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Features and optimization approaches of the entrance section cooling gas flow of the IFMIF High Flux Test Module
Fusion Engineering and Design 83 (2008) 1536–1542
Contents lists available at ScienceDirect
Fusion Engineering and Design journal homepage: www.elsevier.com/locate/fusengdes
Features and optimization approaches of the entrance section cooling gas flow of the IFMIF High Flux Test Module F. Arbeiter a,∗ , S. Gordeev a , V. Heinzel a , T. Ihli a , D. Leichtle a , A. Möslang b , V. Slobotchouk a a b
Institut für Reaktorsicherheit, Forschungszentrum Karlsruhe, Hermann-von-Helmholtz-Platz 1, 76344 Eggenstein-Leopoldshafen, Germany Institut für Materialforschung I, Forschungszentrum Karlsruhe, Hermann-von-Helmholtz-Platz 1, 76344 Eggenstein-Leopoldshafen, Germany
a r t i c l e
i n f o
Article history: Available online 3 September 2008 Keywords: IFMIF High flux test module Helium Local heat transfer Entrance effects Minichannel Gas cooling
a b s t r a c t The International Fusion Materials Irradiation Facility (IFMIF) is devised to contribute experimental evidence to an irradiated material properties database for candidate materials exposed to irradiation spectra and doses relevant for future fusion power reactors. Due to neutron fluxes generated by high-energy deuterons reacting in a liquid lithium target, damage rates of 20–50 displacements per atom in one full power year can be achieved in steel specimens inside a volume of approximately 0.5 L. The design of the high flux test module developed at the Forschungszentrum Karslruhe (FZK) allows for maximizing the space available in the high flux neutron field for material irradiation, while at the same time allowing precise adherence of the irradiation temperature of the specimen stacks. Since enhancement of the neutron irradiation requires placement of the specimens as close as possible to the neutron source, the design proposes thin container structures (obeying mechanical constraints) and flat coolant channels between the rigs. A helium gas flow is designated to remove the heat from the rigs to keep the required irradiation temperature, which may be chosen between 250 and 650 ◦ C. As a result of the thin container walls and the small channel dimensions, the helium cooling gas flow is characterized by low pressure, transitional Reynolds numbers and intermediate Mach numbers. Dedicated experimental investigations on such minichannel cooling gas flows have been conducted with the ITHEX helium loop facility. Results obtained by laser Doppler anemometry indicate a complex three-dimensional evolution of the transitional laminar-turbulent flow field in the hydraulic entrance section. In the short cooling channels, a relevant portion of the flow alongside the rigs is influenced by this developing region. Detailed knowledge of the flow development and the resulting heat transfer coefficients is necessary to optimize the flow channel inlet design and to avoid inhomogeneities of the temperature field inside the specimen stacks, which otherwise could be caused by varying local heat transfer coefficients and mass flux redistributions (in the axial and the lateral coordinates). Experimental results are presented. Specific features of the minichannel entrance flow are identified, and conclusions are drawn for an optimized design of the entrance geometry, to enhance the utility of the HFTM for fusion materials research. © 2008 Elsevier B.V. All rights reserved.
1. Introduction The International Fusion Materials Irradiation Facility (IFMIF) is devised to contribute experimental evidence to an irradiated material properties database for candidate materials exposed to irradiation spectra and doses relevant for future fusion power reactors. Due to neutron fluxes generated by high-energy deuterons reacting in a liquid lithium target, damage rates of 20–50 displacements per atom in one full power year can be achieved in steel
∗ Corresponding author. Tel.: +49 7247 823452. E-mail address: [email protected] (F. Arbeiter). 0920-3796/$ – see front matter © 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.fusengdes.2008.07.014
specimens inside a volume of approximately 0.5 L. The design of the high flux test module (the abbreviation HFTM will be used) developed at the Forschungszentrum Karslruhe (FZK) allows for maximizing the space available in the high flux neutron field for material irradiation, while at the same time allowing precise adherence of the irradiation temperature of the specimen stacks. Since enhancement of the neutron irradiation requires placement of the specimens as close as possible to the neutron source, the design proposes thin container structures (obeying mechanical constraints) and flat coolant channels between the rigs. A helium gas flow is designated to remove the heat from the rigs to keep the required irradiation temperature, which may be chosen between 250 and 650 ◦ C. As a result of the thin container walls and the small channel dimensions, the helium cooling gas flow is characterized
F. Arbeiter et al. / Fusion Engineering and Design 83 (2008) 1536–1542
Nomenclature channel cross-sectional area AQ b rectangular channel width cp specific heat capacity for constant pressure dh = 4AQ /SW hydraulic diameter fs (x1 ,x2 ) apparent Fanning friction factor ˙ m mass flux through the channel N number of samples from LDA p static pressure q˙ w specific wall heat flux r radius s channel gap width SW channel wetted perimeter T Kelvin temperature u instantaneous velocity urms root mean square velocity uShift,LDA virtual velocity of the LDA fringe pattern U time-averaged axial velocity ˙ Um = m/( 1 AQ ) bulk fluid velocity x, y, z Cartesian coordinates axial distance from the beginning of the heater xth Greek symbols thermal conductivity kinematic viscosity gas density w wall shear stress Indices 1 2 i fl,m w
conditions in the entrance plenum conditions in the exit plenum integer bulk fluid property property at wall
by low pressure, transitional Reynolds numbers and intermediate Mach numbers. Since the gas cooling, together with the nuclear heat release and electrical compensation heaters, determines the temperature homogeneity within the specimen stacks, as well as the level of thermomechanical stresses in the thin container structures, the thermal–hydraulic layout and the reliability of the cooling system during operation is a crucial point within the ongoing optimization process. In this paper new experimental results on the thermal–hydraulics within the HFTM cooling ducts are discussed. Improved insight into the relevant local flow phenomena was obtained. The experimental validation of the numerical tools employed in the layout process is an ongoing task; first results were already published by Arbeiter [3–5]. Optimization approaches for the inlet sections are based on the experimental results presented in this paper. 2. Thermal–hydraulic situation of the high flux test module As indicated in Fig. 1, the material specimens are stacked in 12 irradiation rigs, which are arranged in a 4 × 3 pattern directly behind the neutron source. Using spacers on the rig surfaces, channels with gap widths of 0.6 mm are created between the rigs and the compartment walls, and gap widths of 1.0 mm remain between adjacent rigs as well as between the rigs and the outer container walls. The cross-sections of these channels are rectangular, L-shaped or T-shaped, and are subject to the ongoing optimization of the HFTM. Helium gas at a moderate pressure of about 3 bars is
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routed upwards through these channels alongside the rig walls as a coolant. The gas enters the channels from a distribution plenum via converging nozzles (Fig. 1). The nozzle structure also serves as a neutron reflector, which results in higher neutron fluxes in the specimen. In this way, an array of parallel minichannels is created and supplied with cooling gas. A similar layout is used for the lateral neutron reflectors. The temperature in the specimen and the structures are determined by the local nuclear heat release, by conduction inside the structures, by the convective heat transfer to the helium gas flow, and by auxiliary electrical heaters with controllable power. The electrical heaters can replace the nuclear heat during beam off periods, and are used to compensate the axial temperature gradient in the specimen stacks. To ensure correct dimensioning of the heaters, the local heat transfer to the gas flow must be known. It is the aim of the thermal–hydraulic optimization to adjust a homogenous temperature field within a single specimen stack, and avoid temperature gradients in the structure which would result in thermomechanical stresses and deformations. Thermomechanical stresses represent the main loads in some parts of the module. The determination of channel deformations, the distribution of the coolant flow and the temperature field are coupled problems; therefore iterative application of CFD and FEM analyses are necessary. The flow is subject to several phenomena which can have considerable impact on the local heat transfer from the walls and therefore are considered in detail. • Entrance effects. Growing hydraulic boundary layer after the entrance nozzle, and growing thermal boundary layer after the beginning of the heated section. Hydraulic and thermal development regions partially overlap. • Laminar to turbulent transition. The flow is laminarized by the acceleration in the entrance nozzle, the channel Reynolds numbers range between 6000 and 10,000. The adiabatic flow is therefore in the laminar-turbulent transitional range. • Turbulent to laminar retransition. The considerable heating rates of the flow effect gradients of the gas properties in axial and wall normal directions. The gas viscosity, which rises with increasing temperature, diminishes the local Reynolds number near the wall and increases the turbulence dissipation. Additionally, the high pressure drop and the rise in mean temperature of the gas decrease its density, and thereby accelerate the flow. This will result in a negative term in the turbulence kinetic energy budget (see Ref. [1]) and could result in relaminarization. • Lateral redistribution of massflow. The nozzle shape determines the massflow distribution at the entrance of the channel. Inside the channel, secondary flows in the corners, as well as growing turbulent spots at random positions within the laminar-turbulent transitional region, exhibit flat velocity profiles and thereby displace massflow to the adjacent laminar regions. This paper aims to present a comprehensive picture of the thermal–hydraulic situation studied by experiments. Where single effects have already been discussed in more details in previous publications, references are given. 3. ITHEX experimental programme To examine the thermal–hydraulic effects listed above, an extensive experimental programme has been conducted. Two different types of testsections were used: the first testsection was an annular minichannel made of stainless steel, instrumented with electrical heaters and thermocouples between the heaters and the cooled surfaces in the inner and the outer cylinder. This annular testsection
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Fig. 1. View on the IFMIF high flux test module, showing the arrangement of specimen irradiation rigs, inlet nozzles and minichannels.
has been used to study the laminar to turbulent transition, and the laminarizing effects of the intense heating. A 3D CAD drawing and a cross-section of this testsection are shown in Fig. 2. The annular channel has a hydraulic diameter dh = 1.23 mm, a complete length of 94 mm and a heated length of 50 mm. As indicated in Fig. 2, the axial coordinate xth originates at the beginning of the heated sec-
Fig. 2. ITHEX annular minichannel testsection with major dimensions and coordinate system.
tion. The second testsection is a rectangular duct with optical access for 1D Laser Doppler Anemometry (LDA) to measure the evolution of the velocity and turbulence fields in the minichannel, behind different shapes of entrance nozzles. A 3D CAD drawing of two of these nozzle configurations is shown in Fig. 3. The orientation of the coordinate system used is also indicated in Fig. 3. The origin y = 0, z = 0 is defined on the channel symmetry axis; x = 0 is defined at the location where the nozzle ends and the minichannel with the constant cross-section 1 mm × 45 mm begins. The LDA probe in backscattering configuration can access the testsection through an optical window (not shown in the drawing) from the side (the
Fig. 3. ITHEX rectangular duct testsections for LDA measurements.
F. Arbeiter et al. / Fusion Engineering and Design 83 (2008) 1536–1542
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laser beams are intersecting in the xz-plane), and thus measure the velocity component in the axial x direction Results will be presented for both testsections, to address the issues of local flow phenomena listed above. To describe the results, the variables defined in Eqs. (1)–(6), which are commonly used in fluid dynamics, are chosen. Re =
Um dh
(1)
= q+ 1
q˙ w AQ ˙ p (T1 )T1 mc
(2)
Nu(x) =
dh q˙ w (x) f,ml Tfl,m (x) − Tw (x)
2 f = w 1/2Um
N
U=
i=1 N
−1
ui w(ui )
i=1
w(ui )
urms =
(3) (4)
N with
w(ui ) =
N w(ui )(ui − U)2 i=1 N w(ui ) i=1
u /N i=1 i
ui − uShift,LDA
(5) Fig. 4. Results of LDA measurements on the channel axis (y = 0, z = 0) in the testsection “17-V-1” at Re = 6000 showing the transitional nature of the flow.
1/2
(6)
The variable q+ defined by Eq. (2) is referred to as heating rate, since it is proportional to (T2 − T1 )/T1 for a perfect gas. To determine the time-averaged velocity U, and the turbulence velocity urms , a weighting factor w(ui ) is applied to the individual LDA samples to reduce systematic errors evoked by the individual realization sampling procedure inherent to LDA, similar to the proposal of McLaughlin and Tiederman [2]. LDA experiments were carried out with ambient air; therefore the density, and subsequently the velocity magnitude of individual experiments adjusted to the same Reynolds number is slightly different for each individual experiment. To cancel this effect, the velocity profiles are presented scaled ˙ to the bulk fluid velocity Um = m/( 1 AQ ) of the individual experiment. The quantity U/Um is referred to as normalized velocity. 3.1. Axial boundary layer development behind the entrance nozzle and laminar to turbulent transition The shape of the entrance nozzle determines the amount of laminarization in the flow due to acceleration, but also the regeneration of turbulence by high shear forces in the region where the nozzle enters the minichannel. The geometry of this region will control the magnitude, but also the spatial distribution of turbulence at the entrance of the minichannel, and thus critically determines the length and the character of the hydraulic development of the boundary layer in the minichannel. Effects such as flow detachment, can also occur at the end of the nozzle. Different generic shapes based on the “17-V-1” nozzle shape shown in Fig. 2 were studied using the LDA setup. For a smooth nozzle of the 17-V-1 geometry and transitional Reynolds numbers 3000 ≤ Re ≤ 12,000, it was found from LDA as well as from friction measurements, that the flow can realize both laminar and turbulent state, induced by very small disturbances. Such results for the normalized velocity U(x)/Um at Re1 = 6000 are shown in Fig. 4, where the results of multiple experiments (identified by a label DDMMYYx), all for similar inlet conditions, are mutually compared. High levels (typical value U(x = 50 mm)/Um = 1.5) of U(x) indicate laminar flow, while low levels (typical value U(x = 50 mm)/Um = 1.3) of U(x) indicate turbulent flow. This effect has been discussed by Arbeiter [3] and Arbeiter et al. [4] in more detail. It is based on the appearance of flow detachment in the entrance, as well as on the laterally distributed onset of turbulence discussed below. As such behaviour must be avoided in all circumstances in the IFMIF HFTM application, different design
modifications dedicated to yield a reproducible level of turbulence of high magnitude to ensure a short hydraulic entrance length were studied in the next step. Of course, the new design features for reliable turbulence production must be compatible with the manufacturing technology available to fabricate the complex nozzle array. So far, the shapes which have been tested were (a) the basic 17-V-1 nozzle (smooth), (b) a T-shaped step instead of a nozzle, (c) a protruding small step mounted on the surface of the 17-V-1 nozzle, and (d) a grid (formed by a spiral of thick wire) inserted into the 17-V-1 nozzle. The axial profiles for the turbulence velocity on the channel axis (y = 0, z = 0) for these four nozzle shapes are plotted in Fig. 5. As already shown in Fig. 4, the smooth 17-V-1 (a) nozzle can produce various modes of flow. In Fig. 5, the laminar/turbulent case has been included for the nozzle 17-V-1 (a). In this case, the turbulence is produced in the shear layer near the wall, and is transported to the channel axis by convection, resulting in an ascending curve for urms (x) measured on the channel axis. The same is true for the nozzle 17-T-1 (b), but the growth of turbulence on the channel axis takes place slightly more rapidly. The variants (c) and (d) additionally produce turbulence by unsteady detachments from the flow around the obstructions (grid or step) upstream. Therefore, turbulence is generated not only near the wall, but also in the core flow. As it can be inferred from the axial profiles of urms (x) in Fig. 5,
Fig. 5. Measurements of the turbulence velocity on the channel axis (y = 0, z = 0) behind different kind of nozzle designs (a)–(d).
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the level of turbulence which can be sustained on the channel axis at Re = 6000 is approximately urms = 3 − 3.5 m/s. Judged by these results, a variant comparable to the case (d) seems most advantageous, but also will cost most production efforts and most pressure loss. 3.2. Laminarization due to intense heating Experiments on local heat transfer have been conducted with the annular minichannel. Lateral effects caused by secondary flow in corners cannot occur here. Since the heated section is positioned more than 18 times the hydraulic diameters downstream of the hydraulic entrance, the hydraulic boundary layer is considered to be largely – but not completely – developed at the begin of the thermal development. This can also be inferred from the plots of the local friction coefficient shown in Fig. 6 which will be discussed below. Local Nusselt numbers can be inferred from the measured temperatures by a procedure using heatflux budgets for the heated structure and the fluid, as described by Arbeiter [3]. In Fig. 6, the local Nusselt numbers at two different heating rates q+ = 0.001 and 0.002, and the friction factors for adiabatic conditions, are plotted for Re = 6000 and Re = 10,500 in two diagrams. The friction factor fs (xA ,x) shown in Fig. 6 corresponds to the mean apparent friction factor between the entrance and the location x. It can be seen, that the changes in fs (xA ,x) are small after the beginning of the heater (xth > 0), indicating an already largely progressed hydraulic development. For the axial development of the Nusselt numbers, a significant effect of the heating rate can be observed. The Nusselt numbers are generally lower for the higher heating rates at the same entrance Reynolds number Re1 . In the case of Re1 = 10,500, the difference between the curves is being approximately maintained. In the case of Re1 = 6000, the curves converge. This can be understood, considering that the deteriorating effect of the heating reaches a lower bound when the flow is completely relaminarized. As seen from Fig. 7 (which has been previously discussed by Arbeiter et al. [5] in detail), the Nusselt number for developed flow at Re = 6000 is showing incipient
Fig. 6. Axial profiles of the Fanning friction factor and the Nusselt number in the annular minichannel for two Reynolds numbers and two heating rates.
Fig. 7. Measured local Nusselt numbers 30 mm downstream of the beginning of the heated section as a function of the local Reynolds number.
effects of turbulence, but is still very close to the value of the laminar correlation. This experimental evidence shows, that the CFD tools employed for the layout of the HFTM must be able to handle both the influence of the Reynolds number, as well as of the heating rate, in order to correctly predict the axial variation of the local heat transfer. 3.3. Lateral flow distribution from the adiabatic rectangular minichannel velocity fields measurements Two different channels with different nozzle types have been examined by this setup. Both channels are shown in Fig. 3. The nozzle “17-V-1” (identified as case (a) in Fig. 5) contracts the channel gap width from 17 to 1 mm by a 26 mm long V-shaped tapering section. The nozzle “17-V-3-T-1” contracts the channel gap width from 17 to 3 mm over a V-shaped tapering section with rounded entry and a length of 78 mm, followed by a T-shaped step reducing the gap width from 3 to 1 mm. The 17-V-3-T-1 design is very similar to the present HFTM nozzle design. The crossection of the minichannel behind the nozzle is 1 mm × 45 mm for both testsections. Fig. 8 shows the profiles of the normalized time-averaged velocity U/Um (z) and urms (z) for one half width 0 ≤ z ≤ 22.5 mm of the channel. The two nozzle geometries produce different lateral distributions of velocity and turbulence. While the nozzle geometry “17-V-1” delivers a nearly constant velocity distribution over the complete channel width, with the expected decrease towards the
Fig. 8. Lateral profiles of mean velocity and turbulence velocity on the channel symmetry plane (y = 0) x = 15 mm behind the nozzle for two different nozzle geometries.
F. Arbeiter et al. / Fusion Engineering and Design 83 (2008) 1536–1542
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Fig. 9. Turbulent streaks become visible by particle precipitation on the testsection wall (disassembled after the experiment), and can be evidenced by LDA measurements, which show a drastic increase of turbulence within the streak regions.
wall only within 1.5 mm distance towards the lateral wall (z > 21), the nozzle “17-V-3-T-1” exhibits a distinguishable velocity defect for 16 mm ≤ z ≤ 21 mm. The turbulence velocity produced by the two nozzle geometries is very similar near the channel axis (z = 0). For the nozzle “17-V-1”, the turbulence velocity increases towards the lateral wall (with a local maximum near z = 18.5 mm), while for the nozzle “17-V-3-T-1”, there is a significant drop for the turbulence velocity in the accelerated region between the velocity defect and the wall. The local peak of turbulence, located within a minimum of velocity, is thought to result in a streak where high local heat transfer to a lower local massflow will result in a locally higher gas temperature. Concerning the specimen temperature, however, the effects of higher gas temperature, but combined with a higher heat transfer coefficient, have mutually contrary effects and tend to cancel each other out. Fine TiO2 powder is used as light scattering tracer particles needed for the LDA method. The maximum of the particle diameter distribution has been measured at 0.2–0.3 m, but in smaller fractions, larger particles are present up to 5 m. Traces of the particles can be found on the testsection walls after an experiment. It is well known, that particle precipitation is fostered, where the centrifugal forces of turbulent eddies can separate the suspended particles from the gas. A typical result of particle precipitation for the testsection with nozzle “17-V-1” is shown in Fig. 9. It can be seen, that the TiO2 deposits in wedge formed streaks near the wall. This observation has motivated comparative LDA measurements in the “streaked” regions as well as in the core region without streaks. The results for normalized time-averaged velocity, as well as turbulence velocity, are shown also in Fig. 9. The profiles are acquired along the wall normal y-axis at x = 50 mm, one at z = 0 on the channel axis, and one at z = 20.5 mm near the lateral wall inside the streaked region. The level of turbulence velocity was measured nearly twice as high at z = 20.5 mm compared to the channel axis at z = 0. In accordance with this, the velocity profile at z = 20.5 mm has a flat top shape typical for turbulent flow, while the velocity profile at z = 0 is similar to the parabola expected for laminar flows. Similar streaks could be observed in some experiments also at random positions away from the wall. It is obvious from the profiles of the normalized
time-averaged velocity U/Um shown in Fig. 9, that the volumetric flow is significantly lower in the turbulent wedges than in the adjacent regions with less turbulence and therefore higher peak velocity. The massflow is displaced from the regions, where turbulence first develops, to the adjacent laminar/transitional regions which are thereby accelerated. This is the reason for U/Um reaching values of approximately 1.6 on the channel axis, while the analytical solution for laminar flow predicts a maximum of only 1.5. Again, the combination of a flow region with low massflow rate, but increased heat transfer coefficient, will result in a hot gas streak. A solution to these lateral inhomogeneities of heat transfer is seen by providing fully turbulent inlet conditions over the whole width of the channel. Introducing turbulence enhancing devices at the entrance as described in Section 3.1 will remove both issues, the unstable inlet turbulence patterns and the variation of heat transfer in lateral direction. 4. Conclusions Experiments have shown that the heat transfer in channel geometries and boundary conditions relevant for the design of the IFMIF high flux test module is influenced by the laminar to turbulent transition, as well as the retransition from turbulent to laminar flow (in the case of intense heating). The effects superimpose, and can result in axial and lateral variations of the turbulence velocity and the massflow distribution, which affect the gas and surface temperature. The effects are especially pronounced, if the nozzle is smooth, and the inlet turbulence level is low. Furthermore, the shape of the nozzle will determine the lateral distribution of the inflow to the minichannel. Based on these results, new design features to stabilize and increase the turbulence at the inlet of the minichannels were studied, whereby a spiral like turbulence generator design, or alternatively a forward facing step at the inlet shows most promising results, and will be considered in the HFTM design evolution. The experimental data will also be used to qualify numerical methods (i.e. turbulence models, discretization schemes, mesh structure) which can be employed in the optimization process in the future. The optimization of minichannel cooling system design
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of the HFTM is of high relevance to achieve a high specimen number in the available irradiation space, combined with precise temperature control, both necessary to guarantee a good quality of the materials testing program in IFMIF. References [1] J.O. Hinze, Turbulence, McGraw Hill, New York, 1987. [2] D.K. McLaughlin, W.G. Tiedermann, Biasing correction for individual realization of laser anemometer measurements in turbulent flows, Phys. Fluids 16 (12) (1973) 2082–2088.
[3] F. Arbeiter, Experimental and numerical investigations on minichannel cooling gas thermal–hydraulics, Proceedings of the 15th International Conference on Nuclear Engineering ICONE15-10514, 2007. [4] F. Arbeiter, S. Gordeev, V. Heinzel, T. Ihli, Laseroptische Untersuchungen des dreidimensionalen Strömungsfeldes von IFMIF-relevanten Gasströmungen in rechteckigen Minikanälen, Jahrestagung Kerntechnik, May 22–25, 2007, Karlsruhe. [5] F. Arbeiter, S. Gordeev, V. Heinzel, D. Leichtle, E. Stratmanns, Mini-channel flow experiments and CFD validation analyses with the IFMIF thermohydraulic experimental facility (ITHEX), Fusion Eng. Des. 82 (15–24) (2007) 2456–2461.
Contents lists available at ScienceDirect
Fusion Engineering and Design journal homepage: www.elsevier.com/locate/fusengdes
Features and optimization approaches of the entrance section cooling gas flow of the IFMIF High Flux Test Module F. Arbeiter a,∗ , S. Gordeev a , V. Heinzel a , T. Ihli a , D. Leichtle a , A. Möslang b , V. Slobotchouk a a b
Institut für Reaktorsicherheit, Forschungszentrum Karlsruhe, Hermann-von-Helmholtz-Platz 1, 76344 Eggenstein-Leopoldshafen, Germany Institut für Materialforschung I, Forschungszentrum Karlsruhe, Hermann-von-Helmholtz-Platz 1, 76344 Eggenstein-Leopoldshafen, Germany
a r t i c l e
i n f o
Article history: Available online 3 September 2008 Keywords: IFMIF High flux test module Helium Local heat transfer Entrance effects Minichannel Gas cooling
a b s t r a c t The International Fusion Materials Irradiation Facility (IFMIF) is devised to contribute experimental evidence to an irradiated material properties database for candidate materials exposed to irradiation spectra and doses relevant for future fusion power reactors. Due to neutron fluxes generated by high-energy deuterons reacting in a liquid lithium target, damage rates of 20–50 displacements per atom in one full power year can be achieved in steel specimens inside a volume of approximately 0.5 L. The design of the high flux test module developed at the Forschungszentrum Karslruhe (FZK) allows for maximizing the space available in the high flux neutron field for material irradiation, while at the same time allowing precise adherence of the irradiation temperature of the specimen stacks. Since enhancement of the neutron irradiation requires placement of the specimens as close as possible to the neutron source, the design proposes thin container structures (obeying mechanical constraints) and flat coolant channels between the rigs. A helium gas flow is designated to remove the heat from the rigs to keep the required irradiation temperature, which may be chosen between 250 and 650 ◦ C. As a result of the thin container walls and the small channel dimensions, the helium cooling gas flow is characterized by low pressure, transitional Reynolds numbers and intermediate Mach numbers. Dedicated experimental investigations on such minichannel cooling gas flows have been conducted with the ITHEX helium loop facility. Results obtained by laser Doppler anemometry indicate a complex three-dimensional evolution of the transitional laminar-turbulent flow field in the hydraulic entrance section. In the short cooling channels, a relevant portion of the flow alongside the rigs is influenced by this developing region. Detailed knowledge of the flow development and the resulting heat transfer coefficients is necessary to optimize the flow channel inlet design and to avoid inhomogeneities of the temperature field inside the specimen stacks, which otherwise could be caused by varying local heat transfer coefficients and mass flux redistributions (in the axial and the lateral coordinates). Experimental results are presented. Specific features of the minichannel entrance flow are identified, and conclusions are drawn for an optimized design of the entrance geometry, to enhance the utility of the HFTM for fusion materials research. © 2008 Elsevier B.V. All rights reserved.
1. Introduction The International Fusion Materials Irradiation Facility (IFMIF) is devised to contribute experimental evidence to an irradiated material properties database for candidate materials exposed to irradiation spectra and doses relevant for future fusion power reactors. Due to neutron fluxes generated by high-energy deuterons reacting in a liquid lithium target, damage rates of 20–50 displacements per atom in one full power year can be achieved in steel
∗ Corresponding author. Tel.: +49 7247 823452. E-mail address: [email protected] (F. Arbeiter). 0920-3796/$ – see front matter © 2008 Elsevier B.V. All rights reserved. doi:10.1016/j.fusengdes.2008.07.014
specimens inside a volume of approximately 0.5 L. The design of the high flux test module (the abbreviation HFTM will be used) developed at the Forschungszentrum Karslruhe (FZK) allows for maximizing the space available in the high flux neutron field for material irradiation, while at the same time allowing precise adherence of the irradiation temperature of the specimen stacks. Since enhancement of the neutron irradiation requires placement of the specimens as close as possible to the neutron source, the design proposes thin container structures (obeying mechanical constraints) and flat coolant channels between the rigs. A helium gas flow is designated to remove the heat from the rigs to keep the required irradiation temperature, which may be chosen between 250 and 650 ◦ C. As a result of the thin container walls and the small channel dimensions, the helium cooling gas flow is characterized
F. Arbeiter et al. / Fusion Engineering and Design 83 (2008) 1536–1542
Nomenclature channel cross-sectional area AQ b rectangular channel width cp specific heat capacity for constant pressure dh = 4AQ /SW hydraulic diameter fs (x1 ,x2 ) apparent Fanning friction factor ˙ m mass flux through the channel N number of samples from LDA p static pressure q˙ w specific wall heat flux r radius s channel gap width SW channel wetted perimeter T Kelvin temperature u instantaneous velocity urms root mean square velocity uShift,LDA virtual velocity of the LDA fringe pattern U time-averaged axial velocity ˙ Um = m/( 1 AQ ) bulk fluid velocity x, y, z Cartesian coordinates axial distance from the beginning of the heater xth Greek symbols thermal conductivity kinematic viscosity gas density w wall shear stress Indices 1 2 i fl,m w
conditions in the entrance plenum conditions in the exit plenum integer bulk fluid property property at wall
by low pressure, transitional Reynolds numbers and intermediate Mach numbers. Since the gas cooling, together with the nuclear heat release and electrical compensation heaters, determines the temperature homogeneity within the specimen stacks, as well as the level of thermomechanical stresses in the thin container structures, the thermal–hydraulic layout and the reliability of the cooling system during operation is a crucial point within the ongoing optimization process. In this paper new experimental results on the thermal–hydraulics within the HFTM cooling ducts are discussed. Improved insight into the relevant local flow phenomena was obtained. The experimental validation of the numerical tools employed in the layout process is an ongoing task; first results were already published by Arbeiter [3–5]. Optimization approaches for the inlet sections are based on the experimental results presented in this paper. 2. Thermal–hydraulic situation of the high flux test module As indicated in Fig. 1, the material specimens are stacked in 12 irradiation rigs, which are arranged in a 4 × 3 pattern directly behind the neutron source. Using spacers on the rig surfaces, channels with gap widths of 0.6 mm are created between the rigs and the compartment walls, and gap widths of 1.0 mm remain between adjacent rigs as well as between the rigs and the outer container walls. The cross-sections of these channels are rectangular, L-shaped or T-shaped, and are subject to the ongoing optimization of the HFTM. Helium gas at a moderate pressure of about 3 bars is
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routed upwards through these channels alongside the rig walls as a coolant. The gas enters the channels from a distribution plenum via converging nozzles (Fig. 1). The nozzle structure also serves as a neutron reflector, which results in higher neutron fluxes in the specimen. In this way, an array of parallel minichannels is created and supplied with cooling gas. A similar layout is used for the lateral neutron reflectors. The temperature in the specimen and the structures are determined by the local nuclear heat release, by conduction inside the structures, by the convective heat transfer to the helium gas flow, and by auxiliary electrical heaters with controllable power. The electrical heaters can replace the nuclear heat during beam off periods, and are used to compensate the axial temperature gradient in the specimen stacks. To ensure correct dimensioning of the heaters, the local heat transfer to the gas flow must be known. It is the aim of the thermal–hydraulic optimization to adjust a homogenous temperature field within a single specimen stack, and avoid temperature gradients in the structure which would result in thermomechanical stresses and deformations. Thermomechanical stresses represent the main loads in some parts of the module. The determination of channel deformations, the distribution of the coolant flow and the temperature field are coupled problems; therefore iterative application of CFD and FEM analyses are necessary. The flow is subject to several phenomena which can have considerable impact on the local heat transfer from the walls and therefore are considered in detail. • Entrance effects. Growing hydraulic boundary layer after the entrance nozzle, and growing thermal boundary layer after the beginning of the heated section. Hydraulic and thermal development regions partially overlap. • Laminar to turbulent transition. The flow is laminarized by the acceleration in the entrance nozzle, the channel Reynolds numbers range between 6000 and 10,000. The adiabatic flow is therefore in the laminar-turbulent transitional range. • Turbulent to laminar retransition. The considerable heating rates of the flow effect gradients of the gas properties in axial and wall normal directions. The gas viscosity, which rises with increasing temperature, diminishes the local Reynolds number near the wall and increases the turbulence dissipation. Additionally, the high pressure drop and the rise in mean temperature of the gas decrease its density, and thereby accelerate the flow. This will result in a negative term in the turbulence kinetic energy budget (see Ref. [1]) and could result in relaminarization. • Lateral redistribution of massflow. The nozzle shape determines the massflow distribution at the entrance of the channel. Inside the channel, secondary flows in the corners, as well as growing turbulent spots at random positions within the laminar-turbulent transitional region, exhibit flat velocity profiles and thereby displace massflow to the adjacent laminar regions. This paper aims to present a comprehensive picture of the thermal–hydraulic situation studied by experiments. Where single effects have already been discussed in more details in previous publications, references are given. 3. ITHEX experimental programme To examine the thermal–hydraulic effects listed above, an extensive experimental programme has been conducted. Two different types of testsections were used: the first testsection was an annular minichannel made of stainless steel, instrumented with electrical heaters and thermocouples between the heaters and the cooled surfaces in the inner and the outer cylinder. This annular testsection
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Fig. 1. View on the IFMIF high flux test module, showing the arrangement of specimen irradiation rigs, inlet nozzles and minichannels.
has been used to study the laminar to turbulent transition, and the laminarizing effects of the intense heating. A 3D CAD drawing and a cross-section of this testsection are shown in Fig. 2. The annular channel has a hydraulic diameter dh = 1.23 mm, a complete length of 94 mm and a heated length of 50 mm. As indicated in Fig. 2, the axial coordinate xth originates at the beginning of the heated sec-
Fig. 2. ITHEX annular minichannel testsection with major dimensions and coordinate system.
tion. The second testsection is a rectangular duct with optical access for 1D Laser Doppler Anemometry (LDA) to measure the evolution of the velocity and turbulence fields in the minichannel, behind different shapes of entrance nozzles. A 3D CAD drawing of two of these nozzle configurations is shown in Fig. 3. The orientation of the coordinate system used is also indicated in Fig. 3. The origin y = 0, z = 0 is defined on the channel symmetry axis; x = 0 is defined at the location where the nozzle ends and the minichannel with the constant cross-section 1 mm × 45 mm begins. The LDA probe in backscattering configuration can access the testsection through an optical window (not shown in the drawing) from the side (the
Fig. 3. ITHEX rectangular duct testsections for LDA measurements.
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laser beams are intersecting in the xz-plane), and thus measure the velocity component in the axial x direction Results will be presented for both testsections, to address the issues of local flow phenomena listed above. To describe the results, the variables defined in Eqs. (1)–(6), which are commonly used in fluid dynamics, are chosen. Re =
Um dh
(1)
= q+ 1
q˙ w AQ ˙ p (T1 )T1 mc
(2)
Nu(x) =
dh q˙ w (x) f,ml Tfl,m (x) − Tw (x)
2 f = w 1/2Um
N
U=
i=1 N
−1
ui w(ui )
i=1
w(ui )
urms =
(3) (4)
N with
w(ui ) =
N w(ui )(ui − U)2 i=1 N w(ui ) i=1
u /N i=1 i
ui − uShift,LDA
(5) Fig. 4. Results of LDA measurements on the channel axis (y = 0, z = 0) in the testsection “17-V-1” at Re = 6000 showing the transitional nature of the flow.
1/2
(6)
The variable q+ defined by Eq. (2) is referred to as heating rate, since it is proportional to (T2 − T1 )/T1 for a perfect gas. To determine the time-averaged velocity U, and the turbulence velocity urms , a weighting factor w(ui ) is applied to the individual LDA samples to reduce systematic errors evoked by the individual realization sampling procedure inherent to LDA, similar to the proposal of McLaughlin and Tiederman [2]. LDA experiments were carried out with ambient air; therefore the density, and subsequently the velocity magnitude of individual experiments adjusted to the same Reynolds number is slightly different for each individual experiment. To cancel this effect, the velocity profiles are presented scaled ˙ to the bulk fluid velocity Um = m/( 1 AQ ) of the individual experiment. The quantity U/Um is referred to as normalized velocity. 3.1. Axial boundary layer development behind the entrance nozzle and laminar to turbulent transition The shape of the entrance nozzle determines the amount of laminarization in the flow due to acceleration, but also the regeneration of turbulence by high shear forces in the region where the nozzle enters the minichannel. The geometry of this region will control the magnitude, but also the spatial distribution of turbulence at the entrance of the minichannel, and thus critically determines the length and the character of the hydraulic development of the boundary layer in the minichannel. Effects such as flow detachment, can also occur at the end of the nozzle. Different generic shapes based on the “17-V-1” nozzle shape shown in Fig. 2 were studied using the LDA setup. For a smooth nozzle of the 17-V-1 geometry and transitional Reynolds numbers 3000 ≤ Re ≤ 12,000, it was found from LDA as well as from friction measurements, that the flow can realize both laminar and turbulent state, induced by very small disturbances. Such results for the normalized velocity U(x)/Um at Re1 = 6000 are shown in Fig. 4, where the results of multiple experiments (identified by a label DDMMYYx), all for similar inlet conditions, are mutually compared. High levels (typical value U(x = 50 mm)/Um = 1.5) of U(x) indicate laminar flow, while low levels (typical value U(x = 50 mm)/Um = 1.3) of U(x) indicate turbulent flow. This effect has been discussed by Arbeiter [3] and Arbeiter et al. [4] in more detail. It is based on the appearance of flow detachment in the entrance, as well as on the laterally distributed onset of turbulence discussed below. As such behaviour must be avoided in all circumstances in the IFMIF HFTM application, different design
modifications dedicated to yield a reproducible level of turbulence of high magnitude to ensure a short hydraulic entrance length were studied in the next step. Of course, the new design features for reliable turbulence production must be compatible with the manufacturing technology available to fabricate the complex nozzle array. So far, the shapes which have been tested were (a) the basic 17-V-1 nozzle (smooth), (b) a T-shaped step instead of a nozzle, (c) a protruding small step mounted on the surface of the 17-V-1 nozzle, and (d) a grid (formed by a spiral of thick wire) inserted into the 17-V-1 nozzle. The axial profiles for the turbulence velocity on the channel axis (y = 0, z = 0) for these four nozzle shapes are plotted in Fig. 5. As already shown in Fig. 4, the smooth 17-V-1 (a) nozzle can produce various modes of flow. In Fig. 5, the laminar/turbulent case has been included for the nozzle 17-V-1 (a). In this case, the turbulence is produced in the shear layer near the wall, and is transported to the channel axis by convection, resulting in an ascending curve for urms (x) measured on the channel axis. The same is true for the nozzle 17-T-1 (b), but the growth of turbulence on the channel axis takes place slightly more rapidly. The variants (c) and (d) additionally produce turbulence by unsteady detachments from the flow around the obstructions (grid or step) upstream. Therefore, turbulence is generated not only near the wall, but also in the core flow. As it can be inferred from the axial profiles of urms (x) in Fig. 5,
Fig. 5. Measurements of the turbulence velocity on the channel axis (y = 0, z = 0) behind different kind of nozzle designs (a)–(d).
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the level of turbulence which can be sustained on the channel axis at Re = 6000 is approximately urms = 3 − 3.5 m/s. Judged by these results, a variant comparable to the case (d) seems most advantageous, but also will cost most production efforts and most pressure loss. 3.2. Laminarization due to intense heating Experiments on local heat transfer have been conducted with the annular minichannel. Lateral effects caused by secondary flow in corners cannot occur here. Since the heated section is positioned more than 18 times the hydraulic diameters downstream of the hydraulic entrance, the hydraulic boundary layer is considered to be largely – but not completely – developed at the begin of the thermal development. This can also be inferred from the plots of the local friction coefficient shown in Fig. 6 which will be discussed below. Local Nusselt numbers can be inferred from the measured temperatures by a procedure using heatflux budgets for the heated structure and the fluid, as described by Arbeiter [3]. In Fig. 6, the local Nusselt numbers at two different heating rates q+ = 0.001 and 0.002, and the friction factors for adiabatic conditions, are plotted for Re = 6000 and Re = 10,500 in two diagrams. The friction factor fs (xA ,x) shown in Fig. 6 corresponds to the mean apparent friction factor between the entrance and the location x. It can be seen, that the changes in fs (xA ,x) are small after the beginning of the heater (xth > 0), indicating an already largely progressed hydraulic development. For the axial development of the Nusselt numbers, a significant effect of the heating rate can be observed. The Nusselt numbers are generally lower for the higher heating rates at the same entrance Reynolds number Re1 . In the case of Re1 = 10,500, the difference between the curves is being approximately maintained. In the case of Re1 = 6000, the curves converge. This can be understood, considering that the deteriorating effect of the heating reaches a lower bound when the flow is completely relaminarized. As seen from Fig. 7 (which has been previously discussed by Arbeiter et al. [5] in detail), the Nusselt number for developed flow at Re = 6000 is showing incipient
Fig. 6. Axial profiles of the Fanning friction factor and the Nusselt number in the annular minichannel for two Reynolds numbers and two heating rates.
Fig. 7. Measured local Nusselt numbers 30 mm downstream of the beginning of the heated section as a function of the local Reynolds number.
effects of turbulence, but is still very close to the value of the laminar correlation. This experimental evidence shows, that the CFD tools employed for the layout of the HFTM must be able to handle both the influence of the Reynolds number, as well as of the heating rate, in order to correctly predict the axial variation of the local heat transfer. 3.3. Lateral flow distribution from the adiabatic rectangular minichannel velocity fields measurements Two different channels with different nozzle types have been examined by this setup. Both channels are shown in Fig. 3. The nozzle “17-V-1” (identified as case (a) in Fig. 5) contracts the channel gap width from 17 to 1 mm by a 26 mm long V-shaped tapering section. The nozzle “17-V-3-T-1” contracts the channel gap width from 17 to 3 mm over a V-shaped tapering section with rounded entry and a length of 78 mm, followed by a T-shaped step reducing the gap width from 3 to 1 mm. The 17-V-3-T-1 design is very similar to the present HFTM nozzle design. The crossection of the minichannel behind the nozzle is 1 mm × 45 mm for both testsections. Fig. 8 shows the profiles of the normalized time-averaged velocity U/Um (z) and urms (z) for one half width 0 ≤ z ≤ 22.5 mm of the channel. The two nozzle geometries produce different lateral distributions of velocity and turbulence. While the nozzle geometry “17-V-1” delivers a nearly constant velocity distribution over the complete channel width, with the expected decrease towards the
Fig. 8. Lateral profiles of mean velocity and turbulence velocity on the channel symmetry plane (y = 0) x = 15 mm behind the nozzle for two different nozzle geometries.
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Fig. 9. Turbulent streaks become visible by particle precipitation on the testsection wall (disassembled after the experiment), and can be evidenced by LDA measurements, which show a drastic increase of turbulence within the streak regions.
wall only within 1.5 mm distance towards the lateral wall (z > 21), the nozzle “17-V-3-T-1” exhibits a distinguishable velocity defect for 16 mm ≤ z ≤ 21 mm. The turbulence velocity produced by the two nozzle geometries is very similar near the channel axis (z = 0). For the nozzle “17-V-1”, the turbulence velocity increases towards the lateral wall (with a local maximum near z = 18.5 mm), while for the nozzle “17-V-3-T-1”, there is a significant drop for the turbulence velocity in the accelerated region between the velocity defect and the wall. The local peak of turbulence, located within a minimum of velocity, is thought to result in a streak where high local heat transfer to a lower local massflow will result in a locally higher gas temperature. Concerning the specimen temperature, however, the effects of higher gas temperature, but combined with a higher heat transfer coefficient, have mutually contrary effects and tend to cancel each other out. Fine TiO2 powder is used as light scattering tracer particles needed for the LDA method. The maximum of the particle diameter distribution has been measured at 0.2–0.3 m, but in smaller fractions, larger particles are present up to 5 m. Traces of the particles can be found on the testsection walls after an experiment. It is well known, that particle precipitation is fostered, where the centrifugal forces of turbulent eddies can separate the suspended particles from the gas. A typical result of particle precipitation for the testsection with nozzle “17-V-1” is shown in Fig. 9. It can be seen, that the TiO2 deposits in wedge formed streaks near the wall. This observation has motivated comparative LDA measurements in the “streaked” regions as well as in the core region without streaks. The results for normalized time-averaged velocity, as well as turbulence velocity, are shown also in Fig. 9. The profiles are acquired along the wall normal y-axis at x = 50 mm, one at z = 0 on the channel axis, and one at z = 20.5 mm near the lateral wall inside the streaked region. The level of turbulence velocity was measured nearly twice as high at z = 20.5 mm compared to the channel axis at z = 0. In accordance with this, the velocity profile at z = 20.5 mm has a flat top shape typical for turbulent flow, while the velocity profile at z = 0 is similar to the parabola expected for laminar flows. Similar streaks could be observed in some experiments also at random positions away from the wall. It is obvious from the profiles of the normalized
time-averaged velocity U/Um shown in Fig. 9, that the volumetric flow is significantly lower in the turbulent wedges than in the adjacent regions with less turbulence and therefore higher peak velocity. The massflow is displaced from the regions, where turbulence first develops, to the adjacent laminar/transitional regions which are thereby accelerated. This is the reason for U/Um reaching values of approximately 1.6 on the channel axis, while the analytical solution for laminar flow predicts a maximum of only 1.5. Again, the combination of a flow region with low massflow rate, but increased heat transfer coefficient, will result in a hot gas streak. A solution to these lateral inhomogeneities of heat transfer is seen by providing fully turbulent inlet conditions over the whole width of the channel. Introducing turbulence enhancing devices at the entrance as described in Section 3.1 will remove both issues, the unstable inlet turbulence patterns and the variation of heat transfer in lateral direction. 4. Conclusions Experiments have shown that the heat transfer in channel geometries and boundary conditions relevant for the design of the IFMIF high flux test module is influenced by the laminar to turbulent transition, as well as the retransition from turbulent to laminar flow (in the case of intense heating). The effects superimpose, and can result in axial and lateral variations of the turbulence velocity and the massflow distribution, which affect the gas and surface temperature. The effects are especially pronounced, if the nozzle is smooth, and the inlet turbulence level is low. Furthermore, the shape of the nozzle will determine the lateral distribution of the inflow to the minichannel. Based on these results, new design features to stabilize and increase the turbulence at the inlet of the minichannels were studied, whereby a spiral like turbulence generator design, or alternatively a forward facing step at the inlet shows most promising results, and will be considered in the HFTM design evolution. The experimental data will also be used to qualify numerical methods (i.e. turbulence models, discretization schemes, mesh structure) which can be employed in the optimization process in the future. The optimization of minichannel cooling system design
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of the HFTM is of high relevance to achieve a high specimen number in the available irradiation space, combined with precise temperature control, both necessary to guarantee a good quality of the materials testing program in IFMIF. References [1] J.O. Hinze, Turbulence, McGraw Hill, New York, 1987. [2] D.K. McLaughlin, W.G. Tiedermann, Biasing correction for individual realization of laser anemometer measurements in turbulent flows, Phys. Fluids 16 (12) (1973) 2082–2088.
[3] F. Arbeiter, Experimental and numerical investigations on minichannel cooling gas thermal–hydraulics, Proceedings of the 15th International Conference on Nuclear Engineering ICONE15-10514, 2007. [4] F. Arbeiter, S. Gordeev, V. Heinzel, T. Ihli, Laseroptische Untersuchungen des dreidimensionalen Strömungsfeldes von IFMIF-relevanten Gasströmungen in rechteckigen Minikanälen, Jahrestagung Kerntechnik, May 22–25, 2007, Karlsruhe. [5] F. Arbeiter, S. Gordeev, V. Heinzel, D. Leichtle, E. Stratmanns, Mini-channel flow experiments and CFD validation analyses with the IFMIF thermohydraulic experimental facility (ITHEX), Fusion Eng. Des. 82 (15–24) (2007) 2456–2461.