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Abstract Fusion Engineering and Design 82 (2007) 249 264 Design and performance study of the helium-cooled T-tube divertor concept T. Ihli a, A.R. Raffray b,, S.I. Abdel-Khalik c, S. Shin d, the ARIES-CS Team a Forschungszentrum Karlsruhe GmbH, P.O. Box 36 40, 76021 Karlsruhe, Germany b Mechanical & Aerospace Department and Center for Energy Research, University of California, San Diego, 458 EBU-II, 9500 Gilman Dr., La Jolla, CA 92093-0438, USA c School of Mechanical Engineering, Georgia Institute of Technology, Atlanta, GA 30332-0405, USA d Department of Mechanical and System Design Engineering, Hongik University, Seoul, South Korea Received 12 August 2006; received in revised form 21 November 2006; accepted 22 November 2006 Available online 25 January 2007 The ARIES-CS study has been launched with the goal of developing through physics and engineering optimization an attractive power plant concept based on a compact stellarator configuration. The study included an effort to characterize the divertor location and corresponding heat load distribution, and to develop a He-cooled divertor concept that could accommodate a heat flux of at least 10 MW/m 2, and that would integrate well with the other power core components. This paper describes the design study of this divertor concept, which, although developed for a compact stellarator, is well suited for a tokamak configuration also. 2007 Elsevier B.V. All rights reserved. Keywords: Divertor; Compact stellarator; High heat flux; Helium coolant; Jet flow; Tungsten alloy 1. Introduction The ARIES-CS study has been launched with the goal of developing through physics and engineering optimization an attractive power plant concept based on a compact stellarator configuration [1]. During its Corresponding author. Tel.: +1 858 534 9720; fax: +1 858 822 2120. E-mail address: rraffray@ucsd.edu (A.R. Raffray). initial phase, the study considered a number of different physics configurations including two and three-field period options, as well as several power core designs and maintenance procedures. These were analyzed in some detail as part of a scoping effort, which led to the preferred choice of a dual coolant (He + Pb 17Li) blanket and a port-based maintenance scheme [2]. The blanket would be connected through a heat exchanger to a helium-driven Brayton power cycle. An example of an NCSX-like three-field period configuration is shown 0920-3796/$ see front matter 2007 Elsevier B.V. All rights reserved. doi:10.1016/j.fusengdes.2006.11.004

250 T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 Fig. 1. Three-field period configuration considered for ARIES-CS (based on NCSX). in Fig. 1 while a schematic of the machine layout is shown in Fig. 2, illustrating the location of the different power core components (including the likely divertor location) and the main maintenance port location per field period. Fig. 2. Layout of the ARIES-CS three-field period power core for port-based maintenance. The physics effort also included an investigation of the divertor whose main functions are to intercept thermal particle and heat fluxes escaping from the plasma, help remove the collected heat, and control the particle density via pumping. The goal is to characterize the divertor location, surface topology and the corresponding heat load distribution. A parallel engineering effort focused on developing a divertor unit cell capable of handling the maximum expected heat load while providing the flexibility of being assembled into plate sizes to cover any required location. A helium-cooled concept was selected as it is compatible with the blanket coolants and its high operating temperature results in good quality heat being transferred to the power cycle fluid through the heat exchanger. This paper covers the design study of the ARIES-CS divertor, which, although developed for a compact stellarator, is well suited for a tokamak configuration also. The physics and power flow considerations are briefly summarized; the development of the divertor design is then described, including the material choices. Next, results of the key supporting analyses are described; the integration of the divertor with the power core components is then summarized. Finally, the remaining issues and required R&D are highlighted. 2. Power flow considerations The problem of characterizing the divertor location, coverage and heat load was approached from two sides. From the physics side, studies were carried out to find out the location of the divertor and the corresponding heat flux profile. Preliminary results indicated that the scrape-off layer (SOL) region near the tips of the crescent-shaped cross-section (shown in Fig. 2) is a desirable location for the target plate for two reasons. First, there is a much larger number of field line crossings than the rest of the circumference, increasing the probability of field line interception. Secondly, there is significant flux expansion in the same region that facilitates spreading of the heat load [3]. These initial studies also indicated that the divertor coverage would include both poloidal poles and would probably not be continuous toroidally (perhaps covering for each field period a span of about ±25 from the toroidal location shown in Fig. 2). These studies are proceeding and the final results will be reported in a future publication [4].

T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 251 From an engineering side, the power flow for the entire machine was considered to help determine the required parameters that would result in an acceptable heat flux on the divertor. Fig. 3 illustrates such a power flow. The core power consists of the alpha power and of added power to drive the plasma (in the ARIES- CS case, we assume no added power in steady state). The core power is then divided into alpha-particle loss, core radiation and particle power. In the case of a compact stellarator the alpha loss can be significant up to 5 10%. The core radiation fraction, f rad,core, includes Bremsstrahlung and synchotron radiations; it would depend on parameters such as the assumed impurities in the plasma and can be set to a certain extent. The particle power would then be divided in an edge radiation fraction, f rad,edge, which can also be set to some extent since it would depend on assumptions about impurity level in the scrape-off layer to enhance radiation, and the conducted power to the divertor. The latter would reach the divertor resulting in a heat load profile governed by factors such as the divertor inclination and the particle intersection footprint on the divertor plates. This would cause a peaking factor, F div,peak (defined as the maximum to average conducted heat flux on the total divertor area), which needs to be included Fig. 3. Example of power flow diagram for ARIES-CS. when estimating the maximum heat load on the divertor. Edge radiation can be further divided into a component radiating from the vicinity of the divertor region mostly to adjacent plasma facing zones, f rad,edge,div, and another component radiating to the entire first wall. For the simple estimates shown here, f rad,edge,div is set to 50% with a further assumption that the edge radiation from the divertor region would also be split evenly between radiation to the divertor only and radiation to the rest of the first wall (f rdr,div is set to 50%) [5]. In a stellarator the magnetic field ripple associated with the non-axisymmetry of the magnetic geometry results in a relatively high fraction of prompt alphaparticle loss, up to 5 10%. These alphas would deposit their energy in regions of the plasma facing components (PFC) which would need to be designed to accommodate the resulting heat flux. An effort is underway to better determine the location and density of the alphaparticle fluxes on the PFC [3]. For the purpose of this paper, it is assumed conservatively that all the alpha loss energy ends up on the divertor. However, in order not to impose unreasonable requirements on the divertor, the coverage area is assumed to be higher than the typical 10% assumed for the divertor function. A 15%

252 T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 Fig. 4. Combination of fractional core radiation fraction, edge radiation and divertor peaking factor required to maintain the maximum divertor heat flux <10 MW/m 2 under the given assumptions. coverage is assumed for the divertor region under the assumption that all the alpha loss energy is deposited there, and the alpha heat flux peaking factor is assumed similar to that of the conducting power to the divertor. This analysis is done with the understanding that its results will need to be updated as more detailed information on the alpha power deposition and divertor footprint become available from the physics effort. As an illustration of the fractional radiation required to maintain the divertor peak heat flux to an acceptable level, Fig. 4 shows the combination of core radiation fraction, edge radiation fraction and peaking factor required to maintain the maximum divertor heat flux <10 MW/m 2 under the above assumptions. As an example, for a fractional core radiation of 55% and edge radiation of 75%, the maximum peaking factor for the conducted power to the divertor should be < 14 in order to maintain the maximum heat flux on the divertor <10 MW/m 2. 3. Divertor design 3.1. Background The engineering effort focused on developing a divertor design with: the capability to accommodate an adequate peak heat flux level; the flexibility of being assembled in plates suited to the coverage requirement of a compact stellarator; and the provision for integration with the core component design, maintenance scheme and power cycle operation. This led to the choice of a high-temperature helium-cooled tungstenalloy divertor design which would fit very well into the overall plant concept with the in-reactor coolants transferring their energy through a heat exchanger to a helium working fluid driving a closed Brayton cycle. The ARIES-CS study is focused on a fusion power plant and it is reasonable to include a degree of extrapolations in the physics and technology assumptions, for example, in regard to the refractory alloy material development. As a reasonable initial goal and in anticipation of the physics modeling results, a maximum heat flux of 10 MW/m 2 was assumed as design requirement. A range of different He-cooled divertor configurations have been considered in the past, including a tungsten plate design [6]. More recently, a finger configuration utilizing tungsten caps has been evolved with the aim of minimizing the use of tungsten as structural material and of accommodating higher heat fluxes through the use of smaller units [7]. It was decided to build on the tungsten cap design and to explore the possibility of a new mid-size configuration with good heat flux accommodation potential, reasonably simple (and credible) manufacturing and assembly procedures, and which could be well integrated in the CS reactor design. An important advantage of the stellarator configuration is the assumption of no major disruption events. Thus, the requirements on the target armor are less demanding than in the case of near-term tokamaks. In the absence of armor melting events due to disruptions the need and thickness of a sacrificial armor layer depend on the sputtering rate resulting from the amount and energy of the different surface hitting particles and the properties of the target armor material. These parameters are not yet determined in detail but it was foreseen that, if necessary, a sacrificial castellated layer ( 1 mm) would be brazed onto the directly cooled target structure. It is assumed that the impact of this castellated layer on the stress levels in the underneath structure would be small and it was not included in the initial design analysis. Thus, the target design effort was focused on the development of the basic target structure made from tungsten alloy and of an efficient gas cooling technique, which is necessary to keep the structure temperature within its design limits.

T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 253 3.2. Basic target structure The allowable heat load on the directly cooled and pressure-carrying target structure is constrained by the total stress levels mainly thermal stresses due to the high heat flux which have to be kept below the design limits. Reduced thermal stress levels can be achieved by: (i) keeping temperature differences small, and (ii) minimizing the restraints against differential thermal expansion and deformation during heat up. The latter requirement can be achieved by segmentation of the target structure in small separate parts and by introducing flexibility in the design. The temperature differences in the cooled walls are proportional to the heat flux through the wall and to the wall thickness. Thinner wall can be achieved in a pressure-carrying component by using a suitably curved shape and a small cooling channel width. These considerations led to the development of a target built from small separate T-tubes offering the advantages of: (i) circular cross-section and small diameter for a small wall thickness and low radial temperature differences in the tube walls, and (ii) reasonably unrestrained bending of the tubes by use of T-connectors in the tube middle position [8]. To avoid shadowing effects, a flat basic armor layer on the heat-loaded side of the tubes is used, whereby the minimum thickness of the armor layer was set to 0.3 mm. An additional castellated layer (e.g. made up of small separate cuboids) could be brazed onto the basic armor layer to accommodate armor lifetime requirements, as described earlier. Schematics of the T-tube concept are shown in Figs. 5 and 6, illustrating the different design components and the He flow configuration. Three flow channels are placed in the T-connector and the tran- Fig. 5. Schematic of the ARIES-CS divertor T-tube concept showing the major components. Fig. 6. Cut of the divertor T-tube concept showing the He flow configuration. sition piece, whereby the coolant enters through the middle section and returns through the outer sections. For distribution of the incoming coolant flow within the tubes a concentric cartridge is placed in the tubes, which is connected with the inlet section of the T-connector. To achieve sufficient heat transfer the coolant is accelerated towards the inside of the tube by passing through a slot in the cartridge (see the cooling technique section). After impingement, the coolant returns through the annular gap between tube and cartridge towards the outlet sections in the T-connector. 3.3. Cooling scheme for high heat flux removal The inside of the tube has to be cooled very effectively to keep the maximum temperature within the tube below the design limit (assumed as 1300 C for W alloy, as discussed in Section 4). Studies of the European HEMJ (or multi-jet) divertor concept [7] indicated that a very high heat transfer coefficient can be reached by using a jet impingement cooling technique. This technique can be employed directly due to its simple design involving a plenum chamber (concentric cartridges in the tubes) and orifices (holes or slots in the cartridges). The outlet velocity of the impinged jets is high enough to result in turbulent flow immediately after impingement. The characteristic coolant flow along the wall after impingement (as illustrated in Figs. 6 and 7) is extremely turbulent with high velocity fluctuations and increased local turbulent mixing. As a result, a significant increase in the heat transfer performance is achieved. A correlation for jet flow was used to perform the initial scaling analysis. It is based on the German VDI

254 T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 Fig. 7. Illustration of the jet flow through the slot showing the jet wall spacing, H, and the slot width, B (the slot length in the longitudinal direction perpendicular to the page is L). Waermeatlas [9] and is shown in Eqs. (1) and (2): 1 m = 0.695 X nd + Hnd 1.33 + 3.06 Nu = hd k He = 1.53 Rem Pr 0.42 X nd + H nd + 1.39 (1) (2) X nd = X D and H nd = H (3) D where m is a geometry-dependent exponent used in Eq. (2); X nd the distance from the jet centerline, X, normalized to the hydraulic diameter, D (D =2B, where B is the slot width shown in Fig. 7); H nd the wall spacing, H (shown in Fig. 7), normalized to the hydraulic diameter; Re, Pr and Nu the Reynolds, Prandtl and Nusselt numbers, respectively; and h and k He are the heat transfer coefficient and He thermal conductivity, respectively. Example results are shown in Fig. 8 to illustrate the effect of changes in mass flow rate and tube geometry required to maintain the maximum W alloy temperature <1300 C under different maximum heat fluxes for a He coolant inlet temperature of 600 C and pressure of 10 MPa. The heat transfer coefficient (based on the He coolant inlet temperature) decreases with the angle from the jet impingement region. For the case of a maximum heat flux of 10 MW/m 2 (with B = 0.5 mm, H = 1.2 mm; the inner T-tube radius, R i = 6.5 mm; and the distance from the jet centerline, X = π R i ), the average heat transfer coefficient is about 17,000 W/(m 2 K). The local value of the heat transfer coefficient increases to high values close to Fig. 8. Heat transfer coefficient as a function of angular position (zero at mid-plane) for different heat fluxes and layout cases of T- tube (q = heat flux; G/A: specific mass velocity; dp = pressure drop; d = outer diameter of T-tube; s: average wall thickness of outer tube). the jet impingement region. However, modeling of the heat transfer there falls outside the bound of the above correlation and computational fluid dynamics (CFD) analysis is required. Detailed 2-D and 3-D CFD analyses (using FLU- ENT [10]) were carried out independently at FZK in Karlsruhe, Germany, and at Georgia Institute of Technology in Atlanta in support of the design. The assumed W alloy and He coolant material properties are shown in Table 1. By symmetry, a quarter of the T-tube was modeled as illustrated in Fig. 9. Typical dimensions and parameters of the T-tube concept are shown in Fig. 10 and summarized in Table 2. An illustration of a typical grid generation is provided in Fig. 11. The standard k ε with wall enhancement model for turbulent flow and compressible flow (perfect gas law) were assumed. The results from FZK and Georgia Tech. were very consistent, indicating a very good heat transfer performance for such a concept, with maximum local heat transfer coefficient values of about 40,000 W/(m 2 K) Table 1 W alloy and He properties assumed in the analysis (p refers to pressure in MPa and T to temperature in K) W alloy He coolant Density (kg/m 3 ) 19,254 487.5p/T Thermal conductivity 115 0.012T 3.5 10 3 (T) 0.67 (W/(m K)) Specific heat (J/(kg K)) 138 5192 Viscosity (Pa s) N/A 4.5 10 7 (T) 0.67

T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 255 Fig. 9. Quarter T-tube used in CFD modeling. or more. Fig. 12 shows an example of the 2-D variation of the heat transfer coefficient with angle from the impingement region, illustrating the sharp increase in value close to the impingement region. Results for the velocity and temperature distributions in the tube based on 3-D FLUENT analysis are shown in Figs. 13 and 14 for a heat flux of 10 MW/m 2, Table 2 Divertor design parameters to accommodate a peak heat flux of 10 MW/m 2 He pressure (MPa) 10 Mass flow rate per heat-loaded surface area (kg/(m 2 s)) 24 Outer diameter of T-tube (mm) 15 Slot width, B (mm) 0.5 Pressure drop for jet (MPa) 0.07 Toroidal tube tube distance (mm) 90 Poloidal tube tube distance (mm) 15.8 Jet wall distance, H (mm) 1.2 Average jet velocity (m/s) 200 Average heat transfer coefficient (W/(m 2 K)) 17,000 Fig. 11. Example grid structure used for the FLUENT analysis. Fig. 10. Typical dimensions (in mm) of the T-tube divertor concept used in the analysis. Fig. 12. Variation of normalized heat transfer coefficient with angle from the jet impingement region for an example case.

256 T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 Fig. 13. Velocity distributions in the divertor tube based on 3-D FLUENT analysis for a heat flux of 10 MW/m 2, a volumetric heat generation of 53 MW/m 3 in the W alloy, a tube length of 90 mm and the tube dimensions shown in Fig. 10. a volumetric heat generation of 53 MW/m 3 in the W alloy, a tube length of 90 mm and the tube dimensions summarized in Fig. 10 and Table 1. The maximum velocity for this case is 216 m/s; the corresponding average inlet and outlet temperature of He are 873 and 950 K, respectively. The maximum W armor temperature is 1699 K; however, the maximum W alloy temperature at the tube/armor interface is 1523 K (less than the assumed limit of 1300 C or 1573 K). This is illustrated in Fig. 15 showing the temperature distribution at the armor/tube interface, as well as the corresponding local heat transfer coefficient distri- Fig. 14. Temperature distribution in the divertor tube based on 3-D FLUENT analysis for a heat flux of 10 MW/m 2, a volumetric heat generation of 53 MW/m 3 in the W alloy, a tube length of 90 mm and the tube dimensions shown in Fig. 10. bution in the tube (with a maximum value close to 40,000 W/(m 2 K) in this case). The corresponding pressure drop, also obtained from FLUENT, is reasonable, 10 5 Pa for the tube. These results need to be verified experimentally. Such a thermal hydraulic experiment has been planned and is being carried out at the Georgia Institute of Technology. Initial results are very encouraging and complete details of the experimental setup, results and of the accompanying modeling analyses of both the test mock-up and divertor T-tube will be reported in a future publication [11]. Fig. 15. Temperature and heat transfer coefficient distributions from FLUENT analysis for an example divertor case with a heat flux of 10 MW/m 2, a volumetric heat generation of 53 MW/m 3 in the W alloy, and a He mass flow rate of 8.5 g/s per tube.

3.4. Parametric analysis of thermal hydraulic performance T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 257 The tube geometry and dimensions described in the previous section and summarized in Fig. 10 and Table 1 are based on accommodating a maximum heat flux of 10 MW/m 2 while maintaining the W alloy temperature <1300 C. A parametric analysis of the impact of a change in the maximum heat flux was performed. Basically, in order to maintain the same maximum W alloy temperature, the T-tube thickness and diameter have to be changed in inverse proportion to the heat flux. For example, to accommodate a heat flux of 5 MW/m 2, the tube diameter could be increased to d = 30 mm and wall thickness to s = 0.8 mm. Under the reduced heat load, the He mass flow rate could be reduced to 12 kg/(m 2 s) per tube with an associated reduction in the He pressure drop to dp = 0.027 MPa. Conversely, an increase in heat flux to 20 MW/m 2 would require d = 7.5 mm, s = 0.25 mm, a He flow rate of 48 kg/(m 2 s) and an increase of dp to 0.28 MPa. These are summarized in Fig. 8, which shows the corresponding variation in heat transfer coefficient. Thus, in principle, the T-tube design could be modified to accommodate a larger heat flux. However, the thickness of the wall becomes very small ( 0.25 mm for the 20 MW/m 2 ) and is probably not realistic from fabrication feasibility and design integrity considerations. A reliable cooling system should be able to accommodate possible geometrical uncertainties, such as deformations, which might occur during the years of operation. In this respect the described T-tube divertor concept is very promising due to the characteristics of the jet cooling system. The influence of the cartridge position in the tube on the cooling performance is very small, as the cooling function is based on wall jet flows which are produced fairly independently of the gap size between the cartridge and the outer tube. The slot nozzles are the only flow contraction region within the flow path and, as a consequence, a major fraction of the total pressure drop is available for the acceleration in the slot nozzles. The slot width can be fabricated to exact specifications, as it does not depend on the relative positions between different parts. Nevertheless, the cooling system design must be sufficiently robust to accommodate possible changes in the slot width during operation. A parametric study of the effect of changing the slot width on the thermal performance of the T-tube was per- Fig. 16. Heat transfer coefficient at the inside of the tube as a function of angular distance from the impingement point for different slot widths. formed. The results are summarized in Fig. 16 in terms of the variation of the local heat transfer coefficient (determined from Eqs. (1) and (2)) with the angular distance from the middle position in the upper half of the tube (distance from impingement point) for three different slot sizes, under the same flow conditions and geometry as previously assumed for the 10 MW/m 2 case. These results indicate that a change in slot width of ±20% (from the 0.5 mm initially assumed) would only slightly influence the overall heat transfer coefficient at the inside tube wall, especially at the angular position of 30, where the highest wall temperature in the tube can be expected. This indicates that, for a given pressure ratio in the circulation system, the heat transfer capacity in the tubes would be quite stable and robust against geometrical uncertainties. The coolant mass flow and temperature rise in the tubes will vary slightly if some deviations in slot width is assumed, but the average coolant mass flow and temperature rise will be fairly stable. 3.5. Stress analysis The ANSYS Workbench software package [12] was utilized to calculate the stresses and deformations within the T-tube basic configuration (see Table 1 and Fig. 10) for an internal helium pressure of 10 MPa and a heat load of 10 MW/m 2. The armor layer was assumed to be bonded to the W alloy tube. The results are summarized in Fig. 17. The highest stress levels

258 T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 Fig. 17. Stress intensity for the basic T-tube layout and a surface heat flux of 10 MW/m 2. occur in the T-junction area as the T-connector restrains the free bending of the tube. The total stress intensity (primary and secondary stresses) is <370 MPa for the entire geometry, which is assumed to be less than the 3Sm limit of an anticipated W alloy at the corresponding temperatures and therefore is regarded to be acceptable (as discussed in Section 4). The effect of asymmetry in the heat loading conditions was also investigated. This could be due to some shadowing effects, which is likely to be self-cured by armor erosion. An assumed scenario was analyzed, whereby the heat load in a limited section of the tube is set to 0 MW/m 2 and the heat load in the remaining Fig. 18. Primary and secondary stress intensity for asymmetric heat load. Fig. 19. Outside temperature distribution of cooled tube for asymmetric heat load. section is set at 12 MW/m 2. The results, summarized in Figs. 18 and 19, show that no local stress peak would occur in the tube and that the maximum temperature at the outer W alloy tube is still acceptable ( 1275 C), which provides another indication of the robustness of this design this time to the possibility of asymmetric heat loading. Finally, an analysis of the tube ends (end caps) was performed to determine whether there could be local peaks in temperature and stress due to the heat load level there. Results for two cases are shown in Fig. 20. The results on the left side correspond to a case where the heat load on the tube ends was set to 10 MW/m 2 (100% of the heat load in the remaining armor region), and on the right side to case with a heat load on the tube ends of 5 MW/m 2 (50% of overall heat flux). In the latter case, both temperatures and stress intensities at the tube ends are lower than the maxima in the remaining geometry. In the first case, the stress intensity at the tube end is in the same range as in the T-connection area, but the temperature at the tube end is slightly higher than in the remaining part of the tube. From this, it could be concluded that lowering the heat flux to the end caps by shielding of the tube ends would be favorable; however, a local heat flux of up to at least 50% of the overall heat flux would not cause any stress or temperature peaks. This is important as a high fraction of radiated energy could result in significant heat load for the tube ends, even if there is a large geometrical step in the armor layer to prevent particles hitting the tube ends.

T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 259 Fig. 20. Stress intensities and temperatures for heat-loaded tube ends (left side 100%, right side 50% of overall surface heat load of 10 MW/m 2 ). 4. Materials and fabrication It is proposed to fabricate the directly cooled heat transfer parts of the gas-cooled divertor target from a tungsten alloy to provide the required high heat flux and high-temperature capabilities. It is known that the low-temperature ductility and the ductile to brittle transition temperature (DBTT) of a tungsten alloy can be improved by alloying (with rhenium, for example). However, there is limited data as to the mechanical properties of irradiated W alloy. Severe embrittlement (DBTT > 900 C) was observed in W and W 10% Re irradiated at 300 C to a fluence corresponding to 1 dpa [13]. Due to this lack of data, it is difficult to make an accurate estimate of DBTT as a function of irradiation temperature for W alloys. To-date, the minimum operating temperature would be in the range of 800 1000 C scaling from the limited data [13]. This would be too high for the divertor design since at some point the W alloy need to transition to a ferritic steel (FS) structure, with the requirement that the interface temperature (and the coolant temperature) be compatible with the minimum allowable temperature for W alloy and the maximum allowable temperature for FS. Optimization of the W alloy material development could lead to lower allowable temperatures, which would provide better application to the design of fusion high heat flux component in conjunction. Clearly, an effort is needed both in the material area and in the experimental area to develop different W alloys and test them under irradiation in the range 600 1000 C to address this issue. The ARIES-CS study is considering long-term power plant application and it seems reasonable to assume a lower operating temperature of 600 700 C for the W alloy to determine the benefit of utilizing such a material and to provide a motivation for its development. The design configuration and parameters would have to be adjusted in the future in light of results from such R&D. The possibility of stress- or radiation-enhanced recrystallization of refractory alloys can severely impact the strength limit. For example, Fig. 21 shows the ultimate tensile strength (UTS) and yield strength (YS) of unirradiated recrystallized W (5 10%) Re [13]. The allowable stress in refractory alloys is generally controlled by the UTS rather than the YS due to their low work hardening capacity [13]. For this particular unirradiated recrystallized W alloy, the UTS at 1300 C is 202 MPa. If one assumes a value of the allowable primary membrane stress intensity

260 T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 Fig. 21. Ultimate tensile strength and yield strength of unirradiated recrystallized W (5 10%) Re [13]. (Sm) equivalent to (UTS/3), and a maximum allowable secondary stress equivalent to 3Sm, the maximum allowable secondary stress would then be 202 MPa. This is very low, below the estimated thermal stress of 370 MPa for the design. Thus, it seems very important to aim at avoiding recrystallization of the W alloy. Mechanical properties of stress-relieved (nonrecrystallized) refractory alloys are superior to those of recrystallized specimen (increase in strength by up to a factor of 2). The recrystallization temperature in commercially pure (undoped) tungsten can be as high as 1200 1400 C. The recrystallization temperature could be increased by alloying the W. However, it could also be reduced by the specific stress and radiation conditions. Radiation-enhanced recrystallization and radiation creep effects need to be further investigated, but for the purpose of this study the recrystallization temperature of the W alloy is assumed to be 1300 C or more. It is planned to fabricate the main structure of the divertor target from an advanced ferritic steel, such as oxide-dispersion strengthened (ODS) FS. The assumed maximum operational temperature of ODS FS is 700 800 C [13,14], which sets a ceiling for the outlet coolant temperature. The tube and its end caps as well as the T-connector may be produced by plasma spraying or as CVDtungsten parts. The cartridge is cooled from inside and outside and could therefore be made from a tungstenbased material like Densimet 18, which would allow mechanical machining [9]. To reduce the thermal stress levels due to the different thermal expansion coefficients of ferritic steel and tungsten alloy, a graded transition piece is foreseen in the T-junction area of the tubes. The transition pieces could be fabricated as a sandwich of slices from different material grades to adjust step by step in terms of thermal expansion. The slices could possibly be diffusion welded under hot isostatic pressure (HIP) utilizing thin brazing foils between each layer to improve the diffusion bonding process. Alternatively, the functional gradient could be achieved by a vacuum plasma spray (VPS) forming technique. The different components that would need to be separately fabricated and then assembled in the manufacture of the T-tube concept are shown in Fig. 5. The bonding technique used to attach the end caps and the T-connector to the tubes must provide good interface properties. High-temperature brazing is a possibility. An R&D effort is underway at Plasma Processes Incorporated (PPI) in Alabama to demonstrate the fabrication of such a W alloy concept, including investigation of techniques for joining the W alloy to the ferritic steel base through a functional gradient [15]. 5. Integration with the machine The T-tube concept has been developed with the idea of having a small unit cell that could be assembled to a common manifold to produce a mid-size target plate on a manifold module, as illustrated in Fig. 22 (with example dimensions). The T-tubes will be favorably aligned with the magnetic field lines (likely toroidal), whereby supporting manifold modules below the tubes will be directed perpendicular to the tubes. The manifold modules will be combined to form the target plates, of radial thickness 0.2 m. Neutronics analysis indicates that an overall TBR of 1.1 can be achieved assuming no breeding in the first 0.2 m of the divertor region for a divertor coverage of up to 15% [16]. To accommodate the relatively high neutron wall loads for a divertor target in the plasma chamber of a compact stellarator (e.g. 2 3 MW/m 2 ) it is desirable to cool the walls of the target steel structure with the low-temperature coolant before flowing the coolant to the T-tubes. Therefore, the coolant is first routed through small rectangular flow channels in the walls of the manifolds and then to the inner flow channel in the middle of the manifolds (which is surrounded by the

T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 261 Fig. 22. Target design with T-tubes (tungsten alloy) and coolant manifold fabricated from ODS FS. cooled walls). These inner flow channels are divided diagonally in a supply section to feed the T-tubes and a collector section for the returning high-temperature coolant flow. Such manifold modules could them be assembled to form larger plates of different dimensions to fit within the particular divertor geometry of the machine. In this regard, the design provides a flexibility which is particularly suitable to the complex compact stellarator geometry, but also applicable to a tokamak configuration. 6. Heat load footprint From Section 2, the heat flow to the divertor is governed by a number of parameters including the fusion power and assumed core and edge radiation fractions. The resulting average heat flux would depend on the divertor coverage and the maximum heat flux by the peaking factor. If one assumes a Gaussian profile to the heat load deposition, the average heat flux (and peak- Fig. 23. Example footprint of target heat load (assuming a Gaussian profile) for a maximum heat flux of 10 MW/m 2 and an average heat flux of 3 MW/m 2 (peaking factor = 3.33). ing factor) can be estimated for a given divertor target plate coverage and maximum allowable heat flux. For example, the Gaussian profile shown in Fig. 23 would correspond to a maximum heat flux of 10 MW/m 2,a peaking factor of 3.33 and a poloidal distance of 1 m, which would be typical of a tokamak. For the compact stellarator, the peaking factor is likely to be higher and the Gaussian profile will be more peaked and thinner. As an illustration of the overall flow configuration required to accommodate a given heat load profile, the example profile of Fig. 23 is used in the next section. The final flow parameters could change somewhat for the actual stellarator heat load profile and the results will be updated as required from the final results of the divertor physics analysis. 6.1. Target plate manifolding Fig. 24 illustrates the flow configuration in the target manifold unit (see also Fig. 22) including example coolant temperatures. The unit consists of outer cooling channels and of inner main manifold and collector sections. The wall between the manifold and collector sections is diagonal to keep the poloidal flow velocity nearly constant in these sections. The coolant enters the outer cooling channels at a temperature below 600 C. At the end of the wall cooling section, the flow turns in the manifold-section of the unit at a temperature of 600 C. Each T-tube within the unit is fed in parallel with coolant of roughly the same inlet temperature, the coolant then returning to the collector sections.

262 T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 Such a target manifold unit would be designed to accommodate a given heat flux. As the heat flux changes poloidally, other units designed for different heat fluxes can be connected in series such that the outlet flow from one unit can then be used to cool the next unit, with the collector section of one unit being connected to the manifold section of the next unit. The overall layout of these units is likely to be more complicated for a compact stellarator divertor with its high heat flux peaking factor than for a tokamak divertor. The exact layout will be based on the divertor physics results regarding the divertor location and heat flux profile. An example of the flow configuration and parameters is provided below for the heat flux footprint shown in Fig. 23. Under this footprint, the target plate is divided into three poloidal zones cooled in series. Each zone is made up of a number of similar manifold units arranged toroidally and cooled in parallel to provide the necessary toroidal width of the plate. In zone 2 (see Fig. 23), small T-tubes, 15 mm in diameter, will be used to accommodate heat fluxes up to 10 MW/m 2 (the parameters for such a T-tube are shown in Table 1). The poloidal length of cooling zone 2 was chosen to be 0.25 m to reach the necessary mass flow required per target surface area (24 kg/(m 2 s), from Table 1) for a coolant temperature rise of about 100 C. If one assumes a toroidal width of 2 m for the whole manifold unit assembly, the mass flow rate would be 12 kg/s. An estimate of the pressure drop in such a manifold unit was performed based on the Re and friction factor for local channel components for an assumed roughness Fig. 24. Illustration of flow configuration in a target manifold unit. of 0.05 mm. The total pressure drop for such a unit in zone 2 of Fig. 23 is estimated as 0.12 MPa. Details of the analysis are given in Ref. [9]. In zones 1 and 3, the maximum heat flux is 6 MW/m 2. The tube dimensions can be scaled accordingly with an outer diameter of 25 mm. The T-tube lengths were set as twice those of the zone 2 tubes in order to keep the design simple. The slot width for the cartridges in zones 1 and 3 (B = 0.97 mm) was adjusted to obtain an average heat transfer coefficient of about 10,400 W/(m 2 K) at the inside of the tubes, sufficient to accommodate heat fluxes up to 6 MW/m 2 without exceeding the assumed material temperature limits. As a result, the pressure drop in zones 1 and 3 was estimated as 0.045 MPa. When adding the pressure drop in the supply channel, the total pressure drop in a divertor target with series flow through zones 1 3 was estimated as 0.34 MPa, with a corresponding isentropic pumping power of 0.7 MW for a target plate assembly of dimension 2 m (toroidal) 1 m (poloidal) (corresponding to 12% of the heat deposited for a region with the heat load footprint shown in Fig. 23) [9]. These numbers will change depending on the final layout of the ARIES-CS divertor and the local heat flux peaking and possible non-uniformities. 6.2. Coupling with blanket coolant and heat exchanger In determining to how to support and feed the divertor target plates, different possibilities were considered including integration of the plate with a blanket mod-

T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 263 ule and inclusion of the divertor manifold as part of the blanket manifold region. However, this approach raises concerns about the lack of space at the back for an additional divertor manifold running next to the blanket manifold (see Fig. 2) and also about the effect of the blanket module thermal expansion on the alignment of the divertor in this integrated configuration. It was decided instead to try to attach the divertor directly to the colder vacuum vessel (VV), with the divertor coolant supply tube/mechanical attachment penetrating the breeding modules and the hot shielding ring. In this case the concentric coolant pipes (with the inlet and outlet flows in the annular region and center tube, respectively) to each of the divertor plates are attached at the back of the VV, and can be cut/rewelded with in bore tools inserted from the outside of the VV. The divertor target unit can then be replaced though the port designed for blanket replacement. Details about this scheme can be found in Ref. [17]. In this configuration, the divertor unit will be independent of the blanket modules. However, the divertor coolant must be integrated with the blanket coolants (He and Pb 17Li) when transferring power through Fig. 25. Illustration of coupling of blanket and divertor coolants to a Brayton power cycle. a heat exchanger (HX) to the Brayton cycle He. This is illustrated in Fig. 25 with typical power core coolant and HX fluid temperatures. The maximum divertor coolant temperature of 700 C fits well with the maximum Pb 17Li coolant temperature of about 720 730 C, and the corresponding cycle efficiency is about 42 44% [18]. 7. Conclusions An attractive new design concept for a gas-cooled divertor was developed as part of the ARIES-CS power plant study. The concept is based on a modular arrangement of T-tubes made of tungsten alloy, which can be assembled to fabricate target plate of different dimensions to fit the particular divertor geometry of the machine. This provides a design flexibility which is particularly suitable to the complex compact stellarator geometry, but also applicable to a tokamak configuration. A slot-based jet impingement cooling technique is applied to provide the heat transfer performance to accommodate heat fluxes of at least 10 MW/m 2 along with reasonable pumping power and pressure drop requirements. The stress levels in the heat-loaded geometry seem acceptable due to the small wall thickness of the tubes, the high W thermal conductivity and the relatively flexible design configuration. A scaling scheme was evolved to modify the dimensions in order to accommodate various heat flux values. Modeling and experimental R&D are being conducted at the Georgia Institute of Technology to verify the thermal hydraulic performance of this concept and will be reported in the future. The design effort included identifying possible fabrication methods and integration of the T-tube. For example, the design showed that a functional gradient in connecting the tungsten components to the steelmanifolds would reduce the thermal stresses. An R&D effort is underway at Plasma Processes Incorporated in Alabama to demonstrate the fabrication of such a T- tube concept, including the use of a functional gradient between W alloy (as presently available) and ferritic steel. Different schemes were considered in integrating the target in the machine. An independent target supported by the vacuum vessel and with its own cooling

264 T. Ihli et al. / Fusion Engineering and Design 82 (2007) 249 264 system is preferred with the target plate assembly being maintained through the blanket ports. The divertor coolant ( 600 700 C) is compatible with the blanket coolant temperatures, and the coupling to a Brayton power cycle through a heat exchanger results in reasonable cycle efficiency values. A key assumption is that the W alloy will be able to operate in the temperature range of about 600 1300 C based on ductility and recrystallization limits, and will be able to accommodate thermal stresses of 370 MPa. Clearly, R&D is required in this area both to develop W alloys that would provide the possibility of operation in this temperature range with acceptable mechanical properties, as well as to test them at temperature and under irradiation, and to demonstrate feasible and cost-effective fabrication methods for the geometry of interest. Acknowledgements The work at UCSD was supported under U.S. Department of Energy Grant No. DE-FC03-95- ER54299. The authors would like to thank colleagues from FZK for the valuable discussions and especially Dr. R. Kruessmann for her support in the CFD analysis of impinging jet flows. The authors would also like to acknowledge Dr. T.K. Mau from UCSD for his physics input. References [1] F. Najmabadi, the ARIES Team, Overview of ARIES-CS compact stellarator study, Fusion Sci. Technol. 47 (April (3)) (2005) 406 413. [2] A.R. Raffray, L. El-Guebaly, S. Malang, F. Najmabadi, X. Wang, the ARIES Team, Major integration issues in evolving the configuration design space for the ARIES-CS compact stellarator power plant, Fusion Eng. Des. 81 (2006) 1159 1168. [3] T.K. Mau, H. McGuinness, A. Grossman, A.R. Raffray, D. Steiner, the ARIES Team, Divertor heat load studies for compact stellarator reactors, in: Proceedings of the 21th IEEE/NPSS Symposium on Fusion Engineering, Knoxville, TN, September 26 29, 2005, (CD-ROM). [4] T.K. Mau, T. Kaiser, A.A. Grossman, A.R. Raffray, X.R. Wang, J.F. Lyon, R. Maingi, L.P. Ku, M.C. Zarnstorff, the ARIES Team, Divertor configuration and heat load distribution for the ARIES-CS power plant, Fusion Sci. & Technol., submitted for publication. [5] J. Lyon, R. Maingi, M. Zandsdorf, Private communication, May 2006. [6] S. Hermsmeyer, S. Malang, Gas-cooled high performance divertor for a power plant, Fusion Eng. Des. 61 62 (2002) 197 202. [7] T. Ihli, R. Kruessmann, I. Ovchinnikov, P. Norajitra, V. Kuznetsov, R. Giniyatulin, An advanced He-cooled divertor concept: design, cooling technology, and thermohydraulic analyses with CFD, Fusion Eng. Des. 75 79 (November) (2005) 371 375. [8] T. Ihli, R. Raffray, the ARIES Team, Gas-cooled divertor design approach for ARIES CS, in: Proceedings of the 21st IEEE/NPSS Symposium on Fusion Engineering (SOFE), Knoxville, Tennessee, USA, September 26 29, 2005 (CD-ROM). [9] T. Ihli, A.R. Raffray, the ARIES Team, ARIES CS Report: Helium Cooled Divertor Design Study, UCSD-CER-06-04, July 2005, available at http://aries.ucsd.edu/cer/reports.shtml. [10] FLUENT, Inc., FLUENT Flow Modeling Software, Version 6.2, 2006, www.fluent.com. [11] S.I. Abdel-Khalik, L. Crosatti, D.L. Sadowski, S. Shin, J.B. Weathers, M. Yoda, the ARIES Team, Thermal-hydraulic studies in support of the ARIES-CS divertor design, Fusion Sci. Technol., submitted for publication. [12] ANSYS Inc., ANSYS Release 9.0 Documentation, 2004. [13] S. Zinkle, S. Majumdar, N.M. Ghoniem, S. Sharafat, Materials consideration and data base, APEX Interim Report, UCLA, November 1999 (Chapter 13) available at http://www.fusion. ucla.edu/apex/interim report.html. [14] S.J. Zinkle, N.M. Ghoniem, Operating temperature windows for fusion reactor structural materials, Fusion Eng. Des. 51 52 (2000) 55 71. [15] S. O Dell, Personal communication, June 2006. [16] L. El-Guebaly, P. Wilson, D. Henderson, M. Sawan, G. Sviatoslavsky, T. Tautges, R. Slaybaugh, B. Kiedrowski, A. Ibrahim, C. Martin, R. Raffray, S. Malang, J. Lyon, L.P. Ku, X. Wang, L. Bromberg, B. Merrill, L. Waganer, F. Najmabadi, the ARIES Team, Designing ARIES-CS compact radial build and nuclear system: neutronics, shielding and activation, Fusion Sci. Technol., submitted for publication. [17] X.R. Wang, S. Malang, A.R. Raffray, the ARIES Team, Configuration design and maintenance approach for the ARIES-CS stellarator power plant, in: The 17th ANS TOFE, Albuquerque, NM, November 2006, submitted for publication. [18] A.R. Raffray, L. El-Guebaly, T. Ihli, S. Malang, X. Wang, the ARIES-CS Team, Engineering design and analysis of the ARIES-CS power plant, Fusion Sci. Technol., submitted for publication.