Scholarly article on topic 'Superheated Water-Cooled Small Modular Underwater Reactor Concept'

Superheated Water-Cooled Small Modular Underwater Reactor Concept Academic research paper on "Chemical engineering"

Share paper
Academic journal
Nuclear Engineering and Technology
OECD Field of science
{"Advanced Reactors" / "Nuclear Design" / Safety / "Small modular reactor (SMR)" / Superheat}

Abstract of research paper on Chemical engineering, author of scientific article — Koroush Shirvan, Mujid Kazimi

Abstract A novel fully passive small modular superheated water reactor (SWR) for underwater deployment is designed to produce 160 MWe with steam at 500ºC to increase the thermodynamic efficiency compared with standard light water reactors. The SWR design is based on a conceptual 400-MWe integral SWR using the internally and externally cooled annular fuel (IXAF). The coolant boils in the external channels throughout the core to approximately the same quality as a conventional boiling water reactor and then the steam, instead of exiting the reactor pressure vessel, turns around and flows downward in the central channel of some IXAF fuel rods within each assembly and then flows upward through the rest of the IXAF pins in the assembly and exits the reactor pressure vessel as superheated steam. In this study, new cladding material to withstand high temperature steam in addition to the fuel mechanical and safety behavior is investigated. The steam temperature was found to depend on the thermal and mechanical characteristics of the fuel. The SWR showed a very different transient behavior compared with a boiling water reactor. The inter-play between the inner and outer channels of the IXAF was mainly beneficial except in the case of sudden reactivity insertion transients where additional control consideration is required.

Academic research paper on topic "Superheated Water-Cooled Small Modular Underwater Reactor Concept"

Nuclear Engineering and Technology xxx (2016) 1—11

Available online at ScienceDirect

Nuclear Engineering and Technology

journal homepage:

Original Article

Superheated Water-Cooled Small Modular Underwater Reactor Concept

Koroush Shirvan* and Mujid Kazimi1

Department of Nuclear Science and Engineering, Massachusetts Institute of Technology, 77 Massachusetts Avenue, Cambridge, MA 02114, USA



Article history:

Received 23 December 2015 Received in revised form 16 April 2016 Accepted 8 June 2016 Available online xxx


Advanced Reactors Nuclear Design Safety SMR


A novel fully passive small modular superheated water reactor (SWR) for underwater deployment is designed to produce 160 MWe with steam at 500°C to increase the thermodynamic efficiency compared with standard light water reactors. The SWR design is based on a conceptual 400-MWe integral SWR using the internally and externally cooled annular fuel (IXAF). The coolant boils in the external channels throughout the core to approximately the same quality as a conventional boiling water reactor and then the steam, instead of exiting the reactor pressure vessel, turns around and flows downward in the central channel of some IXAF fuel rods within each assembly and then flows upward through the rest of the IXAF pins in the assembly and exits the reactor pressure vessel as superheated steam. In this study, new cladding material to withstand high temperature steam in addition to the fuel mechanical and safety behavior is investigated. The steam temperature was found to depend on the thermal and mechanical characteristics of the fuel. The SWR showed a very different transient behavior compared with a boiling water reactor. The inter-play between the inner and outer channels of the IXAF was mainly beneficial except in the case of sudden reactivity insertion transients where additional control consideration is required.

Copyright © 2016, Published by Elsevier Korea LLC on behalf of Korean Nuclear Society. This is an open access article under the CC BY-NC-ND license (


1. Introduction

Traditionally, the pursuit of a higher temperature and harder spectrums are the two key paths taken for improving the current light water reactor (LWR) technology. While the latter is focused on improving fuel utilization, the former is focused on improving the economics of the reactor through an increase in power conversion efficiency. With the current vast uranium reserves, improving the economics of a nuclear

reactor is of more importance than fuel utilization. In the initial phases of the development of the current LWR technology, the concept of a superheat water reactor (SWR) was explored in the 1950s and 1960s, with only a few years of operational experience accumulated in the US, Sweden, Soviet Union, and Germany. The main motivation behind the SWR concept is to produce superheated steam at approximately 500-600°C, to increase the thermodynamic efficiency of LWRs that produce steam at approximately 300°C. A 200°C

* Corresponding author.

E-mail address: (K. Shirvan). 1 Deceased.

1738-5733/Copyright © 2016, Published by Elsevier Korea LLC on behalf of Korean Nuclear Society. This is an open access article under the CC BY-NC-ND license (

2 Nuclear Engineering and Technology xxx (2016) 1—11


AEC US atomic energy commission

BWR Boiling water reactor

CRD Control rod drive

DHRHE Decay heat removal heat exchanger

EBT Emergency boron tank

iPWR Integral pressurized water reactor

ISP Internal suppression pool

IXAF Internally and externally cooled annular fuel

LBFR Lead-bismuth fast reactor

LOCA Loss of coolant accident

LWR Light water reactor

MSIV Main steam isolation valve

PWR Pressurized water reactor

RIP Reactor internal pump

RPV Reactor pressure vessel

SMR Small modular reactor

SWR Superheated water reactor

increase in steam temperature results in an approximately 4% increase in efficiency (or 10% relative increase) with an ideal simple steam Rankine cycle [1].

Superheating steam has been used in the power industry for many decades. In addition to an increase in efficiency, the turbine performance is improved by avoiding the presence of droplets in saturated steam. The turbine blades can go through erosion and pitting when exposed to water droplets at high speeds, thus reducing their lifetime. The higher steam temperature will also have broader applications such as industrial process heat and liquid fuel production. The former US Atomic Energy Commission funded many prototype

superheat reactors. Other countries also started similar programs. These nuclear systems can be divided into two categories: (1) superheating within the core; and (2) superheating outside of the core. In the latter design, superheating was achieved by use of fossil fuels. The pressurized water reactors (PWRs) or boiling water reactors (BWRs) were coupled to a coal- or oil-fired power plant for increased thermal efficiency. In the former case, the steam from a BWR was used as the input to the reactor core which was cooled by the superheated steam and moderated by light water. The superheat core was either integrated in the same core or a separate reactor. In both categories, the operation reliability (e.g., capacity factor) of both nuclear and fossil power plants was not as high as today [1]. Table 1 lists the various nuclear reactor designs that produced superheated steam as a product. In order to withstand high temperature steam, many of these designs adopted steel alloys which were popular by their construction time instead of zircaloy. However, no significant operational experience was gain and fuels were taken to very low burnups.

Improving the economic performance of small modular reactors (SMRs) is key for allowing their deployment. The Felxblue, a 160-MWe SMR designed by DCNS, is an off-shore underwater reactor using the traditional PWR technology and layout [15]. While Flexblue utilizes technology-ready and proven systems, the design is within a large expensive hull with the LWR pedestrian thermodynamic efficiency of 33%. In order to increase both the compactness and efficiency of the Flexblue design, the integral superheater class of SWRs is considered for this study. Specifically, the use of a SWR conceptual design by Ko and Kazimi in 2010 [1] is investigated. The reason the earlier integral nuclear superheater concepts or other conceptual integral SWR designs discussed in literature were not chosen for such an application is that they all

Table 1 - Design characteristics of the nuclear power plants with superheat [1-14].

Reactor Designer Moderator Coolant Thermal Efficiency Steam Material

power (%) exit Fuel Clad

(MWt) temp.fC)

A. Nuclear power plant with fossil-fired superheater

Elk River AC H2O H2O 58.2 (N) 14.8(F) 30.8 441 UO2- ThO2 SS

Indian Point I B&W H2O H2O 585 (N) 215(F) 32.0 538 UO2- ThO2 SS

CVTR West. D2O D2O 65 29.2 385 UO2 Zr-4

B. Nonintegral nuclear superheater

EVESR GE H2O Steam 17 — 493 UO2 SS

C. Integral nuclear superheater

BORAX-V ANL H2O H2O & STEAM 35.7 (T) 454 UO2 (B) 304 SS (B) 304 SS (S)


BONUS GNEC H2O H2O & steam 50 32.6 482 UO2 (B) UO2 (S) Zr-2 (B) 316 SS (S)

Pathfinder AC H2O H2O & steam 203 30.5 441 UO2 (B) Zr-2 (B) 316L SS (S)


APS-1 USSR Graphite H2O & steam 30 (T) 299 U-alloy SS

Beloyarsk-1 USSR Graphite H2O & steam 285 33 500 U-alloy SS

Beioyarsk-2 USSR Graphite H2O & steam 457 35 500 U-alloy SS

Marviken Sweden Graphite D2O 593 33.7 475 UO2 (B) Zr-2 (B) Inconel (S)

UO2 (S)

HDR Germany H2O H2O & steam 100 25.0 457 UO2 (B) SS (B) Inconel (S)

UO2 (S)

B, boiler; F, fossil fuel power plant; N, nuclear power plant; S, superheater; T, test reactor.

Nuclear Engineering and Technology xxx (2016) 1—11

suffered from power/flow mismatch stability issues, since they isolated boiling in one region of the core and superheating the steam in another region.

In the Ko and Kazimi design [1], this is solved by using internally and externally cooled annular fuel (IXAF). The coolant boils in the external channels throughout the core to approximately the same quality as a conventional BWR and the steam is separated from the liquid with the use of traditional separators. This saturated steam, instead of exiting the reactor pressure vessel (RPV), turns around and flows downward in the central channel of some IXAF fuel rods within each assembly and then flows upward through the rest of the IXAF pins in the assembly and exits the RPV as superheated steam as shown in Fig. 1. The Ko and Kazimi [1] study can be referred to for a more detailed discussion on the design's operational performance, including startup. The IXAF fuel concept has been extensively studied in the 2000s for both PWRs and BWRs [16] and is currently being irradiated in Korea to assess its application for providing power uprates to the OPR1000 PWR design [17]. In addition to compactness and an increase in efficiency relative to a PWR technology, SWR technology cannot be readily used for propulsion because of power oscillations induced by ship motion due to operation with a two-phase flow and susceptibility to flow instabilities. This will give the SWR an edge in terms of dual use resistance of a nuclear technology for offshore seabed deployment.

Fig. 1 - Flow configuration of the superheated water reactor design by Ko and Kazimi [1].

2. Assembly design

For this study, the assembly of the SWR has the same dimensions as an earlier Massachusetts Institute of Technology (MIT) design [1], as shown in Fig. 2. While the dimensions are not fully optimized, at the power density of 50 kW/L, the design meets the critical power limits on its outer channels and produces a steam temperature of approximately 510°C (783 K) in the internal channels. In this study, a lower power density of 40 kW/L is assumed. This is to account for the expected higher local peaking factors which results from the smaller core along with a desired 5-year fuel cycle length compared with the original design of a 400-MWe reactor core by Ko and Kazimi [1].

At 40 kW/L, the steam outlet temperature is estimated to be at 750 K using the RELAP5 code which is approximately 40 K lower than the previous analysis by Ko and Kazimi [1]. In the previous analysis, a constant pressure in the steam channels was assumed in the calculations, resulting in an over-estimation of the exit temperature by 16 K. The other reason for such a discrepancy is that changes in the fuel thermal conductivity, as well as the gap conductance results in a different estimation of the steam outlet temperature, since it changes the power split between the outer and inner channels. Two increases in the inner or outer gap conductance results in an approximate 50-K change in the steam outlet temperature. While the steam in each stage of superheating is collected in a single plenum region, the changes in fuel and gap temperatures need to be carefully calculated and predicted in order to accurately assess the system outlet temperatures as well as the local safety margin of such core design. Fig. 3 shows the calculated axial temperature profile of steam and the inner channel outer surface cladding for the average powered assembly and a hot assembly with radial peaking factor of 1.45. For these simulations a cosine shaped power profile with peak of 1.55 is assumed.

Fig. 3 shows that at the beginning and end of the axial length of the fuel, there is energy transferred from the steam to the fuel, since the fuel heat generation rate is so low. This also implies that the outlet steam temperature is also sensitive to the axial power profile. Additionally, with regards to the dimensions of the gap, fuel thickness can be adjusted to minimize the heating of the top of the fuel by steam. Since the RELAP model assumes a constant gap conductance and does not account for material dimension changes as a function of temperature (e.g., thermal expansion), the FRAPCON-Annular code [18] was utilized to compare with RELAP calculations. FRAPCON-Annular, which is able to model dual cooled fuel, simulated SWR fuel geometry for both the steam downflow and steam upflow pins to understand the evolution of cladding-gap and other fuel pin thermomechanical properties as a function of burnup. A conservative simulation using a linear heat generation of 32.6 kW/m with the RELAP assumed axial power profile over 5 years was simulated. The cladding in this simulation was FeCrAl alloy, which was chosen due to its superior corrosion performance as discussed in the following section. As shown in Fig. 4 (left), the inner gap remains relatively stable while the outer gap decreases, similar to what is observed in a regular PWR. As seen, the maximum annular fuel

Nuclear Engineering and Technology xxx (2016) 1—11

Fig. 2 - The superheated water reactor assembly design specifications. D, diameter; P/D.

temperature is lower than a typical solid pin PWR/BWR fuel (>1,200 K), leading to lower fission gas release and low plenum pressures (approx. 5 MPa at end-of-life). The fuel temperatures calculated by RELAP5 are also in the same range as calculated with the FRAPCON code that accounts for more detailed material properties. Sensitivity of the outlet temperature over time to the assumption of a cosine and bottom-peaked cosine power distribution and a fuel-clad gap was found to be within approximately 10 K and considered negligible. The effect of a gap evolution on the steam/cladding temperature was more significant for the steam downflow pins compared with steam upflow pins, as shown in Fig. 4 (right).

3. Cladding options

3.1. Corrosion

In this section, the options for cladding material for the SWR design are discussed. The current LWR cladding is required to meet the limit of 17% equivalent cladding oxidation at the end of its core residence life. Also, there are transient limits that dictate the allowable cladding oxidation thickness. Therefore, under a steady state operation, it is desirable for the cladding

to go through oxidation at levels well below the 17% accident limit. From Figs. 3 and 4, it can be inferred that the cladding material needs to at least withstand 1,000-K steam for a 5-year fuel cycle. Alloys T91 and Inconel 718 options were considered by Ko and Kazimi [1], due to a reported favorable experience in a radiation environment. Three other materials are considered as cladding in this work: SiC/SiC ceramic composite and FeCrAl, which are currently under development by the accident tolerant fuel program in the US for LWR applications [19] and Type-310 stainless steel that is known to have very high oxidation resistance. Table 2 lists the percent cladding consumed by oxidation during 5 years of operation in steam at their respective conditions. The table also lists the references for the oxidation models which are empirically based on a few day's oxidation data used for each cladding. For SiC oxidation, the data from Robinson and Smialek [20] goes up to 15 atm, while the data for Pint et al. [19] is only for atmospheric pressure. The flow velocity is only approximately 0.1 m/s in the Pint et al. [19] data, while in Robinson and Smialek [20] it is up to 24 m/s. The model from Robinson and Smialek [20] predicts the oxidation rate reported by Pint et al. [19] for alpha phase SiC with reasonable accuracy. Unlike metals, the SiC oxidation rate is dominated by its volatilization by forming silicon hydroxide species. This volatilization rate is dependent

Fig. 3 - The axial temperatures for the peak assembly for the superheated water reactor design.

Nuclear Engineering and Technology xxx (2016) 1—11

Fig. 4 - The evolution of fuel-cladding gap (left) and maximum fuel and cladding temperature for the downflow/upflow/ superheated water reactor pins.

on velocity as well as pressure and the partial pressure of oxygen in the steam. The SWR has pressures up to 73 atm and a steam velocity of 35 m/s. Therefore, in the absence of experimental data, the model from Robinson and Smialek [20] was used outside its range of applicability. As shown, SiC-based cladding will most likely not be feasible to be used in a SWR for 5 years. Even though the oxidation data for metals in Table 2 are at atmospheric pressure, it is expected that they would go through a similar rate of oxidation at the high pressures of the SWR. This is at least true for the FeCrAl alloy, since the necessary oxygen potential for the oxidation of Al is so low that at even very low oxygen partial pressures it is expected to oxidize. Although there were no data found on high pressure oxidation of Type 310 stainless steel cladding, the data for other steel alloys such as T91, shows no effect of steam pressure on oxidation at the temperature range of a SWR [21], though T91 is a ferritic material.

As listed from Table 2, the FeCrAl and Type 310 stainless steel show the most promise as a cladding material to withstand high degrees of oxidation at the conditions of interest. No decisive conclusion can be drawn for SiC in the absence of experimental data at conditions of interest, although it is expected that higher pressures and a higher velocity increase its corrosion rate in steam.

3.2. Enrichment requirements and cost

The assembly shown in Fig. 2 was modeled in a commercial reactor physics tool, CASMO4 [24]. CASMO4 is able to produce a reactivity decrement per unity energy extracted from a SWR assembly. Assuming neutron leakage of 6% [25] for the SWR core, the enrichment required for each cladding option to meet a 5-year cycle with 40 kW/L power density can be calculated. The FeCrAl and Type 310 stainless steel cladding results in average core enrichments of 8.83% and 9.4%, respectively. The Type 310 stainless steel, due to its nickel content, requires 0.6% higher enrichment compared with FeCrAl. Compared with a PWR of similar power output and cycle length with standard zircaloy cladding, the Type 310 stainless steel enrichment is 4% higher. The higher enrichments results in fuel costs of 3,566 $/kgU and 3,814 $/kgU for FeCrAl and Type 310 stainless steel cladding, respectively, estimated with the Ux fuel cost calculator (http://www.uxc. com/tools/FuelCalculator.aspx). This is a 70% and 80% increase in fuel costs for FeCrAl and SS-310 compared with a PWR fuel. The levelized cost is not as high as the fuel cost due to the higher thermal efficiency and specific power. The FeC-rAl has a levelized cost of 12.4 million/kW/h and the Type 310 stainless steel has levelized cost of 13.3 million/kW/h, compared with 9.5 million/kW/h for a PWR of a similar once-through fuel cycle. The 30% and 40% higher levelized cost for the claddings could result in a 6% or 8% higher overall lev-elized cost of electricity, if the fuel cycle cost is assumed to be approximately 20% of the total levelized cost of the plant.

Since no SMRs of this type or land based have been built, the capital cost (45%) and operation and maintenance cost (35%) typically assumed for LWRs [26], might not be valid and more uncertain than the fuel cycle cost. The core internals of the SWR also need to be rated at approximately 1,000 K, most likely made of Type 310 stainless steel, since it is a more mature material than FeCrAl. According to AK Steel Corporation (, the cost of SS-310 is approximately

Table 2 — The superheated water reactor cladding option

oxidation during 5 years of assumed operation.

Material Conditions Cladding lost (%) Reference

SiC 1,000 K; atm 0.01 [19,20]

SiC 1,000 K; 7.2 MPa 5 [20]

SiC 1,273 K; atm 1.10 [19,20]

SiC 1,273 K; 7.2 MPa 80 [20]

FeCrAl 1,273 K; atm 2.25 [19]

SS-310 1,273 K; atm 5.50 [19]

Inconel 718 1,000 K; atm 100 [22]

T91 1,000 K; atm 100 [23]

Nuclear Engineering and Technology xxx (2016) 1—11

twice as much as SS-316 used in PWRs. The cost savings from elimination of the steam generators as well as having half as thick pressure vessel and piping should offset the higher price of Type 310 stainless steel. Also, operation at approximately 500°C could enable nonelectricity applications such as more efficient industrial process heat.

4. Plant layout

Fig. 5 displays the SWR layout for an offshore seabed setting along with some selected design specifications compared with an integral PWR (iPWR) SMR [27]. The hull size has a diameter of 14 m and a height of 20 m with similar elevated internal suppression pool, used for decay heat removal as well as safety functions. The core can be depressurized using the decay heat removal heat exchangers (DHR HEs) in the internal suppression pools. There is also additional water around the vessel that can flood the core cavity to keep the core covered in case of steam line breaks. Alternatively, this space can be used for auxiliary systems. There is also an emergency boron tank to ensure shutdown in case of anticipated transients without Safety Control Rod Axe Man (SCRAM). As shown in Fig. 5, while the power density of the SWR is lower than the iPWR, the RPV diameter is smaller since no heat exchanger or steam generator is present. Also, the RPV thickness is approximately half of a typical PWR, due to the lower operating pressure. The pumps are connected to the RPV in a manner similar to the advanced BWR (ABWR) design. The core pressure drop per unit length is higher than a typical PWR and much higher than

an iPWR as shown in Fig. 5. However, the pump head is actually relatively low, ~16 kPa, since the average coolant density of the SWR is small resulting in a lower gravity pressure drop. The RPV also has to be situated at a higher elevation relative to the bottom of the hull in order to accommodate the standard BWR control rod drives (CRDs) from the bottom. It is noted that for the SWR, the steam dryers may not be necessary, since superheating occurs in the core before the steam reaches the inlet to the turbine. As a last resort, there is a heat exchanger to cool the containment with the seawater on top of the hull, if desired, though the safety analysis of similar hull geometry implies that such a feature is not necessary [27].

The noted 40% thermal efficiency is calculated by assuming the power cycle technology is based on SST-700 turbine model from Siemens [28] and the turbo-machinery isentropic efficiency will improve to 90% by the 2030-time frame. The SST-700 features both a high pressure and low pressure turbine with a feedwater heater system.

5. Safety assessment

The safety of the 400-MWe MIT SWR design was assumed to be similar to typical BWRs [1]. In this section, the transient response of the SWR for a few selected transients is simulated. It is expected that the behavior during longer term transients would be the same as for other BWRs with similar passive safety systems, such as the economic simplified BWR. The RELAP5 nodalization of the SWR with specifications of Fig. 5 is shown in Fig. 6. The core internal structures were assumed to

Fig. 5 - The superheated water reactor (SWR) small modular reactor layout along with some design specifications. ave, average; CRD, control rod drive; D, diameter; DHR HE, decay heat removal heat exchanger; EBT, emergency boron tank; iPWR, integral pressurized water reactor; ISP, internal suppression pool; N/A, not applicable; RPV, reactor pressure vessel.

Nuclear Engineering and Technology xxx (2016) 1—11

Fig. 6 - The RELAP5 nodalization of the superheated water reactor.

be perfectly insulated for a conservative response. The boundary conditions utilized were the feedwater inlet temperature and turbine inlet pressure. The core is divided into two sections: (1) average assembly; and (2) hot assembly. The hot assembly has 1.45 times the average assembly power rating with a similar assumption as the Fig. 2 analysis. Before the accident scenarios are simulated, the SWR RELAP5 model was first run for 200 s, in order to reach steady state.

A point kinetics model was used to model neutronic feedback of the core. The void coefficient used in the point kinetics model was based on single assembly CASMO4 neutronics calculations. The magnitude of the void coefficient was calculated to be 60% of a nominal BWR (—100 pcm/%void) and negative. The smaller void coefficient is an advantage during an overpressurization transient as well as for flow thermal hydraulic stability. The control rod worth was assumed to be

the same as a typical BWR. This is conservative as the SWR rod worth is required to be higher than typical BWRs to meet the minimum shutdown margin of 1% reactivity. The SWR rods require enriched boron or effectively one CRD per assembly as oppose to one CRD per four assemblies in a typical BWR. The disadvantages of a larger number of CRDs are the additional cost as well as more CRD penetrations in the bottom of the vessel. The disadvantage of an enriched-boron CRD is its limited lifetime, as it requires more frequent replacements. A detailed stability analysis of the reactor was not performed since the outlet quality of the SWR is similar to a regular BWR. The less negative void coefficient and shorter height of the active fuel also provides stabilizing effects. Therefore, it is expected that the two-phase flow stability performance will be better than a typical BWR during steady state and anticipated operation occurrences. Table 3 lists the considered

Table 3 - List of accident sequences modeled in RELAP5.


Sequence (sec)

Total loss of feedwater (decrease in RPV water level)

Turbine trip without bypass (increase in RPV pressure)

Total pump trip (decrease in core flow rate)

Main steam line break (loss of coolant accident)

0: Trip of all feedwater pumps initiated. 5: Feed water flow decays to zero

7.5: Reactor SCRAM due to low water level and trip of four RIPs.

49.5: DHR HE flow enters vessel (end of simulation).

0: Turbine trip initiates closure of main stop valves.

0: Turbine by pass valves fail to operate.

0.01: Main turbine stop valves reach 85% open position.

0.01: Turbine stop valves are closed.

0.15: Bypass valves fail, SCRAM, and four RIP trips are initiated.

1.7: Safety/relief valves open due to high pressure.

10: End of simulation.

0: Trip of all RIPs initiated.

1.22: Reactor SCRAM.

1.85: Feedwater flow pump trip.

1.97: Turbine trip initiates bypass operation.

30: End of simulation

0: 40 cm diameter pipe break which triggers the SCRAM signal. 5: MSIV is closed, which stops the loss of coolant. 50: End of simulation

DHR HE, decay heat removal heat exchangers; MSIV, main steam isolation valve; RIP, reactor internal pump; RPV, Reactor Pressure Vessel; SCRAM, Safety Control Rod Axe Man.

Nuclear Engineering and Technology xxx (2016) 1—11

Water level

220 230

Time (s)

550! 200

. 1 .„ 1 ' 1 1 1 1 -1-1

— Core inlet

— Steam upflow inlet

— Core outlet

220 230

Time (s)

Fig. 7 - The reactor pressure vessel (RPV) water level (left) and the average temperatures in core regions (right) of the superheated water reactor design during the loss of total feedwater accident.

accidents and brief descriptions of the SWR's ability to respond to a decrease in RPV water level and core flow rate with an increase in RPV pressure and loss of coolant in four separate accidents.

5.1. Total loss of feedwater

In this accident, the feedwater decays to 0 in 5 s as listed in Table 3. Following the loss of feedwater, core safety requires it to be covered until the DHR HE condensate flow enters the core. In the ABWR, this sequence results in a RPV water level decrease by 2.5 m [29]. As shown in Fig. 7 (left), the core remains covered as the water level is only decreased by approximately 1 m, which is lower than the ABWR design [29]. Fig. 7 (left) also displays the water level decreased by 0.5 m before the reactor SCRAM occurs after 7.5 s. The steam temperature does not increase significantly as the increased boiling on the outer fuel surface results in a higher steam flow rate as shown in Fig. 7 (right). Therefore, the SWR design performs satisfactorily in this accident.

5.2. Turbine trip without bypass

During the turbine trip without bypass transient, the turbine valve suddenly closes while the bypass valve fails to open. This results in a pressure wave traveling back to the reactor core, resulting in the collapse of the voids. Since typical BWRs, including the SWR design, have negative void coefficients of reactivity, the collapse of the voids results in positive reactivity and a power spike. The failure of the bypass valve sends a SCRAM signal. The power spike also results in a high RPV pressure and opening of the safety relief valves. The sequence is summarized in Table 3, which is similar to the ABWR sequence [29]. Fig. 8 displays the power and the maximum cladding temperatures along with the core outlet steam temperature during the transient for two different cases. In Case 1, a SCRAM is initiated due to the failure of the bypass valve to open, while in Case 2, a SCRAM is initiated due to the closure of the trip valves (0.15 s sooner). The peak power reached in Case 2 is closer to that of the ABWR during this accident sequence [29] and it does not result in any increase in peak cladding temperature. In Case 1, the higher power spike

Fig. 8 - The core power (left) and maximum cladding and core outlet steam temperatures (right) during a turbine trip without bypass transient.

Nuclear Engineering and Technology xxx (2016) 1-11 9

results in the deposition of power in the steam side of the fuel. Since a small amount of enthalpy is required to raise a superheated steam temperature, the cladding temperature increases rapidly. This is unique to the SWR design due to the presence of superheating. If the SWR void coefficient was the same as the ABWR, the peak cladding temperatures would have reached approximately 1,800 K, since more reactivity is inserted resulting in reaching a higher power.

The maximum pressure rise in the RPV is approximately 0.6 MPa for Case 1, while no significant rise is observed for Case 2. These values are less than for the ABWR pressure rise of 1.3 MPa and the pressure stays below the RPV design pressure of 8.5 MPa [29]. The hot assembly outlet quality for both cases did not significantly increase from its steady state value, implying that there are no critical power concerns for both cases. Based on this initial analysis, for this transient, the SWR design performance is acceptable, though it is desirable to have the SCRAM signal transmitted upon the turbine valve closure to improve performance.

5.3. Total pump trip

During the pump trip transient, all the pumps are tripped and the SCRAM signal is initiated following a 1.22-second delay, consistent with the sequence for ABWR licensing [29]. The feedwater pumps are also tripped following the SCRAM signal, as listed in Table 3. Fig. 9 (left) shows that the water level does not decrease below zero (e.g., top of the fuel rods) and does not result in core uncovery. In the ABWR design, the core does briefly uncover resulting in higher cladding temperatures. However, in the SWR design the cladding temperature never rises above its steady state temperature. When the core flow rate decreases, it results in a higher flow rate in the inner channels, as shown by the increase in the equilibrium quality on the outer channel in Fig. 9 (right). This is a unique behavior of the SWR as the increase in boiling on the outer channel results in a lower power fraction toward the inner channel. This, combined with the higher mass flow rate of steam escaping the vessel, results in eventual inner side condensation due to low pressure. Since the inner side experiences

condensation, two-phase flow instabilities could develop under certain conditions. However, such analysis is not covered in this work and further analysis is needed in this area. The analysis suggests that the SWR design behaves safely during the total pump trip transient, mainly due to the core remaining covered. This is unique to the SWR design due to the ability of the inner channel to increase its cooling power of the fuel if an increase in boiling occurs in the outer channel (hence, higher steam flow rate). This design feature is similar to the negative moderator temperature/density feedback of reactivity but for thermal hydraulics.

5.4. Main steam line break

Unlike the ABWR Nuclear Regulatory Commission design certification, where only one of the four main steam lines is assumed to break [29], a double guillotine break of all of SWR's steam lines is assumed for conservatism. It is noted that in this simulation, only one steam line was modeled, which could be the case in a real plant design. The break is modeled as a 40-cm diameter pipe break which triggers the SCRAM signal. The main steam line isolation valve is closed after 5 s of break development, which stops the loss of coolant. Fig. 10 displays the steam dome pressure and RPV water level for this transient. Similar to the pump trip transients, the core does not uncover due to the "negative thermal hydraulic feedback" of the SWR design described in the previous section. The initial decrease in steam dome pressure seen in Fig. 10 (left) is due to the break, while the pressure goes back up to the opening set point pressure of the relief valves. Due to the decay heat, water continues to boil which adds steam to the RPV that cannot escape because of the main steam line isolation valve closure.

Fig. 11 (left) shows that the maximum outer cladding temperature remains at approximately its steady state temperature, while the inner temperature decreases. This is consistent with Fig. 10 (right), where no core uncovery is observed. Lastly, using the free volume and suppression pool volume, shown in Fig. 5, as well as the hull heat transfer coefficient on the ocean side calculated with the Churchill-Chu

Fig. 9 - The reactor pressure vessel (RPV) water level (left) and hot assembly exit equilibrium quality (right) during the all pump trip transient.

10 Nuclear Engineering and Technology xxx (2016) 1—11

Timéis) Time (s)

Fig. 10 - The steam dome pressure (left) and reactor pressure vessel (RPV) water level (right) during the steam line break accident.

natural convection correlation [30], the peak pressure in the containment is estimated and shown in Fig. 11 (right). The peak pressure is below the 0.8 MPa reached in the standard Flexblue design with the larger hull volume [15]. Therefore, a further reduction in the hull diameter may be possible in the SWR design, although the indefinite coolability has not been determined in this study. Based on the initial analysis, the SWR has a safe response during the simulated steam line break accident. It is expected that the DHR HE in the suppression system, which will act as an isolation condenser and be able to passively cool the core indefinitely as the heat transfer around the hull to the ocean, should be sufficient to support the decay heat removal [27].

6. Conclusion

The SWR technology provides higher compactness (30% by hull volume) and an increase in efficiency (10% assuming ideal steam Rankine cycle) and more resistance to be used as a propulsion device compared with a PWR for offshore seabed

deployment. This study outlined a 160-MWe SWR design with recommended cladding material of either 310 SS or FeCrAl and passive safety systems to be able to produce an approximate 500° C steam temperature for a 5-year fuel cycle. While the steam temperature produced by the SWR depends on the fuel thermo-mechanical performance and core power distribution as displayed by Fig. 4, its sensitivity is expected to be manageable during operation. The safety analysis in Section 5 also highlights the unique features of the SWR. The inherit "negative thermal hydraulic feedback" of the design was considered advantageous in most transients. While during fast transients, unless proper modifications to the current BWR control system is accommodated, such as actuation of SCRAM signal upon turbine valve closure, a small amount of power can increase the steam temperature. Overall, the SWR displays a promising option to reduce the cost of SMRs compared with PWRs. However, many challenges remain that need to be addressed including: (1) licensing of cladding material; (2) beyond design basis accident performance; (3) experimental demonstration of fuel thermo-mechanical performance under irradiation; and (4) feasibility of a reliable

Fig. 11 - The maximum cladding temperatures on the inner channel and outer channel cladding outer surface (left) and containment (hull) pressure (right) during the steam line break accident.

Nuclear Engineering and Technology xxx (2016) 1—11

operation with 500°C superheated steam. A detailed economic analysis needs to be performed to see if the gain in power conversion efficiency can overcome the added cost of the fuel and high temperature steam-resistant piping and structures.

Conflicts of interest

All authors have no conflicts of interest to declare. Acknowledgments

We would like to acknowledge the financial support for this work provided by DCNS. The authors would like to thank the DCNS members who provided feedback throughout this work. Dr Yingwei Wu is also acknowledged for his work in updating the MIT FRAPCON-ANNULAR with steel cladding properties. Lastly, the original SWR design is based on the MTI PhD thesis under the supervision of the late Professor Mujid Kazimi, whom the first author is eternally thankful to have had the pleasure of working with him for many years.


[1] Y.-C. Ko, S.M. Kazimi, Conceptual design of an annular-fueled superheat boiling water reactor MIT-ANP-TR-130, Massachusetts Institute of Technology, Center for Advanced Nuclear Energy Systems. Advanced Nuclear Power Program, Massachusetts, 2010.

[2] IAEA Power Reactor Information System—PRIS database [Internet]. Operational data of the Elk River Reactor. Available from:

[3] P.H. Margen, L. Leine, R. Nilson, The design of the Marviken boiling heavy-water reactor with nuclear superheat, Proceedings of the Third International Conference on the Peaceful Uses of Atomic Energy, Geneva, August 31, 1964 to September 9, 1964.

[4] N.A. Dollezhal, Uranium-graphite reactors for power stations using superheated steam, Atom. Energy+ 3 (1957) 1249—1256.

[5] D.H. Lennox, D.R. MacFarlane, R.C. Brubaker, An evaluation of reactor concepts for use as separate steam superheaters, Argonne National Laboratory, Illinois, 1961.

[6] US Department of Energy, Environmental Assessment for Authorizing the Puerto Rico Electric Power Authority (PREPA) to allow Public Access to the Boiling Nuclear Superheat (BONUS) Reactor Building, DOE/EA-1394, Ricon, Puerto Rico, 2003.

[7] Technical Reports Series No. 407, Heavy Water Reactors: Status and Projected Development, International Atomic Energy Agency, 2002.

[8] M.B. Bader, G.L. O'Neilla, EVESR power distribution and thermal-hydraulic analysis, Nucl. Eng. Des. 6 (1967) 115—133.

[9] OECD Nuclear Energy Agency, Radioactive Waste Management Committee (RWMC), Management Board of the Co-operative Program for the Exchange of Scientific and Technical Information concerning Nuclear Installations Decommissioning Projects, Progress During 1995—2005, NEA/ RWM/CPD(2006)3, 2006.

[10] Nuclear facilities in Germany [Internet], Heavy Water Reactors: Status and Projected Development, Technical Reports Series No. 407, International Atomic Energy Agency,

2010, p. 2002. Available from: kerntechnik/Kernanlagen_Stilllegung_Jan_2010engl.pdf.

[11] L. Ritz, Development of Steam-Cooled Fast Breeders, Atomwirt Atomtech 14 (1969) 2-200 [In German].

[12] R.L. Loftness, Nuclear power plants: design, operating experience, and economics, D. Van Nostrand Company, New York, 1964.

[13] N.A. Dollezhal, P.I. Aleshchenkov, Y.V. Evdokimov,

I.Y. Emel'yanov, B.G. Ivanov, L.A. Kochetkov, M.E. Minashin, Y.I. Mityaev, V.P. Nevskii, G.A. Shasharin, V.N. Sharapov, K.K. Orlov, Operating experience with the Beloyarsk Nuclear Power Station, Atom. Energy+ 27 (1969) 1153-1161.

[14] D.J. Stoker, L.S. Mims, S. Siegel, Steam superheat boiling water nuclear reactor, United States Patent # 3150052, Sep. 1964.

[15] G. Haratyk, C. Lecomte, F.X. Briffod, Flexblue: a subsea and transportable small modular power plant, Proceedings of ICAPP'14, Charlotte, USA, 2014.

[16] D. Feng, P. Hejzlar, M.S. Kazimi, Thermal-hydraulic design of high-power-density annular fuel in PWRs, Nucl. Technol. 160 (2007) 16-44.

[17] Y.S. Yang, D.H. Kim, J.G. Bang, H.K. Kim, T.H. Chun, K.S. Kim, K.W. Song, C.G. Seo, H.T. Chae, Irradiation test of dual-cooled annular fuel pellets, Proceedings of WRFPM/Top Fuel, Paris, France, 2009.

[18] Y. Yuan, M.S. Kazimi, P. Hejzlar, Thermomechanical performance of high-power-density annular fuel, Nucl. Technol. 160 (2007) 135-149.

[19] B. Pint, K.A. Terrani, M.P. Brady, J.R. Keiser, High temperature oxidation of fuel cladding candidate materials in

steam-hydrogen environments, J. Nucl. Mater. 440 (2013) 420-427.

[20] R.C. Robinson, J.L. Smialek, SiC recession caused by SiO2 scale volatility under combustion conditions: I, Experimental results and empirical model, J. Am. Ceram. Soc. 82 (1995) 1817-1825.

[21] A. Fry, S. Osgerby, M. Wright, Oxidation of alloys in steam environments: a review, NPL Materials Centre, NPL Report MATC(A) 90, 2002.

[22] G.A. Greene, C.C. Finfrock, Oxidation of Inconel 718 in air at tempratures from 973 K to 1620 K, Energy Sciences and Technology Department, Brookhaven National Lab., BNL-52620, 2000.

[23] D. Laverde, T. Gomez-Acebo, F. Castro, Continuous and cyclic oxidation of T91 ferritic steel under steam, Corros. Sci. 46 (2004) 613-631.

[24] Studsvik, CASMO-4E: A Fuel Assembly Burnup Program User's Manual, Studsvik SSP-09/443-U Rev 0 Proprietary, 2012.

[25] M. Erighin, A 48 month extended fuel cycle for the B&W mPowerTM Small Modular Nuclear Reactor, Proceedings of PHYSOR, Knoxville, USA, 2012.

[26] IAEA, Status of small and medium sized reactor designs, IAEA, Geneva, 2011.

[27] K. Shirvan, R. Ballinger, J. Buongiorno, C. Forsberg, M. Kazimi, N. Todreas, Advanced offshore seabed reactors, MIT-ANP-TR-155, MIT, Cambridge, 2014.

[28] Siemens, Industrial Steam Turbines: The comprehensive product range from 2 to 250 megawatts, 2013. http://www.

[29] ABWR Design Control Document Tier 2 [Internet]. [updated 2013 Mar 11] General Electric, 2007. Available from: http://

[30] F. Incropera, Fundamentals of heat and mass transfer, fifth ed., John Wiley and Sons, New York, 2002.