- Email: [email protected]
Experimental results of a compact thermally driven cooling system based on metal hydrides
i n t e r n a t i o n a l j o u r n a l o f h y d r o g e n e n e r g y 3 5 ( 2 0 1 0 ) 7 6 2 3 e7 6 3 2
Available at www.sciencedirect.com
journal homepage: www.elsevier.com/locate/he
Experimental results of a compact thermally driven cooling system based on metal hydrides Marc Linder*, Rainer Mertz, Eckart Laurien Institute of Nuclear Technology and Energy Systems, University of Stuttgart, Pfaffenwaldring 31, 70569 Stuttgart, Germany
article info
abstract
Article history:
In this paper a detailed experimental analysis of a metal hydride based cooling system is
Received 19 March 2010
presented. For the high temperature side an AB5 type alloy (LmNi4.91Sn0.15) was chosen,
Received in revised form
whereas an AB2 type alloy (Ti0.99Zr0.01V0.43Fe0.09Cr0.05Mn1.5) is used for the low temperature
22 April 2010
side. Due to very good heat and also mass transfer characteristics (among others, large heat
Accepted 23 April 2010
transfer surface area) of the utilized capillary tube bundle reaction bed, very short half-
Available online 3 June 2010
cycle times in the order of 100 s have been reached. Consequently, the specific cooling
Keywords:
temperature boundary conditions. The system was experimentally analyzed for different
Metal hydride
cooling and ambient temperatures, whereas the heating temperature was fixed to 130 C.
power of the system is up to 780 W per kg desorbing metal hydride e depending on the
Sorption system
ª 2010 Professor T. Nejat Veziroglu. Published by Elsevier Ltd. All rights reserved.
Heat pump Capillary tube bundle reaction bed
1.
Introduction
The thermally driven generation of cold seems to be a reasonable way to utilize waste heat to increase the overall efficiency of energy conversion systems. In case of automotive cooling, a thermally driven cooling system could supply the cooling demand using waste heat from the coolant or the exhaust gases. The substitution of mechanical energy (generated by the engine) by waste heat as main driving energy for the cooling system would consequently lead to reduced fuel consumption along with reduced CO2 emissions. Additionally, the climate damaging refrigerants of current automotive air-conditioning systems could be replaced. However, main obstacles for an automotive application of thermally driven sorption systems are their current weight and volume constraints. Due to the fast intrinsic reaction kinetics of metal hydrides, compact and light-weight sorption systems can be realized if limitations due to insufficient mass or heat transfer within the
reaction bed can be minimized. Therefore, this work focuses on the experimental investigation of a metal hydride sorption system using capillary tube bundle based reaction beds in order to increase the weight specific power of the system.
1.1.
Working principle
The basic configuration of metal hydride sorption systems consists of two reaction beds coupled by a hydrogen connection pipe. Each reaction bed contains a different metal hydride, indicated as A and B (Fig. 1). As the equilibrium pressure of metal hydride A is higher than the equilibrium pressure of metal hydride B at the same temperature, metal hydride A tends to release hydrogen and cools down. If the coupled metal hydride B is able to absorb the released hydrogen, a hydrogen exchange along with the corresponding cooling effect is obtained (cooling half-cycle). The absorption of hydrogen in the coupled low-pressure hydride B generates heat (QAmb_cool) that must be released to a heat sink to prevent
* Corresponding author. E-mail address: [email protected] (M. Linder). 0360-3199/$ e see front matter ª 2010 Professor T. Nejat Veziroglu. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2010.04.184
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Fig. 1 e Van’t Hoff diagram of sorption cooling system.
a temperature increase of the alloy. As the environment is generally used as heat sink its temperature (TAmb) determines the characteristic of the coupled alloy B. As soon as the maximum amount of hydrogen is exchanged, the flow direction must be reversed. Therefore, a thermal energy input (QHeat) is necessary to supply the desorption enthalpy of metal hydride B. As the absorption heat during regeneration (QAmb_reg) is released to the environment, the absorption pressure of metal hydride A at ambient temperature determines the necessary regeneration pressure level (pReg). The described batch system is not able to supply a continuous cooling effect. As the reaction bed pairs working either in regeneration or cooling mode, two identical reaction bed couples working in opposite direction are necessary to realize a quasi-continuous cold output. Therefore, the necessary time for regeneration must be shorter or at least identical to the cooling half-cycle time.
1.2.
inner artery was intensively used to investigate different sorption system prototypes [1e5]. The second design for hydrogen distribution was recently investigated by Qin et al. [6] and Ni et al. [7] in 2007. Instead of a central filter tube, the filter surrounds the metal powder bulk. Therefore, hydrogen enters the reaction bed from one side, is distributed in the annular gap around the filter and enters in radial direction into the powder bulk. A finned copper tube is inserted in the bulk to divide the bed into several sections and to improve the heat transfer characteristics. Kang et al. [8] used a comparable design: The reaction bed consisted of two identical copper tubes with aluminum fins working serially in an U-shaped heat exchanger and hydrogen distribution was realized by a gap between the metal hydride part of the reaction bed and its pressure vessel. Although, heat recovery measures of the systems and different working temperatures complicate a direct comparison of the systems, the relevant weight-specific characteristics for coupled metal hydride reaction beds realized within the last 15 years can be summarized as follows: The half-cycle time ranges between 10 and 20 min, depending on the applied reaction bed design. The maximum achieved specific power is around 140e170 W per kg desorbing alloy of the respective reaction bed. As most of the intended applications were stationary, the necessity to increase the specific power by means of heat transfer enhancements was mainly based on the reduction of metal powder and not implicitly on weight and volume optimizations. However, taking into account weight and volume constraints of automotive applications, the need to increase the weight specific performance of the metal hydride sorption technology is obvious. Therefore, this work focuses on the experimental analysis of a sorption system using a capillary tube bundle based reaction bed with a large heat transfer surface.
State of the art
In order to compare different designs of metal hydride sorption systems regarding their weight specific performance, two characteristic values can be deduced from published data, the specific power (per kg desorbing alloy in the respective reaction bed) and the half-cycle time. Although the systems may be designed for different applications, like combined heat/ cold generation or heat transforming, for each system the same principle applies: thermally driven hydrogen exchange between two reaction beds. The limiting factors (quantified by specific power and half-cycle time) are therefore independent of the designed application, the total number of reaction beds or the amount of different metal hydrides. In general, metal hydride reaction beds for sorption systems are of cylindrical shape with two different possible hydrogen distribution designs: Hydrogen enters the metal hydride bulk either through an inner filter artery or through a surrounding filter tube. Both cases enable an easy hydrogen distribution along the axial direction, whereas the radial hydrogen distribution characteristic depends on the consistency of the bulk material. The reaction bed design with an
Fig. 2 e Schematic flow diagram of test bench for coupled reaction beds.
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Table 1 e Measurement equipment of experimental set-up for coupled reaction beds. Transducer
Measurement range
Resistance thermometers Pt 100 Absolute pressure transducer Keller Druck, PA 23 Mass flow meter (Hydrogen) Hastings HFM-201 Mass flow meter (Water) Krohne.IFS6000F
Accuracy
Max. Deviation
100 to þ 450 C 0 to 50 bar
(0.15 þ 0.002*T) 0.5% F. S.
0.55 K (<200 C) 0.25 bar
0 to 100 sl/mina
1.3%F. S.
1.3 sl/min
0 to 14.97 1/min
0.3%F.S.
0.045 l/min
AGILENT Data Acquisition. Switch Unit (34 970 A). a standard liter, for T ¼ 0 C and p ¼ 1.013 bar.
2.
Experimental details
2.1.
Experimental set-up
The test bench shown in Fig. 2 consists of a hydrogen part (blue), the two reaction beds (A and B including reaction bed valves V1 and V2) and two separate water cycles (fully drawn and dashed lines, respectively). The main component of the hydrogen part is the mass flow meter implemented in the hydrogen connection pipe. Additionally, a pneumatically actuated hydrogen valve (V3) allows the separation of the reaction beds. Each water cycle consists of a pump, a mass flow meter for water, two resistance thermometers (T1-T4), two electromagnetic 3-way valves (EMV1-EMV4) and two heat exchangers. Additionally, each reaction bed contains an internal heat exchanger that is connected to the water cycle in order to transfer heat from the water current to the metal hydride during desorption or vice versa during absorption. The connected thermostats and electrical heaters, not shown in Fig. 2, control the set-point temperatures of the heat exchangers. During regeneration mode, the necessary driving heat (QHeat) is supplied by the left heat exchanger in the dashed water cycle, whereas the left heat exchanger in the full-drawn cycle rejects heat (QAmb_reg) at simulated ambient temperature. The switching of the electro-magnetic valves leads to a temperature (and correspondingly pressure) change of the reaction beds. During the following cooling half-cycle, the heat input at cooling temperature (QCool) and heat rejection at ambient temperature (QAmb_cool) are realized by the second heat exchangers in the respective water cycles. As pure water is used as heat carrier fluid, the hot water cycle is pressurized to around 4 bar to prevent boiling at temperatures of around 130 C. For a given half-cycle time (time between switching of the electro-magnetic valves) and adjusted heat exchanger
Fig. 3 e Capillary tube bundle reaction bed (length: 135 mm, diameter (max.): 65 mm).
temperature levels, the system is working automatically. A software program based on HP VEE Pro (Agilent) realizes the controlling of the test bench. The description of the implemented measurement equipment is summarized in Table 1.
2.2.
Reaction beds
The chosen capillary tube bundle reaction bed design has already been mentioned in publications [9e11] but up to now without detailed experimental results. Recently, a “zero-dimensional” model was developed and published based on selected experimental results obtained from the system described in this work [12]. The main part of the reaction bed is the capillary tube bundle heat exchanger shown in Fig. 3, left. It consists of 372 small tubes with an inner diameter of 1.4 mm made of stainless steel. The heat carrier flows through the tubes that are surrounded by the metal powder. To prevent discharge of the metal powder, a sintered metal filter tube surrounds the tube bundle (Fig. 3, middle). A cladding tube (Fig. 3, right) encloses the reaction bed and forms an annular gap around the filter tube that is used for uniform hydrogen distribution around the metal hydride bed. The total length of the assembled reaction bed is 135 mm. With an outer diameter of 76 mm the total volume of the reaction bed is around 0.6 l. Its weight without metal powder is 1.7 kg. The volume within the sintered metal tube (Vi z 0.23 l) is sufficient to charge 800e1000 g of powder, depending on the density of the alloy and its volume expansion during hydriding.
Table 2 e Properties of applied metal hydrides. Alloy LmNi4 91Sn0.15 Reaction enthalpy (948 g in reaction [J/molH2] Reaction entropy bed B) [J/molH2K] Plateau slope [-] a Ti0.99Zr0.01V0.43Fe0.09 Reaction enthalpy Cr0.05Mn1.5 [J/molH2] Reaction entropy (800 g in [J/molH2K] reaction bed A) Plateau slope [-] a
Absorption Desorption 27 411
29 849
103.9
109.1
0.43 17 922
0.48 22 606
89.2
101.1
0.42
0.22
a slope of PCI plateau calculated for mid of plateau region.
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2.3.
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Metal hydride working pair
Based on a detailed experimental analysis of available metal hydrides, LmNi4.91Sn0.15 was chosen for the high temperature side and Ti0.99Zr0.01V0.43Fe0.09Cr0.05Mn1.5 for the generation of the cooling effect. Respective PCI measurements were performed on a volumetrically measuring set-up (described in [13,14]). Based on the obtained PCI data, the characteristic properties of the respective alloys were deduced and are summarized in Table 2. Additionally, the respective van’t Hoff plots are shown in Fig. 4.
3.
Experimental results
The set-up consists of the reaction bed couple and four heat exchangers at three different temperature levels (THeat, TAmb and TCool). These temperature levels correspond to the respective simulated boundary conditions of the cooling system. However, the temperature sensors T1-T4 are close to the reaction beds and measure the respective in- and outlet temperature of the water current. In Fig. 5 these in- and outlet temperatures (T1-T4) are shown for 4 complete cycles. The temperatures of the heat exchangers were adjusted to THeat ¼ 130 C, TAmb ¼ 28 C and TCool ¼ 20 C. The blue lines correspond to the cold reaction bed A which is cycled between ambient and cooling temperature. As the temperatures of the reaction bed and the metal hydride are around ambient temperature, the reaction enthalpy at the beginning of the cooling cycle is used to cool down the thermal masses of the reaction bed. Therefore, the generated cooling effect starts slightly later than the actual changing of the half-cycle. After this short period, the endothermic reaction cools the water flow. This heat capacity effect is even more obvious for the hot reaction bed (red). Due to the higher temperature difference between cooling and regeneration
cycle, the necessary amount of thermal energy to reach the temperature level of the following half-cycle is higher. Therefore, the thermal masses, especially the parasitic thermal masses like pressure vessel, heat exchanger surface or filter tube should be minimized in order to maximize the performance of the system. In Fig. 6 the corresponding hydrogen pressure and the exchanged amount of hydrogen during the four cycles are shown. The fully drawn line represents the accumulated amount of exchanged hydrogen and is shown on the left y-axis, whereas the right y-axis gives the values of the pressure progression drawn as dashed line. The total amount of exchanged hydrogen (nH2) is around 3.6 mol. Based on the desorption enthalpy of the metal hydride in reaction bed A, the corresponding amount of thermal energy is around 81 kJ. In order to be able to open the reaction bed and to change the alloys, two massive brass constructions for the distribution of water at the in- and outlet are necessary. Whereas the alloy mass is around 900 g and the reaction bed itself weighs around 1.7 kg, the massive brass in- and oulets weight together additional 2 kg. These components are not necessary for possible applications but reduce the efficiency of the test bench due to the almost doubled parasitic mass of the reaction bed. Their impact is shown in Fig. 7, where the thermal power (measured in the water flow, fully drawn line) is compared with the calculated “hydrogen power” (dotted line). “Hydrogen power” is calculated by the measured mass flow rate of hydrogen and the respective reaction enthalpy of the metal hydride. It is the maximal achievable value (without thermal losses) as a temperature change of thermal masses also leads to a hydrogen exchange. Positive values correspond to the cooling half-cycle, whereas negative values represent the regeneration cycle. Although the temperature conditions are favourable (small difference between TCool and TAmb), the effective cooling power measured in the water flow is clearly lower than the theoretically achievable “hydrogen power” (without thermal losses).
Fig. 4 e Van’t Hoff plots of applied metal hydrides.
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Fig. 5 e In- and outlet temperature progression during cycling experiment.
3.1.
Variation of cooling temperature
The adjusted cooling temperature influences the behavior of the coupled reaction beds. In Fig. 8, this influence on the thermal power (PWater A, fully drawn line) and the calculated “hydrogen power” is shown (PH2,A, dashed line). As the heating energy is not changed and the cooling temperature influences the cooling half-cycle only, the regeneration half-cycles (negative values) are comparable. The main difference during
the cooling half-cycle is the higher measured thermal power for a cooling temperature of 20 C in comparison to the lower cooling temperatures of 15 C and 10 C. The reason is due to the thermal masses described earlier. With an increasing temperature difference during regeneration and cooling cycle, the fraction of exchanged hydrogen necessary to change the temperature of the reaction bed itself increases. Therefore, less hydrogen is available to cool to water flow. Additionally, the total amount of exchanged hydrogen (nH2) decreases with
Fig. 6 e Pressure and hydrogen exchange during cycling experiments.
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Fig. 7 e Cooling power based on hydrogen exchange rate and measured in the water flow.
Fig. 8 e Variation of the cooling temperature e effect on thermal and calc. power.
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Fig. 9 e Variation of the cooling temperature e effect on total hydrogen exchange.
decreasing cooling temperature (see Fig. 9, fully drawn lines, left y-axis). The comparably small difference in the measured thermal power between 15 C and 10 C cooling temperature is due to the instable inlet temperature during the cooling cycle. Additionally, the set point of 10 C is not reached and therefore the difference between adjusted 10 C and 15 C cooling temperature is reduced (Fig. 9, dotted lines, right y-axis). Both effects are due to the applied thermostat that
is not able to balance properly the high heat loads during regeneration and cooling mode for high temperature differences.
3.2.
Variation of ambient temperature
The ambient temperature (TAmb) is the most critical boundary condition for thermally driven sorption systems as it affects the regeneration as well as the cooling half-cycle. An
Fig. 10 e Hydrogen exchange and pressure development for different ambient temperatures.
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Fig. 11 e Pressure progression during the regeneration half-cycle for different ambient temperatures e absorption in reaction bed A (T3 [ 130 C).
increased ambient temperature increases the corresponding equilibrium pressure of the absorbing metal hydride and consequently the system pressure p2 during both half-cycles (see dashed lines in Fig. 10). The half-cycle time of the sorption system was set to 120 s. As the heating and the cooling temperature were fixed to THeat ¼ 130 C and TCool ¼ 20 C, respectively, the available desorption pressures are identical for all measurements. Therefore, the exchanged amount of hydrogen nH2 shown in Fig. 10 as fully drawn lines depends
clearly on the ambient temperature (TAmb). Around 4 mol of hydrogen are exchanged for an ambient temperature of 20 C, whereas only 3.8 and 3.4 mol are exchanged for TAmb ¼ 28 C and 35 C, respectively. In Fig. 11 and Fig. 12, the respective system pressure progressions (p2) are shown for one half-cycle and compared to the corresponding van’t Hoff plots of the absorbing alloy (non-logarithmic scale). Although, the desorbing alloy generates the thermally driven compression of hydrogen
Fig. 12 e Pressure progression during the cooling half-cycle for different ambient temperatures e absorption in reaction bed B (T1 [ 20 C).
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Fig. 13 e Achievable specific cooling power (related to the desorbing metal hydride).
during the regeneration half-cycle, the system pressure p2 at the end of the regeneration phase corresponds to the absorption pressure of metal hydride A with a load situation of around 75% of the maximum storage capacity (see nonlogarithmic van’t Hoff plots in Fig. 11). However, the load situation of metal hydride A at the end of the regeneration half-cycle decreases with increasing ambient temperature TAmb. Whereas the metal hydride is charged to more than 75% of the maximum hydrogen storage capacity for a water inlet temperature of T1 ¼ 20 C, the load situation is reduced to less than 70% at the end of the regeneration half-cycle for T1 ¼ 35 C (even for increased system pressure, about 47 bar vs. about 37 bar). Fig. 12 displays the progression of the system pressure p2 during the respective cooling half-cycle (T1 ¼ 20 C). A comparison with the van’t Hoff plot absorption lines of metal hydride B shows a similar dependency of the hydrogen load situation on the ambient temperature. The metal hydride B is charged to less than 85% for T3 ¼ 35 C, whereas a load situation of almost 90% of the maximum storage capacity is reached for a water inlet temperature of 20 C. Taking the used metal hydride masses and the maximum storage capacity of the respective alloys into account, 5% of the maximum storage capacity corresponds to around 0.31 mol of hydrogen (calculated for LmNi4.91Sn0.15) and additional 0.32 mol of hydrogen calculated for T0.99Zr0.01V0.43Fe0.09Cr0.05Mn1.5. Therefore, the exchangeable amount of hydrogen based on the van’t Hoff plots is reduced by around 0.63 mol for an ambient temperature increase from 20 C to 35 C. This value agrees with the value derived from the mass flow meter (MFM H2) shown in Fig. 10. The cooling capability of the sorption system depends therefore considerably on the ambient temperature as the reduced amount of exchangeable hydrogen represents the desorption heat of metal hydride A that is not available to generate the cooling effect.
4.
Summary
Based on the capillary tube bundle reaction bed, a metal hydride sorption system was built-up and experimentally investigated. Due to the large heat transfer surface of the reaction bed, the optimum half-cycle time of the system is in the order of 100e120 s. The influence of different simulated ambient and cooling temperatures on the hydrogen exchange and the cooling performance of the system were analyzed. Based on these measurements, the achievable specific cooling power of the system (related to the desorbing metal hydride) was deduced and is summarized in Fig. 13 for different temperature boundary conditions. Two main reasons for the reduced cooling power for higher ambient temperatures or lower cooling temperatures have been experimentally identified: 1) The amount of hydrogen necessary to cool down the passive mass (e.g. reaction bed cladding tube, capillary tubes) determines the available hydrogen to generate the useful cooling effect. Therefore, higher temperature differences between regeneration and cooling half-cycle reduce the achievable cooling power of the sorption system. 2) The total amount of exchangeable hydrogen is reduced for higher differences between ambient and cooling temperature.
Acknowledgment This work is financially supported by the European Commission, Brussels, in the frame of the Sixth Framework Program e Sustainable Surface Transport, EU-Contract TST4-CT-2005012471 ‘Thermally Operated Mobile Air-Conditioning Systems (Topmacs)’.
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references
[1] Isselhorst A, Groll M. Two-stage metal hydride heat transformer laboratory model. J Alloy Compd 1995;231:888e94. [2] Klein HP, Groll M. Development of a two-stage metal hydride system as topping cycle in cascading sorption systems for cold generation. Appl Thermal Eng 2002;22:631e9. [3] Ram Gopal M, Srinivasa Murthy S. Experiments on a metal hydride cooling system working with ZrMnFe/MmNi4.5Al0.5 pair. Int J Refrigeration 1999;22:137e49. [4] Willers E, Wanner M, Groll M. A multi-hydride thermal wave device for simultaneous heating and cooling. J Alloy Compd 1999;293e295:915e8. [5] Chernikov AS, Izhvanov LA, Solovey AI, Frolov VP, Shanin YI. An installation for water cooling based on a metal hydride heat pump. J Alloy Compd 2002;330-332:907e10. [6] Qin F, Chen J, Lu M, Chen Z, Zhou Y, Yang K. Development of a metal hydride refrigeration system as an exhaust gasdriven automobile air conditioner. Renew Energy 2007;32: 2034e52. [7] Ni J, Liu H. Experimental research on refrigeration characteristics of a metal hydride heat pump in auto airconditioning. Int J Hydrogen Energy 2007;32:2567e72.
[8] Kang BH, Park CW, Lee CS. Dynamic behavior of heat and hydrogen transfer in a metal hydride cooling system. Int J Hydrogen Energy 1996;21:769e74. [9] Groll M, Isselhorst A, Willers E, Dehouche Z. Evaluation of Concepts for Metal Hydride Air-Conditioning Systems for Cars, Proceedings of Ab-Sorption Heat Pump Conference 1996, Montreal, 785e793. [10] Willers E, Groll M. Evaluation of metal hydride machines for heat pumping and cooling applications. Int J Refrigeration 1999;22:47e58. [11] Dehouche Z, de Jong W, Willers E, Isselhorst A, Groll M. Modelling and simulation of heating/air conditioning systems using the multi-hydride-thermal-wave concept. Appl Thermal Eng 1998;18:457e80. [12] Paya´ J, Linder M, Laurien E, Corbera´n JM. Dynamic model and experimental results of a thermally driven metal hydride cooling system. Int J Hydrogen Energy 2009;34: 3173e84. [13] Friedlmeier G, Schaaf M, Groll M. How to measure pressureconcentration-isotherms representative for technical applications. Z Phys Chem 1994;183:185e95. [14] Muthukumar P, Linder M, Mertz R, Laurien E. Measurement of thermodynamic properties of some hydrogen absorbing alloys. Int J Hydrogen Energy 2009;34:1873e9.
Available at www.sciencedirect.com
journal homepage: www.elsevier.com/locate/he
Experimental results of a compact thermally driven cooling system based on metal hydrides Marc Linder*, Rainer Mertz, Eckart Laurien Institute of Nuclear Technology and Energy Systems, University of Stuttgart, Pfaffenwaldring 31, 70569 Stuttgart, Germany
article info
abstract
Article history:
In this paper a detailed experimental analysis of a metal hydride based cooling system is
Received 19 March 2010
presented. For the high temperature side an AB5 type alloy (LmNi4.91Sn0.15) was chosen,
Received in revised form
whereas an AB2 type alloy (Ti0.99Zr0.01V0.43Fe0.09Cr0.05Mn1.5) is used for the low temperature
22 April 2010
side. Due to very good heat and also mass transfer characteristics (among others, large heat
Accepted 23 April 2010
transfer surface area) of the utilized capillary tube bundle reaction bed, very short half-
Available online 3 June 2010
cycle times in the order of 100 s have been reached. Consequently, the specific cooling
Keywords:
temperature boundary conditions. The system was experimentally analyzed for different
Metal hydride
cooling and ambient temperatures, whereas the heating temperature was fixed to 130 C.
power of the system is up to 780 W per kg desorbing metal hydride e depending on the
Sorption system
ª 2010 Professor T. Nejat Veziroglu. Published by Elsevier Ltd. All rights reserved.
Heat pump Capillary tube bundle reaction bed
1.
Introduction
The thermally driven generation of cold seems to be a reasonable way to utilize waste heat to increase the overall efficiency of energy conversion systems. In case of automotive cooling, a thermally driven cooling system could supply the cooling demand using waste heat from the coolant or the exhaust gases. The substitution of mechanical energy (generated by the engine) by waste heat as main driving energy for the cooling system would consequently lead to reduced fuel consumption along with reduced CO2 emissions. Additionally, the climate damaging refrigerants of current automotive air-conditioning systems could be replaced. However, main obstacles for an automotive application of thermally driven sorption systems are their current weight and volume constraints. Due to the fast intrinsic reaction kinetics of metal hydrides, compact and light-weight sorption systems can be realized if limitations due to insufficient mass or heat transfer within the
reaction bed can be minimized. Therefore, this work focuses on the experimental investigation of a metal hydride sorption system using capillary tube bundle based reaction beds in order to increase the weight specific power of the system.
1.1.
Working principle
The basic configuration of metal hydride sorption systems consists of two reaction beds coupled by a hydrogen connection pipe. Each reaction bed contains a different metal hydride, indicated as A and B (Fig. 1). As the equilibrium pressure of metal hydride A is higher than the equilibrium pressure of metal hydride B at the same temperature, metal hydride A tends to release hydrogen and cools down. If the coupled metal hydride B is able to absorb the released hydrogen, a hydrogen exchange along with the corresponding cooling effect is obtained (cooling half-cycle). The absorption of hydrogen in the coupled low-pressure hydride B generates heat (QAmb_cool) that must be released to a heat sink to prevent
* Corresponding author. E-mail address: [email protected] (M. Linder). 0360-3199/$ e see front matter ª 2010 Professor T. Nejat Veziroglu. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2010.04.184
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Fig. 1 e Van’t Hoff diagram of sorption cooling system.
a temperature increase of the alloy. As the environment is generally used as heat sink its temperature (TAmb) determines the characteristic of the coupled alloy B. As soon as the maximum amount of hydrogen is exchanged, the flow direction must be reversed. Therefore, a thermal energy input (QHeat) is necessary to supply the desorption enthalpy of metal hydride B. As the absorption heat during regeneration (QAmb_reg) is released to the environment, the absorption pressure of metal hydride A at ambient temperature determines the necessary regeneration pressure level (pReg). The described batch system is not able to supply a continuous cooling effect. As the reaction bed pairs working either in regeneration or cooling mode, two identical reaction bed couples working in opposite direction are necessary to realize a quasi-continuous cold output. Therefore, the necessary time for regeneration must be shorter or at least identical to the cooling half-cycle time.
1.2.
inner artery was intensively used to investigate different sorption system prototypes [1e5]. The second design for hydrogen distribution was recently investigated by Qin et al. [6] and Ni et al. [7] in 2007. Instead of a central filter tube, the filter surrounds the metal powder bulk. Therefore, hydrogen enters the reaction bed from one side, is distributed in the annular gap around the filter and enters in radial direction into the powder bulk. A finned copper tube is inserted in the bulk to divide the bed into several sections and to improve the heat transfer characteristics. Kang et al. [8] used a comparable design: The reaction bed consisted of two identical copper tubes with aluminum fins working serially in an U-shaped heat exchanger and hydrogen distribution was realized by a gap between the metal hydride part of the reaction bed and its pressure vessel. Although, heat recovery measures of the systems and different working temperatures complicate a direct comparison of the systems, the relevant weight-specific characteristics for coupled metal hydride reaction beds realized within the last 15 years can be summarized as follows: The half-cycle time ranges between 10 and 20 min, depending on the applied reaction bed design. The maximum achieved specific power is around 140e170 W per kg desorbing alloy of the respective reaction bed. As most of the intended applications were stationary, the necessity to increase the specific power by means of heat transfer enhancements was mainly based on the reduction of metal powder and not implicitly on weight and volume optimizations. However, taking into account weight and volume constraints of automotive applications, the need to increase the weight specific performance of the metal hydride sorption technology is obvious. Therefore, this work focuses on the experimental analysis of a sorption system using a capillary tube bundle based reaction bed with a large heat transfer surface.
State of the art
In order to compare different designs of metal hydride sorption systems regarding their weight specific performance, two characteristic values can be deduced from published data, the specific power (per kg desorbing alloy in the respective reaction bed) and the half-cycle time. Although the systems may be designed for different applications, like combined heat/ cold generation or heat transforming, for each system the same principle applies: thermally driven hydrogen exchange between two reaction beds. The limiting factors (quantified by specific power and half-cycle time) are therefore independent of the designed application, the total number of reaction beds or the amount of different metal hydrides. In general, metal hydride reaction beds for sorption systems are of cylindrical shape with two different possible hydrogen distribution designs: Hydrogen enters the metal hydride bulk either through an inner filter artery or through a surrounding filter tube. Both cases enable an easy hydrogen distribution along the axial direction, whereas the radial hydrogen distribution characteristic depends on the consistency of the bulk material. The reaction bed design with an
Fig. 2 e Schematic flow diagram of test bench for coupled reaction beds.
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Table 1 e Measurement equipment of experimental set-up for coupled reaction beds. Transducer
Measurement range
Resistance thermometers Pt 100 Absolute pressure transducer Keller Druck, PA 23 Mass flow meter (Hydrogen) Hastings HFM-201 Mass flow meter (Water) Krohne.IFS6000F
Accuracy
Max. Deviation
100 to þ 450 C 0 to 50 bar
(0.15 þ 0.002*T) 0.5% F. S.
0.55 K (<200 C) 0.25 bar
0 to 100 sl/mina
1.3%F. S.
1.3 sl/min
0 to 14.97 1/min
0.3%F.S.
0.045 l/min
AGILENT Data Acquisition. Switch Unit (34 970 A). a standard liter, for T ¼ 0 C and p ¼ 1.013 bar.
2.
Experimental details
2.1.
Experimental set-up
The test bench shown in Fig. 2 consists of a hydrogen part (blue), the two reaction beds (A and B including reaction bed valves V1 and V2) and two separate water cycles (fully drawn and dashed lines, respectively). The main component of the hydrogen part is the mass flow meter implemented in the hydrogen connection pipe. Additionally, a pneumatically actuated hydrogen valve (V3) allows the separation of the reaction beds. Each water cycle consists of a pump, a mass flow meter for water, two resistance thermometers (T1-T4), two electromagnetic 3-way valves (EMV1-EMV4) and two heat exchangers. Additionally, each reaction bed contains an internal heat exchanger that is connected to the water cycle in order to transfer heat from the water current to the metal hydride during desorption or vice versa during absorption. The connected thermostats and electrical heaters, not shown in Fig. 2, control the set-point temperatures of the heat exchangers. During regeneration mode, the necessary driving heat (QHeat) is supplied by the left heat exchanger in the dashed water cycle, whereas the left heat exchanger in the full-drawn cycle rejects heat (QAmb_reg) at simulated ambient temperature. The switching of the electro-magnetic valves leads to a temperature (and correspondingly pressure) change of the reaction beds. During the following cooling half-cycle, the heat input at cooling temperature (QCool) and heat rejection at ambient temperature (QAmb_cool) are realized by the second heat exchangers in the respective water cycles. As pure water is used as heat carrier fluid, the hot water cycle is pressurized to around 4 bar to prevent boiling at temperatures of around 130 C. For a given half-cycle time (time between switching of the electro-magnetic valves) and adjusted heat exchanger
Fig. 3 e Capillary tube bundle reaction bed (length: 135 mm, diameter (max.): 65 mm).
temperature levels, the system is working automatically. A software program based on HP VEE Pro (Agilent) realizes the controlling of the test bench. The description of the implemented measurement equipment is summarized in Table 1.
2.2.
Reaction beds
The chosen capillary tube bundle reaction bed design has already been mentioned in publications [9e11] but up to now without detailed experimental results. Recently, a “zero-dimensional” model was developed and published based on selected experimental results obtained from the system described in this work [12]. The main part of the reaction bed is the capillary tube bundle heat exchanger shown in Fig. 3, left. It consists of 372 small tubes with an inner diameter of 1.4 mm made of stainless steel. The heat carrier flows through the tubes that are surrounded by the metal powder. To prevent discharge of the metal powder, a sintered metal filter tube surrounds the tube bundle (Fig. 3, middle). A cladding tube (Fig. 3, right) encloses the reaction bed and forms an annular gap around the filter tube that is used for uniform hydrogen distribution around the metal hydride bed. The total length of the assembled reaction bed is 135 mm. With an outer diameter of 76 mm the total volume of the reaction bed is around 0.6 l. Its weight without metal powder is 1.7 kg. The volume within the sintered metal tube (Vi z 0.23 l) is sufficient to charge 800e1000 g of powder, depending on the density of the alloy and its volume expansion during hydriding.
Table 2 e Properties of applied metal hydrides. Alloy LmNi4 91Sn0.15 Reaction enthalpy (948 g in reaction [J/molH2] Reaction entropy bed B) [J/molH2K] Plateau slope [-] a Ti0.99Zr0.01V0.43Fe0.09 Reaction enthalpy Cr0.05Mn1.5 [J/molH2] Reaction entropy (800 g in [J/molH2K] reaction bed A) Plateau slope [-] a
Absorption Desorption 27 411
29 849
103.9
109.1
0.43 17 922
0.48 22 606
89.2
101.1
0.42
0.22
a slope of PCI plateau calculated for mid of plateau region.
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Metal hydride working pair
Based on a detailed experimental analysis of available metal hydrides, LmNi4.91Sn0.15 was chosen for the high temperature side and Ti0.99Zr0.01V0.43Fe0.09Cr0.05Mn1.5 for the generation of the cooling effect. Respective PCI measurements were performed on a volumetrically measuring set-up (described in [13,14]). Based on the obtained PCI data, the characteristic properties of the respective alloys were deduced and are summarized in Table 2. Additionally, the respective van’t Hoff plots are shown in Fig. 4.
3.
Experimental results
The set-up consists of the reaction bed couple and four heat exchangers at three different temperature levels (THeat, TAmb and TCool). These temperature levels correspond to the respective simulated boundary conditions of the cooling system. However, the temperature sensors T1-T4 are close to the reaction beds and measure the respective in- and outlet temperature of the water current. In Fig. 5 these in- and outlet temperatures (T1-T4) are shown for 4 complete cycles. The temperatures of the heat exchangers were adjusted to THeat ¼ 130 C, TAmb ¼ 28 C and TCool ¼ 20 C. The blue lines correspond to the cold reaction bed A which is cycled between ambient and cooling temperature. As the temperatures of the reaction bed and the metal hydride are around ambient temperature, the reaction enthalpy at the beginning of the cooling cycle is used to cool down the thermal masses of the reaction bed. Therefore, the generated cooling effect starts slightly later than the actual changing of the half-cycle. After this short period, the endothermic reaction cools the water flow. This heat capacity effect is even more obvious for the hot reaction bed (red). Due to the higher temperature difference between cooling and regeneration
cycle, the necessary amount of thermal energy to reach the temperature level of the following half-cycle is higher. Therefore, the thermal masses, especially the parasitic thermal masses like pressure vessel, heat exchanger surface or filter tube should be minimized in order to maximize the performance of the system. In Fig. 6 the corresponding hydrogen pressure and the exchanged amount of hydrogen during the four cycles are shown. The fully drawn line represents the accumulated amount of exchanged hydrogen and is shown on the left y-axis, whereas the right y-axis gives the values of the pressure progression drawn as dashed line. The total amount of exchanged hydrogen (nH2) is around 3.6 mol. Based on the desorption enthalpy of the metal hydride in reaction bed A, the corresponding amount of thermal energy is around 81 kJ. In order to be able to open the reaction bed and to change the alloys, two massive brass constructions for the distribution of water at the in- and outlet are necessary. Whereas the alloy mass is around 900 g and the reaction bed itself weighs around 1.7 kg, the massive brass in- and oulets weight together additional 2 kg. These components are not necessary for possible applications but reduce the efficiency of the test bench due to the almost doubled parasitic mass of the reaction bed. Their impact is shown in Fig. 7, where the thermal power (measured in the water flow, fully drawn line) is compared with the calculated “hydrogen power” (dotted line). “Hydrogen power” is calculated by the measured mass flow rate of hydrogen and the respective reaction enthalpy of the metal hydride. It is the maximal achievable value (without thermal losses) as a temperature change of thermal masses also leads to a hydrogen exchange. Positive values correspond to the cooling half-cycle, whereas negative values represent the regeneration cycle. Although the temperature conditions are favourable (small difference between TCool and TAmb), the effective cooling power measured in the water flow is clearly lower than the theoretically achievable “hydrogen power” (without thermal losses).
Fig. 4 e Van’t Hoff plots of applied metal hydrides.
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Fig. 5 e In- and outlet temperature progression during cycling experiment.
3.1.
Variation of cooling temperature
The adjusted cooling temperature influences the behavior of the coupled reaction beds. In Fig. 8, this influence on the thermal power (PWater A, fully drawn line) and the calculated “hydrogen power” is shown (PH2,A, dashed line). As the heating energy is not changed and the cooling temperature influences the cooling half-cycle only, the regeneration half-cycles (negative values) are comparable. The main difference during
the cooling half-cycle is the higher measured thermal power for a cooling temperature of 20 C in comparison to the lower cooling temperatures of 15 C and 10 C. The reason is due to the thermal masses described earlier. With an increasing temperature difference during regeneration and cooling cycle, the fraction of exchanged hydrogen necessary to change the temperature of the reaction bed itself increases. Therefore, less hydrogen is available to cool to water flow. Additionally, the total amount of exchanged hydrogen (nH2) decreases with
Fig. 6 e Pressure and hydrogen exchange during cycling experiments.
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Fig. 7 e Cooling power based on hydrogen exchange rate and measured in the water flow.
Fig. 8 e Variation of the cooling temperature e effect on thermal and calc. power.
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Fig. 9 e Variation of the cooling temperature e effect on total hydrogen exchange.
decreasing cooling temperature (see Fig. 9, fully drawn lines, left y-axis). The comparably small difference in the measured thermal power between 15 C and 10 C cooling temperature is due to the instable inlet temperature during the cooling cycle. Additionally, the set point of 10 C is not reached and therefore the difference between adjusted 10 C and 15 C cooling temperature is reduced (Fig. 9, dotted lines, right y-axis). Both effects are due to the applied thermostat that
is not able to balance properly the high heat loads during regeneration and cooling mode for high temperature differences.
3.2.
Variation of ambient temperature
The ambient temperature (TAmb) is the most critical boundary condition for thermally driven sorption systems as it affects the regeneration as well as the cooling half-cycle. An
Fig. 10 e Hydrogen exchange and pressure development for different ambient temperatures.
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Fig. 11 e Pressure progression during the regeneration half-cycle for different ambient temperatures e absorption in reaction bed A (T3 [ 130 C).
increased ambient temperature increases the corresponding equilibrium pressure of the absorbing metal hydride and consequently the system pressure p2 during both half-cycles (see dashed lines in Fig. 10). The half-cycle time of the sorption system was set to 120 s. As the heating and the cooling temperature were fixed to THeat ¼ 130 C and TCool ¼ 20 C, respectively, the available desorption pressures are identical for all measurements. Therefore, the exchanged amount of hydrogen nH2 shown in Fig. 10 as fully drawn lines depends
clearly on the ambient temperature (TAmb). Around 4 mol of hydrogen are exchanged for an ambient temperature of 20 C, whereas only 3.8 and 3.4 mol are exchanged for TAmb ¼ 28 C and 35 C, respectively. In Fig. 11 and Fig. 12, the respective system pressure progressions (p2) are shown for one half-cycle and compared to the corresponding van’t Hoff plots of the absorbing alloy (non-logarithmic scale). Although, the desorbing alloy generates the thermally driven compression of hydrogen
Fig. 12 e Pressure progression during the cooling half-cycle for different ambient temperatures e absorption in reaction bed B (T1 [ 20 C).
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Fig. 13 e Achievable specific cooling power (related to the desorbing metal hydride).
during the regeneration half-cycle, the system pressure p2 at the end of the regeneration phase corresponds to the absorption pressure of metal hydride A with a load situation of around 75% of the maximum storage capacity (see nonlogarithmic van’t Hoff plots in Fig. 11). However, the load situation of metal hydride A at the end of the regeneration half-cycle decreases with increasing ambient temperature TAmb. Whereas the metal hydride is charged to more than 75% of the maximum hydrogen storage capacity for a water inlet temperature of T1 ¼ 20 C, the load situation is reduced to less than 70% at the end of the regeneration half-cycle for T1 ¼ 35 C (even for increased system pressure, about 47 bar vs. about 37 bar). Fig. 12 displays the progression of the system pressure p2 during the respective cooling half-cycle (T1 ¼ 20 C). A comparison with the van’t Hoff plot absorption lines of metal hydride B shows a similar dependency of the hydrogen load situation on the ambient temperature. The metal hydride B is charged to less than 85% for T3 ¼ 35 C, whereas a load situation of almost 90% of the maximum storage capacity is reached for a water inlet temperature of 20 C. Taking the used metal hydride masses and the maximum storage capacity of the respective alloys into account, 5% of the maximum storage capacity corresponds to around 0.31 mol of hydrogen (calculated for LmNi4.91Sn0.15) and additional 0.32 mol of hydrogen calculated for T0.99Zr0.01V0.43Fe0.09Cr0.05Mn1.5. Therefore, the exchangeable amount of hydrogen based on the van’t Hoff plots is reduced by around 0.63 mol for an ambient temperature increase from 20 C to 35 C. This value agrees with the value derived from the mass flow meter (MFM H2) shown in Fig. 10. The cooling capability of the sorption system depends therefore considerably on the ambient temperature as the reduced amount of exchangeable hydrogen represents the desorption heat of metal hydride A that is not available to generate the cooling effect.
4.
Summary
Based on the capillary tube bundle reaction bed, a metal hydride sorption system was built-up and experimentally investigated. Due to the large heat transfer surface of the reaction bed, the optimum half-cycle time of the system is in the order of 100e120 s. The influence of different simulated ambient and cooling temperatures on the hydrogen exchange and the cooling performance of the system were analyzed. Based on these measurements, the achievable specific cooling power of the system (related to the desorbing metal hydride) was deduced and is summarized in Fig. 13 for different temperature boundary conditions. Two main reasons for the reduced cooling power for higher ambient temperatures or lower cooling temperatures have been experimentally identified: 1) The amount of hydrogen necessary to cool down the passive mass (e.g. reaction bed cladding tube, capillary tubes) determines the available hydrogen to generate the useful cooling effect. Therefore, higher temperature differences between regeneration and cooling half-cycle reduce the achievable cooling power of the sorption system. 2) The total amount of exchangeable hydrogen is reduced for higher differences between ambient and cooling temperature.
Acknowledgment This work is financially supported by the European Commission, Brussels, in the frame of the Sixth Framework Program e Sustainable Surface Transport, EU-Contract TST4-CT-2005012471 ‘Thermally Operated Mobile Air-Conditioning Systems (Topmacs)’.
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references
[1] Isselhorst A, Groll M. Two-stage metal hydride heat transformer laboratory model. J Alloy Compd 1995;231:888e94. [2] Klein HP, Groll M. Development of a two-stage metal hydride system as topping cycle in cascading sorption systems for cold generation. Appl Thermal Eng 2002;22:631e9. [3] Ram Gopal M, Srinivasa Murthy S. Experiments on a metal hydride cooling system working with ZrMnFe/MmNi4.5Al0.5 pair. Int J Refrigeration 1999;22:137e49. [4] Willers E, Wanner M, Groll M. A multi-hydride thermal wave device for simultaneous heating and cooling. J Alloy Compd 1999;293e295:915e8. [5] Chernikov AS, Izhvanov LA, Solovey AI, Frolov VP, Shanin YI. An installation for water cooling based on a metal hydride heat pump. J Alloy Compd 2002;330-332:907e10. [6] Qin F, Chen J, Lu M, Chen Z, Zhou Y, Yang K. Development of a metal hydride refrigeration system as an exhaust gasdriven automobile air conditioner. Renew Energy 2007;32: 2034e52. [7] Ni J, Liu H. Experimental research on refrigeration characteristics of a metal hydride heat pump in auto airconditioning. Int J Hydrogen Energy 2007;32:2567e72.
[8] Kang BH, Park CW, Lee CS. Dynamic behavior of heat and hydrogen transfer in a metal hydride cooling system. Int J Hydrogen Energy 1996;21:769e74. [9] Groll M, Isselhorst A, Willers E, Dehouche Z. Evaluation of Concepts for Metal Hydride Air-Conditioning Systems for Cars, Proceedings of Ab-Sorption Heat Pump Conference 1996, Montreal, 785e793. [10] Willers E, Groll M. Evaluation of metal hydride machines for heat pumping and cooling applications. Int J Refrigeration 1999;22:47e58. [11] Dehouche Z, de Jong W, Willers E, Isselhorst A, Groll M. Modelling and simulation of heating/air conditioning systems using the multi-hydride-thermal-wave concept. Appl Thermal Eng 1998;18:457e80. [12] Paya´ J, Linder M, Laurien E, Corbera´n JM. Dynamic model and experimental results of a thermally driven metal hydride cooling system. Int J Hydrogen Energy 2009;34: 3173e84. [13] Friedlmeier G, Schaaf M, Groll M. How to measure pressureconcentration-isotherms representative for technical applications. Z Phys Chem 1994;183:185e95. [14] Muthukumar P, Linder M, Mertz R, Laurien E. Measurement of thermodynamic properties of some hydrogen absorbing alloys. Int J Hydrogen Energy 2009;34:1873e9.