PROGRESS OF ITER FIRST WALL DESIGN ABSTRACT 1. INTRODUCTION

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1 PROGRESS OF ITER FIRST WALL DESIGN R. Mitteau(*), R. Raffray(*), P. Chappuis(*), M. Merola(*), D. Loesser(**) on behalf of the ITER Organization and the Blanket Integrated Product Team(1) (*) ITER Organization, Route de Vinon sur Verdon, Saint Paul Lez Durance, France. First Author : Raphael.mitteau@iter.org (**) Blanket Integrated Product Team, Plasma Physics Laboratory, Princeton University, Princeton, NJ USA ABSTRACT The blanket first wall concept has substantially evolved since the ITER design review of 2007, toward an improved accounting of the constraints of high power steady state operation. A major incentive for the design change has been the need to account for large steady state plasma heat fluxes to the first wall. The current first wall design is described, with emphasis on specific accounts for the needs of steady state operation : power exhaust, bond health, avoidance of critical heat flux and plasma control. 1. INTRODUCTION The design of the ITER blanket first wall has substantially evolved since the project design review of Major incentives for the design changes have been the need to account for large plasma heat fluxes to the first wall and the need for efficient maintenance of first wall components. In addition, integration of the vertical stability coils and possibly of the ELM control coils causing large cut-out in the shield blocks have also justified an intense redesign activity. Changes to the heat load specification have led to the introduction of shaping of the first wall components, in order to recess the sharp edges. Thanks to the shaping of the first wall components, direct plasma contact to the first wall at moderate power (7.5 MW) is possible, which led to the introduction of wall start-up and ramp down. The problematic moveable port limiters are abandoned. The change toward easier maintenance of the first wall has led to plasma facing components that are independent from the shield block. The first wall is now made of panels having the same size of the shield block, that are in-vessel separable. The current baseline concept is described in section 2. One of the most challenging design change concerns the need to handle plasma heat loads to the first wall. This requirement is a strong design driver to the first wall, and actually, the shape design of the first wall panels is determined by keeping moderate loads for all steady state conditions. An engineering heat load envelope was established. It is addressed by 1

2 employing actively cooled plasma facing components, which are designed according to the heat load magnitude they have to extract. This requirement is closely linked to steady state operation. An update of the plasma heat load specification has been released, with first wall heat load increased by typically one order of magnitude. Key new requirements are listed in section 3. Section 4 gives more detailed examples on how the first wall design addresses the specific needs of steady state operation. Reliable operation requires a sound first wall that can transmit the heat to the cooling circuit. This is obtained through ensuring a high bond quality at the interface between armor tile and heat sink. A dedicated experimental test program validates the bond behavior at the proper heat flux and lifetime. The continuous heat exhaust in adequate thermo-hydraulic conditions is done by a proper design of the cooling circuit, and cooling conditions have been reviewed. Stable operation is achieved by avoiding critical heat flux in the cooling channel, and current working parameters are listed. Finally, insights about the risks associated to plasma wall gap control are given. 2. OVERVIEW OF FIRST WALL DESIGN The first wall is composed of panels having the same segmentation as the 2004 blanket modules (Figure 1). The needs for an in-vessel separable first wall lead to the definition of a mechanical structure fully independent from the one of the shield block. Evenness of the finger heating power distribution and minimization of electromechanical efforts drives the design toward a central poloidal beam, with toroidal overhanging plasma facing component (PFC) units attached to it. The beam is attached to the shield block through a single bolt, located deep inside the shield block for acceptable nuclear heating (Figure 2). The bolt is preloaded, and the load is reacted by four compression pads are pressed against the shield bock. The compression pads are also used to react the torques created by the electromagnetic loads during VDEs and disruptions. The contact pressure also helps keep a good thermal transfer of the pads to both beam and shield block, for keeping the pad temperature within allowable. The first wall beam serves also as a water manifold, distributing the water to the plasma facing units. The beam is fed through branch pipes from and to the shield block. A multilam electrical connection provides electrical contact to the shield block. The beam is serviced from the front face through holes giving access to the bolt, water connection, and gripping holes (as well as some diagnostic parts for some units). The holes are made possible because of recess of the shaping in the central region of the first wall panel, allowing the tile edges to receive acceptable heat load. The plasma facing surface is shaped to reduce the power deposition on the panel edges down to acceptable values, while keeping acceptable heat loads on the beryllium amour tiles. 2

3 Handling these high heat fluxes has necessitated the use of enhanced heat flux panels capable of accommodating an incident heat flux of 5 MW/m² in steady state, utilizing a technology similar to the one of the former port limiters. The port limiters have now been abandoned and the enhanced heat flux panels are also used on three inboard and four outboard toroidal rows of first wall panels, allowing wall-limited steady state plasmas to a current of 7.5 MA (both inboard and outboard), consistent with ramp-up and ramp-down scenarios. 3. KEY STEADY STATE RELATED REQUIREMENTS While the 2004 wall design took no credit for convection/conduction heat loads to the main chamber wall, the heat load associated to charged particles are now listed as the principal surface heat load to the first wall. The heat flux is oriented along the field line, which means that the engineering heat load (incident heat flux) is strongly design dependant. The load depends on the incidence angle of the field line on the component surface, meaning that both detailled geometrical arrangement of the wall and magnetic surfaces determine the heat load. The shadowing pattern as well as the heat flux radial decay in the scrape off layer are also important elements. For overcoming this difficulty, the heat and nuclear load specification only describes the plasma heat loads, parallel to field lines. A description of these heat load is published in [1,2], including main justifications. The heat load specification has been extended to include all phases of the plasma scenario (start-up and ramp down, flat top, offnormal situations, localized heat loads). The specification of the parallel heat load assumes an axisymetric scrape off layer (SOL); the parallel heat flux decays with radial distance into the SOL and does not depend on the toroidal angle. The four main heat loads are: 1. Conduction/convection at the top panel during the burn phase. The upper X-point at the top of the magnetic configuration diverts flux tubes to the ceiling of the chamber during full power pulses, starting with flux surfaces which are 4 cm outside the separatrix at the OMP (outside midplane). The combination of static heat load (8 MW/m² locally) and time-averaged heat load from ELMs (25 MW/m²) gives a parallel heat flux of up to 33 MW/m². Proximity to a null point of the poloidal magnetic field means that the field lines can run almost in the toroidal direction, with a very small pitch angle. The design of protection of gaps that run in the toroidal direction is challenging because flux tubes can penetrate deeply into the gap. The need to accommodate both positive and negative helicity makes this a design case of special complexity. 3

4 2. Far SOL heat loads load Converging results from all tokamak indicate large tails of plasma density and temperature that extend into the far scrape off layer, due to turbulence driven convective (filamentary) transport. Although the heat flux is moderate (a maximum of 8 MW/m² along the field lines, after having added the time averaged ELMs), the long heat flux decay length (up to 170 mm) makes it a design challenge, because the distance from the separatrix does not provide very effective protection against the heat load. This means basically that all toroidal facing edges need be shadowed during this operation phase. 3. Local heat loads caused by the additional heating systems requires special care for some panels. These include a maximum heat flux of 4 MW/m² ( ) to the panels located in front of the neutral beam (shine-through), and of 20 MW/m² (//) for the panels on the side of the ion cyclotron resonance heating system. 4. The heat load during limiter operation for start up and ramp down are reviewed, for wall limiter operation. Both inboard and outboard operation are envisioned. Rows are designed for plasma contact on the inboard, with a parallel heat load up to 25 MW/m². On the outboard, contact is allowed for modules 14 to 17, and the parallel heat load reaches 40 MW/m². Rows 16 and 17 are involved for a first plasma contact during the reduction of elongation of the plasma at ramp down. Based on this parallel heat flux specification, a shape design was designed for the first wall panels, that gives engineering heat loads that are within capability of the current available technology [3,4]. A detailed heat load envelope was established for each FW panel. The obtained engineering heat loads are larger than the one that would have been obtained on a perfectly smooth wall, however the design ensure that the panel edges receive heat loads that are compatible with steady state operation. 4. ACCOUNTS OF STEADY STATE REQUIREMENTS ON THE DESIGN 4.1. Health of bond between armor tile and heat sink The bond between armor tiles and heat sink is a crucial point of actively cooled plasma facing components. Due to differences of material properties between armor tile and heat sink (Young s modulus, coefficient of thermal expansion), mechanical stresses are created around the bond during thermal cycling. They are of the same order of material resistance, and this makes that bonds are a location of increase fatigue and damage. A common failure 4

5 mechanism for such bonded structures is the debonding of the tile. There is no widely accepted technique for describing by analysis the behavior of heterogeneous bonds. This is still a domain of dedicated research activities. (An insight of semi-quantitative modeling is given in section 4.4). For practical engineering purposes such as the design of the beryllium to copper bonds for the first wall, the design activity relies on design by experiment. This means the functionality of the bonding is demonstrated by experimental tests. A broad R&D program has been carried out since more than a decade in several laboratories around the world [5-9]. The bonding of beryllium tile to copper heat sink has been demonstrated up to about 10 MW/m² (The former port limiter where planned to operate at about this heat load), and other bonds have sustained more than 10 4 heat cycles at a lower heat flux (at 0.8 MW/m²). The design heat flux & lifetime planned for the ITER first wall are compatible with current technological capabilities. While R&D activity is pursued (tile size optimization, determination of acceptance criteria), a larger part of the experimental program is oriented toward prototype testing. After completion of the qualification program, all domestic agencies fabricating first wall panels (CN, EU, RF) have qualified for the next step, which is qualification of semi-prototypes mock-ups. The semi-prototype poses new challenges in term of bonding quality, as a large number (90%) of tiles shall resists cycling at the design heat load (Figure 3). Additionally, the mock us design has key technical features of the baseline first wall panels (shaped PFC units, attachment of PFC units to the first wall beam including water connect). The future test plan includes tests of full scale prototypes Heating power removal Section 3 deals mainly about the engineering heat load at the tile scale. This load is relevant for the local structure design, including tile/structure bonding and margin to critical heat flux (see section 4.3.). The heat load at the scale of the component (PFC unit, blanket module) is of less importance for the local structure design, but it is the major driver for defining the cooling circuit (series/parallel design, mass flow rate requirement). The design driver is the outlet temperatures of the various circuits, which is limited to an allowance dictated by structural limits and water circuit design (+50 C maximum at the outlet of the tokamak cooling water loop). The total heating power to the blanket during the DT inductive phase is of 736 MW. This heat is distributed between 596 MW of nuclear heating (neutron heat load, neutron energy multiplication factor and design error about fusion power measure) and 140 MW of surface heat load that is deposited mainly on the first wall. In the 140 MW design situation, the heat flux to the first wall is mainly carried by radiation and charge exchange particles. It is well distributed spatially, and the wall loading is about 0.21 MW/m² on the 660 m² of the first wall. A first wall panel of 1.5 m² (typical for inboard wall) receives a surface 5

6 power of 320 KW. There is however another more constraining design case regarding the surface power, because the maximum heat flux to the surface can reach 0.5 MW/m². The 1.5 m² panel receives then 750 KW, a power larger by a factor 2.4 than the one given by the 140 MW case. This illustrates the quite different design situations that the wall cooling circuit design needs to accommodate: High total power case, where the total heating power to the blanket is large but well distributed. This design case gives the highest temperature at the outlet of the first wall loop, when all individual circuits have rejoined. High local power case, where the total heating power might be moderate, but the heating power to one given blanket module is large. This design case gives the highest temperature (+62.5 C) at the outlet of a given blanket module. Due to the remixing of hot and cold water coming from different blanket modules, the temperature increase at the outlet is smaller than +50 C. The case of the panels of rows is singular, because they are loaded by the intense heating power around the second null (plasma legs at the 4 cm flux surface, increased radiation). The maximum surface power per module is defined here to a total load of 15 MW for rows 8&9, 10 MW for rows 7&10, hence the maximum surface heating power to a top FW panel is defined as 1/18 of it. A toroidal peaking factor of 1.5 is evaluated based on a geometrical approximation, giving a maximum surface power of 1.25 MW for the panels of rows 8 and 9. The blanket mass flow rate distribution accounts for all design cases, so that the maximum outlet temperature of each module is the same. The design numbers are given in Table 1 (Design total surface power and the design mass flow rate for regular panels. Regular panels are toroidal sections of 10 -outboard- and 20 -inboard.) Prevention of critical heat flux Heat transfer into the water is usually the most critical step of transmitting the plasma heat to the cooling circuit. For the ITER first wall, this is especially the case for the enhanced heat flux panels. Because of the intense heat flux passing from the heat sink to the water, the heat sink wall temperature is larger than the water saturation temperature. For first wall enhanced heat flux components which are loaded at 4.7 MW/m² and cooled by water that reaches locally 170 C, the wall can be at 260 C while the saturation temperature is of 222 C at 2.5 MPa. At the cooling tube wall, the water does not remain in liquid phase flow and local boiling occurs. The vapor bubbles transit quickly to the bulk water stream because of the highly turbulent flow. Vapor bubbles recondense in the bulk stream because the bulk temperature is less than the saturation temperature. This phenomenon is called subcooled boiling, and provides very efficient heat transfer from a hot wall to a cooling fluid. The heat 6

7 transfer coefficient increases by about one order of magnitude. This is a standard operating regime of high heat flux components like the divertor. However, the phenomenon is bounded by the boiling crisis, which occurs when bubble transit is not fast enough to evacuate the heat as quickly as it arrives. The wall may then dry out and heat up, until the maximum stress resistance is exceeded due to falling material strength with temperature. The structure of the plasma facing component then fails catastrophically; this is the critical heat flux (CHF). The blanket system requirement document requires a margin of 1.4 to this event (the margin is defined as q design /q CHF ). Most of the CHF experimental points are for high heat flux application (ΔTsat > 50 C, V = 5-15 m/s), and are barely usable for first wall PFC units. Tests are on-going to refine the CHF database, for parameters relevant to the blanket operation (ΔTsat < 50 C, V = 2-4 m/s). Early results show that bare rectangular channel do not meet the required margin of 1.4. Two methods have been implemented toward meeting the required margin of 1.4: (1) the use of hypervapotron (Figure 4), and ( 2) running the coolant inseries over two PFC units per circuit (for doubling the local velocity). Current on-going tests indicate that these actions would provide the required CHF margin Separatrix-wall gap control Heat load to the wall are defined on the quasi-static time scale, expressing the notion that the heat load are averaged over ~10 seconds for design purposes. There is however a large variety of short transient heat loads to the first wall, and they impact essentially lifetime and plasma operation. The heat loads associated to off-normal events such as vertical displacement events, disruptions and runaways are well described in the heat and nuclear load specification and are not detailed here. They impact mainly the lifetime. Other transients are related to plasma movements, a topic that interfaces with plasma control, bringing operational risks. Plasma movements are quantified by the geometrical gap between magnetic surfaces and the wall. During flat top operation, the gap between the 4 centimeter flux surface and the wall shall be larger than 8 centimeter on quasi-static time scale. Plasma movements occur on times scale shorter than 10 seconds. Due to the distance from the poloidal field coils to the plasma, and the shielding effect of the vacuum vessel and the blanket, plasma position feedback operates with a time constant of about one second. The consequence is that the overall time of the plasma movement spans over several seconds, and the plasma come close to the wall over similar durations compared to the thermal time constant of the PFC units (3 seconds). As the plasma heat flux increases rapidly as a function of the distance to the separatrix in the scrape of layer, large heat loads are applied to the wall during such plasma movements. Depending on the initial condition, heat loads larger than the design heat flux of the first wall panel can be applied to some PFC units. Conventional devices using radiation cooled massive tiles are 7

8 not sensitive to such events, because the heat is stored in the tile whatever the rate it arrives. Hence the heat arrival rate is not a strong driver to the operation. This is not the case for actively cooled plasma facing components, which undergo strong damage above the design heat flux, and may even fail in about 1 second. As mentioned in section 4.1, only experimental tests would allow characterizing the component behavior under higher than design conditions (q > q design). For the purpose of quick evaluation, semi-quantitative transient thermal analyses are run to evaluate the effect of q > q design transients on the bond. A triangular wave profile was employed, for various heat flux excess (130% of q design, 160% of q design ) and various durations (Full width at mid height = 0.5 s, 1s, 2s ). The thermal gradient at the interface (expressed in C/mm) is plotted as a function of time. This quantity is used as a relative indicator of the stress at bond. A thermal gradient of 7.7 C/mm is calculated at the bond for q design = 1 MW/m². At 2 MW/m² (double the design heat flux, representative of a 8 centimeter plasma movement toward the outboard) for a thermal pulse wave of 4 seconds, the thermal gradient is of 12 C/mm, indicating a typical 55% increase of the stress at the bond. Such a figure is not acceptable; a figure of 10% increase seems closer to the allowable, for avoiding excessive reduction of lifetime. For a 2 seconds movement, the heat load should remain lower than 1.3x q design and the allowed time at 1.6 x q design is only 0.5 seconds. Careful plasma control will be needed to avoid damaging the wall. For extreme movement, e.g. as might occur under some circumstances at an H-L transition, the disruption mitigation system might need to be employed for preventing wall failure. Such consideration will need to be included in plasma operating instructions which can be developed in the course of operating experience with ITER plasmas. 5. CONCLUSION Following important design changes, and a new issue of the heat and nuclear load specification, the blanket first wall is redesigned. The new design corresponds to better accounting for the specific needs of steady state operation. First wall panels have been shaped in order to keep engineering heat load under allowable in term of stress and temperatures. This has led to the use of enhanced hat flux panels, capable of 5 MW/m², on about 50% of the blanket modules. The wall has also included the function of start-up and shut down of the plasma. A conceptual design is finalized, and the preliminary design review is planned for late REFERENCES [1] A. Loarte, et al, nd IAEA Fusion Energy Conference, Geneva, Switzerland, IT/P6-1 8

9 [2] R. Pitts, tproc. Of 13 th SOFT conference, to be published in Journal of Nuclear materials [3] R. Mitteau, Fusion Engineering and Design 85 (2010) [4] P Stangeby, et al, J. Nucl. Mater , (2009) [5] Assessment of water cooled beryllium components for plasma facing applications, E.B. Deksnis, HD Falter, AT Peacock et Al., Jour. Nucl. Mater (1994) [6] Beryllium armored Mockups for fusion high heat flux applications, A. Gervash, R. Giniyatulin, I. Mazul, R. Watson, Proc. of 20 th SOFT conference, Marseille, 1998 [7] Comparative thermal cycling test of different beryllium grades previously subjected to simulated disruption loads, A. Gervash, R. Giniyatulin, I. Mazul, Fus. Eng. Des. 46(1999) [8] Low cycle thermal fatigue testing of beryllium, R.D. Watson, DL Youchison, DE Dombrowski et Al., Fus. Eng. Des 37(1997) [9] Application of beryllium as first wall armour for ITER primary, Baffle and limiter modules, A. Cardella, V. Barabash, K. Ioki et Al, Fusion technology Vol. 38 (2000), pp

10 FIGURES Table 1 : Nuclear and surface heating power to the blanket modules, and required water mass flow rate. Figure 1 : ITER plasma chamber with in vessel components. 10

11 Figure 2 : An ITER First wall panel (inboard side) Figure 3 :First wall semi prototype design 11

12 Figure 4 : test mock up with hypervapotron cooling channel Figure, 5: Thermal gradient at Be/CuCrZr interface vs. time: q_transient=2 MW/m2 and Δt=4 s 12