ARIES Project Meeting Minutes

7-9 January 1998 1997

University of California, San Diego

Prepared by: L. Waganer

Attendees: M. Billone, L. El-Guebaly, S. Jardin, C. Kessel, H. Khater, S. Malang, T.K. Mau, Ron Miller, Bob Miller, E. Mogahed, F. Najmabadi, W. Reiersen, R. Stambaugh, D. Steiner, D-K. Sze, M. Tillack, L. Waganer, X. Wang

Attachments: ARIES Meeting Agenda (7-9 January 1998)


The main purpose of the meeting was to determine the readiness of the physics and engineering baselines to select the parameter space for the final strawman Spherical Tokamak (ST) design. The project had hoped to be able to finalize the strawman parameters at this meeting, but neither the physics nor engineering groups were confident that the optimal parameter space had been found. The current intent is to determine the correct physics regime within three weeks, which would enable the definition of a new set of physics and engineering parameters by 15 February. This would facilitate a new strawman being generated by the end of February 1998. The next project conference call is scheduled for 22 January (Thursday) to critique our progress toward these milestones.

Physics Team Results - Spherical Tokamak Design

Equilibrium Results - Steve Jardin summarized the PPPL work on LAR equilibrium configurations that are "stable," id est, stable to ballooning modes (not affected by conducting walls) and stable to kink modes (strongly influenced by nearby conducting walls). For long plasma times compared to the resistive time of the wall, active feedback may be required (and/or plasma rotation). If active feedback is required, a resistive wall will be required to slow the plasma movement sufficiently that the active feedback can be accomplished within time and energy constraints (see following talk by Kessel for more details).

It was recommended that two equilibrium classes be investigated, with and without conducting walls. The wall-stabilization cases will have both high beta (b) and a high ratio of bootstrap current to plasma current (IBS/IP). The cases without wall stabilization will have significantly lower values of b and/or IBS/IP). and will probably not lead to an attractive power plant design. Nevertheless, the data should be examined and compared to the other cases. Steve illustrated several early results which indicated the general parameter space and the trends expected. However, he cautioned that the data are preliminary. Final optimized and stable plasma parameters and related wall characteristics are expected in three weeks for input in the ARIES system code.

Stability Results - Chuck Kessel described his approach to obtain an optimal PF coil set with the fewest number of coils. He first evaluates the PF system energy with a large number of PF coils placed along a defining coil envelope. Then, in turn, each of the coils is removed, the equilibrium current conditions for the remaining coils are recalculated, and the energy is recomputed. The eliminated coil which produces the lowest increase in system energy is chosen to be permanently removed. Then the process is repeated until the lowest number of coils is obtained, typically a six-coil set (three above midplane and three below). The top and bottom coils, which define the diverted plasma position, might be placed inside the TF coils, but there would not be a significant benefit from a stored energy standpoint.

Plasmas with elongations (k) greater than the natural elongation will be vertically unstable. For example, for A=1.6, li ~0.18 and bp~1.83, the k (natural) is approximately 1.8 to 1.9, while the desired elongation (k) is 3.2. Hence some nearby conducting structure is needed to slow the vertical instability from tens of microseconds to tens of milliseconds. A representative plasma case was evaluated with the TEQ code which yielded a reasonable (plasma) vertical growth rate around a wall distance ratio to minor radius of 0.4 although the TEQ code cannot produce the actual plasma equilibrium conditions. Additional cases were run on the higher fidelity (and time intensive) TSC code, but it was found that the k = 3.25 plasmas were ideally unstable and the position of the conducting wall had little stabilizing effect. Moreover, the plasmas were non-rigid and deformed quite easily. It was felt that lowering the elongation (k) would be beneficial in finding a plasma that would be more easily controlled. It was suggested that the same case run with the TSC code be run on the TEQ code. Also it was suggested that the conducting wall would be assumed as continuous to compare to a similar GA case.

Current Drive and Startup - TK Mau presented a graph of possible RF frequencies for an on-axis seed current drive in a 50/50 DT plasma using the A=1.8 ARIES-ST strawman. ICRF (ion cyclotron resonance frequency) has a window through the plasma at 28-32 MHz, but it cannot penetrate due to high-beta damping and propagation effects. LFFW (low frequency fast wave) at less than 11 MHz can penetrate the plasma, but requires a large antenna. Little current drive experimental evidence exists for LFFW systems. [Later in the meeting, TK sketched a preliminary concept for a LFFW antenna.] Direct ECH (electron cyclotron heating) current drive cannot be used because of accessibility problems. However, ECH might be coupled with EBW (electron Bernstein wave) or the EBW could be generated directly to drive the plasma current, but more theoretical work is required.

Neutral beams could also be used to drive plasma currents at the center depending on the beam energy (>2 MeV may be required), using the NFREYA code on the A=1.8 ARIES-ST strawman. The current cannot selectively be deposited only at the center, rather plasma current will be deposited on all layers up to the edge. A typical NB current deposition should be superimposed onto the bootstrap current profile in order to obtain the complete current profile. Also NB current drive will be more efficient driving the plasma current. Negative ion beams with RFQ accelerator technology will be required.

TK also noted that recent theoretical work (Shaing, et al.) has established that the bootstrap current on axis is actually not zero. It was shown that magnetically trapped particles close to the axis do not move in banana orbits, rather they move in orbits described as potato orbits. This effect of the potato orbits and non-zero bootstrap current on axis will be included in the physics analyses.

In summary, LFFW can drive plasma currents on axis with a peaked profile; but large antenna structure may be problematic, which will be addressed. ECH has the potential for current drive, but much research work is needed. Neutral beams can provide peaked on-axis current drive, but some current drive will also occur off-axis - need to evaluate the combined effect with the predicted bootstrap current profile. TK Mau will determine reference current drive systems (on-axis and off-axis) and develop current drive scaling over a range of aspect ratios.

GA Physics Study Results - No work had been accomplished by GA for the project during the intervening period so no presentation was given. There was some concern during the meeting about the size and location of TF coil power supplies, so Ron Stambaugh showed a slide of the ST pilot plant power supplyconfiguration. This configuration affirmed the requirement to locate the TF coil power supplies as close as possible to the outer perimeter of the coils. Ron also showed some general charts of plasma elongation versus aspect ratio with various stability limits. GA calculations of vertical stability using GATO have found ideally stable plasmas in the presence of a conducting wall with elongations over 3. He presented scoping charts of global performance parameters for the general class of superconducting tokamak commercial power plants.

Physics Summary of Issues and Action Items .

Systems Study Results

Ron Miller reviewed the most recent historical data of US electrical generation capacity, fuel prices/trends, and new plant cost projections. Ron recently completed a new systems code run with the new physics parameters, but did not show the data because the input data would likely change. The higher joule losses in the CP did push the COE higher. Currently he is modeling the straight centerpost but can flare the centerpost if needed. He discussed the plot of COE versus the major core radius for various aspect ratios. An aspect ratio of 1.6 appeared to be the best of the current data sets. The curves for constant aspect ratios are truncated at lower values of aspect ratio because of TF field limitations (superconducting limit?). The systems code will be modified to incorporate new concepts and data from both physics and engineering (due Feb 15), after which, a new strawman will be developed by the end of February.

Systems Summary of Issues and Action Items .

Engineering Team Results - Spherical Tokamak Design

Stress Analysis of TF Coils for a Low Aspect Ratio Tokamak- Leslie Bromberg was unable to attend the meeting, but forwarded his slides to Mark Tillack for presentation. Since the analysis was done on an obsolete TF and PF coil configuration, it was decided it would not be of value to present the slides.

Helium-Cooled Blanket and Divertor Design and Analysis - Mark Tillack summarized the UCSD efforts to develop a first wall, blanket, and divertor system that is compatible with the ARIES-ST design constraints. The European Union extensively analyzed a dual coolant first wall design to meet a 0.5 MW/m2peak surface heat flux. However, the ARIES-ST has an average heat flux of 0.65 MW/m2 (0.8 peak MW/m2) on the first wall (assuming 50% of the total transport heat is deposited uniformly on the first wall). Mark conducted his thermal analysis using some total power data that seemed to be in error, which will be corrected in the next iteration.

The general configuration of the helium-cooled ferritic steel blanket with a flowing lithium lead blanket is shown in the attached figure. Helium cools the first (front) wall and side walls with the cooler helium (inlet temperature 350°C) to keep the first wall structural temperature below 550°C. Helium also cools the interior structural elements. This first wall coolant is heated to approximately 450°C and is sent to the first intermediate heat exchanger. The breeding lithium lead coolant absorbs the nuclear heat behind the first wall and is allowed to attain a higher temperature (~700°C) than the first wall helium coolant. The higher temperature lithium lead is sent to a second intermediate heat exchanger to increase the helium up to a temperature (~650°C) necessary for efficient Brayton cycle energy conversion. The lithium lead is thermally insulated from the ferritic steel so that the structural steel will not exceed its allowable temperature limits. The insulating material is tentatively SiC. There was concern about lithium lead being between the SiC insulator and the ferritic steel structure, which could allow tritium to accumulate in the interspace. If the space was wide enough, the lithium lead would slowly migrate out of the narrow space.

Ferritic steel has a nominal upper temperature limit of 550°C (decrease in tensile and creep strength) and a lower limit of 300°C ± 50°C (ductile-to-brittle transition temperature). Use of an oxide dispersion in the ferritic steel and /or a low primary stress system may allow the use of higher temperatures. Anticipated improvements in the development of low activation steels may result in lower DBTT values and, hence, lower operating temperatures.

A low film drop inside the helium coolant channel on the first wall is important to maintain acceptable temperatures. Single-sided roughness allows a doubling of the heat transfer coefficient with a modest and acceptable increase in the pumping power.

The attached figure shows the preliminary wall thickness and channel dimensions which seemed reasonable; however, mechanical and thermal stress analyses have not been completed. During the later discussion on achievable tritium breeding ratio, it was suggested that the first wall be thinned from 4 mm to 3 mm and the second wall be reduced in thickness by 25 to 50%. Moreover, it was suggested that the amount of lithium lead breeder be increased from three cells deep to four as shown in the above sketch. Accompanying the change to a thicker breeder section will be a thinner shield.

Divertor Design Options - Mark Tillack discussed the use of empirical "compression" factors between the divertor region and the core plasma to improve the accuracy of the predictions. There was concern if factors from existing experiments are appropriate, since they depend on several conditions, including edge temperature and particle flux - R. Stambaugh agreed to investigate. Even with a liquid metal target, a radiative divertor will be necessary for ARIES-ST due to the extremely high power flow along the scrape-off layer.

On a solid surface divertor, there is a need to incorporate an erosion-resistant surface, such as tungsten. In order to minimize thermal stresses at the tungsten/substrate interface, the surface should to be castellated (divided into small elements after joining to base structure). Milling or sawing the tungsten surface is possible but costly. An alternative technology is being developed by the ITER Divertor group that uses tungsten rods mechanically attached to a copper base structure as shown below courtesy of Boeing.

Mark presented the amount of power to be handled by the divertor, but the total power seemed to be low for the cases being considered (need to include bremsstrahlung radiation and higher recirculating power for straight centerpost). The requirement to include inboard divertor plates essentially prevents centerpost flaring, and so the possibility to eliminate inboard particle fluxes was debated. The team suggested that the inboard slots be retained. The divertor geometry and power balance have not been finalized.

Power Core Configuration and Maintenance Options - Xueren Wang summarized the major configuration parameters of the ARIES-ST power core. The CAD drawings are not yet in exact agreement with the systems code data, as both are evolving. This was especially evident with regard to the scrape-off region and the nearby first walls. But conceptually, they are in agreement. One of the major maintenance assumptions is that there is no hands-on maintenance inside the primary vacuum vessel, which is outside the TF and PF coils. There is also an inner vacuum boundary (defined to be the inboard, divertor, and outboard shield) to contain plasma vacuum conditions and a majority of gaseous DT. Pumping ports through the two vacuum barriers should be shown near the core midplane, and the cryopumps should be located outside the TF coils. Wayne Reiersen suggested moving the busbar connections from beneath the core to a position closer to the midplane. This would free up valuable room at the bottom of the core.

One of the main objectives of the configuration and maintenance task is to select the better maintenance approach. The horizontal approach has been eliminated since it would require the TF coils' outer legs to be too large a radius to extract a core segment. The main emphasis has been to evaluate the vertical maintenance from either above or below. The core geometry modifications for both are roughly similar, either on the top or bottom. But the space for the removed core elements and the lifting or lowering capability is quite different. The mass of the removable core components totals 2345 tonnes. A reasonable extrapolation of crane capability is 1500 tonnes. So the replaceable unit, when lifted with a crane, should be divided into two elements. Still the concern is about the lifting and positional capabilities of such a large component with a crane. The volume needed to contain the radiation contamination is very large. The building is quite high with large wall thickness. Thus, the vertical maintenance from above (the top) was rejected.

The vertical maintenance from below (the bottom) requires power core maintenance features similar to maintenance from above. The entire core is supported from below with a removable structure. When maintenance is required, several hydraulic lift cylinders support the core and centerpost while the operational supports are removed. Then the core and/or centerpost is removed into a sealed antechamber below the power core. After the lower vacuum door immediately below the reactor is closed, the removed core and/or centerpost will be transferred to the hot cell and the replacement core and/or centerpost is transferred to the antechamber and subsequently into the reactor. The distance the components are moved is the same in both approaches, but the hydraulic lifts are capable of heavier lifts and more precise movements. This heavier lift capability allows the entire power core to be replaced as a unit, as opposed to two separate elements. The walls can be thinner for the latter approach. Hence, the vertical maintenance from below was adopted.

One concern was that only the components which need to be replaced actually are removed. Hence, there should be a parting plane between the permanent shield and the replaceable shield.

Magnet Systems Design Studies - Wayne Reiersen developed a conceptual design for the TF coil sets including the capability for maintenance. He developed a concept featuring vertical maintenance from above. With minor modifications, his concept will be incorporated into X. Wang's CAD model for the ARIES-ST. Wayne's concept featured a nearly spherical shell for the outer legs of the TF coils, with a straight centerpost. Busbar connections were located at 300 below the midplane. The busbar leads were shown leaving the machine vertically, but it was suggested to have them leave horizontally and have the power supplies placed in close proximity. Vacuum ports should be on the midplane, with vacuum pumps also placed in close proximity.

There was a concern that the currents in the outer legs of the TF coils could not be properly balanced if the shell were a continuous, common sphere as opposed to separate parallel outer legs. Wayne will determine if this is a problem. The schematics generally describe the two approaches.

The decision to flare the centerpost has not been made yet. The incentive to flare is to decrease the centerpost resistance due to the inclusion of more copper. However, a flared centerpost effectively traps the inboard shield during assembly and maintenance. Moreover, if an inboard divertor slot is required (it likely will be required), the actual copper increase due to flaring will be minimal.

Wayne stressed the importance of establishing envelopes and openings for the vacuum pumping system.

The decision to remove the core vertically downward establishes the position of the structural support connections. The envelope of the removable portion of the power core should be provided to Reiersen ASAP to complete the TF design.

Wayne discussed the assembly and disassembly of the power core. This is useful information but the maintenance procedure has been determined to be modular replacement with assembly and disassembly taking place in the hot cell.

The thermal analysis of the central post does not indicate any superior water-coolant condition. A cold-water cooled centerpost is possible but not desirable due to copper embrittlement. Moreover, a straight centerpost has a high resistance power loss. Warm-water cooling (to retain ductility) does not provide enough cooling at the predicted current densities. Wayne recommended a constrained parameter analysis of aspect ratios from 1.6 to 1.8 and current densities appropriate for the new strawman case for both the cold and warm water cooling cases.

There was a renewed interest in assessing if a superconducting TF coil system might be feasible. This low aspect ratio regime had never been assessed in the previous ARIES systems studies, but the trends at A=3 indicated the COE would increase for further reductions in the aspect ratio. These trends were supported by the previous Ron Stambaugh charts. The space requirement for protection of the central TF coil legs mandates a much larger machine size (more costly but lower recirculating power) than a comparable copper machine.

Neutronic Analyses of Power Core - Laila El-Guebaly discussed her 3-D neutronics results with the MCNP code as compared to her prior 1-D code results. The core geometry and the neutron source distribution were from the 9/97 ASC code strawman. The modeling of the neutron source distribution seemed to be coarse, but Laila assured the group that the model had adequate fidelity. Homogenized, but segmented, zones were established for the first wall, blanket, manifolds, shields, divertors, and centerpost with no toroidal gaps or ducts.

A comparison of the neutron wall loading results showed a good correlation between the MCNP and the ASC code. The nuclear heat loads were reasonable except, that the outboard blanket and manifold should be thickened to reduce the heating in the outboard LT shield. The first wall and blanket intercept 80% of the nuclear heating. There was an excellent agreement between the 3-D and the 1-D codes on the energy multiplication values. The 1-D code slightly overestimated the tritium breeding ratio (TBR) results. So the TBR calculation with 1-D code should be renormalized to the new 3-D data. It was felt the best design approach was to add more depth in the outboard blanket (add a fourth row of cells). The group asked if varying the thickness of SiC could be assessed. It was also decided to reduce the first and second wall thicknesses to improve the TBR and reduce the inboard shield thickness to 20 cm.

The predicted radiation damage to the inboard and outboard first walls was overestimated by 40-80% and 40-60%, respectively, by the 1-D code. However, the radiation damage to the centerpost showed reasonable correlation. All 1-D code results will be normalized to the 3-D code results.

LOCA Analyses - E. Mogahed described his 2-D thermal analysis finite element model for the ARIES-ST power core. He has modified the model so that he can evaluate the three dimensional effects of a heat sink (copper bus bar) by incorporating special conductivity links to the bus bar for both the 30-cm-thick inboard shield and the outboard blanket. The analysis should be rerun with the current 20-cm-thick inboard shield. Another suggestion was to maintain the ultimate heat sink at 35°C rather than allow it to rise over time. This would eliminate the gradual thermal rise in component temperature after long duration.

One of the key design assumptions is the inclusion of a nominal 10% steel conductivity factor between the inboard shield and the centerpost and a 5% conductivity factor between the outboard manifold and the outboard shield. Both structural and non-structural solutions were proposed at the meeting but the negative factor is the undesired conduction during normal operations. L. Waganer maintained there should be novel, passive thermal connections that would only be employed during an abnormal thermal excursion. Actually, E. Mogahed conducted a parametric analysis over a range of conduction factors of 0, 1, 5, and 10%. The results indicated a knee in the curve near 1%. L. Waganer suggested assessing more points near zero in case a trivial conduction area would provide adequate protection yet not impose an operational penalty.

Safety Analysis - H. Khater conducted a safety analysis based upon the L. El-Guebaly 3-D neutronic data just presented. The 20 MWy/m2 fluence should yield a lifetime of 4 full power years (FPY) as opposed to the previous estimate of 2.4 FPY. The outboard low temperature shield lifetime remains at less than or equal to 40 FPY.

The radiation damage in the copper centerpost is most pronounced at the outer radius of the post. The magnitude is highly dependent on the thickness of the inboard shield. For the strawman data, the centerpost resistivity, averaged over the centerpost volume, increases approximately 20% at the end of life. There was significant debate about the correctness of converting the discrete resistivity into a volume averaged resistivity as the current will redistribute to areas with lower resistance. R. Miller took the action to validate the centerpost average resistance for incorporation into the systems code.

The Waste Disposal Ratings (WDR) for Class C were calculated using 10 CFR61 and the Fetter disposal limits, using the proposed ORNL ferritic steel, 9Cr-2WVTa, and evaluating the results after one year following removal from the reactor or shutdown. The dominant nuclide for both limits was niobium. M. Billone said if this was a waste problem, the niobium could be reduced or eliminated from the steel at a moderate cost. After 2.6 FPY, the centepost would meet the Class C waste disposal limits. However, extra shielding would be needed to allow for its disposal as Class C waste after 4 FPY. The polonium radiological hazard arising from the LiPb blanket can be greatly minimized by removing bismuth (which transmutes into polonium) from the initial inventory. Since lead also transmutes into bismuth during operation, the created bismuth will have to be removed from the coolant stream to maintain a minimal amount of bismuth in the coolant. With these assumed modifications, the off-site dose will be less than the 10 mSv required for evacuation in case of an accident.

Evaluation of ST Blanket Options - Dai-Kai Sze affirmed the need to design a blanket that is compatible with the centerpost, is able to handle the high heat flux and neutron wall loading, can operate at high temperature for high thermal efficiency, and can breed sufficient tritium. A lithium/vanadium blanket meets all the criteria except that it is incompatible with the chosen water-cooled centerpost design. A helium-cooled, solid breeder design with ferritic steel (FS) has problems meeting the heat transfer criteria. A LiPb/He/V blanket has a problem with compatibility between LiPb and vanadium due to trace amounts of oxygen and tritium in the coolant. The blanket with the most promise for the ARIES-ST design is the LiPb/He/FS/SiC, which employs both LiPb and He as coolants. Helium maintains the FS first wall and internal structures below the 550°C (or maybe 600°C) structural limit for ferritic steel. Then the lithium-lead coolant is heated to a higher temperature (~700°C) to obtain efficient thermal conversion. Both of these coolants are sent to intermediate heat exchangers to transfer their energies to a helium working fluid for the Brayton conversion cycle. SiC is used to thermally insulate the 700°C lithium-lead from the ferritic steel, which should not exceed 550-600°C.

There are three design issues that are of great concern: (1) the capability of the SiC insulator to thermally isolate the ferritic steel in an effective manner, (2) a high tritium solubility in the SiC insulator (hence a large tritium inventory), and (3) the material compatibility issues of the LiPb/He heat exchanger. There is also design and analysis work to improve the heat transfer capability of the first wall to accommodate the anticipated surface heat loads.

Revised Material Guidelines for Advanced Power Reactors - Mike Billone's slides addressed all the material questions that should be answered for definition of an advanced power reactor. But the discussions centered around the design guidelines for structural materials; namely, ferritic steels and vanadium. He offered some preliminary guidance and promised a more in-depth investigation of several other material issues, such as the tritium solubility in SiC.

The maximum operating temperature of ferritic steel is nominally taken as 550°C. The precise value is design-dependent, with higher values allowed for low primary stress systems. Oxide-Dispersion-Strengthened (ODS) ferritic steels offers the possibility of higher tensile and creep strength at temperatures above 550°C, but these materials need to be further developed, tested and optimized to determine what engineering price is paid (e.g., higher DBTT, joining problems, etc.) for the higher strength. Mike said that the design code guidelines of membrane stresses should be less than Sm (one third of the ultimate stress), the membrane stresses + bending stresses be less than 1.5xSm, and the membrane stresses + bending stresses + the thermal stresses be less than 3xSm. These are to be taken as general rules. If they cannot be met, the particular design and stress conditions should be evaluated. In the case of accident conditions, the upper temperature limit is dependent on the material properties and heat treatment conditions. For instance, if the heat treatment was at 750°C, this might be the upper limit for reuse. But time and temperature exposure determine crystal regrowth, so no general rules should be applied.

There is no consensus in the materials community regarding the design rules for vanadium. The maximum operating temperature is quoted as 650, 700, and up to 800°C, depending upon the source. A lower operating temperature for the onset of ductile-to-brittle conditions (DBTT) is questionable at 400°C ± TBD°C. Different heats of the same nominal composition (e.g., V-4Cr-4Ti) behave differently, and so there is not enough data to build confidence and concensus.

Mike said he would define the ferritic steel design temperature and stress limits. He would work with Sze on the SiC tritium permeation issue and possible replacement material. They would also work on the IHX material issue. Mike and Wayne Reiersen will collaborate on the copper alloy selection and its design properties.

Engineering Summary of Issues and Action Items .

Assessment of Markets and Customers for Fusion Applications

Les Waganer reviewed the purpose and methodology to assess alternate fusion products. The general approach and specific attributes were affirmed in a meeting with Dr. Terry Bahill of Arizona University, who has done similar assessments. In light of the Kyoto environmental recommendations, Les Waganer modified the product attribute weightings and reassessed the product set being evaluated. The results now favored the production of hydrogen fuel and improved the ratings of electricity generation, process heat, and desalination. Transmutation of nuclear wastes and dissociation of chemical wastes remained high on the prioritized list.

Les also presented a list of criteria that should be used to decide which application the ARIES project would choose to investigate upon the conclusion of the ARIES-ST assessment phase. The criteria were:


The next project conference call will be in two weeks, 22 January (Thursday) at 11:30 CST. The new phone number will be 314-232-7034. The next project meeting will likely be in March 1998, but this depends on the progress in the development of the next strawman. Inputs for the strawman from the physics and engineering groups are due at the middle of February so that the strawman can be developed by the end of February.

ARIES Project Meeting Agenda

Wednesday, 7 January 1998

University of California, San Diego

Room 479


1:00 Welcome and Administrative Details......................Miller 1:05 Discussion of Agenda....................................Waganer 1:10 DOE Update..............................................Dove 1:20 Program Overview and Plans..............................Najmabadi

Spherical Tokamak Physics Session

1:30 A Tokamak Stability with and without Kink Stabilizing Shells.................................Jardin 2:15 Low-A Free-Boundary Equilibrium and Vertical Stability Studies..............................Kessel 2:45 Physics Understanding for Divertor Region..............Stambaugh/Miller 3:15 Break 3:30 Current Drive Requirements and Recommendations..........Mau Systems Studies 4:15 Discussion of the New ARIES-ST Strawman and Test Cases..Miller Spherical Tokamak Engineering Session 5:00 Poloidal Field Considerations for ST Reactors...........Bromberg 5:30 Adjourn for Project Dinner 5:30 Project Dinner (location to be announced)

ARIES Project Meeting Agenda

Thursday, 8 January 1998

University of California, San Diego

Room 479
8:00 Coffee and Socializing..................................UCSD hosts

Spherical Tokamak Engineering Session (Continued)

8:30 Implications of Engineering Design Choices..............Tillack 9:15 Power Core Configuration/Maintenance Option Assessment..Wang 9:45 Extra Discussion Period 10:00 Break 10:15 CP Design Assessment................................... Reiersen ........Conductivity Enhancement Design Approach ........Evaluation of TF Coil Strengths for Vertical Downward Maintenance ........Straight vs Flared Centerposts ........Proposal for Vertical Upward Maintenance ........Assessment of Strawman Design 11:00 Divertor Design Options and Blanket Analyses........... Tillack 11:45 Lunch 12:45 3-D Neutronics for the ARIES-ST Power Core............. El-Guebaly 1:30 LOCA Results........................................... Mogahed 2:00 Results of Preliminary Safety Studies.................. Khater 2:30 Blanket Design Issues and Recommendations.............. Sze 3:15 Break 3:30 Revised Material Guidelines for Advanced Power Reactors.. Billone 4:00 Extra Discussion Time...................................All 5:00 Adjourn

ARIES Project Meeting Agenda

Friday, 9 January 1998

University of California, San Diego

Room 479
8:00 Coffee and Socializing.................................UCSD hosts

Spherical Tokamak Engineering Session (Continued)

8:30 Review of Progress in Alternate Applications Study.....Waganer 9:00 Preliminary Evaluation of Levitated Dipole Configuration...Bromberg

Overview of ARIES-ST

(9:35 to 11:00) Summary of Physics Results and Recommended Actions...........Jardin Summary of Engineering Results and Recommended Actions.......Tillack Project Next Actions to Prepare for Baseline Design..........Najmabadi 11:00 Adjourn