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Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles
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Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles R. Mitteau b,∗, R. Eaton b, A. Gervash a, V. Kuznetcov a, V. Davydov a, R. Rulev a a b
Efremov Research Institute, 189631, St. Petersburg, Russia ITER Organization Route de Vinon-sur-Verdon, CS 90 046, 13067 St Paul Lez Durance Cedex France
a r t i c l e
i n f o
Article history: Received 11 July 2016 Revised 6 December 2016 Accepted 1 February 2017 Available online xxx
a b s t r a c t Plasma facing components are usually qualified to a given heat load density applied at the top face of the armour tiles with normal incidence angle. When employed in tokamak fusion machines, heat loading on the tile sides is possible due to optimised shaping, that doesn’t provide edge shadowing for all design situations. An edge heat load may occur both at the tile and component scales. The edge load needs to be controlled and quantified. The adequate control of edge heat loads is especially critical for water cooled components that uses armour tiles which are bonded to the heat sink, for ensuring the longterm integrity of the tile bonding. An edge heat load allowance criterion of 10% of the top heat load is proposed. The 10% criterion is supported by experimental heat flux tests. © 2017 Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license. (http://creativecommons.org/licenses/by-nc-nd/4.0/)
1. Introduction In tokamaks, a large plasma heat travels along the magnetic field lines in the scrape of layer. The heat flux density parallel to the field lines can reach 1 GW/m² in ITER, a heat flux density far above the engineering capability of any steady state thermal shield. The main chamber wall receives a parallel heat flux density of up to 50 MW/m², which is attenuated down to 2 - 4.7 MW/m² thanks to the grazing angles of incidence of the magnetic field lines onto the wall. The grazing angle is a consequence of the strong toroidal component of the magnetic field. The grazing angles of incidence are controlled by careful component shaping design, to 1 to 5° typically. There are however inevitable leading edges at PFC ends (Fig. 1), because of port openings in the wall, or because of components or tiles gaps, possibly associated to assembly tolerances [1,2]. The leading edges have larger incidence angles (15–90°) than the overall plasma facing surface. They are then loaded with a large local heat load, and possible overloading of the component edges. Such leading edges occur on the global (component to component) or local (tile to tile) scale. The shaping of the wall components mitigates the heat load to edges, by recessing edges, and by having these edges in locations shadowed by neighbour components. Full shadowing of all edges for all plasma cases is however not realistic. It would require large recesses at component edges. Large edge re∗
Corresponding author. E-mail address: [email protected] (R. Mitteau).
cesses increase manufacturing constraints, due to the increase curvature of the component. In a nuclear machine like ITER, recessing the edges reduce the wall shielding capability because of the blanket thickness reduction, compared to the space allocation that is normally available for an unshaped blanket. At the local scale of the tile, complete shadowing of the tile edges would require local tile shaping, hardly realistic for industrial scale components. A better design trade-off involves accepting some limited heat load on component and tile edges [3]. Accepting the edge heat load raises the question about how much heat load is acceptable on the edge. The edge heat load allowable is a key design constraint, which strongly affects component design, such as global or local shaping. The edge heat load allowable needs be firmly established, which is the object of this paper. In ITER, the first wall (FW) is fitted with panels of dimensions about 1.5 m (toroidal) x 1 m (poloidal). Heat loads on component edges are accepted on toroidal-facing and poloidal facing edges, notably at port openings, or as the result of the design 5 mm radial assembly tolerance between panels. Section 2 lists the main drivers to the definition of the allowable edge heat load, and proposes an edge heat load criterion. Section 3 provides experimental evidence about the relevance of this criterion for ITER first wall PFC. 2. Edge heat load allowance A heat load on the tile edge (Fig. 2) is associated with various possible adverse consequences for a PFC. The occurrence of an
http://dx.doi.org/10.1016/j.nme.2017.02.001 2352-1791/© 2017 Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license. (http://creativecommons.org/licenses/by-nc-nd/4.0/)
Please cite this article as: R. Mitteau et al., Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles, Nuclear Materials and Energy (2017), http://dx.doi.org/10.1016/j.nme.2017.02.001
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Fig. 3. Plastic strain in an EHF mock up under top heat loading. The perspective view is oriented from the top to the bond, the tiles are visible in half tones. The colour scale is reversed for readability in B&W, the white colour indicate large strain zones. The white arrow points to the node with the largest strain, located at the free tile to heat sink interface, in the gap separating two tiles.
Fig. 1. Schematic picture of the outer ITER first wall at the global scale, showing components gaps and port openings which are potential leading edges.
Fig. 2. Beryllium armoured first wall mock-up with usual heat flux testing direction (bolt arrows) and possible edge load.
uncontrolled edge heat load can evolve to a serious issue for machine operation. A higher surface temperature at the tile edge and thus the possibility for evaporation or melting of material, together with increased sputtering, may increase plasma contamination. The enhanced loss of material is associated to faster component wear, hence reduced lifetime – or restricted operational space if the issue is detected and tentatively managed. A maximum allowable temperature of 800 °C is adopted for the tile edge, for preventing these risks. Beyond these two constrains, the additional constraint of the tile bonding needs to be addressed. Most recent fusion machines are fitted with water cooled components for enabling long pulse operation. This is also planned in ITER. Water cooled components make use of intimate bonding such as brazing or diffusion bonding between the tile and the cooled heat sink. The issue of edge heat load is even more critical with these bonded components, compared to passively cooled tiles, because ensuring longterm integrity of the tile armour bonding is essential for maintaining the machine in a good operational state. This requires all armour tiles to be present, and well bonded to the cooling structure. The tile-to-heat sink bond is a high-tech and critical technical feature. The main issue is the differential thermal expansion, between the heat armour tiles (Beryllium, Tungsten, or Carbon) and the heat sink (copper alloy, steel). When a heat load is imposed to a bonded armour tile, the heat is evacuated by heat conduction to
the cooled substrate. The heat conduction is obtained - and governed - by the build-up of a thermal gradient through the sandwich structure. In parallel to driving the evacuation of the heat, the thermal gradient generates also internal mechanical stresses, especially at and close to the bond where differential thermal expansion occurs. Most of the bond operates under shear stresses. But the stress pattern is actually far more complicated than mere shear, especially at the extremity (or free edge) when the bonding ends. At this location, the stress changes dramatically both in magnitude and direction over short distances, of the order of the material grain size and bond thickness (10–100 μm). Usual damage or failure criterion like Tresca or Von-Mises equivalent stress fail at being accurate yardstick of the mechanical state. Many studies have been done [4–8], but have so far failed at obtaining a simple and widely recognised bond damage criterion. The consequence is that analysis alone is insufficient for supporting design and guarantying lifetime of bonded components. Hence design justification relies on bond qualification by experiment: defining the allowable of a given bonding technique is obtained by extensive experimental heat flux testing. Water cooled PFCs are normally qualified to a given power load density on the armour top face (Normal heat flux deposition in Fig. 2). The qualification tests are done in dedicated heat flux facilities [9,10]. All these facilities involve a heat loading direction being close to perpendicular to the component surface. Component mounts and masks allow to restrict the irradiation to the test surface, while shadowing the component edges and protecting the test chamber from the shine-through beam. When an edge heat load is superimposed to the normal heat flux deposition at the top of the tile, additional stresses occur in the bond. While the mechanical analyses are unable to predict accurately the damage and eventual failure, they are still useful for giving a qualitative assessment of the added stress in the bond, when an edge heat load is superimposed to the normal heat load. This is illustrated by the results of a simple finite element analysis. The edge heat load test mock-up (similar to Fig. 2, fitted with two different tiles sizes, namely 12 and 16 mm) is modelled in 3D using Ansys workbench. The thermo-hydraulic parameters are those of the heat flux tests in Section 3 (inlet 140 °C, 3 MPa and 2 m/s). The heat transfer coefficient is obtained from EUPITER heat transfer scheme [11], using the Sieder-Tate heat transfer correlation below the Onset of nucleate boiling temperature (Tonb ), and the Thom-CEA correlation above Tonb . The mechanical boundary conditions are minimal, so that the stress and strain are only those caused by the internal thermal stress of the bonded assembly, and no stress or strain is superimposed from external constraints. In Fig. 3, the plastic strain in the CuCrZr heat sink structure is plotted for a typical mock –up, that is subjected to normal heat load of 4.7 MW/m² on the top face of the tiles. A strain of 0.3% at the copper alloy (CuCrZr) heat sink is calculated at the interface of the tile. When the same calculation is redone, adding an edge heat load of
Please cite this article as: R. Mitteau et al., Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles, Nuclear Materials and Energy (2017), http://dx.doi.org/10.1016/j.nme.2017.02.001
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Fig. 4. Test set-up for edge HHF testing, illustrating mock-up rotation and side mask translation allowing to impose variable heat flux density on the edge. a) perspective view, b) side view.
1.7 MW/m², over a side exposure of 6 mm depth, the plastic strain at the same location increases to 0.43%. The increase of bond strain of 0.13% with an added edge heat load, that is about 33% increase compared to the case without edge load, is clearly illustrated by this analysis. The stress increase is mainly localised at the free edge of the tile to heat-sink interface. When the same calculation is reproduced with another couple of edge heat load parameters (2 mm, 5.2 MW/m², triple the heat load density, exposure depth divided by three, hence same heat power to the edge), the CuCrZr strain is again 0.43%. This identical strain to the previous analysis shows that the edge power is a good criterion of the added stress in the tile-to-structure bonding: The added stress at the interface do not depend whether it is a large heat flux density on a small edge penetration or a moderate heat flux on a large penetration, as long as the total power to the edge is kept constant. Hence, it is adequate to define a bond edge heat load criterion based on the tile edge power. A edge heat load allowable criterion is proposed: the tile shall accept up to 10% additional heat load at the edge. The criterion can be expressed at the tile scale or as a line integrated power load. An example is given here at the tile scale: A 12 × 12 mm² tile on a component designed for 4.7 MW/m² accepts a top heat load of 4.7 × 12 × 12 = 680 W. The tile edge load shall then be less than 0.1 × 680 = 68 W. The same criterion can also be expressed as a line power in watt/meter. Then 68 W over 12 mm defines a criterion of 5.7 kW/m. The rationale for this 10% criterion is that the added stress at the bonding interface remains small, compared to the existing interfacial stress that is already present because of the heat load passing through the structure. While this analyse and the testing described in the next section are done for the tile system of the ITER first wall, the criterion proposed above is expected to have some form of generality and could be applied to other flat tile armour bonding. The experimental verification for this criterion is provided in Section 3, describing edge heat loading tests on a dedicated mock-up. 3. Experimental validation of edge heat load allowance The proposed 10% edge heat load criterion is confirmed experimentally on mock ups. The edge test mock-ups are the one in Fig. 2, namely the adopted design of the enhanced heat flux plasma facing components [12]. A dedicated mock up jig is designed, for allowing a controlled heat loading both on the top and edge face. The set-up employs a rotating actuator and moving mask (Fig. 4). The mock-up rotation allows to adjust the ratio of heat flux density to the edge, relative to the flux density of the top face. The translation of the mask allows to adjust the heat flux wetting pattern on the edge, hence allowing to reproduce various heat load exposure depths on the mock-up side. The test is done with water conditions matching those of the actual ITER first wall during plasma operation: inlet water of 140 °C (representing ITER
3
Fig. 5. Infrared image of the mock-up, at the 770th cycle for heat flux step 3. White colour is 800 °C. The camera field of view is indicated in Fig. 4, the white horizontal line is the leading edge. The tile top face appears above the edge in the picture, and the side face below the edge.
FW outlet temperature), a water pressure of 3 MPa and a velocity of 2 m/s inside the water channel. The testing is delicate, because the tile edge temperature approaches test facility limit. Beryllium evaporation is occasionally large, bringing increased maintenance constraints (facility maintenance more frequent and more complicated than usual). The control of the edge heat load is also delicate, because the edge heat load is a small fraction of the total heat power to the mock-up. The measure of the edge heat load is done by subtracting the mockup power with, from without edge heat load. The power signal is close to the noise over ratio of the calorimetric measurement. This publication cannot address in details all these challenges, and how they have been resolved. This is reported in a another contribution [13]. The objective of this paper is to focus on the testing protocol and its rationale, and on the main achievements of the heat flux test program. The testing protocol is designed so that: •
•
•
•
Three edge heat load are tested: 6%, 10% and 14% of the top heat load. The load levels allow testing the proposed criterion and slightly above. Two different combinations of edge power density and exposure depth. This provides an experimental confirmation that the proposed criterion is independent from the combination of exposure depth and side heat load density. Two tile sizes are tested (12 and 16 mm). These two sizes match the various tile size options in the ITER FW manufacturing parties. The objective is to confirm that the proposed criterion is independent from the tile size, in the size range of 12 to 16 mm. Three tiles are tested for each case, providing some redundancy in the test results.
The testing matrix is provided in Table 1. The tile edge power for 12 and 16 mm tiles is given in columns 8 and 10, followed by the power ratio to the tile power on the top face. For each test case, one thousand heat cycles are done, for checking that the edge temperature is stable. Stable edge temperature is used as an indication that there is no degradation of the bonding, as progressive debonding would be associated to reduced heat transfer, hence larger edge temperature. Indeed, in case tile bonding damage occurs, it is usually detected by a tile temperature increase during the first hundreds of cycles, corresponding to the largest probability of damage detection for this kind of deterministic damage [14]. Fig. 5 shows the infrared image of the edge, at the 770th cycle for the heat flux step 3. The temperature pattern is unchanged compared to the initial thermal cycle, indicating no degradation of the bond. A similar result of no edge temperature change is also obtained for the other cases. This confirms that a 10% additional heat load on the component edge is acceptable. Having also no damage at the 14% edge loading test level provides a 1.4 margin with respect the edge load, a design margin which is small but usual
Please cite this article as: R. Mitteau et al., Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles, Nuclear Materials and Energy (2017), http://dx.doi.org/10.1016/j.nme.2017.02.001
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Table 1 Testing matrix. Tile power for a 12 × 12 tile at 4.7 MW/m² is 677 W; Tile power for a 16 × 16 tile at 4.7 MW/m² is 1203 W. Step
Mock-up
Zone
Penetration (mm)
Tile size [mm x mm]
Edge heat flux level, [MW/m²]
Number of cycles
Edge power 12 × 12
%
Edge power 16 × 16
%
1 2 3 4 5 6
1
1 2 3 4 5 6
6
12 × 12 16 × 16
0.7 1.2 1.7 2.2 3.7 5.2
10 0 0 10 0 0 10 0 0 10 0 0 10 0 0 10 0 0
50.4 86.4 122.4 52.8 88.8 124.8
7.4 12.8 18.1 7.8 13.1 18.4
67.2 115.2 163.2 70.4 118.4 166.4
5.6 9.6 13.6 5.9 9.8 13.8
2 3
2
in PFC design. Given the geometry is very similar for the toroidal facing edge, it is estimated the results is also valid for the toroidal edge. The 10% edge heat load criterion is hence confirmed experimentally. 4. Conclusion There is some heat load on the edges of plasma facing components or tiles, but this needs to be controlled in order to come up with a value which is acceptable. The resulting allowable edge heat load is a key parameter for both local and global surface shaping definition. For ITER enhanced heat flux First wall armoured with beryllium tile, there is the constraint of a local maximum edge temperature of 800 °C based on safety considerations; in addition, another criterion needs to be defined to avoid overstress at the beryllium to copper joint. From the analysis results, it is proposed that the heat power allowed to the side of the tile shall remain under 10% of the power allowed on the main tile face. This acceptable power is confirmed by the results from edge heat load testing of mock-ups, validating the ITER design choices. Disclaimer
References [1] R. Mohanti, E. Deksnis, P. Lomas, M. Pick, Edge loading of plasma facing components in fusion devices, Joint Europ. torus (1993). http://www. euro- fusionscipub.org/wp- content/uploads/2014/11/JETR93001.pdf. [2] Y. Corre, R. Dejarnac, J.L. Gardarein et Al.; Jour. Nucl. Mater. 463(2015)832–836 (10.1016/j.jnucmat.2014.10.036). [3] P. Stangeby, Analysis for shaping the ITER first wall, Jour. Nucl. Mater. 390–391 (15 June 2009) 963–966. [4] D.B. Bogy, J. Appl. Mech., Vol. 35, Trans. ASME, Vol.90, Series E, pp. 460–466 [5] W.T. Chen, C.W. Nelson, IBM J. Res. Dev. 23 (2) (1979). [6] D. Munz, T. Fett, Y. Yang, Eng. Fract. Mech. 44 (2) (1993) 185–196. [7] G. Lovato, Ecole Nationale Supérieure des Mines de Paris, France PhD Thesis, 1995. [8] L. Moncel, Etude des Mécanismes d’endommagement d’un assemblage Cuivre composite Carbone-Carbone, Etude des Mécanismes d’endommagement d’un assemblage Cuivre composite Carbone-Carbone, 1, Thèse de l’Université Bordeaux, 1999. [9] V.K. Gagen-Torn, I.R. Kirillov, V.L. Komarov, et al., Experimental complex for high flux material interaction research, Fus. Technol. (1994). [10] R. Duwe, W. Kühnlein, H. Münstermann, in: Proc. 18th Symposium on Fusion Technology (SOFT), Karlsruhe Germany, 1994, pp. 355–358. [11] M. Merola, G. Vieider, J. Schlosser, EUPITER: the European thermohydraulic package for plasma facing components, in: Proceedings of the Specialists’ Workshop on High Heat Flux Cooling for ITER PFCs, JET, September, 1996. [12] R. Mitteau, Fus. Eng. Des. 88 (6-8) (Oct. 2013) 568–570. [13] A. Gervash, submitted to SOFT2016. [14] G. Pérez, R. Mitteau, R. Eaton, Fus. Eng. Des. 101 (2015) 28–41.
The views and opinions expressed herein do not necessarily reflect those of the ITER Organization
Please cite this article as: R. Mitteau et al., Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles, Nuclear Materials and Energy (2017), http://dx.doi.org/10.1016/j.nme.2017.02.001
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[m5G;February 9, 2017;23:58]
Nuclear Materials and Energy 0 0 0 (2017) 1–4
Contents lists available at ScienceDirect
Nuclear Materials and Energy journal homepage: www.elsevier.com/locate/nme
Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles R. Mitteau b,∗, R. Eaton b, A. Gervash a, V. Kuznetcov a, V. Davydov a, R. Rulev a a b
Efremov Research Institute, 189631, St. Petersburg, Russia ITER Organization Route de Vinon-sur-Verdon, CS 90 046, 13067 St Paul Lez Durance Cedex France
a r t i c l e
i n f o
Article history: Received 11 July 2016 Revised 6 December 2016 Accepted 1 February 2017 Available online xxx
a b s t r a c t Plasma facing components are usually qualified to a given heat load density applied at the top face of the armour tiles with normal incidence angle. When employed in tokamak fusion machines, heat loading on the tile sides is possible due to optimised shaping, that doesn’t provide edge shadowing for all design situations. An edge heat load may occur both at the tile and component scales. The edge load needs to be controlled and quantified. The adequate control of edge heat loads is especially critical for water cooled components that uses armour tiles which are bonded to the heat sink, for ensuring the longterm integrity of the tile bonding. An edge heat load allowance criterion of 10% of the top heat load is proposed. The 10% criterion is supported by experimental heat flux tests. © 2017 Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license. (http://creativecommons.org/licenses/by-nc-nd/4.0/)
1. Introduction In tokamaks, a large plasma heat travels along the magnetic field lines in the scrape of layer. The heat flux density parallel to the field lines can reach 1 GW/m² in ITER, a heat flux density far above the engineering capability of any steady state thermal shield. The main chamber wall receives a parallel heat flux density of up to 50 MW/m², which is attenuated down to 2 - 4.7 MW/m² thanks to the grazing angles of incidence of the magnetic field lines onto the wall. The grazing angle is a consequence of the strong toroidal component of the magnetic field. The grazing angles of incidence are controlled by careful component shaping design, to 1 to 5° typically. There are however inevitable leading edges at PFC ends (Fig. 1), because of port openings in the wall, or because of components or tiles gaps, possibly associated to assembly tolerances [1,2]. The leading edges have larger incidence angles (15–90°) than the overall plasma facing surface. They are then loaded with a large local heat load, and possible overloading of the component edges. Such leading edges occur on the global (component to component) or local (tile to tile) scale. The shaping of the wall components mitigates the heat load to edges, by recessing edges, and by having these edges in locations shadowed by neighbour components. Full shadowing of all edges for all plasma cases is however not realistic. It would require large recesses at component edges. Large edge re∗
Corresponding author. E-mail address: [email protected] (R. Mitteau).
cesses increase manufacturing constraints, due to the increase curvature of the component. In a nuclear machine like ITER, recessing the edges reduce the wall shielding capability because of the blanket thickness reduction, compared to the space allocation that is normally available for an unshaped blanket. At the local scale of the tile, complete shadowing of the tile edges would require local tile shaping, hardly realistic for industrial scale components. A better design trade-off involves accepting some limited heat load on component and tile edges [3]. Accepting the edge heat load raises the question about how much heat load is acceptable on the edge. The edge heat load allowable is a key design constraint, which strongly affects component design, such as global or local shaping. The edge heat load allowable needs be firmly established, which is the object of this paper. In ITER, the first wall (FW) is fitted with panels of dimensions about 1.5 m (toroidal) x 1 m (poloidal). Heat loads on component edges are accepted on toroidal-facing and poloidal facing edges, notably at port openings, or as the result of the design 5 mm radial assembly tolerance between panels. Section 2 lists the main drivers to the definition of the allowable edge heat load, and proposes an edge heat load criterion. Section 3 provides experimental evidence about the relevance of this criterion for ITER first wall PFC. 2. Edge heat load allowance A heat load on the tile edge (Fig. 2) is associated with various possible adverse consequences for a PFC. The occurrence of an
http://dx.doi.org/10.1016/j.nme.2017.02.001 2352-1791/© 2017 Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license. (http://creativecommons.org/licenses/by-nc-nd/4.0/)
Please cite this article as: R. Mitteau et al., Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles, Nuclear Materials and Energy (2017), http://dx.doi.org/10.1016/j.nme.2017.02.001
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Fig. 3. Plastic strain in an EHF mock up under top heat loading. The perspective view is oriented from the top to the bond, the tiles are visible in half tones. The colour scale is reversed for readability in B&W, the white colour indicate large strain zones. The white arrow points to the node with the largest strain, located at the free tile to heat sink interface, in the gap separating two tiles.
Fig. 1. Schematic picture of the outer ITER first wall at the global scale, showing components gaps and port openings which are potential leading edges.
Fig. 2. Beryllium armoured first wall mock-up with usual heat flux testing direction (bolt arrows) and possible edge load.
uncontrolled edge heat load can evolve to a serious issue for machine operation. A higher surface temperature at the tile edge and thus the possibility for evaporation or melting of material, together with increased sputtering, may increase plasma contamination. The enhanced loss of material is associated to faster component wear, hence reduced lifetime – or restricted operational space if the issue is detected and tentatively managed. A maximum allowable temperature of 800 °C is adopted for the tile edge, for preventing these risks. Beyond these two constrains, the additional constraint of the tile bonding needs to be addressed. Most recent fusion machines are fitted with water cooled components for enabling long pulse operation. This is also planned in ITER. Water cooled components make use of intimate bonding such as brazing or diffusion bonding between the tile and the cooled heat sink. The issue of edge heat load is even more critical with these bonded components, compared to passively cooled tiles, because ensuring longterm integrity of the tile armour bonding is essential for maintaining the machine in a good operational state. This requires all armour tiles to be present, and well bonded to the cooling structure. The tile-to-heat sink bond is a high-tech and critical technical feature. The main issue is the differential thermal expansion, between the heat armour tiles (Beryllium, Tungsten, or Carbon) and the heat sink (copper alloy, steel). When a heat load is imposed to a bonded armour tile, the heat is evacuated by heat conduction to
the cooled substrate. The heat conduction is obtained - and governed - by the build-up of a thermal gradient through the sandwich structure. In parallel to driving the evacuation of the heat, the thermal gradient generates also internal mechanical stresses, especially at and close to the bond where differential thermal expansion occurs. Most of the bond operates under shear stresses. But the stress pattern is actually far more complicated than mere shear, especially at the extremity (or free edge) when the bonding ends. At this location, the stress changes dramatically both in magnitude and direction over short distances, of the order of the material grain size and bond thickness (10–100 μm). Usual damage or failure criterion like Tresca or Von-Mises equivalent stress fail at being accurate yardstick of the mechanical state. Many studies have been done [4–8], but have so far failed at obtaining a simple and widely recognised bond damage criterion. The consequence is that analysis alone is insufficient for supporting design and guarantying lifetime of bonded components. Hence design justification relies on bond qualification by experiment: defining the allowable of a given bonding technique is obtained by extensive experimental heat flux testing. Water cooled PFCs are normally qualified to a given power load density on the armour top face (Normal heat flux deposition in Fig. 2). The qualification tests are done in dedicated heat flux facilities [9,10]. All these facilities involve a heat loading direction being close to perpendicular to the component surface. Component mounts and masks allow to restrict the irradiation to the test surface, while shadowing the component edges and protecting the test chamber from the shine-through beam. When an edge heat load is superimposed to the normal heat flux deposition at the top of the tile, additional stresses occur in the bond. While the mechanical analyses are unable to predict accurately the damage and eventual failure, they are still useful for giving a qualitative assessment of the added stress in the bond, when an edge heat load is superimposed to the normal heat load. This is illustrated by the results of a simple finite element analysis. The edge heat load test mock-up (similar to Fig. 2, fitted with two different tiles sizes, namely 12 and 16 mm) is modelled in 3D using Ansys workbench. The thermo-hydraulic parameters are those of the heat flux tests in Section 3 (inlet 140 °C, 3 MPa and 2 m/s). The heat transfer coefficient is obtained from EUPITER heat transfer scheme [11], using the Sieder-Tate heat transfer correlation below the Onset of nucleate boiling temperature (Tonb ), and the Thom-CEA correlation above Tonb . The mechanical boundary conditions are minimal, so that the stress and strain are only those caused by the internal thermal stress of the bonded assembly, and no stress or strain is superimposed from external constraints. In Fig. 3, the plastic strain in the CuCrZr heat sink structure is plotted for a typical mock –up, that is subjected to normal heat load of 4.7 MW/m² on the top face of the tiles. A strain of 0.3% at the copper alloy (CuCrZr) heat sink is calculated at the interface of the tile. When the same calculation is redone, adding an edge heat load of
Please cite this article as: R. Mitteau et al., Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles, Nuclear Materials and Energy (2017), http://dx.doi.org/10.1016/j.nme.2017.02.001
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R. Mitteau et al. / Nuclear Materials and Energy 000 (2017) 1–4
Fig. 4. Test set-up for edge HHF testing, illustrating mock-up rotation and side mask translation allowing to impose variable heat flux density on the edge. a) perspective view, b) side view.
1.7 MW/m², over a side exposure of 6 mm depth, the plastic strain at the same location increases to 0.43%. The increase of bond strain of 0.13% with an added edge heat load, that is about 33% increase compared to the case without edge load, is clearly illustrated by this analysis. The stress increase is mainly localised at the free edge of the tile to heat-sink interface. When the same calculation is reproduced with another couple of edge heat load parameters (2 mm, 5.2 MW/m², triple the heat load density, exposure depth divided by three, hence same heat power to the edge), the CuCrZr strain is again 0.43%. This identical strain to the previous analysis shows that the edge power is a good criterion of the added stress in the tile-to-structure bonding: The added stress at the interface do not depend whether it is a large heat flux density on a small edge penetration or a moderate heat flux on a large penetration, as long as the total power to the edge is kept constant. Hence, it is adequate to define a bond edge heat load criterion based on the tile edge power. A edge heat load allowable criterion is proposed: the tile shall accept up to 10% additional heat load at the edge. The criterion can be expressed at the tile scale or as a line integrated power load. An example is given here at the tile scale: A 12 × 12 mm² tile on a component designed for 4.7 MW/m² accepts a top heat load of 4.7 × 12 × 12 = 680 W. The tile edge load shall then be less than 0.1 × 680 = 68 W. The same criterion can also be expressed as a line power in watt/meter. Then 68 W over 12 mm defines a criterion of 5.7 kW/m. The rationale for this 10% criterion is that the added stress at the bonding interface remains small, compared to the existing interfacial stress that is already present because of the heat load passing through the structure. While this analyse and the testing described in the next section are done for the tile system of the ITER first wall, the criterion proposed above is expected to have some form of generality and could be applied to other flat tile armour bonding. The experimental verification for this criterion is provided in Section 3, describing edge heat loading tests on a dedicated mock-up. 3. Experimental validation of edge heat load allowance The proposed 10% edge heat load criterion is confirmed experimentally on mock ups. The edge test mock-ups are the one in Fig. 2, namely the adopted design of the enhanced heat flux plasma facing components [12]. A dedicated mock up jig is designed, for allowing a controlled heat loading both on the top and edge face. The set-up employs a rotating actuator and moving mask (Fig. 4). The mock-up rotation allows to adjust the ratio of heat flux density to the edge, relative to the flux density of the top face. The translation of the mask allows to adjust the heat flux wetting pattern on the edge, hence allowing to reproduce various heat load exposure depths on the mock-up side. The test is done with water conditions matching those of the actual ITER first wall during plasma operation: inlet water of 140 °C (representing ITER
3
Fig. 5. Infrared image of the mock-up, at the 770th cycle for heat flux step 3. White colour is 800 °C. The camera field of view is indicated in Fig. 4, the white horizontal line is the leading edge. The tile top face appears above the edge in the picture, and the side face below the edge.
FW outlet temperature), a water pressure of 3 MPa and a velocity of 2 m/s inside the water channel. The testing is delicate, because the tile edge temperature approaches test facility limit. Beryllium evaporation is occasionally large, bringing increased maintenance constraints (facility maintenance more frequent and more complicated than usual). The control of the edge heat load is also delicate, because the edge heat load is a small fraction of the total heat power to the mock-up. The measure of the edge heat load is done by subtracting the mockup power with, from without edge heat load. The power signal is close to the noise over ratio of the calorimetric measurement. This publication cannot address in details all these challenges, and how they have been resolved. This is reported in a another contribution [13]. The objective of this paper is to focus on the testing protocol and its rationale, and on the main achievements of the heat flux test program. The testing protocol is designed so that: •
•
•
•
Three edge heat load are tested: 6%, 10% and 14% of the top heat load. The load levels allow testing the proposed criterion and slightly above. Two different combinations of edge power density and exposure depth. This provides an experimental confirmation that the proposed criterion is independent from the combination of exposure depth and side heat load density. Two tile sizes are tested (12 and 16 mm). These two sizes match the various tile size options in the ITER FW manufacturing parties. The objective is to confirm that the proposed criterion is independent from the tile size, in the size range of 12 to 16 mm. Three tiles are tested for each case, providing some redundancy in the test results.
The testing matrix is provided in Table 1. The tile edge power for 12 and 16 mm tiles is given in columns 8 and 10, followed by the power ratio to the tile power on the top face. For each test case, one thousand heat cycles are done, for checking that the edge temperature is stable. Stable edge temperature is used as an indication that there is no degradation of the bonding, as progressive debonding would be associated to reduced heat transfer, hence larger edge temperature. Indeed, in case tile bonding damage occurs, it is usually detected by a tile temperature increase during the first hundreds of cycles, corresponding to the largest probability of damage detection for this kind of deterministic damage [14]. Fig. 5 shows the infrared image of the edge, at the 770th cycle for the heat flux step 3. The temperature pattern is unchanged compared to the initial thermal cycle, indicating no degradation of the bond. A similar result of no edge temperature change is also obtained for the other cases. This confirms that a 10% additional heat load on the component edge is acceptable. Having also no damage at the 14% edge loading test level provides a 1.4 margin with respect the edge load, a design margin which is small but usual
Please cite this article as: R. Mitteau et al., Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles, Nuclear Materials and Energy (2017), http://dx.doi.org/10.1016/j.nme.2017.02.001
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R. Mitteau et al. / Nuclear Materials and Energy 000 (2017) 1–4
Table 1 Testing matrix. Tile power for a 12 × 12 tile at 4.7 MW/m² is 677 W; Tile power for a 16 × 16 tile at 4.7 MW/m² is 1203 W. Step
Mock-up
Zone
Penetration (mm)
Tile size [mm x mm]
Edge heat flux level, [MW/m²]
Number of cycles
Edge power 12 × 12
%
Edge power 16 × 16
%
1 2 3 4 5 6
1
1 2 3 4 5 6
6
12 × 12 16 × 16
0.7 1.2 1.7 2.2 3.7 5.2
10 0 0 10 0 0 10 0 0 10 0 0 10 0 0 10 0 0
50.4 86.4 122.4 52.8 88.8 124.8
7.4 12.8 18.1 7.8 13.1 18.4
67.2 115.2 163.2 70.4 118.4 166.4
5.6 9.6 13.6 5.9 9.8 13.8
2 3
2
in PFC design. Given the geometry is very similar for the toroidal facing edge, it is estimated the results is also valid for the toroidal edge. The 10% edge heat load criterion is hence confirmed experimentally. 4. Conclusion There is some heat load on the edges of plasma facing components or tiles, but this needs to be controlled in order to come up with a value which is acceptable. The resulting allowable edge heat load is a key parameter for both local and global surface shaping definition. For ITER enhanced heat flux First wall armoured with beryllium tile, there is the constraint of a local maximum edge temperature of 800 °C based on safety considerations; in addition, another criterion needs to be defined to avoid overstress at the beryllium to copper joint. From the analysis results, it is proposed that the heat power allowed to the side of the tile shall remain under 10% of the power allowed on the main tile face. This acceptable power is confirmed by the results from edge heat load testing of mock-ups, validating the ITER design choices. Disclaimer
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The views and opinions expressed herein do not necessarily reflect those of the ITER Organization
Please cite this article as: R. Mitteau et al., Allowable heat load on the edge of the ITER first wall panel beryllium flat tiles, Nuclear Materials and Energy (2017), http://dx.doi.org/10.1016/j.nme.2017.02.001