Feasibility And Scaling Analysis For Simulation Of Star-lm Using Maslwr Test Facility, J. Groome, T. Becker, A. Aldrich, T. Wagnon, Q. Wu, J. N. Reyes Jr.

Feasibility and Scaling Analysis for Simulation of STAR-LM using MASLWR Test Facility

J. Groome, T. Becker, T. Wagnon, A. Aldrich, Q. Wu, and J.N. Reyes, Jr. Department of Nuclear Engineering, Oregon State University, Corvallis, OR 97331 J.J. Sienicki Nuclear Engineering Division, Argonne National Laboratory, 9700 S. Cass Ave, Argonne, IL 60493 ABSTRACT – Few large-scale integral system test facilities exist that are capable of conducting a single loop natural circulation fluid analysis for Liquid Metal Fast Reactors. Oregon State University currently has a Multi-Application Small Light Water Reactor (MASLWR) test facility that is capable of performing steady-state and transient natural circulation analysis. A scaling and feasibility analysis has been performed to determine the suitability of using MASLWR to model the Secure Transportable Autonomous Reactor Liquid Metal (STAR-LM) concept developed by Argonne National Laboratory. Based on cost, availability of property data, and material compatibility, lead was selected for the liquid metal system option while a water system was also analyzed for scaling. The natural circulation scaling analysis revealed that MASLWR facility could be used to scale steady state natural circulation in STAR-LM with either Pb-Bi eutectic or water as the working fluid. I. INTRODUCTION The U.S. Department of Energy (DOE) generation IV nuclear energy system initiative has selected six advanced nuclear reactor systems for future research and development. One is the Lead Fast Reactor (LFR). The LFR systems employ either a Pb or a Pb-Bi eutectic (LBE) alloy liquid metal coolant and a fast-neutron spectrum. The Secure Transportable Autonomous Reactor – Liquid Metal (STAR-LM) underdevelopment by Argonne National Laboratory (ANL) is a 400 MWt LFR that uses a lead or Pb-Bi eutectic (LBE) coolant and advanced power conversion system. The power conversion incorporates a gas turbine Brayton cycle utilizing supercritical carbon dioxide to achieve higher plant efficiency. In this study, feasibility and scaling analysis is presented for the simulation of the STAR-LM thermal hydraulic system using the existing MASLWR test facility at Oregon State University, which is a U.S. DOE sponsored integral test facility for natural circulation cooled light water reactors. The STAR-LM is a 400 MWt natural circulation reactor concept (Figure 1). It employs several design features that provide autonomous load following, passive safety, and proliferation resistance. The design utilizes lead as the primary coolant, which allows it to forgo intermediate heat transport medium requirements. STARLM achieves autonomous load following whereby the core power adjusts itself according to the heat removed without user intervention. This design feature makes it passively safe, in that it provides an inherent core shutdown in the event of a loss of heat sink accident.

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In order to fully understand the operation mechanism of a fast reactor cooled by liquid metal natural circulation, certain thermal hydraulics characteristics need to be examined, and the relevant experimental investigations carried out in properly scaled test facility. In particular, the following thermal hydraulic areas have been identified as of significant importance: • Heat transfer in core and heat exchanger; • Liquid metal natural circulation startup, steady-state operation, and shutdown; • Coupling of reactivity feedback and natural circulation; • Thermal stratification in lower/upper plenum regions.

Figure 1

STAR-LM concept

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These phenomena could be simulated using the integral test facility of OSU scaled to the MultiApplication Small Light Water Reactor (MASLWR, Figure 2) designed by INEEL, Bechtel, and Oregon State University, if the test system is properly modified and the relevant scaling analysis is performed. MASLWR was designed to be a small modular natural circulation light water reactor with enhanced safety and economics. The design uses an integrated reactor and steam generator, both housed in the primary pressure vessel. The MASLWR design employs a high-pressure containment vessel that sits in a pool of water as the ultimate heat sink. A one-third height scale model of the MASLWR power reactor was built to assess the steady-state and transient operating characteristics. This model is capable of simulating full power operations at prototypic pressures and temperatures. Its pressure vessel is constructed from SS 304 with a shell side design pressure of 11.38 MPa (1650 psi) at 316°C (600°F) and a Core Simulator consisting of a 56-rod heater bundle capable of producing 670KW of thermal power. A 14-Tube Helical Coil Steam Generator (boiling inside the tubes) has been incorporated into the reactor vessel. The tube side design pressure and temperature are 2.07 MPa (300 psia) and 288°C (550°F) respectively. The test facility also includes a containment tank capable of 2.76 MPa (400 psia) pressurization during ADS and steam vent valve operation and an external pool that acts as the ultimate heat sink for the containment. Vent valve

Water

Containment Reactor pressure vessel Steam Turbine generator

Steam Depressurization valve

Gen

Feedwater Core

Sump makeup valve

Condenser

Steam generator tube bundle Water Feedwater pump

Figure 2

MASLWR concept

For the conversion of the MASLWR model into a STAR-LM thermal hydraulics integral test facility, this investigation focuses on the facility modification needs and the relevant scaling analysis, with the following specific tasks: • Identifying a suitable fusible alloy, preferably of low or no lead content, to use as the liquid metal for this system. The selection of the liquid metal will be based on cost, availability of property data, material compatibility, and regulatory concerns.

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•

•

Plant modifications to support work including an external liquid metal fill and drain system, an external heat trace system, piping and valve replacement, and liquid metal compatible instrumentation for diagnostics. A natural circulation analysis of MASLWR and STAR-LM including a heat transfer analysis for MASLWR’s helical coil steam generator. Using the results to carry out a scaling analysis for the liquid metal and water systems. II. FACILITY MODIFICATIONS

There are several major modifications that must be done to the existing MASLWR system. These include the working fluid, possible structure enhancement, an external fill and drain system, a heat trace system, piping and valve replacements, and instrumentation. II.A. MATERIALS STAR-LM is a liquid metal reactor concept utilizing either Pb or LBE as the working fluid. As is well documented, the regulatory concerns dealing with lead and most notably with the melting of lead, tend to discourage the consideration of Pb or LBE in the modeled facility. MASLWR utilizes water as the working fluid and being able to accurately scale STAR-LM while maintaining the water environment was considered optimal. However, an appropriate liquid metal option was also researched to not only ensure scalability to STARLM, but also due to the attraction of a liquid metal test facility with the ability to test phenomena that otherwise could not be accomplished with a water test facility. Liquid metals considered included Sn-Bi eutectic, Woods metal, Fields metal, and LBE. Of particular interest during the research was an examination of the cost of the metal, availability of property data relevant to the scalability of the metal to the properties of LBE, material compatibility, and any regulatory compliance issues. The cost of metals varies rather dramatically per kilogram depending upon the material. LBE is a relatively inexpensive material and it would be undesirable to replace it with a very expensive material for the purposes of a test facility. Fields metal is a mixture of bismuth, tin and indium of which 51% is indium. Indium is an extremely expensive metal with a cost nearly twice that of silver. Woods metal contains 50% bismuth, 25% lead, 12.5% tin and 12.5% cadmium, none of which are prohibitively expensive; however, cadmium is toxic (as is lead) and is thought to be a cancer risk. The cost of tin is more expensive than lead but not by an appreciable amount. In order to analyze a theoretical model there must be sufficient data to perform the analysis. It was quickly established that there was a considerable lack of property

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data for every liquid metal option except LBE. Specifically temperature-dependent data on density, viscosity, thermal conductivity, and the coefficient of thermal expansion were unavailable in the liquid metal literature for all metals except LBE. Argonne National Laboratory has provided temperature-dependent correlations for the following LBE properties: density, viscosity, thermal conductivity, heat capacity and the coefficient of thermal expansion. If another liquid metal is to be analyzed then fundamental research must be conducted to remedy the lack of available property data. The existing MASLWR facility is made of Type 304 SS. While this is acceptable for water applications, when dealing with liquid metals, significant difficulties are encountered. Dissolution corrosion, the dissolving of the various components of the alloy steel into the liquid metal, is the main type of corrosion encountered. Operating temperatures will range from 300-500°C, which from Table 1 [1] seems to indicate that LBE will corrode the steel the least. As shown in the table, good is defined as a corrosive rate of attack of less than 2.54 cm/yr, limited as 2.54-25.4 cm/yr, and poor as greater than 25.4 cm/yr. It is clear from the table that at high temperatures, the addition of tin results in higher rates of steel corrosion. Wood’s metal, Field’s metal and Sn-Bi all contain tin to some extent. Table 1

800 °C 600 °C 300 °C

Corrosion Properties for various metals with Austenitic Stainless Steels Sn Pb Bi-Pb Bi-Pb-Sn Unknown Poor Poor Poor Poor Poor Limited Poor Limited Good Limited Limited

Dissolution of the steel metals can be countered by creation of an oxidation layer on the steel. This effect has been studied [2] for use in LBE environments. The oxidation layer serves two purposes depending on oxygen potential: oxides are insoluble in LBE and thus the outer oxidation layer of the steel will remain intact, and the oxide layer impedes diffusion of the metal components toward the steel/LBE boundary, thus reducing dissolution of these metals in the LBE. Another possibility is that given the fill and drain tank system, which will be discussed later, the liquid metal will not spend significant amounts of time inside the vessel and therefore will not have the opportunity to significantly degrade the SS-304 components of MASLWR. Based on the preceding information - cost, availability of property data and material compatibility— it was elected to proceed with LBE as the liquid metal option in the test facility. From a regulatory perspective lead has been well documented. The two main references used here are 29CFR1910.1025 and 29CFR1926.62. The regulations therein are well documented and detailed, and

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would require multiple facility modifications to the MASLWR building at OSU. It would require a training program be established for all individuals who work around or might feasibly come into contact with the lead. Adequate ventilation systems must be established and maintained, and respirators must be worn if any airborne concentration of lead is present. To this effect air monitoring must be established throughout the facility to monitor an employee’s exposure to lead in the atmosphere. Coveralls to protect an individual from physical contact with the lead must be worn at all times in the facility and OSU would be required to provide services to clean said clothing. Many of these requirements while cumbersome would not be extraordinarily difficult given the nature of OSU’s facility and its attachment to the Radiation Center. The radiation center already is required to have many similar systems in place for dealing with radioactive material. In addition the local fire department is already provided additional training in dealing with accidents involving the Radiation Center and would require more training for dealing with lead-related accidents. II.B. Fill and Drain System Some of the features of an external fill and drain system are as follows: a storage tank for holding the entire system volume of liquid metal, a storage tank design that allows for liquid metal expansion upon cooling, and a method of filling and draining volume into and out of the MASLWR system. The external fill and drain tank is designed to hold the entire primary system volume (0.265 m3 ,70 gal.) and utilizes inclined sides to allow for the expansion of the liquid metal upon cooling. Pressurized argon gas is utilized to provide the necessary motive force to fill and drain liquid metal from MASLWR while providing an inert atmosphere to minimize corrosion. To calculate the pressure needed to pump the volume of liquid metal into the system, the pressure drops across the system need to be found. As illustrated in Table 2, the main pressure drop is due to pumping the LBE to the height of the MASLWR vessel (4.27m). Calculations were based on filling the MASLWR pressure vessel (0.265 m3) in a half-hour. Table 2

Pressure drop calculations for pumping LBE into MASLWR system Pressure Drop kPa (psi) Vessel fluid height: 4.57m (15') 475.3 (68.93) Piping: 6.1m (20') of 2.54cm (1") piping 1.9 (0.28) 8 elbows & 3 globe valves 12.5 (1.81) TOTAL 489.7 (71.0) The design utilizes 15.6-MPa (2265-psi) tanks of argon reduced to 0.79 MPa (115 psi) to pump the LBE. The argon is preheated with an inline heater such that it

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will not cause solidification of the LBE. For filling the MASLWR system, LBE will be forced from the storage tank and up into the MASLWR vessel from the bottom. For draining, the argon will enter the vessel from the top, forcing the liquid metal out the bottom of the vessel and into the storage tank. When the system is shut down and drained of the liquid metal, argon will be utilized to keep the system pressurized slightly above atmospheric pressure to minimize corrosion. II.C. Heat Trace System. The heat trace system is designed to heat both the steel and LBE to above the liquid metal melting temperature for both the test facility and all external piping systems of the fill and drain system. The heat tracing is designed to be used during both shut down and non-operational conditions and to maintain system boundary conditions during testing. Heat tracing will be installed around all liquid metal external piping, the pressure vessel, and the storage tank. Calcium silicate insulation will be utilized to minimize heat loss on all heat-traced systems. The heat tracing system is designed such that the heat transfer at the wall of the piping, vessel, and storage tank is negligible (i.e. the liquid metal is not losing heat to the surroundings). The heat-tracing requirement for each of the cylindrical components can be calculated using the following equation or the integral form of the equation for the fill tank:

QD =

T H − TC ln(r2 r1 ) ln (r3 r2 ) + 2πk ss L 2πk cs L

(1)

where TH is the temperature of the liquid metal taken to be 175 °C (approximately 50°C above the melting temperature of LBE), TC is ambient temperature (20°C), r1 is the inner radius of the cylindrical component, r2 is the outer radius of the cylindrical component, r3 is the radius of the cylindrical component including the calcium silicate insulation, L is the length of the component, kss is the coefficient of thermal conductivity for the steel (20 W/m-K), and kcs is the coefficient of thermal conductivity for the calcium silicate (0.047 W/m-K). An estimate of the heat loss rate through the fill tank caps was made by treating the pipe caps as slabs of steel with 10.2cm (4”) of insulation. The following equation is relevant to the fill tank caps. ∆T TH − TC = QD = L L RT ss + cs k ss AT k cs AT

(2)

where TH, TC, kss, and kcs are defined as above; Lss is the thickness of the steel pipe caps, Lcs is the thickness of the

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calcium silicate insulation, and AT is the total surface area of both pipe caps. Finally, the power required to heat the MASLWR system and fill tank, 1896kg (4180 lbm) steel, from ambient temperature (20°C) to 175 °C as well as melt the LBE, 2700kg (5952 lbm), and heat it to 175 °C were calculated based on a four hour heating period with the following general equation:

D =  Cp ⋅ m⋅ ∆T + Q   t solid 

Cp ⋅ m⋅ ∆T  +    t liquid 

∆h⋅ m t

(3)

where Cp is the heat capacity, m is the mass, t is the heating time, ∆T is the change in temperature and ∆h is the latent heat of melting (38.8kJ/kg for LBE). Note that for the MASLWR system and the fill tank only the first term of the equation is applicable since the system does not melt. See Table 3 for heat trace requirements. Table 3 Heat Trace Requirements Heat input to component MASLWR & fill tank heat up to 175°C LBE melt and heat up to 175°C Maintaining LBE at 175°C MASLWR vessel Piping Fill Tank Side Fill Tank Caps

kW 9.59 11.17 ~ 0.83 0.75 3.72 0.19

II.D. Piping and valve replacement

Most of the current MASLWR system is composed of Type 304 stainless steel (SS-304). Because this type of stainless steel (austenitic) is corroded fairly easily in an LBE environment, pre-treating the SS-304 with an oxidation layer is required to minimize corrosion. For external piping systems that are easily replaced Type 405 Stainless Steel (SS-405), a ferritic-martensitic stainless steel that is more corrosive-resistant in a LBE environment would be installed. In addition to corrosion, the expansion of the LBE as it cools must be dealt with. Many of the valves that are currently in the MASLWR system trap liquid when closed. If these valves were left in the system when the plant is cooled down, the liquid metal will be trapped in the valves possibly causing the valves to rupture upon expansion. These valves would be replaced with valves that will not trap liquid metal. II.E. Instrumentation

The pressure, temperature, and flow detectors that are currently in the MASLWR primary loop will have to be replaced with instruments that are compatible with the high temperature liquid metal environment. The current

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temperature thermocouples are easily replaced with thermocouples that are made for non-ferrous metals. High temperature pressure transducers are rarely made anymore. Thus, there would need to be a separation between the transducer and the liquid metal. A bellows type system can be used for this separation. The sensing line will contain the liquid metal, and a bellow will separate that from another material, such as NaK, that will be in direct contact with the transducer. The sensing line, however, cannot be allowed to cool down; thus, insulation and heat trace will need to be applied to the sensing lines. In order to measure primary system loop flow, a VCone® differential pressure element system will be employed using the same pressure measurement process described above. III. SCALING ANALYSIS

The scaling analysis is to examine the similarity between the primary loop natural circulations in STARLM and the MASLWR model for single-phase fluid flow condition. The results of this analysis will be employed for the selection of the core power and loop resistance values in MASLWR model that would best simulate the primary loop natural circulation phenomena in STARLM. For the natural circulation cooling of STAR-LM, Figure 3 provides a diagram that describes the scaling analysis process. Single-Phase Loop Natural Circulation Phenomena

mD = ρ iui ai = ρ c uc ac

ρ

du c dt

 ac  ai

N

∑ li  i =1

(4)

  = β g ρ (TH − TC )H th  N   a fl 1 − ρuc2 ∑  + K   c i =1 2  dh  i  ai

Cv M sys

d Tsys dt

  

(5)

2

= mD C P (TH − TC ) − (ua∆T )HEX − QD loss

(6)

The nomenclature for these equations is provided at the end of this paper. For the purpose of scaling study, the following assumptions were made: 1. The flow is one-dimensional along the loop axis. 2. Negligible axial heat conduction. 3. The fluid is incompressible. 4. Boussinesq approximation is applicable, i.e., To obtain the dimensionless form of those balance equations, the following nominal parameters are used as the scaling references: Mass flow rate: mD 0

(T

− TC )0

Temperature: Fluid velocity through heat exchanger:

u HEX , 0

Loop heat loss rate:

Qloss , 0

H

Substituting those parameters into the control volume momentum and energy balance equations, we have:

Bottom-Up Scaling • Loop Resistance • Core Power • HEX Heat Transfer

Top-Down Scaling • Volumetric Flow Rate • Core/Heat Exchanger Energy exchange

balance equations of mass, momentum and energy of the closed loop are given by:

N/C Similarity Criteria

τ loop

τ loop

 duc+ (u + )2  + = ε ne (TH − TC ) Π Ri − c Π F  dt 2   d Tsys+ dt

+ + = mD + (TH − TC ) − Π HEX (u + ∆T + )HEX − Π loss QD loss (8)

N l τ loop = ∑   , i =1  u  i

Evaluate Scaling Distortion

(7)

ε ne =

τ N

2 ∑ (l u )i (ac ai ) i =1

β g(TH − TC )0 Lth , Π Ri = u c2, 0

Specify Necessary Modifications

Figure 3

Scaling analysis flow diagram

III.A. Scaling Criteria

The loop being considered consists of core that serves as a heat source, an in-vessel heat exchanger that functions as the heat sink, and the riser and down-comer. The fluid in the loop is simply divided into a hot fluid side having an average temperature TH, and a cold fluid side having an average temperature TC. The control volume Global 2003

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Π HEX =

(ua∆T )

HEX , 0

Qcore , 0

,

N  fl  a Π F = ∑  + K   c i =1 D  h  i  ai Q Π loss = loss , 0 Qcore , 0

  

2

It is necessary to obtain an expression for the loop natural circulation flow rate (or velocity) for the closure of the dimensionless groups. This can be done by assuming the momentum transient in the control volume is small comparing to the buoyancy and frictional/form pressure loss terms. Thereafter, the fluid velocity in the core section can be expressed as

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1

3

(9)

In terms of the simple one-dimensional analysis, for perfect simulation with identical time scale, a model system should have the same non-dimensional groups as in the prototypic system. Such kind of scaling is extremely difficult for complicated systems. Certain scaling distortion is inevitable. Therefore, a rational scaling approach should preserve the phenomena of high priority in the bottom-up scaling analysis. Πloss and ΠHEX are adjustable and thus may not be crucial for the system scaling. Also, εne, the ratio of the loop transport time to an area averaged loop transport time may be of low priority for relatively slow transient. In this investigation, the MASLWR integral test facility already exists. The following geometric scaling ratios are fixed:

(A )

c R

(D )

= 1 322 ,

h, c R

= 1 2.38 ,

(l )

c R

= 1 3.35 ,

(l ) HEX

R

= 1 4.09 ,

(L )

th R

(∀)

R

= 1 3.62 ,

= 1 51.7 .

The subscripted “R” denotes a MASLWR to STAR-LM ratio. In the following sections, scaling criteria for both LBE and water to simulate the prototypic LBE fluid will be discussed.

III.B. Scaling Ratios for MASLWR LBE Loop Using LBE in the MASLWR model at roughly the same system temperature, the properties are the same as the prototypic system. With a power ratio QD c R = 1/530, the ratio of the fluid velocity in the core is 1/1.81, which makes the scaling ratios of friction number (ΠF) and Richardson number (ΠRi) close to unity as presented in Table 4. The fluid transport time (τloop) ratio is only ½, which suggests that the fluid particle in the MASLWR model takes only half of the prototypic time to travel through the entire loop. When comparing the transient data, stretching the time axis of the model test data by a factor of 2 will produce excellent match to the prototypic transient. Using the results from the scaling analysis a comparison was made to determine how well the scaling factors match the data. That is, the scaling factors can be used along with the MASLWR data to predict STAR-LM operating conditions. In Figure 4, the STAR-LM coolant velocity in the core is compared with the MASLWR scaled results with the velocity and power being up-scaled by appropriate factors. Figure 5 provides the similar comparison for coolant temperature change through the core. These plots indicate that the scaled MASLWR

( )

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results matches reasonably well to the calculated STARLM results. However, when consider the bottom-up scaling, certain local processes may not be properly scaled. One of such processes is the heat transfer in the heat exchanger section. Perhaps the helical coil heat exchanger in the MASLWR test facility needs to be replaced with U-tube type heat exchanger in order to preserve the prototypic process.

III.C. Scaling Ratios for MASLWR Water Loop Using water to simulate LBE present a greater challenge for the scaling analysis, due to the significant differences in fluid property. Through trivial iterations, it was found that a power ratio of 1/1148 would result in a very close to unity ratios of friction number (ΠF) and Richardson number (ΠRi). All the relevant ratios are also summarized in Table 4. Because of the high thermal expansion coefficient of water comparing to LBE ( β R =19.14), much less core power is needed to simulate the STAR-LM. The scaling ratio of the core coolant velocity is about 1/1.21, and the corresponding loop time scale is roughly 1/3. In other words, the fluid transport time through the entire MASLWR loop is only one third of the prototypic value. The major distortion may come from the heat transfer process in core and heat exchanger section, because of the fluid property difference. This is inevitable due to the selection of water as the working fluid. Nevertheless, the transient phenomena of the natural circulation in the prototypic STAR-LM system can be simulated. The comparisons of in-core fluid velocity and temperature change through core are presented in Figure 4 and 5. a close match is achieved when the MASLWR parameters are scaled up to the prototypic conditions. 0.7

0.6

Core Velocity (m/s)

 2 βQD core Lth g   uc =   a C ρΠ  F   c p

0.5

0.4

0.3

STAR-LM

0.2

LBE: MASLWR Scaled STAR-LM

WATER: MASLWR Scaled STAR-LM

0.1

0.0 0

50

100

150

200

250

300

350

400

450

500

Core Power (MW)

Figure 4 STAR-LM and MASLWR-scaled core coolant velocity versus power

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MASLWR system. Heat tracing can be used to maintain the liquid metal temperature, as well as warm up the system although more extensive heating units may be desired for melting the LBE. Piping that will carry the LBE liquid metal should be replaced with a ferriticmartensitic stainless steel for better corrosion resistance. Diagnostic instrumentation that is compatible with LBE can replace the existing instrumentation.

200 180

Delta T--Core (°° C)

160 140 120 100 80

STAR-LM 60

LBE: MASLWR Scaled to STAR LMMASLWR Scaled to WATER:

40

NOMENCLATURE

STAR LM

20 0 0

50

100

150

200

250

300

350

400

450

A a Cp Dh f g h K l, L Lth m Msys QD t T u β µ Π ρ

500

Core Power (MW)

Figure 5 STAR-LM and MASLWR-scaled temperature change in core velocity power Table 4:

Scaling ratios (model : prototype) Scaling Ratios LBE Water 1:2 1:3 Time (τloop)R Velocity (uc)R 1 : 1.81 1 : 1.21 1 : 530 1 : 1148 Power ( QD )R c

Core Reynolds number (Rec)R Richard Number (ΠRi)R Core Length (lc)R Core Hydraulic Diameter(Dh,c)R Core Area (ac)R Thermal Length (Lth)R Volume (∀)R Friction coefficient, (ΠF)R Thermal Expansion Coefficient βR Density ρR Dynamic Viscosity µR Heat Capacity (Cp)R

1 : 4.31 1 : 1.00 1 : 3.35 1 : 2.38 1 : 322 1 : 3.62 1 : 51.7 1 : 1.00 1:1 1:1 1:1 1:1

1 : 2.38 1 : 1.01 1 : 3.35 1 : 2.38 1 : 322 1 : 3.62 1 : 51.7 1 : 1.01 19.4 : 1 1 : 13.7 1 : 16.6 36.3 : 1

V. CONCLUSION

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Area [m ] flow area [m2] heat capacity [J/kg Co] hydraulic diameter [m] Darcy friction factor local gravitational constant [m/s2] convective heat transfer coefficient [J/m2 Co] minor pressure loss coefficient length [m] Thermal center height [m] LBE density [kg/m3] Total mass of the system [J/m2 Co] heat transfer rate (J/s) time [s] temperature [Co] coolant average velocity [m/s] coefficient of thermal expansion fluid dynamic viscosity [kg m/s] dimensionless group fluid density [kg/m3] REFERENCES

1. 2.

The MASLWR facility, with a few changes, can be used to model STAR-LM’s natural circulation behavior. The preliminary scaling analysis suggests that the test facility with LBE or water can be scaled with proper power selection. For a MASLWR LBE system, a ½ time scale is necessary, whereas a 1/3 time scale for MASLWR water system. Certain local distortions may be inevitable, especially for water system where local heat transfer mechanism could be affected by the fluid property differences. The facility structure itself is adequate as built. For use of LBE, more research into using active oxygen control to maintain an oxidation layer that protects the structure from corrosion should be pursued in addition to determining what substantive corrosive damage could occur in the MASLWR system during the limited LBE exposure allowed by the external fill and drain system. The fill and drain system meets the requirements for moving the LBE into and out of the

2

3.

4. 5. 6.

Liquid Metals Handbook (2d ed. revised), Atomic Energy Commission, Department of the Navy, Washington, D.C., 1954, pp. 184 – 212. D. KOURY, “Investigation of the Corrosion of Steel by Lead-Bismuth Eutectic (LBE) Using Scanning Electron Microscopy and X-Ray Photoelectron Spectroscopy.” Thesis for Graduate College, University of Nevada, Las Vegas, December, 2002. C.F. COLEBROOK, “Turbulent Flow in Pipes with Particular Reference to the Transition Region between the Smooth and Rough Pipe Laws.” Proc. Inst. Civil Eng. 11, 133 (1939) I.E. IDELCHIK, Handbook of Hydraulic Resistance (2nd ed.). New York: Hemisphere, 1986. F.P. INCROPERA and D.P. DEWITT, Fundamentals of Heat and Mass Transfer. (5th ed.), New York: John Wiley & Sons, 2002, p. 419 S. KALISH and O.E. DWYER, “Heat Transfer to NaK Flowing through Unbaffled Rod Bundles”, Int. J. Heat Mass Transfer, 10, 1533-1558 (1967)

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