ML19208B028

Forwards Responses to ACRS Subcommittee 790725 Questions Re Core Ladle & TMI-2.Response Modifies Design Presented in Rept 36A59 Floating Nuclear Plant Core Laddle Design & Safety Evaluation. Requests ACRS Meeting in Oct

ML19208B028
Person / Time
Site:	Atlantic Nuclear Power Plant
Issue date:	09/14/1979
From:	Haga P OFFSHORE POWER SYSTEMS (SUBS. OF WESTINGHOUSE ELECTRI
To:	Baer R Office of Nuclear Reactor Regulation
References
FNP-PALP050, FNP-PALP50, NUDOCS 7909180502
Download: ML19208B028 (106)

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                                                     .: 1 September 14, 1979 Mr. Robert L. Baer, Chief Light Water Reactors Branch No. 2 Division of Project Management U.S. Nuclear Regulatory Commission                               -

7920 Norfolk Avenue  ! Bethesda, Maryland 20852 E B. Ha ga Re: Docket STN 50-437; ACRS Questions 1- , c. t < on Core Ladle and TMI-2

Dear Mr. Baer:

Transmitted herewith are 20 copies of the Offshore , Power Systems responses to the ACRS Subcommittee questions  : contained in R. F. Fraley's letter to H. R. Denton dated July 25, 1979. Please note that we have not offered responses to part d. of Mr. Fraley's letter as these requests were made specifically to the NRC Staff. By copy  ! of this letter, 20 copies of our responses are being transmitted directly to Mr. Fraley for distribution within l ACRS. l Certain materi:1 in the attached responses reflects modifi-cation to the design presented in OPS Report 36A59, "FNP Core Ladle Design and Safety Evaluation". The principal changes are increased ladle volume and increased refractory l insulation on the walls of the reactor cavity. Both of these changes resulted from our ongoing evaluation of radiant upheating from the pool surface. The analyses of r diant upheating, which are described in the attached responses, are believed to be adequately conservative to show feasibility ar.d therefore to support the issuance of the Manufacturing License. Following NRC Staff review those responses which affect the present content of Report 36A59 will be retrans-mitted in the form of a revision to that report. We ask that these responses bc reviewed on an expedited basis le~ading to an ACRS Subcommittee meeting as early as October Q'O 9

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76: Ft , 2909180 5'o a y

e Page Two September 14, 1979 1979. To this end, we are prepared to offer any assistance the Staff may require. Ve truly y urs, s . Jr L P. B. Haga

       / lei Attachments CC:   R. F. Fraley (ACRS)

V. W. Campbell A. R. Collier 9 8

Offshore Power Systens Responses to ACRS Letter Dated July 25, 1979 Contents Page Introduction 2 Question a.1 6 Question a.2 (a) 12 Question a.2 (b) 13 Question a.2(c) 14 Question a.3 (a) 15 Question a.3(b) 16 Question a.3 (c) 17 Question a.3(d) 20 Question a.4 (a) 21 Question a.4 (b) 23 Question a.4 (c) 25 Question a.4(d) 26 Question a.5 27 Question a.6 29 Question a.7 32 Question b.1 33 Question b.2 36 Question b.3 37 Question b.4 44 Question b.5 46 Question b.6 48 Question b.7 52 Question c.1 55 Question c.2 60 Question c.3 63 Question c.4 66 References 67 Tables 69 - 76 Figures -

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INTBODUCTION In the Final Environmental Statement for the Floating Nuclear Plant (FNP) , Part III (Reference 1), the NBC Staff concluded that alterations of the design of the concrete shielding pad beneath the reactor vessel would be required to provide increased resistance to melt-through to mitigate the environmental consequences of the highly unlikely core-melt accident. This requirment stated:

    "The applicant shall replace the concrete pad beneath the reactor vessel with a pad constructed of magnesium oxide or other equival-ent refractory material, that will provide increased resistance to melt-through by the molten reactor core in the event of a highly unlikely core-melt accident and which will not react with core-debris to forra a large volume of gases. The pad should be as thick as practical, taking into account space availability and applica-ble design and operatire considerations, bo' not thinner than the concrete pad currently proposed. The prov a refractory material arx3 pM design should not emprmise sa .ety requirements and the applicant shall obtain NRC approval of the selected material an3 pad design prior to the start of construction of major elements of the FNP hull structure."

As a result of that requirement, OPS set forth functional and design requirements for the core ladle, and proposed a design in the OPS 'Ibpical Report No. 36A59, (Reference 2) . 'Ihe design proposed in OPS Report No. 36A59 was for the purpose ._ establishiry that a ladle coulo be in-corporated into the ENP which met the specified functional and design requirments as derived frm the FES-III. Af ter preliminary review by the NRC Staff, the proposed design was discussed with the ACRS subcommittee on June 27,1979. As a result of that meeting, ACPS transmitted questions to NRC related to the core ladle design, to core ladle material selection, and to the 'IMI-2 event. The ACRS letter (Reference 3) requested that OPS and T .I NBC address the questions in writing prior to another ACRS subconmittee meeting. As mted above, the design proposed in the core ladle design report (Reference 2) was developed specifically in response to FES-III require-ments. Ongoing evalcations of radiant upheating from the debris pool surface have shown that a substantial fraction of the reactor vessel and its contents may melt during the debris retent ion period of the ladle (about 2 days) . As a result OPS has mMified the ladle configuration to accommMate the Mditional volume of steel associated with the reactor vessel aM its internals. Figures 1 and 2 illustrate this configuration. The ladle configuration shown in Figures 1 and 2 has a volme in excess of 3 3 4,000 ft as ccnpared with the ladle volume of approximately 1000 ft for the configuration shown in Reference (2). In this configuration there is 5-1/4 feet of MgO on the bottcm ad side walls of the ladle. The additional volume has been largely provided by making the ladle deepar.

be final thickness of both the ladle aM refractory material for pro-tection of cavity side walls will te based on a refined estimate of the temperature history of the molten debris. The thermal-hydraulic processes occurring in the trolten pool and the heat transfer between the pool ard the structures above the pool are coupled ard constitute a ccmplex analytical problem. Methods are being developed to analyze this coupled problem and to provide a refined estimate of the pool surface terrperature history. To establish a configuration of the ladle and sidewalls, two pool surface temperature histories have been assumed. One is based on the black body 9 h() 6JE taperature of the pool surface needed to r&liate the generated decay heat as a function of time to cooled external su-faces. The second higher temperature history is an estimate made by Sandia Laboratories based on their testing experience.

For the pool surface temperature histories assumed, an acceptable con-figuration is one in which the concrete side walls of the reactor cavity are protected by a 2 foot thick layer of Mgo. The side wall thermal analyses in the question responses which follow are based on this re-fractory material and thickness. Concrete walls of the primary srield adjacent to the reactor vessel (where space is limited) can also be protected by an 8" layer of Zr02 backed by a 3" air gap and a 4" layer of ceramic fire as shown in Figures 1 and 2. The analytical resulte presented in the gaestion responses which follow are believed to be conservative ard show that the concrete and steel structures in the cavity side walls and in the region adjacent to the reactor vessel can be adequately protected. Substantial flexibility exists within the space beneath the reactor vessel (bounded by the m isting containment pressure boundary) to alter the configte ation of the ladle in nmerous ways such as to optimize performance as additional analytical and exp3rimental information becmes available. OPS regards this flexibility as a sig-nificant advantage in view of incomplete information regarding interactions between the molten core debris ard the ladle material at high taperatures.

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Should future information, developnents or test results indicate a need for additional ladle capacity, other configurations would be investigated such as providing additional ladle depth by bricking of the incore instrunenta-tion tube penetration or expanding the ladle into an adjacent platform canpartment. 0 ,. ( n,i7 dU  ;}:1l guestion a.1 Calculate the fraction of decay heat radiated from the pool for the propose 3 design.

Response

Fraction of Decay Heat Radiated From the Pool Following a postulated core melt accident a molten heat generating pool of core debris will fall into the core ladle beneath the reactor vessel. The walls above the ladle ard the renaining steel in the reactor vessel assembly and reactor internals will absorb heat radiated from the surface of the molten pool beirg held in the ladle. With a larger fraction of decay heat absorbed by the structures above the molten pool, less energy will be available to erode the M30 ladle. Figure 3 is a sketch of the region above the core ladle. The dimensions are approximate ard are intended to indicate the relative size of the walls and reactor vessel. Figure 4 is a ketch to show the various heat and mass transfer processes occurring in the pool and between the pool and sur-rounding structures. The thermal-hydraulic processes occurring in the molten pool and t% heat transfer between the pool and the structures above the pool are coupled and their analysis constitutes a complex problem. Cmputer prograns are presently being developed to solve the coupled problem. For present purposes the radiation losses from the pool were treated as decoupled fran the pool heat transfer processes ard were considered independently in two parts; the first to the walls above the f L ,

ladle and the second to the steel remaining in the reactor vessel and reactor internals. These are discussed in the following sections. Walls The radiant heat loss from the pool surface to the surrounding walls depeMs in part on the view factors between the walls and pool surface. The walls above the ladle will intercept approximately 80% of the energy raliated fran the pool surface, the reactor vessel ard internals inter-cepting the remaining 20%. Figure 5 is a sketch showing the numerical value of the view factors for the various regions above the pool. The decay heat stored in the walls above the pol was calculated for two postulated molten pool surface tenperature histories. The first one corresponds to that estimated by Sandia personnel in their critique of the FNP Core Ladle Design and Safety Evaluation (Reference 2) and represents the higher pool surface temperature history. The second was determined by calculatire the effective black body tenperature for the pool surface required to radiate 100% of the decay heat from the pol surface and results in the lower pool surface temperature history. Table 1 indicates pool surface temperature as a function of time for the assumed temperature histories. Initially the walls above the pool are cool relative to the pol surface ard they act as a good heat sink. As the surfaces of the walls heat up, however, they begin to radiate heat back to the pool so that the net heat transfer fran the pcol to the walls decreases. Even though the pool may be cna hh LU/ radiatirg a significant fraction of the available decay heat from its surface, eventually the majority of this heat will be radiated back to the pool if the surface temperature of the walls approaches the surface temperature of the pool. For this condition there will be very little net heat transfer frcrn the pool to the walls. Estimates of the heat stored in the walls were made by utilizing the TAP- A Camputer Code (Reference 4). A section of the wall was modeled as a one dimensional slab being heated by radiation from the pool on one surface and cooled by natural convection on the other surface. The results of the analyses for both pool surface temperature histories are shown in Figures 6 and 7. Curves for the decay heat rate ard the integrated decay heat as a function of time are shown in Figure 8. Figure 6 shows the amount of decay heat stored in the walls as a function of time for wall surfaces above the top of the ladle and below the reactor vessel. Initially, the amount of heat stored rises quite rapidly, but as the walls heat up the rate of heat absorption decreases. Figure 7 snows the fraction of the integrated decay heat released that is stored in the walls for both pool surface temperature histories. 'Ihe wall geometry is also shown on the figures. It can be seen that after two days the fraction of the decay heat stored in the walls is about 0.12 for the higher pool temperature history and about 0.08 for the lower pool temperature history. After six days these fractions are 0.062 and and 0.040 respectively. The fractions decrease with time since less of the decay heat can be deposited in the walls due to the high surface temperatures of the walls. 0 6(; Ok

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Steel The radiant t. eat from the pool that reaches those parts of the reactor vessel and appurtenances aM the reactor vessel internals remainirg af ter the postulated core melt accident will bring the steel surfaces to the meltirg noint. This molten steel will add to the core debris as it falls into the ladle. A simplified heat transfer nodel was formulated in which the pool surface temperature histories used for the wall analysis were utilized to predict the rate of melting of exposed steel. Since the exposed steel surface will be rapidly heated to its meltiry point, it was assumed in the model that the steel was radiating heat to the pool at the melting temperature of steel. Figures 9 and 10 show the result of this analysis. Figure 9 shows the decay heat absorbed in melting the steel estimated to be lef t in the reactor vessel cavity af ter the postulated core melt accident. It can be seen that, depending upon the assumed initial average temperature of the steel ad the pool surface tmperature history, the entire mass of steel could melt in the order of 0.5 to 4.0 days. Figure 10 shows that the fraction of the decay heat absorbed at the time when all the steel nas melted ranges from approximately 0.14 to 0.37. Another calculation was made in which the fraction of the decay heat absorbed by both the walls and melting of the reactor vessel steel vas estimated as a function of time for the two pool surface temperature histories. The results are shown in Figure 11. Af ter one day the heat stored in the walls and molten steel ranges from 0.24 to 0.46. Af ter six

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days these fractions have decreased to 0.11 and 0.17 respectively. The branches in the curves occur because of the assumptions of two different initial average vessel steel tmperatures. The lower vessel steel temper-ature results in a greatu fraction of the decay heat being required to melt the steel. Figures 12 and 13 show the Mao erosion rates for the ladle configuration shown in Figures 1 and 2. Figure 12 has also ins uded the effect of increasing interface area between the molten pool and MgO bed as the bed erodes. This effect was not considered in the calculations done for Figure IV-1 of Reference 2 and is the major reason for the decreased vertical erosion shown in Figure 12. The curves in Figures IV-1 and IV-2 of Reference 2 were determined by using a constant value for the volmetric heat capacity (V.H.C.) of Mg0 equal to 5 4.86x10 BTU /FT . It is the heat required to raise the temperature of one cubic foot of MgG from 100 F to the melting point of the UO -MgO 2 eutectic and melt it. We density of MgG was assumed to be 227#/Pr (no void fraction) . Figures 9 and 10 utilize a V.H.C. which is a function of pool composition. he initial value is 3.9x10 5 3TU/Pr3 and is based upon a density of 189#/Pr 3and a eutectic mel* i.ng temperature of 4136 F (2280 C). As Mgo is eroded and the pwl becomes rich in Mgo the melting point of the LD -Mg0 mixture ard the V.H.C. increase. The melting point eventually 2 approaches that of Mg0 and the V.H.C. approaches a limiting value of 4.5x105 BIU/Pr . The use of a lower density in evaluating the V.H.C. and e

the snaller mat area for the new ladle configuration are the primary reasons for the erosion rates shown in Figure 10 being higher than those shown in Figure IV-2 of Reference 2. Ej b b () { .';

Question a.2. (a) Calculate the effects of beat radiation in Item 1 on the ra'a of disin-tegration and collapse of . rosed concrete.

Response

Concrete that could be exposed to the thermal radiation from the molten pool will be protected by a high temperature insulating brick such as MgO, ZrO 2 I ^1 23 0 and possibly ceramic fibres (Fibrefrax, consisting primarily 23 and SiO of Al O 2

                    ) . This will prevent the concrete fran disintegrating, collapsing or melting. See 2e Response to Question a.2(b) .

v' // ('s tt OO V }I 'k Question a.2. (b) Calculate the effects of heat radiation in Item 1 on the rate of disin-tegration and collapse or meltire of concrete behind the 6 inch magnesite brick wall.

Response

The concrete walls above the core ladle will be protected from disintegra-tion, collapse or meltirg. Protection of concrete from disintegration is discussed in the response to Question a.4. (c) . A recent test by SANDIA indicated that concrete with basalt aggregate melts at 2200 F. ACI publica-tion, " Refractory Concrete-ACI 547R", (Reference 5) estimates the melting point of a high alumina concrete to be 3600 F. In order to lin.it concrete temperature to 2200 F, concrete walls may require some thermal insulation other than M30, such as bricks utilizing Zro nd insulating fiberboard or 2 blanket to limit temperatures at the concrete surface. Several designs appear feasible to protect the walls above the ladle frm thermal radiation from the Enol for higher of the two assumed pool surface temperature histories. The calculations in this response are for a wall configuration consisting of 24" MgO, 21" Basaltic concrete and 1 1/8" steel. This con-figuration is shown in Figure 1. Figures 14 and 15 show the tmperature distributions in the wall at 1 and 2 day intervals after the start of the Mgo-melt interaction for the two pool surface temperature histories given in Table 1. It is seen that the maximun temperature in the basaltic con-crete is less than 2200 F at two days for the higher of the assumed pool surface temperature histories. Q . Question a.2. (c) Calculate the effects of heat cadiation in Item 1 on the rate of collapse of steel fran the reactor cavity.

Response

The collapse of steel from the reactor vessel was discussed as part of the answer to Question a.1 which indicated that the vessel steel rm aining after the postulated core melt accident may be melted and a3ded to the core ladle within 1/2 to 4 days after the accident. The volume of reactor vessel steel nielted as a function of time deperx3s upon the assumptions maSe for the pool surface temperature history arx3 the average initial tenpera-ture of the steel. Remaining steel in the reactor cavity will be ther; tally shielded to prevent collapse, see the Response to Question a.3. (d) .

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OD Question a.3. (a) Discuss the consequences of Item 2 with respect to loss of integrity of superstructures.

Response

Structures above the reactor vessel which may be affected by upheating followirg vessel melt-out can be protected in the manner octlined in the answer to Question a.2 h) , if necessary. Specific protection, if required, wi: be identified durire detailed design.

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UU n17 Ui/ Question a.3. (b) Discuss the consequences of Item 2 with respect to loss of hearth capacity.

Response

The ladle volune was originally established as 980 cubic feet based on core melt constitutents given in Table IV-1 of Reference 2. Ongoing evaluations have shown that a substantial fraction of the reactor vessel and internals may melt durire the debris retention period. Table 2 lists the total weight and estimated (molten) vol ume of steel in the reactor vessel and appurtenrices. The ladle volune has been increased to approximately 4000 cu ft to accommo-date the volume of the postulated molten debris shown in Table 2 and to satisfy the assumption that horizontal melting of the ladle would proceed at approximately the same rate as vertical melting. The new ladle con-figuration has been developed within the existing constraints of the reactor cavity steel structure by altering the bend radius of the incore instrumentation from a 12 to an 8 foot radius. Offshore Power Systens believes the change in configuration of the ladle within the major constraints of the existing FNP design demonstrates considerable latitude and flexibility to respord to changes that may be dictated in the more detailed design pnase. u,

                                                                     ,     n\Q ui-Question a.3. (c)

Discuss the consequences of Item 2 with respect to impact resistance of the hearth ard its supports. Rest:g.g As discussed in the responses to Questions a.2, the only items which may impact the ladle are portions of the reactor vessel and internals. Offshore Power Systems believes that the most likely failure mode is localized melti g or rupturirg of the bottm vessel head through which fuel pellets and other core debris will slowly spill into the ladle. This failure mode would impose relatively low impact loading on the ladle. In order to evaluate nore severe physical impacts on the ladle, two extreme cases were analysed as described below.

1) Bottom Head Imoact Gross failure of the reactor vessel bottom head requires a scenario in which the bottan hemispherical head is heated to 2200 F - 2300 P. At these temperatures the bottom steel shell will be white hot with inadequate strergth to support the core debris. The bottom head was assumed to fail arourd the periphery just below the cylinder-to-head transition. When the sof tened white-hot bottan head, filled with core debris (with a total weight of approximately 200 tons) , strikes the surface of the ladle it will impact against the cold steel liner plate coverire the surface brick of the ladle. The steel liner plate will minimize damage to the magnesite refractory.

Ebb 0 l () The impact was analysed considering the bottom head as a crushable missile impacting a rigid barrier. The crushable missile model con-sists of a series of concentric cylinders starting fran the center line of the RV iJwer head ard progressing outward. Each cylinder is stopped by the ladle upon impact. The impact forcinc, function on the ladle was a step function of 2800 kips with diainishing magnitude as the impact progresses. Under this impact foicing function, the core ladle support structure will renain within its elastic rr.sistance capacity. The resulting impart pressure on the ladle was 5000 psi which is less than the col 6 crushing strength of the magnesite brick. The top layer of brick of the ladle is "1bpex S" chemically-bonded Magnesite brick, with a cold crushire strength between 6500 psi and 8500 psi.

2) Upper vessel Imoact After the bottom head has melted the upper reactor vessel could fail in such a manner that large pieces may drop into the ladle. An analy-sis was performed for the nozzle region including the head flanges on the assumption that failure occur by shearing of the nozzles. This resulted in a missile with a weight of approximately 200 tons which was assumed to fall freely and impact the ladle. The assumption of free fall is extremely conservative since the clearance between the outside of the vessel flange ard the primary shield is approximately one inch. This postulated missile is more likely to slide down inside the prima;y shield thus being slowed down by friction forces.

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The impact forcing function for this missile was a ramp function havirg a maximun force of 2300 kips which is less severe than that associated with the bottoa head impact.

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                                                         /00     is, c Question a.3. (d)

Discuss the consequences of Item 2 with respect to integrity of structurn steel members.

Response

The primary structural steel members within the reactor cavity (bulkheads, floor ard deck above) will be thermally shielded fran energy radiated from the pool. 'Ihe primary considerations are elevated temperatures at which streryth of steel deteriorates rapidly or thermally induced stresses becane a major concern. The American Institute of Steel Construction considers steel buildings to be fire resistive it a fire will not produce an average temperature in steel members greater than 1000 F (Reference 6). Ccnsequently, primary steel structures necessary to maintain the integrity of the reactor cavity will be shielded such that temperatures at the steel surface will not exceed 1000 F for the duration of the core debris retention period. For those areas where high induced thermal stresses are likely, the tempera-tures will be limited to a lesser magnitude. c ri r** [-) ( , L - ' Question a.4. (a) Discuss the stability of the 6 inch magnesite brick wall above the hearth leva aith respect to loss of brick by spalling. Respot e Any spalling -# the refractory brick wall around the ladle would oe pri-marily thermal spalling due to thermal stresses developed by unegaal rates of contraction or expansion in different parts of the refractory. The resistance of refractory brick to spalling is enhanced by proper design of tne shapes, by optimun sizing of the grains, and by close control of each step in the manufacturire process. The refractories Wich have a uniform expansion rate generally present the least difficulties when tenperatures fluctuate widely, and of these, those with the lowest total expension, as a general rule, are less mbject to thermal spalling. Both magnesite and zirconia brick exparris at a fairly uniform rate over their entire working range of temperature; however, magnesite brick does have a relatively high rate of expansion among refractories. High purity magnesite brick consisting almost entirely of pure periclase, have good spallirg resistance resulting fran optimum grain sizing, the high temperature at which they are fired and good strength at both cold and hot corditions. The Material Sciences Laboratory of Aerospace Corporation conducted small scale tests of molten UO 2 and stainless steel interactions

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with Harklase (a high purity magnesite) brick and reported in Reference 7 that " essentially no crackirg or spallation was observed." Considering the properties of high purity Mgo and Zr0 2 brick, industry experience ard small-scale testire corducted to date, Offshore Power Systems concludes that spalling of the brick wall surrounding the ladle by one " heat-up" (i.e. if any spallire does occur) will be minimal. G/'

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Question a.4. (b) Discuss the stability of the 6 inch magnesite brick wall above the hearth level with respect to differential motion with respect to the hearth, concrete walls, and anchors.

Response

The thermal expansion of the refractory brick walls above the ladle and the concrete shield walls behind them can be accomodated by interlocking expansion joints strategically located so that the walls will be free to move vertically and lorgitudinally with little or no resistance. For lateral support of the refractory brick walls both metal and ceramic type anchcrs which are standard in the construction of furnace walls can be used. The metal anchors will be used to support the walls against all lateral seismic an3 plant motion loadings throughout the life of the plant, while the ceramic anchors will be relied upon only during a reactor core melt accident af ter the temperature of the metal anchors has risen above 2000 F. The ceramic anchors could consist of dovetailed and keyed anchor bricks made of hard burned MgG and/or Zr02 anchor bricks. The anchors would extend across the air gaps behind the refractory brick and anchor them laterally to the concrete walls by fitting into vertically positioned dovetailed grooves in the face of the concrete walls.

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Service temperatures for various steels suitable for metal ties and anchors area as follows: Max. 'Ibmperature of Type of Steel Metallic Component _F_ Carbon Steel 1100 304 Stainless 1800 310 Stainles3 2000 Cr-Ni Castings 2000 Ceramic type anchors can be used for temperatures above 2000 F.

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Question a.4. (c) Discuss the stability of the 6 inch magnesite brick wall -bove the hearth level with respect to loss of concrete behird the wall by spalling, dis-integration, and melting at calculated temperatures, or at temperatures indicated in Fig. IV-6 of OPS 'Ibpical Report No. 36A59.

Response

As previously discussed in the response to Question a.2. (b) the concrete behird the refractory walls in the reactor cavity will be thennally shield-ed such that it will not disintegrate or melt during the approximate 2-day design time of the core ladle. Ioss of concrete by spalling is expected to be minimal for concrete pro-tected by walls of refractory brick because of the low heatup rate of the protected concrete. Calculations have shown the maximtra heatup rate of concrete is 75 F/hr (the corresporr3ing maximum thermal gradient in the concrete is about 150 F/in) . For cmparison , in the ASTM E119 fire test used to qualify concrete walls as fire resistant, the taperature of the exposed concrete surface expas.ed to the test flame rises from ambient to 1000 F in five minutes ard to 1700 F in one hour, a substantially larger heatup rate. p'.7 f  ;

                                                                              \I Question a.4. (d)

Discuss the stability of the 6 inch magnesite brick wall above the nearth level with respect to slagging reaction between the brick walls and melted concrete.

Response

As indicated in the introduction and in preceding Responses to Questions a.4, a layer of refractory material or other means will be provided to prevent melting of the concrete side walls of the reactor cavity thereby eliminatire the potential for slagging reaction between the refractory walls and melted coxtete. (j. !} 6 0 .b Question a.5 Discuss the fluxing of magnesite brick by siliceous material falling into the hearth.

Response

Siliceous material falling into the core ladle after core debris has entered the ladle would melt. The melted siliceous material has potential for attac'cing and dissolving the Mgo at temperatures above about 1550 to 1600 C as discussed below. The design approach is to protect the concrete side walls with refractory materials such that melting of concrete on the side walls is minimized or prevented altogether for the approximate two day core debris retention period required for the ladle. Exsmination of the Mgo-SiO2 phase diagram, shown in Figure 16, shows that the system exhibits an eutectic with a melting point of approximately 1550 C at composition of about 65 mole percent SiO 2

                                                       . At 1700 C, the liquid phase exists over a composition range from about 55 mole percent to about 70 mole percent SiO2    (Reference 8). A temperature in the range 1550 to 1600 C would therefore be required before melting of MgO would commence.

Sandia Laboratories has very recently performed a furnace test in which a sample of crushed basaltic concrete was melted in an Mgo crucible, at lower temperatures, reference (9). The furnace temperature was 1400 C. In the tests sltnping of the concrete sample started at about 1100 C and the [r

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9 sample melted at about 1200 C. Post-test examination showed the melted concrete sample ha3 invaded ard filled the pores of the F40 producity a glass like matrix. On a weight basis, at least 2 pounds of silica would be required to dis-solve a pound of Mgo. Since basaltic concrete is at most 50% SiO by 2 weight, melting of at least 4 pounds of concrete would be necessary to dissolve a pound of Mg0 frm the ladle at lower temperatures. If F40 were to dissolve at bbO-SiO 2 eutectic temperature rather than the eutectic tmperature of 2280 C assured for the Mgo-UO2 system, the energy absorbing capability of that part of the solid Mgo bed material would be reduced by about 20%. c, ( 6 b Question a.6 Discuss the properties and merits of basalt as a concrete aggregate.

Response

To discuss the properties and merits of basalt as a concrete aggregate, two other aggregates are used for cmparison purposes. These aggregates are granite and limestone, both of which are commonly used in power plant construction. Basalt and granite aggregates are classified as igneous rocks. Granite was formed in Batholiths (large intrusive bodies of magma) and underwent slow cooling resulting in large crystals, whereso basalts were formed in surface flows and as a result of rapid cooling are a very fine-grained rock. These igneous rocks are predominantly made up of SiO 2

                                                   , see Table 3 for compari-son. Granite typically contains a higher percentage of SiO    2 nd in addition usually contains quartz (SiO2 ) as one of its free crystall'.e components.

Igneous rocks are cmposed of three primary crystall ae components as follows:

1) Quartz (SiO2)

2) Feldspars which can be one or a mixture of the following:

KO 2

              . Al 0 23   . 6 SiO2  - orthoclase Na y0 . Al 0 23   . 6 SiO2  - albite Ca0   . Al 0 23   . 2 SiO2  - anorthite                                  .,
                                                                      ,ao     o-

3) MICA which is emplex silicates of A10 with K 0, MgO aM FeO.

23 2 Granite is a very hard rock. However, it exhibits a propensity to crack when exposed to high temperatures or thermal cycling largely due to the different coefficients of expansion of its crystalline constituents. Basalts are fine-grained and so tougher than granites. Basalt contains considerably less silica than granite and little or no quartz. Limestone is a sedimentary rock containing large quantities of calcium carbonates formed mainly by the shells and skeletal materials of lime-secreting plants ard animals. Limestone is a widely used building material because of its strength and good fire-resisting properties. At high temper-atures calcian carbonate breaks down to calcium oxide and carbon dioxide gas is given off. This reaction absorbs large quantities of heat and forms ar. insulatirg layer on the surface of the concrete. Concrete made from each of the above aggregates exhibits an acceptable modulus of rupture, canpressive strength, thermal conductivity and specific heat. The concrete thermal coefficient of expansion varies somewhat depeMing on the aggregate type (see Figure 17) and the canent used. Granite aggregate is more sensitive to disintegration than the others because of a rapid volume expansion of the quartz at about 573 C. Quartz at this temperature undergoes an inversion of -quartz to -quartz which causes cracking or shatterirg of the rock structure. For aggregate procerties see references 11 and 12). , 4 ', a-

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Relative to gas generation, the undesirable canponent in rock is carbonate, (either calcium carbonate or magnesium carbonate) which forms CO2 upon heatirg. As can be seen from Table 3, the equivalent CO2content of lime-stone is quite high while the others are essentially void. Thus from a gas generation starx3 point, granite and basalt are equivalent. Basalt aggregate is considered the best choice for aggregate to be used in concrete for the reactor cavity area fran the standpoint of thermal properties and gas generation. The melting point of baslat concrete is approximately 1200 C. C,.a> I f. Obb Question a.7 Discuss the pssibility of the heat flux being higher on the sides of the molten mass than on the bottm (Em conclusion for concrete melt) with melting going horizontally faster than vertically.

Response

To our Knowledge, there has never been a test which utilized Magnesia (Mgo) rather than concrete to model a sacrificial bed, which has resulted in higher lateral erosion rates. The EE results mentioned in the qJestion were obtained frm tests performed on concrete, ard it is our understanding that the reason offered by the experimenters for the observed higher lateral erosion rate was gas generation. In experiments performed on limestone concre'e at the Sandia Laboratories an increase in horizontal erosion rates has also been observed due to gas generation. No gas gen-eration has ever been observed due to the thermal erosion of Mg0, and thus any concrete test resul as involving gas releases are not considered appli-cable. A testing program designed to investigate the lateral and vertical erosion of magnesite by core debris has been subnitted to the NRC (Reference 2) . Should the results of these tests indicate unexpectedly high horizontal erosion rates, changes in the configuration of the core ladle can be made to accmodate them. niA

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e C: M l f) f r - Question b.1 Discuss the pssibility of the Upper Head Injection System releasing nitrogen into the primary systs and impedire the ability to establish or maintain natural circulation.

Response

The Upper Head Injection (UHI) system has b^en specifically designcd to terminate upper head injection reliably before N is injected into reactor 2 coolant piping. This function is accomplished by four hydraulicly operated isolation valves located outside of the contairrnent, and is illustrated schematically ?n Figure 18. The arrangement of isolatim valves and piping consists of two valves in series in each of two parallel pipes to ensure that a single failure of any component will not compromise the protective function of the system. Each of the redundant isolation valves is activated to close (from its normally open position) upon a low-water level signal from a separate instrument on the water-filled accunulator. These four level instrunents pro /ide re-dundant isolation signals to the four isolation valves. The actuation devices for the isolation valves are supplied frm redundant Class lE instrument busses that are supplied from the Class lE batteries. Each UHI accumulator isolation valve is autmatically actuated to the closed position, upon demand, by a hydraulic actuator supplied from its own independent cmpressed gas accunulator, precharged to approximately 3000 PSIG. The combination of these redundant and independent components of the () i L.

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UHI system provides assurance of proper system functional operation to prevent the nitrogen gas frm being discharged into the reactor vessel. The WI system will be tested in-plant in accordance with Regulatory Guides 1.68 and 1.79. These pre-operational tests, conducted in a manner more severe with respect to isolation valve closure than would be experienced durirg actual ac 'ent injection conditions, provide data which verify the proper function of the isolation valves. In order to minimize the amount of non-condensibles introduced to the Reactor Coolant System (RCS) durirg the normal WI delivery process, the WI system has been designed to eliminate significant water-gas surface contact area and thus minimize the absorption of N2in the water. This is accomplished by isolating the pressurized gas frm the water usirg one accumulator to hold the water and another accumu-lator to hold the gas. A membrane is inserted in the crossover line between the accunulators to eliminate gas-water contact. 'Ihe effectiveness of this design has been demonstrated by preoperational tests of the UHI 'qstem on existing plants. During these tests, water samples were taken to verify that nitrogen concentrations in the delivered water were within acceptable limits. The acceptance criterion for the water samples in the UHI preoperational tests is that they contain less than 4.38% by volume total entrained ard dissolved gas at STP. Since the total injected water vol une for UHI type plants is approximately 1000 FT , cmpliance with this N concentration criterion assures that less than 45 2 Pr of gas (at STP) is introduced to the RCS. Actual test performance has shown less than 2% total gas in test samples. This indicates that less than 25 Pf3of gas (at STP) would be introduced to the RCS oy the UHI system. p i (. mii D () UbU This voltrne will be substantially less at the prevaili.ng conditions during injection. Based on the above considerations, it is clear that the UHI system design specifically precludes the introduction of significant amounts of non-condensibles into the FCS by isolating the UHI system from the reactor vessel ef ter it.jection of the specified water volume. Further, the amount of gas introduced during UHI water delivery is small enough so as to teve a netligible impact on natural circulation. f] j }' c)({

Question b.2 Discuss the acceptability of the single failure criterion.

Response

Since its inception, the Floating Nuclear Plant has been designed to meet the single failure criterion, as currently defined in ANSI N658-1976, " Single Failure Criteria for PdR Fluid Systems', and its References, including 10CFR50, Appendix A. Of fshore Power Systems considers the application of the single failure criterion an adequate degree of con-sewatism to ensure that systems important to safety as required by Title 10, CFR Part 50 perform satisfactorily. Adherence to the single failure criterion is demonstrated throughout the Plant Design Report in the various Failure Mode and Effects Analyses.

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                                                                    ,. S o $3 Ouestion b.3 Discuss the timed sequence of events upon the loss of all AC power before core danage will result.

Resoonse Offshore Power Systems believes that a sustained loss of all AC power is an incredible event of such low probability it need not be considered in design. Furthermore, the regulations of the Nuclear Regulatory Cmrnission are explicit regardire the extent to which AC power soarces are assumed to be unavailable. These regulations (10 CFR 50, Appendix A) provide that safeguards systems shall be designed for acceptable ,oerformance with either onsite or offsite power unavailable. Even though the situation postulated by ACRS is outs: . che bourds of existing Ccanission regulations and very unlikely, Offshore Power Systems has evaluated the effects of sustained AC power loss for those areas considered important in terms of avoiding core damage. AC Power Systems Two independent offsite power sources each with capacity to supply all of the protection ard safety systens, provide imediate access circuits to the engineered safeguards busses. Emergency electric power is provided by four independent onsite diesel generators powering four independent safeguards busses. It is a design feature of the Floating Nuclear Plant that in the event of a loss of offsite A.C. power, any two of the four diesel genera-tors and associated safeguards busses are sufficient to maintain the plant

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in a safe shutdown conditicn. 'Ihese aspects of the design greatly reduce the potential for the Floating Nuclear Plant havirs to sustain a loss of both offsite and onsite A.C. power which might lead ultimately to core di Je. DC and Instrument Power Systems Safety grade D.C. power is provided by four independent 125 volt battery systens each with its own charger. Each Class lE battery is sized in accordance with a load profile based upon supplying loads for a period of two hours af ter loss of its battery charger. In the event of loss of all AC pwer, three of the four batteries could be manually stripped of all loa 3 as soon as possible. These batteries can subsequently be placed back in service one at a time as each one is depleted and at least one channel of instrumentation can be provided for approximately six or seven hours using the battery stored energy in this sequential manner. Important instrumentation, such as steam generator level, would then be available for this extended period since each of the four redundant channels is powered by a separate battery through an inverter. Battery capacity increases slightly at elevated temperatures. Hydrogen generation ceases when AC power to the charger is lest since it is directly proportional to charger current.

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                                                                        ,/   jHU fj U Auxiliary Feedwater Supply The Floating Nuclear Plant's Auxiliary Feedwater System consists of two feedwater storage tanks, one steam turbine driven pump, four motor driven pumps and associated piping, valves and instrumentation.        We four motor driven pumps are designed to meet all the safety requirments of 10 CFR 50 assuming a single active failure. We turbine driven pump duplicates the total capacity of the motor driven pumps.

In the event of a total loss of A.C. power, auxiliary feedwater is supplied autmatically to the four steam generators. The flow path frm the auxiliary feedwater storage tanks to the steam generators is open in the standby corr 3ition. The turbine driven pmp discharges to all four steam generators through a manifold arrangement with each branch of the manifold containirg a flow regulating valve, which fails to the open position. Other manual valves can be operated by local manual control. The turbine driven pmp is cooled by by-pass flow frm the auxiliary feedwater t*nrage tanks. During a total loss of A.C. Power , it is expected that the plant operators would regulate the supply of auxiliary feedwater to the steam generators while the safety grade batteries are available to equal or exceed the auxiliary feedwater requirements. We auxiliary feedwater storage tanks are capable of replacing the decay heat boil-off in the steam generators for at least eighteen hours. The initial inventory of water in the steam generators would then supply at least two more hours of boil-off. Thus, the available feedwater capacity i D o i\ h

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and mtean generator inventory are sufficient to preclude core damage for a minimtzn of twenty hours following a total loss of offsite and onsite A.C. power. Reactor Coolant System and Safe Shutdown Considerations A complete loss of both offsite aM onsite power will result in a reactor trip. Ioss of power to the control rod drive mechanisms will result in control rods dropping by gravity into the core. The reactor core will then remain in an undamaged condition as long as it remains subcritical. Adequate heat transfer and heat transport exists to enable the fuel elements to be cooled and decay heat to be carried to the steam generators. On the assumption that an extended period of operation at full power has preceded the total loss of A.C. power, the xenon transient will not recult in a loss of negative reactivity within the first 35 hours after shutdown. A total loss of A.C. power will result in loss of both component cooling water flow ard seal water injection to the reactor coolant ptnps. Seal water leakage during operation is approximately 5 gpn per pump. It is asstrned for the purpose of this discussion that an initial seal leakage flow rate of 5 gpn per pump will exist following the loss of A.C. power and will decrease as the RCS pressure decreases. This gradual loss of primary system inventory will result in falling pressurizer ptessure and level until the pressurizer is enptied and saturation corditions occur in the hot legs. 'Ihese conditions exist approximately at eight hours after the ss of all A.C. power. An analysis has been performed by Westinghouse to 7 e ;_ ( OS' determine that adequate heat transfer ard heat transport mechanisms exist for the removal of decay heat during the phase of the transient where saturation conditions exist ard boiling occurs in the core region. The results show that the core remains cooled and undamaged prior to core uncovery. Existing analyses predict that the core will rmain covered for at least 17 hours. Ventilation Considerations On a loss of AC power, the containment fan-coil air cooling units are de-energized ard not available for removirg heat frm the containment. The temperature within containment will rise as a consequence of heat input fran the RC3 and leakage through the punp seals. As long as ice remains in the ice condenser, however, this temperature cannot substantially exceed 212 F. For 5 gpn leakage through each putp seal, in excess of 60 hours would be required to completely melt the ice bed. Thus, temperature sensitive instrumentation within contairrnent which is qualified to approxi-mately 280 F would not be affected during this period. The following approximate equilibrium temperature levels were calculated for the assumed complete loss of AC power: Equili Area Temp.grium F Control Room 110 F Computer Room 150 F Process Rack Poom 110 F Turbine Driven Pump - SG 100 F Compartment #4

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The tmperature responsen were calculated assmirg cmplete loss of the air conditioning in the control building and fan-coil cooling in the safeguards empartment building housire the turbine driven pump. The rejected heat is dissipated via the rom interface boundaries and by opening the doors except in the case of the cmputer rom where the door opens to the control room and does not provide a path for the release of heat. The calculations were based on the total heat transport coefficient for the interfacing room boundaries and included the exchange of air wtlere doors can be used. Evaluation Summary In association with the assumptions mMe, the above considerations of electric power, auxiliary feedwater supply, core cooling and ventilation, lead to the following conclusions: a) Duti% the first six or seven hours following the event the core can be cooled adequately aM auxiliary feedwater flow to the steam generators can be regulated by the operator. DC power and vital instrumentation will be available to the operator durity this phase. b) Frm six or seven hours to about 20 hours, the plant will remain in a safe undamaged cordition with a continuous supply of auxiliary feed-water which is either preset or manually controlled by the operator. G

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is c) After 20 hours loss of heat sink is likely to occur due to dry out of the steam generators. d) Reactivity considerations and loss of inventory from the reactor coolant systart are not the limiting corrlitions for core damage for at least 17 hours.

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Question b.4 Discuss the reliability of the auxiliary feedwater system.

Response

A probabilistic analysis of the reliability of the Auxiliary Feedwater Systs (AFW), shown in Figure 19, has been performed utilizing the same data base used for the study recently presented to the ACRS by the NRC's Probabilistic Analysis Group. This group recently investigated the AFW reliability of 33 operat ng PWR's. As in the NRC's study, three accident scenarios were investigated for the FNP: (a) loss of main feedwater with offsite power available, (b) loss of main feedwater combined with loss of offsite power, ard (c) loss of main feedwater cmbined with total loss of AC power. For the first two scenarios above, the reliability of the FNP AEW was fourn to be in the rarge 10 -5 - 10

                                              -4 failures per daard, and for the total loss of A.C. case in the 10 -2 failures per demand category. The difference in reliability between the first two cases and the last one, is due to the fact that during a total loss of AC power, only the steam driven train is available, and thus no credit can be taken for the redundancies available in the diesel driven trains. Neverth,eless, in all three cases the FNP AEW is on the high end of the reliability scale relative to AEW Systes previously analyzed for operating PWRs.

Because of the unique redundancies present in the FNP AFW, independent failures make virtually no contribution to the overall unreliability of the v Question b.5 Discuss how H 7 buildup in the ice condenser containment is dealt with following a 'IMI event and followirg a core melt. Rc3ponse Frm the limited information availr51e at this time, Offshore Power Systems has concluded that a 'IMI event results in periodic releases of hydrogen to the containment during venting of the pressurizer while at the same time significant amounts of hydrogen may collect at high points in the reactor coolant system or remain dissolved in the primary coolant. It is further concluded that rapid ventirg followirg a hydrogen buildup in the pressur-izer will result in sufficient hydrogen mixing with oxygen in the lower empartment to form a fla:Tmable mixture which is highly likely to ignite just outside the vent. Analysis of the 'IMI event by the NRC Staff indicates that the 26 psi pressure rise resulted fran burnire approximately 226 pound-moles of hydrogen. Following a 'IMI-type event in an FNP it is likely that hydrogen muld burn just outside the pressurizer relief tank due to rapid venting of hydrogen accumulated in the pressurizer vapor space. Burning of approximately 226 pound-moles of hydrogen in the Floating Nuclear Plant containnent would result in a pressure transient of about 40 psi if burning occurs after all the ice in the ice condenser has melted. If ice melt is delayed through operator action or other events so that as much as one million punds of ice remain at the time of hydrogen burnirs, the result would be an in-significant pressure rise in the containment.

                                                                              ,.t0 {0 h During core melt the amount of hydrogen that can be produced and subse-quently burned is limited by the amount of oxygen in the contaiment. The amount of hydrogen that can be burned by the oxygen available in the contaiment is equivalent to the amount of hydrogen generated by a 100 percent zirconium water reaction. If all this hydrogen were to burn, it would release approximately 1 x 10 8 BW. Stem formed by the burning would absorb approximately 2.4 x 10 7 BW. The remaining energy would produce an excessively high pressure transient ard the contaiment would fail. If the hydrogen does not burn, the maximtra containment pressure following a core melt with 100 percent metal water reaction is estimated to be approximately 40 psig. Although this is abcne the containment design pressure, contain-ment failure would not be expected.

All portions of the containment boundary including the shell, baseplate and penetrations have been evaluated for internal failure pressure. Figure 20 shows the failure pressure estimated at twelve locations on the containment boundary. Irternal failure pressure is considered to be that pressure beyond which large deformations could occur and hence cause local con-taiment bourdary failure. In arriving at the failure pressures a maximtra limiting stress equal to 120% of the minimtn guaranteed yield stress was used. Existing normal operating stresses were incitded in computing the failure pressures.

c. ;a (> (') ,N Question b.6 Discuss how the FNP cmpensates for the difficulty, due to the remote location arti the lack of space available, in improvisire new systms and techniques in case of an accident.

Response

Offshore Power Systems disagrees with the premise that there are unusual difficulties in imprcvisirs new systms ard techniques in case of an accident on an FNP. Floating Nuclear Plants may be sited inshore, either along the coast or on a river bank. These sites have the usual motor vehicle access that land-based plants have as well as access to marine conveyances. Offshore sites do not have direct access by the motor vehicle transport mode but can not be regarded as lacking adequate transportation. Waterborne transportation offers the FNP a mode of delivery not available to many other plants. Barges can be used to transport to the FNP many bulky and heavy items that could not be transported to sme land-based plants in one piece. Among items which could be barged or towed to the FNP site are: o Fuel oil in bulk o Fresh water in bulk (for auxiliary feedwater, for example) o Barge mounted pumpirg stations o Barge mounted generators o large equignent of almost any anticipated size and weight o Tank barges for liquid wastes in bulk (ii\9 (3b The Floating Nuclear Plant is therefore more accessible than many plants because of its access to waterborne transportation. Transportation of nearly any conceivable improvised systen or cmponent is a routine opera-tion in the marine transport industry. Any accident involvirq an unforeseen sequence of events requires a degree of ad hoc reaction. Because Offshore Power Systems is a single source designer ard manufacturer of the Floating Nuclear Power Plants, superior control of design docunentation is achieved and, consequently, retrieval of needed information regardire as-built plants is much easier. Because the Floating Nuclet.r Plant is standardized there will be relatively more operating people faniliar with the total plant design (i.e. , people from several utilities) . This will obviously assist in reacting properly to an accident. Extensive models exist for every portion of the plant. This is a definite advantage because it provides greater flexibility for jury-rigging or use of improvised equipnent with less reliance on checks of the physical plant to confirm locating dimensions or measurements. The Floating Nuclear Plant layout provides considerable laydown (main-tenance) space and storage space. This space can be made available to house additional equipnent. Space is available on all levels in close proximity to the containment building which can be used for improvised systems or equipnent. Table 4, which presents a comparison of available laydown area in the ENP with the Sequoyah plant, shows that the FNP conpares favorably. It also irdicates that over 2-1/2 acres of indoor or ventilated area is available for component storage, temporary work areas, etc. Table 5 and Figures 1.2-2 through 1.2-6A present information on space iEg ()bu, availability in various buildiras and levels on the FNP. These figures which are reproduced from Figures 1.2-2 through 1.2-6A of the Plant Design Report, Amenchent 24, show the available storage ard laydown space in yellow. The modes of access to the various laydown and storage spaces (as iMicated in Table 5) are shown on the figures by bold arrows. The ENP design provides greater storage capacity tnan most land-based plants aM additional storage capability can be generated by using laydown spaces in the plant. Furthermore,17 compartments ranging from 1300 to 3000 square feet and ca prisirg approximately 25 percent of the plat'.orm volme remain unoccupied. With relatively minor structural changes to the platform in the event of an accident, these cmpartments could also be used for storage or installation of equipnent. The Floating Nuclear Plant design provides for the removal of all equipnent except for the reactor vessel aM generator stator without major cutting operations. This of course means that access is generally available throtshout the plant to accomodate additional ecuipnent. Additionally, because of its largely steel construction, additional accessways are more easily cut in the FNP than in plants with concrete walls aM floors. Redundant safeguards trains and egaipment are located in four separate canpartments, three of which are watertight up to plant elevation 136, 80 feet above the bottom of the plant. Radiation shielding is provided for each canpartment. Thus, a leak or indirect transfer of contaninated liquid in one compartment will not prohibit access to the others. c +J,6 0)hk I Transportation of People The exact number of people to be transported to the plant during an mergency depends on plant staffing level at the time of the accident and the tasks required to mitigate the consequences of the accident. Normal plant operation requires approximately 130 persons. However, following the accident there would be more than just the normal crew; for example, technical advisors, representatives of the suppliers of sme of the more critical components such as the NSSS and representatives of the relevant regulatory bodies. Both marine and air transportation modes (for example , ferries, supply vessels, crew boats and helicopters) are available for personnel transport. The capacity of typical marine vessels range from 10 people for boats to perhaps a few hundred for ferries. Any of these vessels could make a round trip to plant 3 miles offshore in an hour or less. Helicopter transport is more rapid (a few minutes) but is limited to a few people. Marine vessels and helicopters can be obtained from several sources, includirg the plant owner, U.S. Coast Guard, U.S. Navy, U.S. Air Force, and Cmmercial Marine Transport Cmpanies. As part of the normal operation of the plant, the plant owner will provide means of transporting the operating crew daily, nL 9L

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jDJ puestion b.7 Diwuss how one faces lack of flexibility for design changes due to the cmpactness and lack of available space on the FNP. Responce The question of compactness and available space has come up periodically in the ACRS review of the Floating Nuclear Plant. In an ACRS meeting on December 8,1971, during the Floating Nuclear Plant pre-application review, the ACRS suggested tt.at normal eouipnent space and clearances might be compromised . Later, during the November 7, 1975 ACRS meeting, it was conjectured by t's ACRS that there might be a reduction both in the level of fire protection and in implementation of the "AIARA" concept due to cmpactness and lack of space. The latter concecn was included in the Committee's letter of December 10, 1975 and was resolved during later meetings in April 1976. In each case our response was the same: FNP building volumes and floor space cmpare favorably with other nuclear plants; therefore, the Floating Nuclear Plant is not unusually conpact. This question suggests again that adequate space might not be available on the Floating Nuclear Plant, this time to acconmodate design changes resulting from an accident. As noted in response to Question b.6, the building volumes and floor area of the FNP cmpare favorably with the Sequoyah land-based nuclear plant. Of particular interest is the amount of unoccupied floor space on the RIP available for cmponent storage, tmporary work areas, etc. during and

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inmediately after an accident. This area is approximately 114,000 square feet, or 2.6 acres. Also, as noted in response to Question b.6, this space is accessible by heavy lift elevators, fork lift trucks and/or large capacity cranes. In tne event of back-fitting requirements, some of the space cited above could be utilized for modifications to existing systems, system additions, or new systems. Furthermere, 17 ccupartments comprising approximately 25 percent of the platform space reain unoccupied ard are presently desig-nated void spaces. These compartments range from 1300 to 3000 sq. feet in area and could be utilized with relatively minor structural changes to the platform. With respect to total plant displacement, modifications to or the addition of safety related systes would have negligible effect. The combined weight of all Nuclear Steam Supply Systems and Auxiliary Nuclear Plant Systems is less than 5 percent of the total FNP displacement. Approximately 400 tons are required to increase the draft of the ENP by one inch. Design changes in floatire structures durirg construction ard subsequent to operation are not unique and in fact is the accepted norm in Navy vessels. The Navy continually retrofits its ships for safety enhancement, mission objectives and habitability. A paramount case in pint was the "SUB-SAFE" program instituted on 19 new construction and 10 operating Fleet Ballistic Missile Subriarines as a result of the loss of the Thresher in 1963. The scope of work was quite extensive ; however, thirteen of the 19 new ships were retrofitted during construction with minimal impact on delivery p L ig c, (_ (U

                                                                               . dJ schedule and the 10 operating ships were retrofitted during the first refuelim activity, all within the existing hull and cmpar tmentation envelope. Further, the change in mission objectives for the Fleet Ballistic Missile Program resulted in approximately 30 subnarines being subjected to major retrofits during conversion from the Polaris to the Poseidon missile launch capability. Finally, the Ship Life Extension Program (SLEP) present-ly underway by the Navy to overhaul the Forrestal Class Aircraft Carrier is The estimated cost (approximately 500 million dollars per ship) of a very extensive retrofitting program to extend the service life of the ship from 25 to 35 years illustrates the magnitude of changes which can be accomp-lished on such vessels.

The above examples provide assurance that the ENP can be retrofitted to accomodate changes without degraSing the safety or arability of the plant When comparing the ratio of equignent and distributive system density between the FNP aM cmplex mar ir.e vessels, Nery vessels in particular, it is felt that the arrangement of the FNP and the segregation and separation of systes afford a high probability that needed design changes can be accomplished to satisfy future reqairements. 6 0D3

Question c.1 Discuss the effects of changing the base mat from concrete to magnesiun oxide on the probability of a major air release dtrirg a core melt accids1 t, Discuss the comparisons of probabilities and dose levels for air reltr I associated with concrete ard magnesium oxide durire a core melt acc.

Response

Sme type of containment failure must occur in order for a major atmos-pheric release to occur durirg a core melt accident. Fi re contairrnent failures modes are discussed la HASH-1400 (Reference 13), steam explosion, isolation leakage, contairment mel t-through, overpressure frm non-con-densibles or steam, and overpressure from hydrogen burning. h%SH-1400 and Appendix A of the OPS LPGS report (Reference 14), reported that the riska from airborne release for both the typical PdR and for the FNP were daninated by the V accident segence (failure of interfacity check valves) which is eqaivalent to isolation leakage. Steps have been taken in most designs, includirg the FNP, to reduce the probability of V sequence failures. For the ice condenser design, investigations subsequent to the LPGS report have shown other accident sequences ard resultire contaiment failure modes to be important contributors to the airborne risks. The most important of these are mall or intermediate pipe breaks with loss of recicculation capability, including loss of recirculation caused by failure to open the drains between the upper ard lower cmpartments after refuelirg. The associated contaiment failure mode is hydrogen burning. Steps will be taken to reduce the probability of leaving the drains closed following refuelirg (such as autmatic status iMication and/or strirgent inspection procedures) . With such an approach, small or intermediate breaks with loss of recirculation capability, caused by other factors are the major con-tributor to airborne risks. Considered below are each of the containment failure modes and the effect of utilizirg an MgO ladle rather than a concrete base mat or the airborne risk associated with that failure mode.

a. Contaiment failure frm hydrogen _urning This containment failure mode is a major contributor to the airborne risk for the ice condenser contaiment.

Contact between hot core melt debris and concrete produces large amounts of gases, primarily CO2 , CO aM water vapor. CO2 and CO are noncoMensible and collect in the containment adding to the non-condensible gas pressure. Water vapor released from the concrete bubb a s up through the melt. At high temperatures the water vapor will react with any iron not yet oxidized to produce noncondensible hydrogen.

    'Ibe use of Mg0 rather than concrete reduces generation of hydrogen and hence reduce the probability of contaiment failure by the hydrogen burning node. However, if significant hydrogen burning occurs, there 966 057

is sufficient hydrogen generated by the zirconium-water reaction to leM to containment failure independent of the material beneath the reactor vessel (See the response to Question b.5) . Iherefore use of an MgO ladle may reduce airborne risk frm this contaiment failure mode but only slightly (this discussion applies to both the loss of recirculation sequences including failure to open the drain following refueling).

b. Failure of containment isolation This containment failure mode is a significant contributor to airborne risk. The risk via this failure mode is unaffected by base mat material choice,

c. Steam explosion This failure mode is a small contributor to airborne risk. The risk via this failure mode is not affected by the choice of base mat material.

d. Containment melt-through This containment failure mode is a small contributor to airborne risk.

The purpose of the ladle is to extend the melt-through time frm a few hours to approximately two days. This delay period, during which rMioactivity is retained within the containment, permits significant D jg

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s. W decay of short lived radioactivity and hence serves to reduce the airborne dose for this containment failure mode for the FNP.

e. Containment overpressure failure mode This contaiment failure mode is a small contributor to the airborne risk. Contact between hot core melt debris and concrete produces large amounts of gases, primarily CO2, CO and water vapor. CO2 and CO are noncondensible and collect in the containment adding to the non-condensible gas pressure. Water vapor released frm the concrete bubbles up through the melt. At high temperatures the water vapor will react with any iron not yet oxidized to produce noncordensible hydrogen. 'Ihe No ladle eliminates these sources of noncondensible gases and hence reduces the likelihood of contairinent failure from noncondensible gas overpressure.

With regard to release of radioactivity to the containment atmosphere frm melt debris, use of MgO rather than concrete produces higher melt temperatures and may therefore increase somewhat the release of volatile species like Cs and Sr still in melt debris. On the other hand, gas sparging of the melt debris associated with use of concrete enhances release of the more volatile species both by strippirg of the radioactivity from the melt and by carry up of small particles of debris. These phenmena ard their role in radioactivity release are discussed on page 2 of the recent Sandia comments (Reference 15) . Sandia implied ard we believe that use of concrete probably leads to a O#"/, l) Ob9 somewhat larger source term in the containment atmosphere although evidence is far fran conclusive. In summary the choice of a concrete base mat or an MgO core ladle does not appear to have a significant effect on the airborne risk for a core melt accident for the FNP. Overall, the risk appears to be slightly less for the case where the Mgo core ladle is included. 9 6 /> 060

Question c.2 Discuss the consideration given to the use of a vented containment. Discuss the consideration given to the use of sea water for venting and/or cooling a molten core.

Response

An evaltation of the use of a vented containment in the FNP design was performed in early 1977 an3 reported in Apperdix F of the OPS LPGS Topical Report, (Reference 14). In the LPGS Report it was concluded that in a relative sense, the cost of the contaiment vent systs per unit benefit is at least 7 times that of the containment itself and its associated engineered safety features. Furthe , the contaiment vent systen would not significantly reduce either the probability or dose commitment for the larger category 1 and category 2 contaiment airborne releases considered in WASH-1400 (these resulting from containment failure via steam explosion or isolation leakage) . The evalua-tion presented in the OPS LPGS Report showed Category 1 and Category 2 releases were the dominant contributors to overall airborne risk. Offshore Power Systems therefore concluded that the benefits provided by the containment vent system did not warrant the additional cost. As discussed in response to Question c.1, subsequc investigation has shown that containment failure due to hydrogen burning (leading to Category 3 contairment airborne releases) is probably the principal contributors to airborne risk. For this case a practical vent system would not have C(f

                                                                          'oJ O L f)l

sufficient vent capacity to accomodate hydrogen burning and protect con-tairment so that the dose comittment reduction afforded by a vent system would also not be very large. The conclusion of the LPGS study thus remains valid. An evaluation of materials and systems for cooling and retaining a molten core (incitdig use of sea water for cooliry) was performed in 1974 and male available to both the NRC and ACRS at that time, (reference 16). Several methods and materials for molten core retention and retention device cooling systems were evaluated in that report which was prepared for OPS by Battelle Columbus Laboratories. The two conclusions from that report pertinent to the question at hand were:

1) The presence of a large heat sink represented by the water surrounding the FNP is not, by itself, a particlar advantage. The ability to arrest ard quench a molten core is determined by heat remeval from within the mass of core and structural debris.

2) Several core catcher concepts have been identified which may be capable of quenching and arresting a molten core for either a floatire or land-based plant. None of these concepts can be demonstrated to be feasible at this time. Each would regaire substantial R&D to demonstrate feasibility without assurance of successful design.

o,vr, tJ 0(32 OPS has reviewed referrnce 16 in light of information developed since 1974 and believe that the two conclusions listed above are still applicable. While substantial new information exists, there are still no concepts demonstrated to be feasible and substantial R&D would still be required to denonstrate feasibility. j' Question c.3 Discuss the change in position for allowing the ENP to be placed on riverine ard estuarine sites. Has the proposed installation of the core ladle changed the NRC Staff's position on this matter, if so why? hhat actions and in what time period, are considered practical to isolate the core for a riverine or estuarine site? Resconse There has been no change in the NRC position relative to siting ENPS at riverine ard estuarine sites. The NRC has in fact consistently held that siting of FNPs at riverine and estuarine sites was feasible. Additional sitire conditions for estuarine and riverine sites were imposed in the Revised Draft Environmental Statement Part III of May,1978, refer-ence (12) . These conditions resulted fran the NRC evaluations contained in the revised and expanded Liquid Pathways Generic Study, reference (18) . The revised Draft Environmental Statenent, Part III, concluded:

    "The siting of Floating Nuclear Plants in estuarine and riverine waters is precluded unless such sites are appropriately modified in an environmentally acceptable manner so as to insure timely source interdiction of radioactive material, and limit the introduction of such material into the surrounding water body in the event of a core-melt accident."

Shh Ob The present policy as stated in the Final Envirormental Statment, Part III, p. xiv (reference (1) is:

    "We staff, therefore, concludes that finding acceptable FNP sites in estuaries, rivers or near barrier islands will most likely be extremely difficult, but cannot conclude that there are no accept-able estuarine, riverine or barrier islands locations for FNP emplacement when appropriate mitigative actions are taken.

Applicants applyirg to the NRC for a license to locate and operate an ENP at such sites would have to demonstrate appropriate mitigative actions that would provide both an acceptable level of environmental impact as well as an acceptable level of core-melt accident risk." The core ladle requirement and the regairement that riverine and estuarine sites have an essentially impermeable enclosure are regarded by OPS as indepervient requirements. We view the core ladle requirement as intended for ocean sites where the an impermeable enclosure is not required. Mitigative actions including post-accident source isolation within the enclosure for an ocean site are treated in Section 8 of the OPS LPGS report, reference (14) . For riverine and estuarine sites, the NRC requirement that the plants be surroun3ed by an essentially impermeable enclosure facilitates source isolation in the event of a core melt accident. We basin can be isolated frm the adjacent surface water body by simply closing the intake conduits for the condenser circulating water system. The intake conduit can be a pipe designed with a closure gate that muld seal off the basin in a relatively short time (hours) . 6 I) 0 6 "a Therefore, source isolation prior to melt-through can easily be accmp-lished without the core ladle in the event of a core melt accident for riverine and estuarine sites. Consequently, the ladle requirment is considered applicable only to ocean siting where a permeable breakwater is utilized.

                                                             'iu>I li i

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Question c.4 Discuss the NRC Staff's position that the ENP Core Ladle is considered an envirormental issue and not a safety issue.

Response

OPS does not believe that the reqJirement for a core ladle is technically justifiable on the basis of either public safety or environmental pro-tection. Further, we do not believe that cost-benefit evaluations of risk reduction support incorporation of the ladle into the FNP design. This position and supporting technical information wcre submitted to NRC in our conments on FES-III, OPS letter from J. A. Nutant to NRC of 6/30/78, reference (19). It is our further belief that consideration of Class 9 accidents is precluded by present NRC Regulations. However , following an appeal by Offshore Power Systems, the Canmission ruled in December,1978 that consideration of Class 9 events in the FNP environmental review was permissible and appropriate under NRC regulations. OPS has therefore chosen to canply with the NRC reauirement of FES-III that a core ladle be included in the ENP design to reduce the environmental risk via ligaid pathways associated with core melt accidents. We support the NRC position that additional design features are not required for reasons relative to public health and safety to reduce the risk associated with core melt accidents. O4fs O. [j 7 References

1) Final Environmental Statement, Part III, Related to Manufacture of Floatini Nuclear Power Plants, NUREG 0502, 12M 8.

2) ENP Core Ladle Design and Safety Evaluation, OPS 'Ibpical Report 30A59, 4/79.

3) Letter from R. Fraley (ACRS Executive Director) , to H. Denton (Direc-

    ' or , NRC) ACRS Review of Floating Nuclear Plant Core Ladle Design, 7/25/79.

4) TAP-A, "A Program for Cmputim Transient or Steady-State Temperature Distributions," WANL 'IME-1872, December,1969.

5) Refractory Concrete, ACI Cmmittee 547 Report No. ACI 517R-79.

6) Steel Construction Manual, AISC, 7th Ed. ,1970, Section 6-9.

7) Annual Progress Report - Ex-Vessel Core Catcher Materials Inter-actions, D.G. Swanson, et al., 10/30/76, Aerospace Corps., Report

    # ATR-77 (7608)-1.

8) E.M. Irvin, C.R. Robins, and H.F. McMurie, " Phase Diagrams for Ceram-ists," The American Ceramic Society, Inc. , Columbus, Ohio (1964) .

9) Private Cmmunication, 7/30/79, Dr. Dana Powers, Sandia Laboratories.

10) SAND-74-0382, " Core Meltdown Experimental Review," Sandia Laboratories, August, 1975; Chapter 9.

11) C.L. Mantell, Engineering Materials Handbook, McGraw-Hill, 1958, Chapter 23.

12) " Thermal Properties of Concrete Under Sustained Elevated Tempera-tures," Nikolai G. Zoldners ACI Publication SP25 Taperature and Concrete, 1971.

13) WASH 1400, Reactor Safety Study, USNBC.

14) OW Liquid Pathways Generic Sttdy 'Ibpical Raport No. 22A60, June 1977.

15) Lett.er from R. L. Baer (NRC) to A. R. Collier (OPS), Comments on Topical Repcrt No. 36A59 by Sandia Laboratories, 6/15/79.

16) Degraded Accident Studies for Floating Nuclear Power Plants, OPS Topical Report No. 'IME 0002, 12/74.

17) Revised Draft Environmental Statement, Part III, related to the manufacture of Floating Nuclear Power Plants, NUPEG-0127, Rev. 1, 5/78.

(;g 068

18) Liquid Pathways Generic Sttdy, NUREG-0440, 2/78.

19) Letter frcxn J.A. Nutant (OPS) to Director , Division of Safety and Enviromental Analysis (NBC) , OPS Ccmnents on Revised Draft Environ-mental Statement, Part III, 6130178.

c f

                                                                          /

Oh TABLE 1 POOL SURFACE TEMPERATURE HISTORIES TIME (DAYS) TEMPERATURE

1. Sandia Estimate 0 4712 F (2600 C) 1 3632 F (2000 C) 0 2 3524 F (1940 C) 4 3308 F (1820 C) 6 3092 F (1700 C)

TIME (DAYS) TEMPERATURE

2. OPS (Black Body) Estimate 0 3641 F (2005 C) 1 2394 F (1312 C) 2 2232 F (1222 C) 4 2092 F (1144 C) 6 1952 F (1067 C) c.

                                                                   /uUu 070 TABIE 2 REACTOR VESSEL AND INTERNALS WEIGHT AND VOLUME 3

Source Weicht (lbs) Volume (ft ), Reactor Vessel 700,920 1752 Reactor Vessel Head 159,500 399 Studs, Nuts, Washers, etc. 45,100 113 Iower Internals 252,000 630 Upper Internals 153,000 382 Bottan Mounted In-Core Instr. 12,000 30 CRDM's 74,000 185 R.V. Insulation 1,000 3 U.H.I. Piping 8,300 21 Lifting Rig & Seismic Support 72,000 180 IUrAL 1,477,820 3695 (*) Steel density of 400 #/ft is assumed. D f(>d f.

                                                                   /

O\ TABIE 3 CHEMICAL COMPOSITION OF AGGREGATE (Weight Percent) GRANITE GRANITE BASALT (GEORGIA) (TEXAS) LIMESKNE SiO 2 52.4 76.0 70.2 14.1 TO i2 'I Al 0 23 15.8 13.1 17.4 1.8 Fe O 23 11.2 .9 3.2 .8 FeO - Ca0 9.9 1.1 1.5 40.6 MgO 6.7 4.5 KO 2 4.7 2.9 .6 Na O 2 2.7 3.9 4.3 .6 HO 2 1.0 1.1 CO 2 35.6 99.7 99.7 99.5 99.8 SOURCE Ref. 10 Ref. 11 Ref. 11 Ref. 10

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                                                                  , i c f r)  -

TABLE 4 SPACE COMPARISON - FNP VERSUS SEQUOYAH I. NUCLEAR AUXILIARY BUILDINGS ( } Sequoyah FNP*

1. Total Floor Space 2.03 x 10 g,2 5

2.88 x 10 5 ft.2

2. Total Volme 4.87 x 10 6 ft.3 5.32 x 10 6 g,3

3. Available Laydown Area 25,700 ft.2 67,700 ft.2 II. cot 7fROL BUILDING ( ' Sequoyah FNP*

4

1. 'Ibtal Floor Space 2.80 x 10 ft.2 1.49 x 10 4 ft.2 6 6

2. Total Volme .95 x 10 ft.3 .23 x 10 ft.2

3. Available Laydown Area N/A 5425 ft.2(5)

III. 'IURBINE BUILDING (4} Sequoyah FNP*

1. 'Ibtal Floor Space 1.70 x 10 5 ft.2 2.08 x 10 5 ft.2 6

2. 'Ibtal Volume 5.5 x 10 ft.3 8.81 x 10 6 ft.3

3. Available Laydown Area 54,700 ft.2 46,000 ft.2 (1)The plants have been separated into three building entities and only spaces under roof have been considered. Sequoyah I & II total flow area and volme were reduced by the following factors to a3just for the two units and make the numbers cmparable to the one unit FNP.

1. Nuclear Auxiliary Buildings - 40%

2. Control Building - 40%

3. Turbine Building - 50%

The "available laydown area" n mber is that area of floor space available for the location of improvised systems and/or equipnent storage in the event of an mergency. (2)For the FNP, the spaces incltx3ed were (a) all space between elevation 100 and 154 between colmn lines 1.5 to 5 and B to J, except for con-tainment, (b) radioactive waste treatment space, upper head injection space, charging space, and safeguards space belos 100, und (c) general stores spaces at 100 and 112. At Sequoyah the Diesel Generators are lo-cated in a separate building; this approximate space has been included. II The battery roms are located in the control building at Sequoyah, this space has been included in the totals. FNP battery roms are located in the safeguards compartments. (4)The FNP floor area includes the area where the CW pumps and intakes are located. At Sequoyah these itens are located separately. (5) Space in Emergency Relocation Area (for personnel) . , , TABLE 5

SUMMARY

OF SPACE AVAILABLE ON THE FNP AVAILABLE FLOOR FLOOR TO CEILING ACCESS ELEVATOR FNP ELEVATION & LOCATION AREA (SQ FT) IIEIGHT, FT. RATING (LBS) AUX & CONTROL BLDG SAFEGUARD AREA #1 SPACE #64-B2-01 790 36'-0" 4000 ELEV 77-C2-01 1150 22'-6" 64 77-C2-01 280 22'-6" AND TOP PIPE TUNNEL 1000 22'-6" 77-6 Y' SAFEGUARD AREA #2 77-F2-01 390 22'-6" 4000 SAFEGUARD #3 77-12-01 1150 22'-6" 4000 280 TOTAL 5040

     .3
  ,h i   's N

4

TABLE 5 - CONT'D. AVAILABLE FLOOR FLOOR TO CEILING ACCESS ELEVATOR FNP ELEVATION & LOCATION AREA (SQ FT) llEIGHT, FT. RATING (LBS) AUX. & CONTROL BLDC. ACCESS BY 100' LOADING DOOR ELEV. 1. CHEM. LAB AREA 1980 NO ELEVATOR REQUIRED 100 2. SPACE BELOW CASK PIT 575

3. SPACE ADJACENT TO CNMNT

                        & CONTROL BLDG                                   18'-0"

4. MAIN STEAM AREA 575 TOTAL 8530 GENERAL STORES 6230 12'-0" TURBINE BLDG. 4680

1. WASTE HANDLING AREA 760

2. SPACE UNDER H. P. 1080 24'-0"

3. SPACE UNDER GEN. 1400

4. SPACE IN MUW AREA 1440 TOTAL 4680 ELECTRICAL BLDG.

s, 1. SHUNT REACTOR AREA 4800 42'-0" <?s CJs 2. XFM BAY 1800 3, TOTAL 6600 v ELEV. GENERAL STORES 12100 12'-0" 14,000 [2. 112

TABLE 5 - CONT'D. AVAILABLE FLOOR FLOOR TO CEILING ACCESS ELEVATOR FNP ELEVATION & LOCATION AREA (SQ FT) IIEICilT, FT. RATING (LBS) AUX. & CONTROL BLDG. ELEV. 1. SPACE ADJACENT TO CNMNT

                       & CONTROL BLDG               5260                 18'-0"         14000 118

2. SOLID WASTE AREA 600 10000

3. CONFER RM 500 14000 TOTAL 6360 ELEV. TURBINE BLDG 8600 34'-0" 4000 & 14000 200 TON TURBINE HALL CRANE CAN 124 ACCESS VIA A liATCll

 ~

m

  '                AUX. & CONTROL BLDG.

ELEV. 136 1. MAIN STEAM AREA 600 14000

2. WASTE IIANDLING AREA 880 18'-0" 10000

3. SPACE @ CNMNT & CONTROL BLDG 4200 1400

4. RECREATION AREA 2520 TOTAL 8200 ELEV. ELECTRICAL BLDG. EL 5500 32'-0" 4000-IIATCH 142 so b\

C', C N. C,'

TABLE 5 - CONT'D. AVAILABLE FLOOR FLOOR TO CEILING ACCESS ELEVATOR FNP ELEVATION & LOCATION AREA (SQ FT) HEIGHT, FT. RATING (LBS) ELEV. AUX. & CONTROL BLDG. 154 1. FUEL BLDG 9320 18'0" TO 77'-0" 14000 250 TON AUX. BLDG CRANE CAN ACCESS 2060 10000 5500 SQ.FT.

2. SPACE ADJACENT TO CNMT &

CONTROL BLDG 3600 18'-0" 14000 5 TON NEW FUEL CRANE CAN ACCESS

3. SPACE OVER S.G. 6250 14000 750 SQ. FT.

TOTAL 21230 TURBINE BLDG. 20600 68'-0" 4000 & 14000 200 TON TURBINE ELEV. HALL CRADE CAN ACCESS 158 c[n so CrN C'. CJ NJ

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FIGURE 15

1900 ~

        -PERICLASE + LIQUID        1890 j

1850 LIQUID 1800 ~ FORSTERITE + LIQUID \ , LIQUID m-TWO \ PERICLASE l 170 0 -

                  +

LIQUIDS"\ 1695 o

                                                                              //_
                                                                                    /

FORSTERITE

                                                    / CRISTOBALITh
                                                           + LIQUID CRIST0BALITE 1600                                                        +   LIQUID FORSTERITE
                 +                       0 CLIN 0ENSTATITE ,s       1557                            0
                                 \             \ CLINOENSTATITE
             '     '           '                \ + CRIST,0BALITE 1500                      '            '                '

O 20 40 60 80 100 Mg0 2 Mgo-SiO2 SiO2 CLIN 0ENSTATlTE

                                                          + LIQUID Ref:  E. M. Levin, C. R. Robbins, H. F. McMurie,
                        " Phase Diagrams for Ceramists," The American Ceramic Society, Columbus, Ohio, 1964
                                                                       'l l. (>       ()'; b Figure 16 Mg0-SiO2 Phase Diagram

4 i i . .

 !             !        l                !             !

f i - i FIGURE 17 .

 !                                 THERMAL EXPANSION OF i             !                   VARIOUS AGGREGATES 3.0                                         ,

1 1 1 I

2.5 I I

I I 2.0 e O I 5 1 g f LIMESTONE 1.5 5 3 GRANITE I E I 1.0 BAS \LT 0.5 Is A 200

                          /            400
                                                  *l 600     800 0                                                              1000 (392)              (752)        (1112)   (1472)           (1832)

TEMPERATURE OC ( F) J6 091

INSIDE OUTSIDE CONTAINMENT CONTAINMENT V HI h 3/4" PI ' 12"

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t dk L mu [S] STRAINER jh I , Hg **W gg 4 ' MAKEUP [% JL " DRAIN [ D H 2 REFUELING WATER MAKEUP

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