4-1-67) . 0H000
- - ~ -- ,_ . - - -
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/% Dockets 50-269 and -270 Q Supplement No. 1 April 1, 1967 QUESTION Please discuss your criteria relating to the qualification test-4.6 ing of instrumentation and associated circuits to ensure their ability to survive an accident environment.
ANSWER Protection system instrumentation will be subject to an accident environmental (qualification) test as required by the proposed l IEEE Standard for Nuclear Power Plant Protection Systems. The tests will establish the adequacy of equipment performance in both normal and accident environments. The qualification tests required will be run on final type equipment. The accident environment tests l will not include the accident radiation environment. The ability of the neutron detectors to perform their intended function will be judged from the detector supplier's typical test data. O'000 28 ( 4.6-1 (4-1-67)
4 Dockets 50-269 and -270 i Supplement No. 1 April 1, 1967 QUESTION Please list those portions of the containment isolation system 4.8 which are not fail-safe upon loss of voltage. Provide justifica-tion for your design basis. ANSWER As described in Section 7, engineered safeguards protective system instrumentation is designed with three independent redundant channels , each of which trips on loss of instrument power. However, other than the instrumentation channels, all electric actuated valves of the containment isolation system require some form of electric power for initiation and operation. Therefure, on loss of voltage they remain in their respective positions. The , isolation system is designed in this manner to prevent single I component failures (eg, relays , voltage transients , etc.) from inadvertently initiating isolation and causing unwarranted plant shutdown and potentially damaging transient upsets. Justification for using this type of containment isolation cor. trol is based on ; the following: 1 (a) No single failura in the engineered safeguards system will prevent Reactor Building isolation. x (b) Diaphragm air operated valves where used for isolation are I held open by air pressure and are closed by a compressed
; spring.
(c) As described in 5.2.2, Type I and II penetration isolation require that the valves be redundant and that one valve be backed up with another of different type. ] (d) Type III penetration isolation valves described in 5.2.2 where singular and electrically actuated are provided with redundant control circuits from redundant power sources. For example, electrically controlled diaphragm operated valves will have two electrical solenoid valves, either of which can independently vent the diaphragm operator on command signals fron the respective redundant circuits causing the valve to close under spring power. (e) Revised material in Section 8 of the PSAR included with Amendment No. I dated 4-1-67 shows the redundancy and re-liability of instrument power, control power and main power supplies necessary to actuate engineered safeguards systems j
; and isolation systems upon an accident. Fail-safe features provided in redundant power supplies assure t,he availability of the necessary power supplies so that emergency systems can A perform their functions. Loss of all power to the containment i ;Q isolation system, which requires the assumption of multiple failures, would result in a loss of function of .many elements of this system.
29 l 4.8-1 (4-1-67) 3000- - .j
! Dockets 50-269 and -270 Supplement No. 1 ! April 1, 1967 QUESTION Describe the capability to flood the reactor cavity after an 7.1 accident. How does the volume required to flood the cavity compare with the primary system volume? l ANSWER Drainage from selected surfaces receiving Reactor Building spray water will be directed to the reactor cavity. Over-flow drains from the cavity, located at an elevation above the bottom head of the reactor vessel, will give positive assurance of water flow out into the Reactor Building and
< thereby into the recirculation sump.
Metal reflective insulation attached to the reactor vessel l will be arranged with an air gap and holes. The reactor i vessel support skirt is provided with holes to permit free flow of water. Outleakage from the bottom of the cavity will be restricted J by the use of seals and plugs where required. i The net volume of the reactor cavity up to the cavity over-l \. flow is approximately 35 per cent of the primary system volume. t i 4 4 4 i
}
l i I I i
-t I, s 0.00.0 J0" 1
1 7.1-1 (4-1-67) i N
O U Dockets 50-269 and -270 Supplement no. 1 April 1, 1967 QUESTION Discuss the requirements for cooling the water recirculated from the 7.2 containcent sump. ANSWER The requirements for cooling the water recirculated from the reactor building sump to the reactor building spray system are set by the de-sign basis of this system. The design basis is to maintain the post-accident reactor building pressure below the design value. This cri-terion can be met by spraying the sump water directly into the reactor building atmosphere without additional cooling. The water temperature in the reactor building sump during the recir-culation phase of a loss-of-coolant accident is maintained below saturation temperature by the low pressure injection coolers. These coolers reduce the temperature of water recirculated to the reactor vessel and returned to the reactor building sump. The L transfer surface of these coolers is set by the nor=al operating conditions under the decay heat renoval operation mode. The cooling capability of this mode of operation vill maintain the reactor coolant at lLO F or less a; 20 hours after extended full power operation and is in ex-j'"N cess of t .at required under accident conditions. The performance of () these coolers at various inlet temperatures is shown in Figure 6-5 of the Preliminary Safety Analysis Peport. The adequacy of the emergency injection coolers to provide cooling of sump water recirculated to the reactor building sprays is demon-strated by an analysis of the reactor building pressure for a 36 in. ID pipe, double-ended rupture using 3,000 spm of spray cooling and ; one low pressure injection cooler in operation. Core flooding tank operation was not assumed, but 6,500 gpm combined high and los pres-sure injection was used to provide for decay heat removal from the reactor vessel and to establish the energy release rates to the re-actor building. The results of this analysis are presented in Fig-ures 1 and 2 which show the reactor building postaccident pressure and te=perature. Figure 1 shows that the reactor building pressure decays to less than 5 psig in 24 hours. For comparison purposes and to show that the ef-feet of spraying cooler water into the reactor building is small, a second curve is presented on Figure } which is based upon a spray re-circulation ecoling rate of 100 x 100 Etu/hr (approxi=ately equivalent to one low pressure injection cooler) at a su=p temperature of 195 F. (This is the temperature of the sump when recirculation to the sprays begins.) Figure 2 shows the te=perature of the reactor building and su=p coolant for the two conditions.
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'q s 7.2-1 (4-1-67) eo m . , . _ . _ _ .
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O O These curves demonstrate that coolin,; of the recirculated sprs-/ vater has no effect on peak building pressure and only a minor effect on the rate of pressure decay during the first 24 hours. Accordingly, it is concluded that no cooling of the recirculated spray water is required for this accident. O 1
~
G 7.2-2 (4-1-67) 0 32
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i 1 ( FICURE 7,2-1 REACTOR BUILDING PRESSURE VS TIME TOR A 36 IN. ID DOUSLE-ENDED RUPTURE WITH AND WITHOUT COOLING OF ME RECIRCUIATED SPRAY WATER
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e FIGURZ 7.2 BIACTOR BUILDING A?0 SPHERE AND SLMP COCLANT TEMPERATUR15 FOLLOWING A % IN. ID DOLSLE-ENDED 3CPTURE 3GG i i i 3 i3 ii i i i i iiiI i i , sii i , ; i ;i;y , , ;,. p ~% _ _- s% 2 N 14 ft Rupture K 8e a ctor Building 6,500 spo Core Injection 3.000 spa Butiding Sprays I
- Teeparature "
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- / (2) With Spray Coolers -
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Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION Discuss the NPSH requirements of the recirculating pumps with re-7.3 spect to the minimum height of water required in the sump. To what water volume inside containment does this correspond? Discuss the sump location considerations , intake design details , and the criterion for redundancy in sump outlet capacity. ANSWER The definition of available NPSH as applied to the emergency recir-culation pumps is as follows: r
+ +
NPS H gy =P t
~
v" a s c
-P t -P y where: P g= pressure at pump entrance P, = Reactor Building air pressure P = Reactor Building steam pressure s
P, = elevational pressure P g = pressure drop in pump suction line Py = vapor pressure of pumped water The lowest availability of NPSH occurs when the recirculated water is at saturated temperature since at this time P = Pv , and the s NPSH availability equation reduces to: NPSH gy,gg) = P, + P e -P g An analysis of each of these factors follows: The partial pressure of the air in the Reactor Building is taken to be equivalent to the dry air pressure at the time the Reactor Build-ing was scaled for operation. Assuming that atmospheric conditions were 90 F and 100 per cent relative humidity at the time of Reactor Building closure and considering the 800-ft elevation of the site, , the partial pressure of the dry air trapped in the Reactor luilding is 13.5 psi. No credit is taken for the fact that upon initial re-circulation tha Reactor Building air is at an increased pressure (due to the increased temperature). The elevational pressure is that produced by the relative elevations of the recirculating pump and the level of the water surface in the Reactor Building. The center line of the pump volute is located at Elevation 768, and the surface of the water in the Reactor Building is assumed to be Elevation 778. This elevation of the water surface assumes only a 2-ft depth of water in the Reactor Building, whereas the total depth will be the sum of the depths given for the sources below: .
~0 }ddo 35 7.3-1 (4-1-67)
(h" e, .
O (a) Reactor coolant loop contents - 6 in. (b) Core flood tanks 1/2 in. (c) Borated water storage tank - 5 ft-6 in. The inlet to the pump suction pipe is at the bottom of the emer-gency sump which will be about 2 ft-6 in, below the lowest Reactor Building floor, ie, Elevation 774. The attached Figure 1 depicts the elevational relationship of the components. The pressure drop in the sump inlet pipe is 17 ft of water based on 9,000 gpm through one of the two 18 in. suction pipes. The 9,000 gpm flow is the sum of 6,000 gpm low pressure injection flow and 3,000 gpm spray flow. If all three LP injection pumps and both spray pumps are operating, the flow is 12,000 gpm, and the pressure drop is 24-1/?. ft of water. Based on the foregoing, the available NPSH is as follows: (a) For 9,000 gpm flow:
=P +P -P = 31.5 + 10 - 17 = 24.5 ft-H O NPSH(avail) a e 1 2 (b) For 12,000 gpm flow:
NPSH ,y,11) = 31.5 + 10 - 24.5 = 17.0 ft-H30 The NPSH requirement for the recirculating pumps is 13 ft of water. This NPSH requirement is smaller than the available, even when all equipment is operating, one section line is assumed unavailable and very conservative assumptions are made regarding the Reactor Building water level and air pressure. The sump shown on Figure 2 is located in the basement (Elevation 776.5) floor on the north-south center line of the Reactor Building. Its position with respect to associated equipment is indicated on Figure 1. The following factors were considered in establishing the location: (a) Proximity to the Recirculating Pumps The sump location will minimize the length of the suction lines. (b) Free Communication with the Largest Portion of the Building Plan Area This will permit the recirculated water to return freely to the sump without significant quantities collecting in isolated parts of the building. Figure 2 shows the preliminary sump arrangement and the following { summarizes its salient design details: 7.3-2 (4-1-67) ( (10.a 36
(a) The top of the sump projects above the basement floor. This projection will serve to reduce the amount of dirt or loose material that might enter the sump during normal and emer?,ency operation. (b) The sump is covered with a steel grating having removable sec-tions. The grating will provide personnel protection, and the removable sections will permit access for inspection. (c) The screening surface is made vertical so that it will have a scif-cleaaing characteristic. A 50 per cent plugging factor was used in calculating the installed surface to allow for per-manent clogging. With 50 per cent plugging, the flow area through the screen is approximately 10 times the flow area through one of the two 18 in, sump outlet lines. (d) The screen will prevent clogging of the decay-heat heat ex-changers and the emergency spray nozzles. The screen is sized , to piss only particles small enough to flow through the various equipment including the Reactor Building spray nozzles. (e) The emergency recirculation lines will have a bell-mouthed entrance. This shape will minimize the entrance shock loss. This shape, in conjunction with the water depth above tre entrance, provides assurance that a vortex will not form at this point. (f) The sump will have a valved drain line that will discharge into the Auxiliary Building sump. This will allow draining of the sump during normal reactor operation. The sump outlet consists of two 18 in. full-capacity pipes, each having its own normally closed automatic isolation valve. In an emergency, both pipes will be placed in operation. The criterion for providing this degree of redundancy is that full suction should be available if one isolation valve fails to open. 7.3-3 (4-1-67) 0000 37
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0000 39
Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION Provide an analysis to show the amount of ti=e available to isolate 7.4 the service water in case of a break in the containment cooler tubu-lar heat exchanger (resulting in the injection of unborated water). ANSWER An analysis has been conducted to determine the a=ount of time avail-able to isolate the service water in case of leakage in the reactor building emergency cooling unit tubular heat exchan;er resulting in the dilution of the borated water in the reactor building sump after a loss of coolant accident. These emergency cooling units are protected against missile damage so that it is highly unlikely that a break in the heat exchanger or cooling watbr line could result fr,m a missile Generated during a loss-of-coolant accident. Periodic hydrostatic tests vill be con-ducted on these units and their associated systems to demonstrate their inte;rity throu;hout the lifeti=e of the reactor. The present arrangement of the emergency cooling units includes a sep e ate fluw indicator at the outlet of each unit. These flow in-dicators will have low-flow alar =s with a mini =um sensitivity of t50 gpm. Thus, any si::able leak from a cooling unit will be quickly de-
' "' " """ "' " " ' "='"""' " " """'"- ^ """" ***
C*J of 50 gpm or less might not be automatically detected and would re-quire operator action to detect the leakage. However, the lower dilution rates provide ample time for operator action. The dilution effect of unborated water leaking into the reactor build-ing sump after a complete loss-of-coolant accident has been evaluated using the following e=ergency coolin3 unit leak rates:(*)
- a. 2,000 3pm (esti=ated maximum attainable coolin; water flow rate per unit).
- b. 1,350 spm (nominal throttled cooling water flow rate per unit).
- c. 50 gpm (estimated maximum leak rate undetected by automatic flow t alarm).
(+)The detailed design of this portion of the service water system has not been co=pleted. An esti=ated maximu= flow rate of 2,000 gpm per unit has been selected to account for increased flow due to the decreased flow resistance associated with a ruptured cool-ing line. The preliminary design nominal throttled'flov has been set at 1,350 gpm per cooling unit. The higher flow rate is ap-prox 1=ately 50 per cent higher than the nominal flow, and final calculations are expected to verify this flow rate as conservative. Q. - 7.4-1 (4-1-67) n. o m, OD00 40
O Table 1 shows the results of the initial reactar operating conditions studied to establish the cost critical diluticn conditions. Case 1 shows tne initial reactor conditions which can tolerate the =inimum quantity of dilution. This case is at the beginning of core life, with the reactor hot at full power, no xenon or sa=ariu= present, and with the control rod assemblies in their appropriate full power configuration. The times at which criticality is approached, as a result of dilution of the borated water in the reactor building sump, are presented in Table 2 for a condition where all control rods are inserted after the accident and for a condition vnere all control rods are withdrawn. The later cases are not considered realistic, but are presented to show the lov i=portance of CCA insertion. As shown in Table 2, witn all CCA's in after the accident, criticality will be approached after 2.6 hours with the e=ergency cooling unit maximum leak rate of 2,000 gp=. ' lith all CCA's out, this ti=e would be reduced to L2.7 minutes. However, as pointed out above, such a leak rate should be readily detected and isolated. For a leak rate of 50 sp= or less whien =ight not be automatically detected by the flow alar =, criticality could be approached after 1CS hours with all CCA's in, and after 28 3 hours with all CCA's out. Tnis provides ample time for operator surveillance of the flow meter for the lea?.- a e rate of 50 @=, or for manual sa:plin.; of the recirculated sump water. O I O' 7.4-2 (4-1-67) 00 d1 1 I 1
m (,m) m (%/ ') N_ ' (v) Table 1 Eoron Cora entration Af ter Loss-of-Coolant Accident Mixed Concentra-Initial Conc entra- tion in Sump; End Concentration Re- Amount of Dilu-tion in Ret.ctor of Injection From quired for Approach tion Water Be-Beactor Conditions Coolant Sy:. tem, Borated Uuter Stor- to Criticality at quired to Approach Case No. ppm boron at;e Tank, ppm boron 125 F, ppm boron Criticality, cul
- 1. Unit Ilo.1 - ICL, full power, no xenon, no samarium. 1,690 2,182 1,206 3 331 x 105
- 2. Unit ilo. 1 - 10L, full power, equi-u librium xenon, no
~
4- samarium 1,540 2,160 1,012 4.669 x 109 w g 3 Unit Ilo. 1 - EOL, 4 full power, equi-a librium xenon and h samarium. 600 2,01'( 29) 23 65 x 10
- 4. Unit !!o. 2 - ICL, full power, no 5
xenon or namarium. 1,310 2,125 932 S.269 x 10 v . C3 - M h
O Table 2 Ti=e Available to Oetect Sys m Leakage Assu=ed Reactor Con-Coolin.; Unit ditions at the Time Po s i +.' of Control Time Available Leak Eate, Loss-of-Coolant Rot t ter Loss-of- EeforeCriticaligr s pm Accident Cccurs Coolant Accident is Approached, hrs / 2,000 3eginnin; of core All CCA's In 2.6 life, hot, full power., no xenon or samarium, appropri-ate full power CCA configuration. 1,350 All CCA's In 3? - 50 All CCA's In 105.0 2,000 All CCA's Out 0.712 (k2 7 min) 1,350 All CCA's Cut 1.05 50 All CCA's Cut 26 3
/.5
' Eased on an assumed moderator te=perature of 125 F.
c 9
. 43 O 7.4-4 (4-1-67)
l Dockets 50-269 and -270 Supplement No. 1
; April 1, 1967 l
l
' QUESTION Discuss the separability and location of the recirculation system ,
7.6 pumps to avoid flooding of the pumps in case of a major system ' l leak. 1
- ANSWER The low pressure injection pumps in the recirculation system will be located in the lowest level of the Auxiliary Building to assure i adequate NPSH for the pumps. The three low pressure injection pumps, two Reactor Building spray pumps and one component drain pump per unit will all be located in a single room, separated only by retaining walls around each individual pump. The pumps will be
; mounted on concrete bases above the floor.
l The only credible sources of system leakage are the pump mechanical seals or gaskets at flanged joints. Even though unlikely, the most plausible and largest source of leakage is failure of a mechan-ical seal. Leakage from the faulty seal would partially fill the retaining sump
- around the pump to an overflow leading to the Auxiliary Building sump located at the low point for this elevation. A level detector located in each individual retaining sump would warn the operator l of leakage of a pump. Remotely operated isolation valves could l then be closed to isolate the leaking pump. A small drain hole in l the retaining sump below the level detector would drain the neces-sary water to clear the alarm and confirm isolation of the leaking pump.
The Auxiliary Building sump pump at this level will be sized to handle at least twice the expected leakage from a comp'ete failure
; of a mechanical seal. The water from this sump would ordinarily be pumped to the waste disposal systen, but a connection will be l provided to pump it to the bleed holdup tanks in an emergency.
1
+
1 i i i 7.6-1 (4-1-67) 4k l
. _ _ . - - _~ __ . _ . _ _ . _ . , - _ __ __ _ _ _ _ _ _ . _ . _ . _ _ _ _ __...I
i Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 i [ QUESTION Provide the loading combination considering the design basis
- 8.1 accident-maximum earthquake combination, and use design basis accident-maximum wind combination.
ANSWER The design basis accident-maximum earthquake loading combination is as follows: Y = h (1.0 D + 1.0P + 1.0T + E ') where Y, 0, D and T are defined in Appendix 5A, PSAR, and f E' = Maximum hypothetical earthquake. ] The design criteria which will be applied to the above loading is that the deformation will be limited to values which will permit a safe and orderly shutdown. t j The response spectra of the design earthquake will be scaled to i the E' value, and acceleration response taken from the curve corresponding to 57 critical damping.
\
4 The design basis accident-maximum wind loading combination is as follows: 1 Y = g (1.05D + 1.25P + 1.0T + 1.25W) i l, where Y, 0, D, P, T and W are defined in Appendix SA, PSAR. i e i l i I l 1 l l j I It l l ~ 0000 45 4 8.1-1 (4-1-67) i
i I l
@ Dockets 50-269 and -270 Supplement No. 1 April 1, 1967
- QUESTION Clarify the design approach in the PSAR allowing limited plastic 8.3 yielding in a working stress design.
ANSWER Paragraph 3.2 in Appendix SA of the PSAR has been revised as follows:
"For Class 2 structures, the working stress design method will 4
be used and stress will be in accordance with ACI 318-63 and I i the AISC Codes." l l 1 l l I l 1 I i I i I l 8.3-1 (4-1-67) ']_Q Q ] - $f; 1 l .
I i i s Dockets 50-269 and -270 Supplement No. 1 , April 1, 1967 l QUESTION Provide the following information relative to seismic design in i 8.4 light of question II.7. I 8.4.1 The damping factors to be used in the various loading com-binations that include a seismic contribution. t i 8.4.2 A statement of the intent of the designer with regard to combination of maximum vertical and horizontal earthquake components in conjunction with the other applicable loadings. 8.4.3 The mathematical model to be used in the seismic design ] analysis. i 8.4.4 The stiffness factors to be used in the design analysis and a detailed basis for the selections. 8.4.5 The design criteria and procedures for design of the piping
, systems and supports for Class I components for seismic loadings in combination with the other applicable loads.
i ANSWER 8.4.1 The following values of per cent critical damping will be used in the analysis together with the natural periods to obtain spectral accelerations . , Per Cent of ) Item Critical Damping
; Welded carbon and stainless steel assemblies (This includes reactor internals, supports and similar weldments. ) 1 ; Steel Frame Structures I
(Both welded and high strength bolted) 2 Reinforced concrete equipment supports 2 Reinforced concrete frames and buildings 5 I Prestressed concrete structures
! Under design earthquake forces 2 l
! Under maximum hypothetical earthquake 5 I Vital Piping 0.5 8.4.2 The horizontal and vertical components of ground motion will be ! applied simultaneously and will be additive. ,
~47.-
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Page SA-3 has been revised to reflect the above changes as follows: O Seismic forces are applied in the vertical and in any horizontal direction. Tne horizontal and vertical components of ground motion are applied simultaneously and the two components considered occurring in such a way that the stresses are directly additive to the stresses resulting from other loadings. 8.4.3 A description of the procedures in making the dynamic analysis for Class I components, systems and structures follows: (a ) Natural Periods The natural periods of vibration of the Reactor Building, steel and reinforced concrete frames, multi-story buildings, and com-ponents and systems whose periods can reasonably be determined, will be computed by use of a simulated mathematical model re- , presenting a lumped-mass system of appropriate number of node ' shapes. (b) Acceleration Response Acceleration response values for the specific periods will be determined from the response spectra, using appropriate damping factors as shown in answer to 8.4.1 above. (c) In those instances where spring coefficients connecting the masses and the rigid base cannot be readily determined and/or vibratory systems are of a highly complex nature, the maximum response value (peak of the curve) corresponding to the appro-priate damping factor will be used in performing the stress analysis. By using this conservative value and demonstrating that the stresses are satisfactory, it will not be necessary to perform any further analysis to determine the natural periods of vibration for the system. 8.4.4 The foundation stif fness factor used in the mathematical model is derived as follows: Borings taken at the site indicate the bedrock to be sound granite gneiss with a measured modulus of elasticity Es = 4.8 x 10 6 lbs/in.2 as shown in Table 2 Appendix 2B PSAR. The base slab will be founded on this rock. Assuming the same depth of zero elastic movement as is I used for the computer program and equal to 65 f t, the modulus of the l subgrade, (stif fness factor, ks ) or the pressure necessary to cause i 1 1 1 01 4 8' " 0.0.0.0 8.4-2 (4 67 )
i' t a unit deflection in the rock, becomes 14,900 kips per . ,_sre foot per foot. 8.4.5 The design criteria for piping systems and supports for Class I com-ponents are contained in Appendix SA of the Preliminary Safety Analysis Report and further described below. Piping i Seismic loads on hot piping systems for Class I components and all piping penetrating the Reactor Building are determined on the basis of dynamic analysis using an acceleration corresponding to the peak of the response spectra curves for ground motion shown in Plate II-4 of PSAR Appendix 2B and in answer to question 8.5 , The natural frequency of segmental lengths of piping that can be held fixed and still meet the Code allowable expansion stress range (by stiffening with the addition of anchors and guides) will be cal- , culated individually and the seismic response obtained from the j spectra curve. l ! The stresses from horizontal and vertical components acting simul-i taneously will be combined with the stresses due to weight, thermal ! and mechanical loads, and internal pressure. These stresses will , [ determine the required yield strength of the piping systems and I component supports. Typical supports at the Reactor Building pene-tracions are shown in Figure 5-2 (Revised 4-1-67). Supports for lines within the building will be designed to withstand all the above loads including positive restraints to prevent whipping associated with pipe fracture of lines containing high internal energy, a i l i Y e e a l 8 \ l b
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(p) Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION With regard to earthquake response spectra provide the following: 8.5 8.5.1 A response spectrum for the maximum earthquake. 8.5.2 The basis for the shape of the proposed response spectra. 8.5.3 Identify, explain and justify the scaling of the re ponse spectra with respect to displacement, velocity, acceleration and frequency on the plots presented. ANSWER 8.5.1 Attached are response spectra curves, Plates 1 and 2, for the 0.10g maximum hypothetical earthquake for Class I structures founded on rock and response spectra curves, Plates 3 and 4, for the 0.15g maximum hypothetical earthquake for Class I structures founded on overburden. 8.5.2
- f. m The basis of the shapes on response spectra curves is explained by
) the following quote from our Seismology Consultant, Dames & Moore:
{G "The general shape is what we believe to be a conservative envelope of possible maximum motion. Theoretically, the velocity spectrum in a sound basement rock should be essentially horizontal over the mid frequency range. We have attempted to utilize the available theory and the one available strong motion record on hard rock in developing the recommended spectrum. The results of this were pre-sented as Plate II-3 of our original report." (See Appendix 2B of PSAR.) 8.5.3 l l The response spectra on the enclosed curves have been scaled in l accordance with the information quoted below from Dames & Moore:
"A response spectrum for an undamped single degree of freedom i oscillator is essentially the same as the ground motion spectrum l at the short and long period ends of the spectrum. The scaling l then increases such that over the intermediate period range, the I velocity, acceleration and displacement on the response spectrum l are approximately five times the corresponding values.on the ground motion spectrum.
"The undamped response spectrum presented for the Oconee Station conforms well to this general case, in that the scaling factor o in the range of frequency ' rom .02 cycles per second to .15 cycles
, per second increases from about 1.5 to 5. Through the intermediate J,
-900( 50 8.5-1 (4 67 )
range of frequencies, from 1.5 cycles per second to 2.5 cycles per second, the scaling factor remains at 5. In the high frequency or low period range from 2.5 cycles per second to about 30 cycles per second the scaling factor decreases from 5 to about L.5. The scaling factors for velocity, acceleration and displacement are general.1y consistent. However, the maximum acceleration indicated on the undamped response spectrum indicates a slight discrepancy which resulted during draf ting of the response spectra for the formal report. The maximum acceleration appears to be about 27 per cent of gravity. The maximum accleration should be 25 per cent of gravity.
"The original of the Tripartite paper used in our report was pro-vided by Dr. Newmark. However, we have made changes in the velo-city, frequency and acceleration scales to fit the amplitudes of motion shown."
O
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'~' / Dockets 50-269 and -270 Supplement No. 1 April 1, 1967 QUESTION Provide analytical studies to support the safety of the dam against 8.6 failure under carthquake loading. If such a failure were to occur, what ef fect would it have on the capability for safe plant shutdown specifically in the areas of shutdown cooling and emergency power?
ANSWER As an independent confirmation of Duke Power Company's analyses, our cubsurface and foundations consultant, Law Engineering Testing Company under the direction of Dr George F Sowers, has analy:cd the Keowee and Little River Dams for seismic loadings. The Law report concludes that the dams can safely withstand a maximum hypothetical earthquake loading of 0.00 g. The Law report, including computa-tion work sheets, is attached. I l The ef fects of a hypothetical failure of the dam under earthquake loading have been examined as requested with respect to shutdown cooling and emergency power. By using stored condensate and by i employing emergency shutdown procedures to conserve the use of this ' water, decay heat cooling can be provided for more than 20 hours. If a dam failure is postulated, the Keowee hydro units would be f-'y lost, but alternative sources of emergency power would continue to (
) be available. Power is not required during an emergency shutdown of this nature untii all stored condensace is expended by use of
]
the emergency steam-driven feed pump as described in 10.2.1 of 1 the Preliminary Safety Analysis Report. l l l l l l l l l l l l l l
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4.%., .* , 2.. L 2,_.y,3, ..s y_ A ,..- ; , . 2 .L x L._ L_ ) . ;'3 - [QrY[%hk - _ ' k MPt.'[' x.' , M, N N. ' ( <k, '['N h' ' ,/< ..!/,f/Y ,k. ' . . .N'&'/f/\ N ' . * * '.' **f.[N.y'%'.;l ' < ' r nG , 't ', N ', ' ? , #'A k, f . . .. ,, ./ f'$ A ' ' s - - g-g. g. g. ~ g g. g. g. & e =, .. *. N -?. Velocity, in /sec . . QUESTION S.5 ' SUPPLEMENT No. 1 PLATE 4 RECOMMENDED RESPONSE SPECTRA Revised 5-25-67 0000-l l l LAW ENGINEERING TESTING COMPANY (c') v SolL AN D POUN DATION INVESTIGATIONS 412 Plasters Ave., N.E. e P. O. Box 13815, Sta. K ATLANTA 24, GEORGIA March 28, 1967 Duke Power Company General Offices Post Office Box 2178 Charlotte, North Carolina 28201 Attention: Mr. L. C. Dail Subj ect : Seismic Stability Keowee Dam Gentlemen: Ci (.) At your request, we have analyzed the safety of Keowee and Little River earth dams under seismic loading to determine their safety. This report summarizes the conclusions we have reached. You ha a informed us that the following accelerations are to be utilized as the basis for seismic analysis: Design Earthquake-----------------------0.05g Maximum Hypothetical Earthquake---------0.lg Maximum Hypothetical Earthquake for s tructures on overburden--------------0.15g FOUNDATION CONDITIONS We have reviewed in detail the foundation conditions - at the Little River and Keowee Dams. The materials encountered are similar:
- 1. A veneer of alluvium and completely weathered rock.
- 2. Partially decomposed rock "saprolite" in alternate hard and sof t seams.
[~.h " \> 3. Continuous unweathered rock. - 2000_ 57 Duke Power Company h March 28, 1967 Page Two The foundation trea tment for the higher parts of both dams will be similar. All alluvium will be removed, as well as the soll ma terials , down to the level of weathered rock.* Although the foundation will not be hard, unweathered rock, this weathered rock in many parts of the world would be considered bedrock. In fact, many laymen describe the " weathered rock" as " sandstone", because it resembles a soft sandstone. Even this weathered material is relatively thin where the embankments have substantial height. At the Keowee site the weathered rock is only 5 to 10 feet thick at the maximum section. On the right abutment (looking downstream) there is a portion of the dam in which the weathered rock is about 40 feet thick; however, at this point, the dam is only 70 feet high, less than half the maximum height. At the Little River site, the greatest thickness of weathered rock is about 20 feet under the maximum dam height of approximately 150 feet. On the left abu tmen t (looking downstream) a portion of the earth dam is underlain by 50 feet of weathered rock; however, in that area , the dam height is only 70 feet. Based on our evaluation of the foundations at both sites, we conclude that the foundations of both dams do not correspond to overburden in any sense of the word. We, therefore, have analyzed the dams utilizing an acceleration of 0.lg. STATIC ANALYSES Static analyses of both dams had been made and the results summarized and submitted with the Contract Drawings. We had previously checked these studies by re-analyzing the most critical circles of failure (fcund by your computer program) independently. The conditions studied, both ups tream and downstream, included " steady state seepage", " sudden drawdown", and " construction"
- Law Engineering Tes ting Company, Little River-Keowee Development, Part I, Vol. I.
0000 ~'58 l Duke Power Company , fT V March 28, 1967 Page Three before the reservoir filled, utilizing the appropriate shear strength data for each condition. The safety factors in all , cases exceed the minimum values established in your design ' criteria. These minimums are consistent with modern practice and were approved by your Consulting Board. SEISMIC ANALYSES The static analyses were extended to include the effect of acceleration and the resulting " inertia forces" on stability. The method utilized is that proposed by N. Newmark (1965)* in the Rankine Lecture at the Institution of Civil Engineers (London). In this analysis a steady acceleration is assumed to be applied to the centroid of the potentially sliding segment of soil in the direction which produces the greatest increase in over-turning moment. The computations for the critical circle on the upstream N slope and two possible circular segments of the downstream slope, /~'/ \s- above the berm and including the berm are attached. Th'e results show that the embankments will have safety factors of 1.0 or more when the steady state acceleration is introduced. Of course, as Dr. Newmark points out, this dynamic approach is not rigorous because earthquakes loadings are transient, not steady, buc the results should be on the safe side. For earthquake loadings, the minimum permissible safety factor considered prudent by such organizations as the Corps of Engineers is 1.0 when combined with steady state seepage. DATA ATTACHED Attached to this letter are the following data that summarize our analyses:
- 1. Summary of static safety factors, seismic loading safety factors and steady state acceleration, N, producing a safety factor of 1.0 for various critical circles.
g-x
- N. Newmark, " Effects of Earthquakes on Dams and Embankments". ,
(_,) Geotechnique, Volume 25, No. 2, June 1965, p. 139-160. s 3000 59 ti til . 'J 's d k, Duke Power Company March 28, 1967 Page Four
- 2. Work sheets showing critical circles.
- 3. Work sheets of safety factor computations for static loading.
- 4. Work sheets of computations for seismic safety factor and steady state acceleration, N, producing a safety factor of 1.0 SHEAR PARAMETERS -
The shear parameters utilized in these analyses were the consolidated-undrained or R values which impose a rapid change in stress upon a soil that has .vusolidated under sustained load. The load change was applied so rapidly that no change in water content could occur even though the soils were saturated. The rate of loading, however, could not be termed " dynamic". In dynamic loading of such clayey soils, viscous forces would be mobil. zed, and therefore, the strength would be somewhat greater. Only cne loading cycle was employed. In loose cohesionless soils or sensitive clays repeated loading can cause a change in s tructure and progressive loss in strength. Previous experience with the undisturbed soils of the region, as well as the compacted soils, shows that the soils do not suffer progressive breakdown with repeated load. Therefore, the static shear parameters should be safe and the N for seismic loading will be substantially the same as for static. We will be pleased to discuss these analyses with you in detail. Very truly yours, LAW ENGINEERING TESTINGs COMPANY c2 c>- - George F -Sowgrs - Vice President Consultant GPS:bjs, '0993 60- l s i l I '
SUMMARY
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- DOWNSTREAM 1.90 1.39 0.15 1
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