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A critical pathway energy efficiency analysis of the thermochemical UT-3 cycle
International Journal of Hydrogen Energy 30 (2005) 559 – 564 www.elsevier.com/locate/ijhydene
A critical pathway energy efficiency analysis of the thermochemical UT-3 cycle E.D. Teoa , N.P. Brandona , E. Vosb , G.J. Kramerb,∗ a Department of Chemical Engineering, Imperial College, London SW7 2AZ, UK b Shell Research and Technology Centre, OGIR, Shell Global Solutions International B.V., P.O. Box 38000, Amsterdam 1030 BN, Netherlands
Available online 21 September 2004
Abstract Numerous thermochemical cycles for the production of hydrogen have been proposed in the literature, and of these the UT3 cycle is a leading exponent. The primary objective of such processes is to be more energy efficient than water electrolysis in converting heat to hydrogen. This paper presents a critical assessment of the actual energy efficiency that could be realised in a thermochemical cycle, taking the UT-3 cycle as the basis for the study. It is concluded that the upper efficiency for this process lies around 12%, on the basis of the lower heating value of the hydrogen product, comparable to current hydrogen production techniques using photovoltaics followed by water electrolysis. The practical upper efficiency may however be lower, even much lower, as several uncertainties exist in key aspects of the process. This in-depth analysis of one thermochemical process brings to the fore issues of whether such cycles can—on the basis of efficiency—compete with electricity generation followed by electrolysis. In view of the renewed interest in thermochemical cycles, analyses such as the one presented in this paper must be carried out and published in the open literature, so that the wider research community can realistically assess the prospects of this intriguing route to hydrogen production. 䉷 2004 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. Keywords: Thermochemical; UT-3 cycle
1. Introduction Numerous thermochemical cycles have been proposed since the publication of the first article on the energy requirements of the water decomposition process by Funk and Reinstrom [1]. The primary objective of such processes is to offer a sustainable route to hydrogen, which is more energy efficient than using photovoltaics (PV) coupled with water electrolysis. Recent developments have focused on the UT-3 process and the iodine–sulphur cycle developed by General Atomics, identified as the two leading candidates at the First Information Exchange Meeting on the Nuclear Production ∗ Corresponding author. Tel.: +31-20-630-2467; fax: +31-20-
630-3964. E-mail address: [email protected] (G.J. Kramer).
of Hydrogen of the Organization for Economic Cooperation and Development in October 2001 [2]. The UT-3 (University of Tokyo, Br, Ca and Fe) process was proposed by Kameyama and Yoshida [3], with the following chemical reactions: reactor 1 :
CaBr 2 (s) + H2 O(g) 973–1023 K
−→
reactor 2 :
(1)
CaO(s) + Br 2 (g) 773–873 K
−→
reactor 3 :
CaO(s) + 2HBr(g),
CaBr 2 (s) + 0.5O2 (g),
(2)
Fe3 O4 (s) + 8HBr(g) 473–573 K
−→
3FeBr 2 (s) + Br 2 (g)
+ 4H2 O(g),
0360-3199/$30.00 䉷 2004 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2004.08.003
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E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
individual heat inputs and the second for the work of each unit operation divided by its associated efficiency [17] the total thermal energy input to the process including the equivalent thermal input of the required electrical energy
Nomenclature UT−3 Qe QH2
Q
reactor 4 :
efficiency of the UT-3 process excess (thermal) energy production heating value of product hydrogen, based on lower heating value (241 kJ/mol) the the heat inputs; specifically sumof all Q = Qi + Wi /i , the first term for the
3 FeBr 2 (s) + 4H2 O(g) 833–873 K −→ Fe3 O4 (s) + H2 (g) + 6HBr(g).
Qi
ΣQi
(4)
The flow diagram of the process is reproduced in Fig. 1. The Model Apparatus for the Study of Cyclic Operation in Tokyo (MASCOT) pilot plant was operated in the late 1980s [4] and a potential energy efficiency of 49.5% (on basis of the higher heating value—HHV—of the product hydrogen) has been suggested [5,6]. The primary heat source could be supplied by a nuclear reactor or by using solar heat. The final energy carrier would be dependent on the location. The terrestrial PV market includes the consumer production, utility generation and remote power generation segments. The remote power generation segment of stand-alone systems has the most rapid growth and contribution to PV sales [7]. This is also the sector where thermochemical cycles would expect to compete, and hydrogen could be the final energy carrier instead of electricity. In this study, the solar to hydrogen efficiency of the UT-3 cycle is compared against the equivalent efficiency of the electrolysis of water with PV-generated electricity. We note that in this field it seems customary to quote efficiencies on the basis of the HHV of hydrogen, which is 286 kJ/mol. Most papers that discuss the use of hydrogen however have a preference to base discussion on lower heating value (LHV), which for hydrogen is 241 kJ/mol. In calculating pathway efficiencies, this may easily lead to confusion and error. In our paper we will standardise on LHV; exceptions, e.g. when discussing literature data, will be explicitly mentioned.
2. Methodology For this study, two “pathway efficiencies” from sunlight to hydrogen will be considered: Route 1: For PV and water electrolysis: Sunlight a → Electricity b → Hydrogen. Route 2: For thermochemical cycles: Sunlight c → Central Receiver System d → Thermal Storage System e → Thermochemical Process producing Hydrogen, where a is the efficiency of PV, defined as the ratio of electricity produced to the amount of sunlight incident on
1033 K 957 K Reactor 1 HX01 861 K
H2O
833 K 725 K Reactor 4
HX1 H2
HX4 O2
HX2
HX3 Reactor 2 864 K
Reactor 3 845 K
576 K
493 K
Qe
Fig. 1. Simplified flow-scheme of the UT-3 cycle.
the PV device [7]. Efficiencies of 11% are attainable using thin-film copper indium diselinide (CIS) solar modules and 14–15% for mono- and multi-crystalline modules, b is the efficiency of water electrolysis, typically ranging from 60% to 70% on LHV basis (equivalent to 70–85% on HHV basis) [8,9]. Electrolysis for hydrogen production is generally used to meet the demand in isolated markets or desert regions where other energy sources are not in abundance. For water electrolysers, the energy efficiency is the ratio of the equilibrium cell voltage to the actual voltage required, taking into account the overpotentials at the electrodes and the ohmic potential drops. c,d is the combined efficiency of the solar furnace/central receiver system (CNS) (29–56%) and the thermal energy storage (98.8%) [10], and e is the reported efficiency of the UT-3 cycle of 49.5% on HHV basis [6], equivalent to 42% on LHV basis. As seen from Fig. 2a, the thermochemical cycle route for hydrogen production (Route 2) is up to three times more efficient than current technologies (Route 1) if hydrogen were the final energy carrier. Note if we use the high-temperature heat to generate electricity with 40–45% efficiency, using high-pressure steam Rankine cycle, this route (indicated by the dashed line in Fig. 2a) would have a 13–25% efficiency, up to two-and-a-half times that of PV.1 If the electric power thus generated is used to produce hydrogen by electrolysis, another relevant comparison results (Fig. 2b). This case, where one starts with high-temperature heat and compares conventional electricity generation 1 Note that the upper efficiency limit corresponds to the as-
sessment by Hamache and Bilgen (1992).
E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
of the UT-3 process [5], UT−3 as
Solar Energy Energy Capture by Photovoltaic cells 11 – 15 % Route 1 7 – 11 %
Electricity
Solar Furnace/ Central Receiver 29 – 55 % High-temperature heat
Electrolysis 60 – 70 %
UT−3 = [QH2 + Qe ]/ Route 2 12 – 23 %
Thermochemical cycle 42 % (suggested)
Hydrogen (LHV)
(a)
High-temperature heat (800oC) High-pressure Rankine cycle 40 – 45% Route 1’ 24 – 32 %
Thermochemical cycle 42 %(suggested)
Electricity
Electrolysis 60 – 70 %
(b)
561
Hydrogen (LHV)
Fig. 2. Pathway efficiencies from solar energy to hydrogen (a) and pathway efficiencies from high-temperature heat to hydrogen (b).
followed by electrolysis (Route 1 ) with the thermochemical route, is the relevant comparison when considering hydrogen production from nuclear heat. Here, the notional 42% efficiency of the UT-3 cycle is about one-and-a-half times as efficient as the conventional Route 1 (24–32%). In both modes of comparison and notwithstanding future developments in CNS and water electrolysis, the attractiveness of the thermochemical route lies in its (potentially) higher process efficiency, which requires studying in greater detail. Cost is often viewed as the only “figure of merit” [11] for the competitiveness of a process, with steam reforming as the dominant route for hydrogen production at 5.22 US$/GJ [2]. Production cost of ‘renewable’ hydrogen via PV and electrolysis is estimated between 50 and 100 US$/GJ [12]. It would be premature to calculate the cost of hydrogen per GJ for the thermochemical cycle given the multitude of assumptions for plant sizing and price quotes that are organisational dependent. In the absence of this, process efficiency is taken to be the best indicator for the prospects of the technology. It is proposed that energy analysis will suffice at this juncture, as the difference in using exergy is a mere 3.4% for the UT-3 process [6]. Exergy analysis could be carried out if the calculated energy efficiency values are significantly more attractive than the reference technology chosen. Due to possible excess thermal energy for power generation, researchers of the UT-3 process defined the efficiency
Qi ,
(5)
the product where QH2 is the (LHV) energy content of hydrogen, Qe is the excess thermal heat and Qi is the heat input. In the original paper, QH2 referred to the HHV of hydrogen. In line with the remarks in the introduction we will use the LHV. In this study the upper bound of the UT-3 process efficiency is revised downwards from the theoretical Carnot value of 55% [2] (LHV basis; the reference quotes 66% on HHV basis) and the 42% (LHV) obtained by the UT-3 researchers [6] by the inclusion of increasingly stringent but practical constraints. The effects on process efficiency of incomplete reactions and separation via steam condensation used in the MASCOT pilot plant are then estimated, with the aid of modelling tools such as Aspen Plus䉸 and HSC Chemistry䉸. The efficiencies of the CNS and thermal storage are subsequently factored to derive the solar pathway efficiencies (i.e. the efficiency of H2 production using a CNS and thermal storage coupled to the UT-3 process) for the production of H2 gas, which are then compared to that of PV and water electrolysis. The analyses of the separations, heat integration, compressor work inputs were based on the stream table given by Sakurai et al. [6].
3. Results 3.1. Upper bound of the UT-3 process efficiency The published process efficiency of the UT-3 process, UT−3 as given in Eq. (5), is 42.1% when expressed on LHV basis. In the case worked out by Sakurai et al., the hydrogen product (25 mol/s) energy content is equivalent to 6.03 MW (QH2 ) and results from an input 15.24 MW of high-temperature heat. This is actually 39.5% efficiency. The remaining 2.6% in Sakurai’s calculation is 0.40 MW of low-temperature (550 K) heat. As there is no use for that heat in the process, its inclusion in the efficiency figure is misleading. The best use is to convert the heat to electricity. For this, Sakurai et al. suggest 17% efficiency, resulting in 0.07 MW of electricity. This could be converted to hydrogen through electrolysis at 70% efficiency, adding 0.05 MW to the product hydrogen stream, resulting in a maximum hydrogen production efficiency of 39.9%. Our independent analysis of the heat streams in the process yielded data that were not significantly different from to those of Sakurai et al. reflecting primarily the uncertainty in the thermodynamic basic data involved. For convenience we will use the heat data of Sakurai et al. in what follows. From the pinch diagramin Fig. 3 it is seen that the hightemperature input heat ( Qi , 15.24 MW) is fully used to heat the process stream (2500 mol/s of recycle water plus
E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
3.2. Effect of incomplete equilibrium conversions While 100% conversion for each reaction has been previously assumed [6], kinetic data indicate a maximum conversion based on bromination of ∼75% for the first two reactions (1) and (2) [13]. HSC䉸 Chemistry was used to predict the thermodynamic equilibrium conversions. This modelling helped to identify the critical reaction steps amongst the four reactions of the UT-3 cycle that require further verification. It was found that the H2 production step (4) has the greatest impact on the process efficiency while incomplete conversions in the other reactions mainly affect the feasibility of flow reversal to regenerate the reactants. In the conceptual design [6] 2500 mol H2 O and 50 mol HBr are fed to the reactor at 833 K and 20 bar. At the reactor exit, at 726 K, 100 mol of water is notionally consumed with 75 mol of solid Fe3 O4 to produce 25 mol of hydrogen and an additional 150 mol of HBr. Such a gas phase composition is however inconsistent with thermodynamic equilibrium. The equilibrium partial pressures of H2 , HBr and H2 O for this reactor are plotted in Fig. 4. It is seen that the hydrogen concentration according to the conceptual design can only be approached when the temperature approaches 1000 K. Such an operating temperature is not feasible however since FeBr2 will sublimate above 850 K. Thus, under practical circumstances—where one would operate the reactor isothermally at a temperature slightly below 850 K, the hydrogen concentration is an order of magnitude below that of the design, leading to ten-fold increase in water circulation. Since the high-temperature heat input to the cycle (Figs. 1 and 3) is essentially determined by the heat duty of
1100 1000 HX1
900
HX01
800 T (K)
25 mol/s of water for thermolysis) from 861 to 1033 K, the temperature at which it enters reactor 1. As suggested by Sakurai, the low-temperature waste heat from HX2 (7.53 MW) can be used to generate power to drive compressors and pumps. At 17% efficiency for the conversion of 550 K heat to power—a figure that we will use here as well—1.35 MW of power can be generated, in excess of the 1.27 MW required. In their assessment however, Sakurai et al. used 100% efficient compressors. When a practical compressor efficiency of 72% was incorporated to calculate the actual work input to the three compressors of the UT-3 cycle we find that 1.75 MW is required to drive the compressors and the water pump. Since only 1.35 MW was available from co-generation from waste heat, 0.40 MW of additional power has to be generated. Assuming 45% efficiency for electricity from high-temperature heat—as suggested above (Section 2)—an additional thermal energy input equivalent to 0.91 MW. At this level of assessment we find that 25 mol/s hydrogen, equivalent to 6.03 MW can be generated from an input 16.15 MW heat, lowering UT−3 to 37.3% (LHV). This should still be considered as an upper bound of the UT-3 process efficiency, which is subject to further reductions due to considerations of the following process uncertainties.
Qe
700
HX2
HX3
ΣQ i
600 500 400
HX4
300 0
10
20
30
40
50
heat flux (MW)
Fig. 3. Pinch diagram of the UT-3 cycle, based on the data of Sakurai et al. [6].
102 H2O
101 partial pressure (bar)
562
100
HBr
10-1
H2
10-2 10-3
726 K 833 K
10-4 10-5 700
800
900
1000
T (K)
Fig. 4. Thermodynamic equilibrium of the hydrogen production reaction (4). While Sakurai et al. [6] used full conversion—indicated by the dashed line—actual equilibrium conversion levels in the operating range (726–833 K) are much lower, making increased water circulation necessary.
the recycle stream from 861 to 1033 K, the input heat requirement would increase 10-fold and the efficiency would drop by the same factor. This is somewhat compensated by the fact that the adiabatic temperature drop in reactor 1 would be 10-fold reduced. Thus, keeping the reactor outlet at 957 K, the inlet temperature would have to be only 965 K, when the increased water flux would need to be raised in HX01 from 861 to 965 K only. Taking this into account, the heat input increases six-fold, rather than 10-fold to 92 MW. This would reduce the hydrogen production efficiency to a value considerably below 10%.2 2 The actual value is 6.6%. However, some additional hydrogen may be generated from the low-temperature heat which is now present in excess.
E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
3.3. Effect of separations via condensation on the process efficiency
Reaction 2
563 Reaction 3
Qe
HX1
O2 HX4
The gaseous HBr and H2 effluent from reactor 4 could be separated using a suitable membrane or bubbled through water for the dissolution of HBr to form hydrobromic acid [14]; the latter route was utilised in the MASCOT plant [15]. In theory, the separation is path-independent and there would be no effect on the process efficiency assuming there were no irreversibilities, and that the heat loss in lowering the inlet stream to the hydrogen separator, i.e. from 742 to the 303 K, the temperature at which hydrogen is separated, is completely recoverable to reheat H2 O (g) and HBr (g) to the expected outlet temperature of the separator at 742 K. In practice however, both membrane separation and condensation do imply the input of additional work or heat. If membrane separation of hydrogen would prove possible, the permeate hydrogen would come out of the membrane at low pressure. We saw in the preceding section that the hydrogen partial pressure in the effluent of reactor 4 is about 10 mbar. The likely permeate pressure will be an order of magnitude lower, so that hydrogen becomes available at 1 mbar. Recompression to 1 bar (of 25 mol H2 per second) would require about 1 MW of power.3 Assuming, as before, that this is generated from high-temperature heat at 45% efficiency, this adds 2.2 MW of thermal power to the process inputs. If membrane separation were not a viable option in the future and the practical process would require condensation as in the pilot plant, this would be an unbearable burden. We saw above that one must expect at least 25,000 mol of water per second to be condensed and reheated to achieve 25 mol/s of hydrogen production. A fantastic 1.3 GW of heat would have to be exchanged between condensing and evaporating water streams. The inevitable losses would render the process both uneconomic and highly inefficient. Thus, the UT-3 process is only viable if high-temperature membrane separation of hydrogen in the presence of HBr proves possible. Even then, an additional 2.2 MW would be required to drive the hydrogen recompression, reducing the efficiency to 25.4%.
efficiency at steady state. However, additional heat may well be necessary to reduce the time required to reach steady state after flow reversal. As a first, and conservative, approximation of the additional heat required during such transient operation, we considered reactor 4, as this requires the highest temperature rise of 270 K per flow reversal. To maintain a residence time in reactor 4 of 10 min, whilst still maintaining a steady output of hydrogen product, additional heat will be needed to induce the 270 K temperature rise over the 10 min period. If this additional heat were not available, the time taken to increase the reactor temperature would be greater than the residence time, decreasing the rate of hydrogen production, and reducing plant capacity. On this basis, the additional heat requirement is ∼24 MW. Such heating (and elsewhere cooling) requirements during this transient operation have not been considered in previous analysis. The additional heat input to reactor 4 would cause UT−3 to drop from the upper bound of 25.4% (LHV) derived earlier to around 12.6% (LHV). Note that the estimation of additional heating during flow reversal presented here assumes that, once the temperature of reactor 4 is equilibrated, all other necessary temperature rises around the system will have taken place, i.e. that no further heat is required. If this assumption proves to be invalid, then additional heat would be required, further decreasing the overall efficiency of the process.
3.4. Effect of flow reversal on the process efficiency
3.5. Solar pathway efficiency
After the completion of a clockwise single pass, the UT-3 process requires that the reactors are switched (R1 ⇔ R2, R3 ⇔ R4) and the flow is reversed for continuous H2 production. The resultant temperature changes at key points of the simplified flowsheet due to this reversal are shown in Fig. 5. The piping in the vicinity of reactors 3 and 4 would undergo the greatest temperature change, being more than 249 K per reversal. If the reversal were achieved without any irreversibilities, there would be no effect on the process
The efficiency of the CNS is 29–56% and that of the thermal energy storage is 98.8% [10]. Therefore, the solar portion of the pathway has efficiencies in the range 28.6–55.3%. This should be multiplied by the process efficiency of the UT-3 process of < 13%. The resultant efficiency of < 7% can be seen to be less than the reference range of 7–11% for PV followed by water electrolysis.
3 Ideal, isothermal compression at room temperature gives 430 kW; 6-stage adiabatic compression with 50% efficient compressors raises the power requirement to 1 MW.
If high-temperature heat is taken as the reference starting point, as is the case when considering thermal hydrogen production from nuclear energy, then the UT-3 process
H2O
HX01
HX2
HX3 ΣQi
Reaction 1 up 181K
up 156K
H2
Reaction 4 up 270 K
up 249 K
Fig. 5. The simplified flowsheet of Fig. 1 after flow reversal. Reactors 2 and 3 of Fig. 1 now carry out reactions 1 and 4, for which the temperature levels have to be raised as indicated. For the other reactors, the temperature levels decrease by the same amounts.
3.6. High-temperature (nuclear) pathway efficiency
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E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
efficiency (< 13%) must be compared with efficiency of electricity production, followed by electrolysis, for which the pathway efficiency is between 24% and 30%. Again, the thermochemical cycle is less efficient than its currently existing alternative. 4. Conclusion The efficiency of hydrogen production by the UT-3 cycle has been shown to be less than 13%, and is likely to be even lower when additional process uncertainties are accounted for. Irrespective of whether solar energy or high-temperature heat is chosen as starting point, this results in pathway efficiencies that are lower than the obvious alternatives that involve electricity generation followed by electrolysis. Unless significant advances are made in improving the efficiency of the CRS, membrane technology for hydrogen separation rather than condensation, or a reduction of operational uncertainties such as flow reversal and incomplete equilibrium conversions, the implication of this work is that the UT-3 process is not a viable alternative to hydrogen generation via electrolysis. 5. Postscript As we circulated a preprint of this article, we received feedback that the poor efficiency of the UT-3 process was well known within the R&D community. We have however not been able to find such a statement or an analysis as the present one in the public domain. Moreover, the UT-3 process generally features as one of the four plausible high-temperature hydrogen processes, next to the sulphur–iodine cycle and the normal and modified Westinghouse cycles. A prominent example is provided by the Generation IV Nuclear Energy Roadmap study [16]. Our in-depth analysis of one process brings to the fore the issue of whether or not a thermochemical cycle can—on the basis of efficiency—compete with electricity generation followed by electrolysis. If not, the incentive for its development seems marginal. In view of the renewed interest in thermochemical cycles, analyses such as the present one should be carried out for other prominent cycles as well and be published in the open literature, so as to allow the wider R&D community to more realistically assess the prospects of this intriguing route for hydrogen production. Acknowledgements The authors wish to thank Mr. Evert Wesker for fruitful discussions. The financial support of Shell Hydrogen B.V. is gratefully acknowledged.
References [1] Funk JE, Reinstrom RM. Energy requirements in the production of hydrogen from water. Ind Eng Chem Process Des Dev 1966;5:336. [2] Doctor RD, Wade DC, Mendelsohn MH. STAR-H2: a calcium–bromine hydrogen cycle using nuclear heat. Paper 139c, Argonne National Laboratory, presented at the Spring National Meeting of the American Institute of Chemical Engineers, New Orleans, Louisiana, USA, March 10–14, 2002. [3] Kameyama H, Yoshida K. Br–Ca–Fe water decomposition cycles for hydrogen production. In: Proceedings of the second world hydrogen energy conference, Zurich, Switzerland. August 21–24, 1978. p. 829–50. [4] Kameyama H, Tomino Y, Sato T, Amir R, Orihara A, Aihara M, Yoshida K. Process simulation of “MASCOT” plant using the UT-3 thermochemical cycle for hydrogen production. Int J Hydrogen Energy 1989;14(5):323–30. [5] Sakurai M, Bilgen E, Tsutsumi A, Yoshida K. Adiabatic UT3 thermochemical process for hydrogen production. Int J Hydrogen Energy 1996;21(10):865–70. [6] Sakurai M, Bilgen E, Tsutsumi A, Yoshida K. Solar UT-3 thermochemical cycle for hydrogen production. Sol Energy 1996;57:51–8. [7] Stokes RA. Solar energy. In: Kroschwitz JI, Grant MH (editors). Encyclopedia of chemical technology, Vol. 22, 4th ed. New York: Wiley; 1996, p. 465–81. [8] Rosen MA. Energy and exergy analyses of electrolytic hydrogen production. Int J Hydrogen Energy 1995;20(7): 547–53. [9] Stoll RE, von Linde F. Hydrogen—what are the costs? Hydrocarbon Process 2000; 42–6. [10] Grainger DR. The solar thermochemical hydrogen pathway. Report to Shell Global Solutions International B.V., Amsterdam, Netherlands, 2001, unpublished. [11] Beghi GE. A decade of research on thermochemical hydrogen at the joint research centre, ISPRA. Int J Hydrogen Energy 1986;11(12):761–71. [12] Jones MD. Presentation at the IEA workshop “Towards Hydrogen”, 3 March 2003. http://www.iea.org/workshop/ /2003/hydrogen/keynotes/bp.pdf . [13] Aihara M, Umida H, Yamaguchi S, Tsutsumi A, Yoshida K. Calcium bromide compounds in the thermochemical production of hydrogen—the design and analysis of calcium compounds. J Chem Soc Jpn 1988;12:2031–6. [14] Hidekazu I. Thermochemical production of hydrogen by iron–bromine cycle. Japanese Patent No. 53034695, 1978. [15] Yoshida K. Present status of R&D for thermochemical hydrogen production. J Fuel Soc Jpn 1991;70(6): 478–86. [16] A Technology Roadmap for Generation IV Nuclear Energy Systems. U.S. DOE, GIF-002-00, December 2002. [17] Beghi GE. Development of thermochemical and hybrid processes for hydrogen production. Int J Hydrogen Energy 1985;10(7/8):431–8.
A critical pathway energy efficiency analysis of the thermochemical UT-3 cycle E.D. Teoa , N.P. Brandona , E. Vosb , G.J. Kramerb,∗ a Department of Chemical Engineering, Imperial College, London SW7 2AZ, UK b Shell Research and Technology Centre, OGIR, Shell Global Solutions International B.V., P.O. Box 38000, Amsterdam 1030 BN, Netherlands
Available online 21 September 2004
Abstract Numerous thermochemical cycles for the production of hydrogen have been proposed in the literature, and of these the UT3 cycle is a leading exponent. The primary objective of such processes is to be more energy efficient than water electrolysis in converting heat to hydrogen. This paper presents a critical assessment of the actual energy efficiency that could be realised in a thermochemical cycle, taking the UT-3 cycle as the basis for the study. It is concluded that the upper efficiency for this process lies around 12%, on the basis of the lower heating value of the hydrogen product, comparable to current hydrogen production techniques using photovoltaics followed by water electrolysis. The practical upper efficiency may however be lower, even much lower, as several uncertainties exist in key aspects of the process. This in-depth analysis of one thermochemical process brings to the fore issues of whether such cycles can—on the basis of efficiency—compete with electricity generation followed by electrolysis. In view of the renewed interest in thermochemical cycles, analyses such as the one presented in this paper must be carried out and published in the open literature, so that the wider research community can realistically assess the prospects of this intriguing route to hydrogen production. 䉷 2004 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. Keywords: Thermochemical; UT-3 cycle
1. Introduction Numerous thermochemical cycles have been proposed since the publication of the first article on the energy requirements of the water decomposition process by Funk and Reinstrom [1]. The primary objective of such processes is to offer a sustainable route to hydrogen, which is more energy efficient than using photovoltaics (PV) coupled with water electrolysis. Recent developments have focused on the UT-3 process and the iodine–sulphur cycle developed by General Atomics, identified as the two leading candidates at the First Information Exchange Meeting on the Nuclear Production ∗ Corresponding author. Tel.: +31-20-630-2467; fax: +31-20-
630-3964. E-mail address: [email protected] (G.J. Kramer).
of Hydrogen of the Organization for Economic Cooperation and Development in October 2001 [2]. The UT-3 (University of Tokyo, Br, Ca and Fe) process was proposed by Kameyama and Yoshida [3], with the following chemical reactions: reactor 1 :
CaBr 2 (s) + H2 O(g) 973–1023 K
−→
reactor 2 :
(1)
CaO(s) + Br 2 (g) 773–873 K
−→
reactor 3 :
CaO(s) + 2HBr(g),
CaBr 2 (s) + 0.5O2 (g),
(2)
Fe3 O4 (s) + 8HBr(g) 473–573 K
−→
3FeBr 2 (s) + Br 2 (g)
+ 4H2 O(g),
0360-3199/$30.00 䉷 2004 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2004.08.003
(3)
560
E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
individual heat inputs and the second for the work of each unit operation divided by its associated efficiency [17] the total thermal energy input to the process including the equivalent thermal input of the required electrical energy
Nomenclature UT−3 Qe QH2
Q
reactor 4 :
efficiency of the UT-3 process excess (thermal) energy production heating value of product hydrogen, based on lower heating value (241 kJ/mol) the the heat inputs; specifically sumof all Q = Qi + Wi /i , the first term for the
3 FeBr 2 (s) + 4H2 O(g) 833–873 K −→ Fe3 O4 (s) + H2 (g) + 6HBr(g).
Qi
ΣQi
(4)
The flow diagram of the process is reproduced in Fig. 1. The Model Apparatus for the Study of Cyclic Operation in Tokyo (MASCOT) pilot plant was operated in the late 1980s [4] and a potential energy efficiency of 49.5% (on basis of the higher heating value—HHV—of the product hydrogen) has been suggested [5,6]. The primary heat source could be supplied by a nuclear reactor or by using solar heat. The final energy carrier would be dependent on the location. The terrestrial PV market includes the consumer production, utility generation and remote power generation segments. The remote power generation segment of stand-alone systems has the most rapid growth and contribution to PV sales [7]. This is also the sector where thermochemical cycles would expect to compete, and hydrogen could be the final energy carrier instead of electricity. In this study, the solar to hydrogen efficiency of the UT-3 cycle is compared against the equivalent efficiency of the electrolysis of water with PV-generated electricity. We note that in this field it seems customary to quote efficiencies on the basis of the HHV of hydrogen, which is 286 kJ/mol. Most papers that discuss the use of hydrogen however have a preference to base discussion on lower heating value (LHV), which for hydrogen is 241 kJ/mol. In calculating pathway efficiencies, this may easily lead to confusion and error. In our paper we will standardise on LHV; exceptions, e.g. when discussing literature data, will be explicitly mentioned.
2. Methodology For this study, two “pathway efficiencies” from sunlight to hydrogen will be considered: Route 1: For PV and water electrolysis: Sunlight a → Electricity b → Hydrogen. Route 2: For thermochemical cycles: Sunlight c → Central Receiver System d → Thermal Storage System e → Thermochemical Process producing Hydrogen, where a is the efficiency of PV, defined as the ratio of electricity produced to the amount of sunlight incident on
1033 K 957 K Reactor 1 HX01 861 K
H2O
833 K 725 K Reactor 4
HX1 H2
HX4 O2
HX2
HX3 Reactor 2 864 K
Reactor 3 845 K
576 K
493 K
Qe
Fig. 1. Simplified flow-scheme of the UT-3 cycle.
the PV device [7]. Efficiencies of 11% are attainable using thin-film copper indium diselinide (CIS) solar modules and 14–15% for mono- and multi-crystalline modules, b is the efficiency of water electrolysis, typically ranging from 60% to 70% on LHV basis (equivalent to 70–85% on HHV basis) [8,9]. Electrolysis for hydrogen production is generally used to meet the demand in isolated markets or desert regions where other energy sources are not in abundance. For water electrolysers, the energy efficiency is the ratio of the equilibrium cell voltage to the actual voltage required, taking into account the overpotentials at the electrodes and the ohmic potential drops. c,d is the combined efficiency of the solar furnace/central receiver system (CNS) (29–56%) and the thermal energy storage (98.8%) [10], and e is the reported efficiency of the UT-3 cycle of 49.5% on HHV basis [6], equivalent to 42% on LHV basis. As seen from Fig. 2a, the thermochemical cycle route for hydrogen production (Route 2) is up to three times more efficient than current technologies (Route 1) if hydrogen were the final energy carrier. Note if we use the high-temperature heat to generate electricity with 40–45% efficiency, using high-pressure steam Rankine cycle, this route (indicated by the dashed line in Fig. 2a) would have a 13–25% efficiency, up to two-and-a-half times that of PV.1 If the electric power thus generated is used to produce hydrogen by electrolysis, another relevant comparison results (Fig. 2b). This case, where one starts with high-temperature heat and compares conventional electricity generation 1 Note that the upper efficiency limit corresponds to the as-
sessment by Hamache and Bilgen (1992).
E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
of the UT-3 process [5], UT−3 as
Solar Energy Energy Capture by Photovoltaic cells 11 – 15 % Route 1 7 – 11 %
Electricity
Solar Furnace/ Central Receiver 29 – 55 % High-temperature heat
Electrolysis 60 – 70 %
UT−3 = [QH2 + Qe ]/ Route 2 12 – 23 %
Thermochemical cycle 42 % (suggested)
Hydrogen (LHV)
(a)
High-temperature heat (800oC) High-pressure Rankine cycle 40 – 45% Route 1’ 24 – 32 %
Thermochemical cycle 42 %(suggested)
Electricity
Electrolysis 60 – 70 %
(b)
561
Hydrogen (LHV)
Fig. 2. Pathway efficiencies from solar energy to hydrogen (a) and pathway efficiencies from high-temperature heat to hydrogen (b).
followed by electrolysis (Route 1 ) with the thermochemical route, is the relevant comparison when considering hydrogen production from nuclear heat. Here, the notional 42% efficiency of the UT-3 cycle is about one-and-a-half times as efficient as the conventional Route 1 (24–32%). In both modes of comparison and notwithstanding future developments in CNS and water electrolysis, the attractiveness of the thermochemical route lies in its (potentially) higher process efficiency, which requires studying in greater detail. Cost is often viewed as the only “figure of merit” [11] for the competitiveness of a process, with steam reforming as the dominant route for hydrogen production at 5.22 US$/GJ [2]. Production cost of ‘renewable’ hydrogen via PV and electrolysis is estimated between 50 and 100 US$/GJ [12]. It would be premature to calculate the cost of hydrogen per GJ for the thermochemical cycle given the multitude of assumptions for plant sizing and price quotes that are organisational dependent. In the absence of this, process efficiency is taken to be the best indicator for the prospects of the technology. It is proposed that energy analysis will suffice at this juncture, as the difference in using exergy is a mere 3.4% for the UT-3 process [6]. Exergy analysis could be carried out if the calculated energy efficiency values are significantly more attractive than the reference technology chosen. Due to possible excess thermal energy for power generation, researchers of the UT-3 process defined the efficiency
Qi ,
(5)
the product where QH2 is the (LHV) energy content of hydrogen, Qe is the excess thermal heat and Qi is the heat input. In the original paper, QH2 referred to the HHV of hydrogen. In line with the remarks in the introduction we will use the LHV. In this study the upper bound of the UT-3 process efficiency is revised downwards from the theoretical Carnot value of 55% [2] (LHV basis; the reference quotes 66% on HHV basis) and the 42% (LHV) obtained by the UT-3 researchers [6] by the inclusion of increasingly stringent but practical constraints. The effects on process efficiency of incomplete reactions and separation via steam condensation used in the MASCOT pilot plant are then estimated, with the aid of modelling tools such as Aspen Plus䉸 and HSC Chemistry䉸. The efficiencies of the CNS and thermal storage are subsequently factored to derive the solar pathway efficiencies (i.e. the efficiency of H2 production using a CNS and thermal storage coupled to the UT-3 process) for the production of H2 gas, which are then compared to that of PV and water electrolysis. The analyses of the separations, heat integration, compressor work inputs were based on the stream table given by Sakurai et al. [6].
3. Results 3.1. Upper bound of the UT-3 process efficiency The published process efficiency of the UT-3 process, UT−3 as given in Eq. (5), is 42.1% when expressed on LHV basis. In the case worked out by Sakurai et al., the hydrogen product (25 mol/s) energy content is equivalent to 6.03 MW (QH2 ) and results from an input 15.24 MW of high-temperature heat. This is actually 39.5% efficiency. The remaining 2.6% in Sakurai’s calculation is 0.40 MW of low-temperature (550 K) heat. As there is no use for that heat in the process, its inclusion in the efficiency figure is misleading. The best use is to convert the heat to electricity. For this, Sakurai et al. suggest 17% efficiency, resulting in 0.07 MW of electricity. This could be converted to hydrogen through electrolysis at 70% efficiency, adding 0.05 MW to the product hydrogen stream, resulting in a maximum hydrogen production efficiency of 39.9%. Our independent analysis of the heat streams in the process yielded data that were not significantly different from to those of Sakurai et al. reflecting primarily the uncertainty in the thermodynamic basic data involved. For convenience we will use the heat data of Sakurai et al. in what follows. From the pinch diagramin Fig. 3 it is seen that the hightemperature input heat ( Qi , 15.24 MW) is fully used to heat the process stream (2500 mol/s of recycle water plus
E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
3.2. Effect of incomplete equilibrium conversions While 100% conversion for each reaction has been previously assumed [6], kinetic data indicate a maximum conversion based on bromination of ∼75% for the first two reactions (1) and (2) [13]. HSC䉸 Chemistry was used to predict the thermodynamic equilibrium conversions. This modelling helped to identify the critical reaction steps amongst the four reactions of the UT-3 cycle that require further verification. It was found that the H2 production step (4) has the greatest impact on the process efficiency while incomplete conversions in the other reactions mainly affect the feasibility of flow reversal to regenerate the reactants. In the conceptual design [6] 2500 mol H2 O and 50 mol HBr are fed to the reactor at 833 K and 20 bar. At the reactor exit, at 726 K, 100 mol of water is notionally consumed with 75 mol of solid Fe3 O4 to produce 25 mol of hydrogen and an additional 150 mol of HBr. Such a gas phase composition is however inconsistent with thermodynamic equilibrium. The equilibrium partial pressures of H2 , HBr and H2 O for this reactor are plotted in Fig. 4. It is seen that the hydrogen concentration according to the conceptual design can only be approached when the temperature approaches 1000 K. Such an operating temperature is not feasible however since FeBr2 will sublimate above 850 K. Thus, under practical circumstances—where one would operate the reactor isothermally at a temperature slightly below 850 K, the hydrogen concentration is an order of magnitude below that of the design, leading to ten-fold increase in water circulation. Since the high-temperature heat input to the cycle (Figs. 1 and 3) is essentially determined by the heat duty of
1100 1000 HX1
900
HX01
800 T (K)
25 mol/s of water for thermolysis) from 861 to 1033 K, the temperature at which it enters reactor 1. As suggested by Sakurai, the low-temperature waste heat from HX2 (7.53 MW) can be used to generate power to drive compressors and pumps. At 17% efficiency for the conversion of 550 K heat to power—a figure that we will use here as well—1.35 MW of power can be generated, in excess of the 1.27 MW required. In their assessment however, Sakurai et al. used 100% efficient compressors. When a practical compressor efficiency of 72% was incorporated to calculate the actual work input to the three compressors of the UT-3 cycle we find that 1.75 MW is required to drive the compressors and the water pump. Since only 1.35 MW was available from co-generation from waste heat, 0.40 MW of additional power has to be generated. Assuming 45% efficiency for electricity from high-temperature heat—as suggested above (Section 2)—an additional thermal energy input equivalent to 0.91 MW. At this level of assessment we find that 25 mol/s hydrogen, equivalent to 6.03 MW can be generated from an input 16.15 MW heat, lowering UT−3 to 37.3% (LHV). This should still be considered as an upper bound of the UT-3 process efficiency, which is subject to further reductions due to considerations of the following process uncertainties.
Qe
700
HX2
HX3
ΣQ i
600 500 400
HX4
300 0
10
20
30
40
50
heat flux (MW)
Fig. 3. Pinch diagram of the UT-3 cycle, based on the data of Sakurai et al. [6].
102 H2O
101 partial pressure (bar)
562
100
HBr
10-1
H2
10-2 10-3
726 K 833 K
10-4 10-5 700
800
900
1000
T (K)
Fig. 4. Thermodynamic equilibrium of the hydrogen production reaction (4). While Sakurai et al. [6] used full conversion—indicated by the dashed line—actual equilibrium conversion levels in the operating range (726–833 K) are much lower, making increased water circulation necessary.
the recycle stream from 861 to 1033 K, the input heat requirement would increase 10-fold and the efficiency would drop by the same factor. This is somewhat compensated by the fact that the adiabatic temperature drop in reactor 1 would be 10-fold reduced. Thus, keeping the reactor outlet at 957 K, the inlet temperature would have to be only 965 K, when the increased water flux would need to be raised in HX01 from 861 to 965 K only. Taking this into account, the heat input increases six-fold, rather than 10-fold to 92 MW. This would reduce the hydrogen production efficiency to a value considerably below 10%.2 2 The actual value is 6.6%. However, some additional hydrogen may be generated from the low-temperature heat which is now present in excess.
E.D. Teo et al. / International Journal of Hydrogen Energy 30 (2005) 559 – 564
3.3. Effect of separations via condensation on the process efficiency
Reaction 2
563 Reaction 3
Qe
HX1
O2 HX4
The gaseous HBr and H2 effluent from reactor 4 could be separated using a suitable membrane or bubbled through water for the dissolution of HBr to form hydrobromic acid [14]; the latter route was utilised in the MASCOT plant [15]. In theory, the separation is path-independent and there would be no effect on the process efficiency assuming there were no irreversibilities, and that the heat loss in lowering the inlet stream to the hydrogen separator, i.e. from 742 to the 303 K, the temperature at which hydrogen is separated, is completely recoverable to reheat H2 O (g) and HBr (g) to the expected outlet temperature of the separator at 742 K. In practice however, both membrane separation and condensation do imply the input of additional work or heat. If membrane separation of hydrogen would prove possible, the permeate hydrogen would come out of the membrane at low pressure. We saw in the preceding section that the hydrogen partial pressure in the effluent of reactor 4 is about 10 mbar. The likely permeate pressure will be an order of magnitude lower, so that hydrogen becomes available at 1 mbar. Recompression to 1 bar (of 25 mol H2 per second) would require about 1 MW of power.3 Assuming, as before, that this is generated from high-temperature heat at 45% efficiency, this adds 2.2 MW of thermal power to the process inputs. If membrane separation were not a viable option in the future and the practical process would require condensation as in the pilot plant, this would be an unbearable burden. We saw above that one must expect at least 25,000 mol of water per second to be condensed and reheated to achieve 25 mol/s of hydrogen production. A fantastic 1.3 GW of heat would have to be exchanged between condensing and evaporating water streams. The inevitable losses would render the process both uneconomic and highly inefficient. Thus, the UT-3 process is only viable if high-temperature membrane separation of hydrogen in the presence of HBr proves possible. Even then, an additional 2.2 MW would be required to drive the hydrogen recompression, reducing the efficiency to 25.4%.
efficiency at steady state. However, additional heat may well be necessary to reduce the time required to reach steady state after flow reversal. As a first, and conservative, approximation of the additional heat required during such transient operation, we considered reactor 4, as this requires the highest temperature rise of 270 K per flow reversal. To maintain a residence time in reactor 4 of 10 min, whilst still maintaining a steady output of hydrogen product, additional heat will be needed to induce the 270 K temperature rise over the 10 min period. If this additional heat were not available, the time taken to increase the reactor temperature would be greater than the residence time, decreasing the rate of hydrogen production, and reducing plant capacity. On this basis, the additional heat requirement is ∼24 MW. Such heating (and elsewhere cooling) requirements during this transient operation have not been considered in previous analysis. The additional heat input to reactor 4 would cause UT−3 to drop from the upper bound of 25.4% (LHV) derived earlier to around 12.6% (LHV). Note that the estimation of additional heating during flow reversal presented here assumes that, once the temperature of reactor 4 is equilibrated, all other necessary temperature rises around the system will have taken place, i.e. that no further heat is required. If this assumption proves to be invalid, then additional heat would be required, further decreasing the overall efficiency of the process.
3.4. Effect of flow reversal on the process efficiency
3.5. Solar pathway efficiency
After the completion of a clockwise single pass, the UT-3 process requires that the reactors are switched (R1 ⇔ R2, R3 ⇔ R4) and the flow is reversed for continuous H2 production. The resultant temperature changes at key points of the simplified flowsheet due to this reversal are shown in Fig. 5. The piping in the vicinity of reactors 3 and 4 would undergo the greatest temperature change, being more than 249 K per reversal. If the reversal were achieved without any irreversibilities, there would be no effect on the process
The efficiency of the CNS is 29–56% and that of the thermal energy storage is 98.8% [10]. Therefore, the solar portion of the pathway has efficiencies in the range 28.6–55.3%. This should be multiplied by the process efficiency of the UT-3 process of < 13%. The resultant efficiency of < 7% can be seen to be less than the reference range of 7–11% for PV followed by water electrolysis.
3 Ideal, isothermal compression at room temperature gives 430 kW; 6-stage adiabatic compression with 50% efficient compressors raises the power requirement to 1 MW.
If high-temperature heat is taken as the reference starting point, as is the case when considering thermal hydrogen production from nuclear energy, then the UT-3 process
H2O
HX01
HX2
HX3 ΣQi
Reaction 1 up 181K
up 156K
H2
Reaction 4 up 270 K
up 249 K
Fig. 5. The simplified flowsheet of Fig. 1 after flow reversal. Reactors 2 and 3 of Fig. 1 now carry out reactions 1 and 4, for which the temperature levels have to be raised as indicated. For the other reactors, the temperature levels decrease by the same amounts.
3.6. High-temperature (nuclear) pathway efficiency
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efficiency (< 13%) must be compared with efficiency of electricity production, followed by electrolysis, for which the pathway efficiency is between 24% and 30%. Again, the thermochemical cycle is less efficient than its currently existing alternative. 4. Conclusion The efficiency of hydrogen production by the UT-3 cycle has been shown to be less than 13%, and is likely to be even lower when additional process uncertainties are accounted for. Irrespective of whether solar energy or high-temperature heat is chosen as starting point, this results in pathway efficiencies that are lower than the obvious alternatives that involve electricity generation followed by electrolysis. Unless significant advances are made in improving the efficiency of the CRS, membrane technology for hydrogen separation rather than condensation, or a reduction of operational uncertainties such as flow reversal and incomplete equilibrium conversions, the implication of this work is that the UT-3 process is not a viable alternative to hydrogen generation via electrolysis. 5. Postscript As we circulated a preprint of this article, we received feedback that the poor efficiency of the UT-3 process was well known within the R&D community. We have however not been able to find such a statement or an analysis as the present one in the public domain. Moreover, the UT-3 process generally features as one of the four plausible high-temperature hydrogen processes, next to the sulphur–iodine cycle and the normal and modified Westinghouse cycles. A prominent example is provided by the Generation IV Nuclear Energy Roadmap study [16]. Our in-depth analysis of one process brings to the fore the issue of whether or not a thermochemical cycle can—on the basis of efficiency—compete with electricity generation followed by electrolysis. If not, the incentive for its development seems marginal. In view of the renewed interest in thermochemical cycles, analyses such as the present one should be carried out for other prominent cycles as well and be published in the open literature, so as to allow the wider R&D community to more realistically assess the prospects of this intriguing route for hydrogen production. Acknowledgements The authors wish to thank Mr. Evert Wesker for fruitful discussions. The financial support of Shell Hydrogen B.V. is gratefully acknowledged.
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